The Pennsylvania State University

The Graduate School

Department of Mechanical and Nuclear Engineering

INFLUENCE OF STEAM ON THE FLAMMABILTY LIMITS OF

PREMIXED NATURAL GAS/OXYGEN/STEAM MIXTURES

A Thesis in

Mechanical Engineering

by

Matthew J. Degges

2010 Matthew J. Degges

Submitted in Partial Fulfillment

of the Requirements

for the Degree of

Master of Science

May 2010

ii

The thesis of Matthew J. Degges was reviewed and approved* by the following:

Kenneth K. Kuo

Distinguished Professor of Mechanical Engineering

Thesis Advisor

Horacio Perez-Blanco

Professor of Mechanical Engineering

Karen A. Thole

Professor of Mechanical Engineering

Head of the Department of Mechanical and Nuclear Engineering

*Signatures are on file in the Graduate School

iii

ABSTRACT

Synthesis gas (syngas), a mixture of CO and H2, is an intermediate in a variety of

industrial processes. Its production is energy and capital intensive. Any improvement of existing

technologies allowing simpler and economic production is of great interest. Recently, a method

known as Short Contact Time – Catalytic Partial Oxidation (SCT-CPO) has been developed into a

commercial technology. SCT-CPO is an entirely heterogeneous catalytic process converting

premixed flammable feedstocks inside a very small reactor. In order to ensure operator safety

with a high selectivity towards CO and H2, it has been important to determine and understand

flammability properties of the gaseous reactant mixtures. A unique test chamber has allowed the

study of ignition, flame propagation, and explosion characteristics of gas mixtures similar to

those used as reactants in the SCT-CPO reactor. The tests were conducted at various pressures

with different mole fractions of steam and two different compositions of natural gas (NG). A

flammability boundary for the mixtures, based on normalized pressure and mole fraction of

steam, was determined. Previous studies indicate that steam can be used to suppress the

flammability of a mixture by both physical and chemical processes. To examine the chemical

processes, Chemkin Code calculations were executed. Similar to the experimentally observed

phenomena, the Chemkin calculations also showed that more steam was required to suppress the

flammability of the mixture with the higher adiabatic flame temperature. The results show highly

non-linear flammability boundaries for both hydrocarbon/oxygen/steam mixtures were very

sensitive to equivalence ratio and pressure, as the mechanism for flammability suppression by

steam is also strongly dependent on these parameters.

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TABLE OF CONTENTS

List of Figures .......................................................................................................................... vi

List of Tables ........................................................................................................................... viii

Nomenclature ........................................................................................................................... ix

Chemistry Nomenclature ......................................................................................................... x

Acknowledgements .................................................................................................................. xii

Chapter 1 Introduction ............................................................................................................ 1

1.1 Overview of Experimentation .................................................................................... 2 1.2 Overview of Chemkin Calculations ........................................................................... 3 1.3 Research Goals ........................................................................................................... 3

Chapter 2 Literature Review ................................................................................................... 4

2.1 Ignition and Flammability .......................................................................................... 4 2.2 Influence of Steam on Flammability Limits .............................................................. 5 2.3 Flammability Limit Dependence on Equivalence Ratio ............................................ 7 2.4 Influence of Pressure on Flammability Limits ........................................................... 8 2.5 Influence of Reaction Zone Temperature on Flammability Limits ............................ 10 2.6 Cool Flame Phenomenon ........................................................................................... 13 2.7 Flammability Experiments ......................................................................................... 16 2.8 Summary of Processes Governing Flammability Limits ........................................... 17 2.9 Chemical Process Safety Characterization of Flammability Limits .......................... 19 2.10 Flammability Modeling ............................................................................................ 22

Chapter 3 Method of Approach .............................................................................................. 24

3.1 Experimental Method of Approach ............................................................................ 24 3.2 Computational Method of Approach ......................................................................... 30

Chapter 4 Experimentation ..................................................................................................... 32

4.1 Instrumentation .......................................................................................................... 32 4.1.1 Pressurization Data .......................................................................................... 32 4.1.2 Photodetector Flame Spreading Data .............................................................. 36 4.1.3 Flow Rate Data ................................................................................................ 39 4.1.4 Ignition Source ................................................................................................ 41

4.2 Summary of Uncertainties.......................................................................................... 42 4.3 Safety Analysis .......................................................................................................... 45

Chapter 5 Chemkin Code Calculations ................................................................................... 48

5.1 Overview of Calculations ........................................................................................... 48

v

5.2 Premixed Laminar Flame Speed Calculations ........................................................... 48 5.3 Homogenous Batch Reactor Calculations .................................................................. 54

5.3.1 Sensitivity Studies ........................................................................................... 55 5.3.2 Flammability Limits Using Homogenous Batch Reactor Model .................... 59

Chapter 6 Discussion of Flammability Results ....................................................................... 61

6.1 Flammability Results Compared with Chemkin Model ............................................. 61 6.2 Summary of Flammability Results ............................................................................. 65

Chapter 7 Conclusions ............................................................................................................ 68

References ................................................................................................................................ 71

Appendix A Experimental Test Matrix ................................................................................... 74

Appendix B Individual Test Summaries ................................................................................. 87

B.1 Test Summaries with Mixture 1 Test Series .............................................................. 87 B.2 Test Summaries with Mixture 2 Test Series .............................................................. 88

Appendix C Test Data Sheet ................................................................................................... 91

Appendix D Premixed Gas Reactor Test Checklist ................................................................ 92

Appendix E Error and Uncertainty ......................................................................................... 96

Appendix F General Calculations ........................................................................................... 98

vi

List of Figures

Figure 1: Effect of temperature and pressure on the flammability of hexane in air [23]. ........ 14

Figure 2: Ignition limits for hydrocarbons [24] ....................................................................... 15

Figure 3: Schematic of chemical and physical processes governing flammability test ........... 19

Figure 4: Flammability Triangle Diagram for a mixture of methane vapor in air [16] ........... 21

Figure 5: A 135o-sectional view of the windowed high-pressure tube reactor ........................ 25

Figure 6: Process flow diagram for flammability study of hydrocarbon/steam/oxygen

mixtures ............................................................................................................................ 26

Figure 7: Liquid petroleum gas (LPG) reservoirs and stock tanks, and methane and

hydrogen manifolds .......................................................................................................... 26

Figure 8: Test reactor experimental setup with flow controls, pre-heater, and steam

generator........................................................................................................................... 27

Figure 9: Three representative P-t traces recorded from the tube reactor ................................ 28

Figure 10: High-speed camera visualization of turbulent premixed flame front [35] ............. 28

Figure 11: Representative photodetector intensity-time traces [35] ........................................ 29

Figure 12: Representative linear fit of flame-propagation speed [35] ..................................... 29

Figure 13: Recorded pressure traces of Test #55 ..................................................................... 32

Figure 14: Recorded dynamic pressure traces of Test #53 ...................................................... 33

Figure 15: Recorded dynamic pressure traces of Test #73 ...................................................... 34

Figure 16: Recorded dynamic pressure traces of Test #56 ...................................................... 35

Figure 17: Photodetector intensity-time traces of Test #43 ..................................................... 36

Figure 18: Photodetector intensity-time traces of Test #55 ..................................................... 37

Figure 19: Photodetector intensity-time traces of Test #53 ..................................................... 38

Figure 20: Linear fit of flame-propagation speed for Test #43 ................................................ 39

Figure 21: Time variations of flow rates of reactants of Test #43 ........................................... 40

Figure 22: Vaporization curves for propane and butane .......................................................... 43

vii

Figure 23: Theoretically calculated maximum pressures for equilibrium reaction of

NG+LPG/steam/O2 as a function of initial chamber pressure with two different fuel

ratios of LPG .................................................................................................................... 47

Figure 24: Conceptual diagram of the Chemkin premixed laminar flame speed model .......... 49

Figure 25: Solution sensitivity to initial number of grids ....................................................... 50

Figure 26: Solution sensitivity to value of curvature and gradient ......................................... 51

Figure 27: Comparison of multi-component diffusion and mixture average transport ........... 52

Figure 28: Chemkin Code calculated evolution of temperature profile with different

amounts of steam addition ............................................................................................... 53

Figure 29: Chemkin calculated evolution of centerline velocity with increasing steam ......... 54

Figure 30: A-Factor sensitivity studies for Mixture 2 ............................................................. 56

Figure 31: Temperature sensitivity study for Mixture 1 and 2 at 15 atm ................................ 57

Figure 32: Chemkin Code Flammability Limit study for Mixture 2 ...................................... 60

Figure 33: Flammability Limits of Mixture 1 ......................................................................... 61

Figure 34: Flammability Limits of Mixture 2 .......................................................................... 63

Figure 35: Glass tube liner post-test from Test #37 ................................................................. 87

viii

List of Tables

Table 1: Composition of Mixture 1 and Mixture 2 without Steam. ......................................... 2

Table 2: Percent uncertainty in various components of the mixture for different tests .......... 42

Table A.1: Mixture 1 test series initial flow rates and chamber pressure ............................... 75

Table A.2: Mixture 1 test series individual reactant species mole fraction ............................ 76

Table A.3: Mixture 1 test series initial conditions: O/C , S/C , temperatures, and φ ............ 77

Table A.4: Mixture 1 test series flow parameters ................................................................... 78

Table A.5: Mixture 1 test series flammability and type of pressurization .............................. 79

Table A.6: Mixture 1 test series steam parameters and orifice type ....................................... 80

Table A.7: Mixture 2 test series initial flow rates and chamber pressure ............................... 81

Table A.8: Mixture 2 test series individual reactant species mole fraction ............................ 82

Table A.9: Mixture 2 test series initial conditions: O/C , S/C , temperatures, and φ ............. 83

Table A.10: Mixture 2 test series flow parameters .................................................................. 84

Table A.11: Mixture 2 test series flammability and type of pressurization ............................ 85

Table A.12: Mixture 2 test series steam parameters and orifice type ...................................... 86

ix

Nomenclature

Symbol Description Units

Pc,i Initial Chamber Pressure [atm or psig]

Pref Reference Pressure [atm or psig]

P Pressure [atm or psig]

t Time [seconds]

ρ Density [kg/m3]

µ Dynamic Viscosity [Pa-s]

ht Total Enthalpy [J/kg]

q Heat Flux [W/m2]

A Area [m2]

Tf Adiabatic Flame Temperature [K]

Equivalence Ratio [-]

Xi Mole Fraction of Species i [-]

Re Reynolds Number on the Diameter of Tube Reactor [-]

V Mean Velocity of Reactants [-]

ST,abs Absolute Turbulent Flame Speed [m/s]

ST,rel Relative Turbulent Flame Speed [m/s]

Nm3/hr Normal cubic meters per hour [-]

SLPM Standard liters per minute [-]

S/C Steam-to-Carbon Mass Ratio [-]

O/C Oxygen-to-Carbon Mass Ratio [-]

Mass Flow Rate [kg/s]

PD Photodetector [-]

x

Symbol Description Units

LPG Liquefied Petroleum Gas [-]

UFL Upper Flammability Limit [% Vol.]

LFL Lower Flammability Limit [% Vol.]

LOC Limiting Oxygen Concentration [% Vol.]

Chemistry Nomenclature

Symbol Description

H Hydrogen Radical

O2 Oxygen

OH Hydroxyl Radical

H Oxygen Radical

M Third Body

HO2 Hydroperoxy Radical

H2O2 Hydrogen Peroxide

CH3 Methyl Radical

C2H5 Ethyl Radical

N2 Nitrogen

H2 Hydrogen

CH4 Methane

C2H6 Ethane

C3H8 Propane

C4H10 Butane

CH2O Formaldehyde

xi

Symbol Description

CH3OH Methanol

CH3HCO Acetaldehyde

xii

Acknowledgements

This research has been supported by Eni Div. R&M. Input from Dr. Luca Basini and Mr.

Andrea Lainati of Eni Div. R&M is greatly appreciated. I would like to also acknowledge Mr.

Patrick Kutzler of PSU for his participation in the early phase of this project. I would personally

like to thank Dr. Kenneth Kuo for his guidance and support throughout my time at the High

Pressure Combustion Laboratory. Dr. Eric Boyer’s hard work and dedication to this research is

greatly appreciated. I would also like to thank Trevor Wachs for many hours spent cleaning,

assembling, and helping run these tests. Also thanks to all of my associates and friends at the

HPCL: Mr. Scott Blakeslee, Mr. Alex Colletti, Mr. Drew Cortapassi, Mr. Jon Essel, Mr. Brian

Evans, Mr. Ryan Houim, Mr. Jeff Krug, Mr. Heath Martin, Mr. Matt Sirignano, and Prof. Bao Qi

Zhang.

Chapter 1

Introduction

As natural gas (NG) is a highly available resource, new methods for processing it into

forms that are more efficient has been of great interest recently. The ability to convert NG into an

easily storable liquid form would benefit the global economy due to the dependence on oil for

liquid fuels. Also, NG can be converted to hydrogen to be used to power fuel cells; another

alternative power generation method. Processing NG into these two useful energetic materials

requires the production of the intermediate material known as synthesis gas or syngas [1,2].

Syngas is composed of H2 and CO.

Several methods for creating syngas are known, but of interest in this research is the

process of Catalytic Partial Oxidation (CPO). CPO is a heterogeneous catalytic process in which

a reactant is flowed over a rhodium coated surface, which acts as a catalyst to initiate a partial

oxidation reaction [2]. The partial oxidation of a natural gas/oxygen/diluent mixture will mainly

produce the products H2 and CO. The CPO method used with the correct mixture composition

will create syngas with a H2/CO ratio of 2, which is a favorable ratio for further processing [2].

Recently a process known as short contact time catalytic partial oxidation, SCT-CPO, has

been developed for use in industry. This process is similar to CPO, except it forces the reactant

mixture over a very hot rhodium surface for only few milliseconds. This short contact time with a

hot catalyst favors the formation of primary reactions and inhibits degradation of their products;

further degradation would cause chain reactions that would lead to loss of production of syngas

and introduce safety issues [3]. SCT-CPO is a process that is very efficient and flexible in the

creation of syngas.

2

The main objective of this study was to define the flammability boundaries of natural gas

(NG), steam, and oxygen mixtures in conditions very close to those that could be adopted in

industrial SCT-CPO processes [3,4,5,6]. The effect of steam on combustion processes is an

increasingly important issue which is not limited to the SCT-CPO case. For other technologies,

steam can affect the combustion processes in: exhaust gas recirculation in I.C. engines; in nuclear

power plant accident suppression systems; and in the operation of combined cycle gas turbine

(IGCC) fed with synthesis gas [7]. It is important to determine the suitable amount of steam used

in the SCT-CPO reactor for the mixture to be non-flammable. Understanding the flammability of

a mixture can be used to enhance safety. The main objective of this study is to characterize the

flammability boundaries of two types of mixtures composed of natural gas simulant combined

with steam and oxygen referred to as Mixture 1 and Mixture 2.

Table 1: Composition of Mixture 1 and Mixture 2 without Steam.

Parameter Mixture 1 Mixture 2

O2 38 % 57 %

CH4 56 % 23 %

C2H6 4 % 1 %

H2 2 % -

C3H8 - 12 %

C4H10

-

3.33

7 %

2.69

1.1 Overview of Experimentation

To simulate industrial applications used in the production of syngas, the experimental

setup of this study required the flow Reynolds number (ReD) of the premixed reactants at levels

above 20,000 and initial chamber pressures of up to 30 atm in the tube reactor. These conditions

implied that the flow was highly turbulent. In order to study the flammability limits of the

mixtures flowing at high Reynolds numbers and elevated pressure, a special test apparatus was

3

designed to sustain any rapid pressurizations caused by ignition. The operating temperature of the

tube reactor was required to be around 450 K so gas mixtures must be pre-heated to within this

temperature range. The heat loss from the gas supply system needed to be minimized so the

reactor was preheated before each flammability test. Another requirement was that the gaseous

components of the mixture needed to be well mixed prior to injection into the reactor. If the

original states of certain chemical ingredients were stored in a liquid form, these ingredients

needed to be vaporized before mixing with other components. All of these requirements for

experimentation were incorporated into a tube reactor test rig at the High Pressure Combustion

Lab at the Pennsylvania State University and are explained in detail in Chapter 3.

1.2 Overview of Chemkin Calculations

In addition to experimental investigation, Chemkin Code [8] calculations were performed

to analyze the flammability limits of these mixtures. Two different types of Chemkin models, the

homogenous batch reactor and the flame speed model, were used to determine which method

would best determine a flammability limit. In addition, the homogenous batch reactor calculations

were coupled with sensitivity studies to understand the reaction paths of the two mixtures. The

calculations were compared to experimental results and were useful as a heuristic study of the

flammability limit of the two mixtures.

1.3 Research Goals

Overall, this study had the following research goals:

1. Experimentally define the flammability boundary for Mixture 2.

2. Compare the flammability boundaries of Mixture 1 and Mixture 2.

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3. Use Chemkin Code calculations to assist in analysis and interpretation.

4. Make conclusions on how steam affects the flammability of the two mixtures.

Chapter 2

Literature Review

2.1 Ignition and Flammability

In determining the flammability limits of Mixture 1 and Mixture 2, tests were performed

to observe the conditions that will allow for ignition to occur. An ignition of a mixture will occur

in a combustion system if the rate of chemical reactions and heat release is sustained or increased

[9]. Every combustion system has different mechanisms for energy loss, for which the energy

addition by chemical reactions must overcome for an ignition to occur. Chemical reactions

generate radicals that collect in a “pool.” As this radical pool grows, reaction rates also increase,

since the radical species can promote reactions. For an ignition to occur, the radical pool growth

has to overcome any loss mechanisms [10]. Losses can be associated with transport mechanisms

for energy and mass, like heat losses by conduction, convection, or radiation. These losses could

also be caused by chain-terminating reactions that remove radicals from the pool.

The process of radical growth occurs in the induction period. This is a time frame in

which it takes the chemical reactions to initiate some stage of chain branching. The induction

time of any ignition experiment can be seen as the combination of two time periods: chemical

time, , and thermal time, . The chemical time is the time to build to a critical concentration of

radicals that would cause ignition. The thermal time is the time it takes subsequent reactions to

heat the mixture to its final stage via exothermic heat release or an extremely fast heat release

process called thermal explosion. Generally, , but in certain cases the thermal time can

5

be longer to bring the system to a thermal explosion, especially if the energy losses in the system

are high. A two-stage ignition can occur; in the first stage, the chemical reactions can produce a

relatively small pressure rise initially and after the accumulation of a greater amount of thermal

energy in the mixture, a second stage more violent ignition follows[10].

2.2 Influence of Steam on Flammability Limits

Despite the relevance of steam in syngas production (particularly with autothermal

reforming and non-catalytic partial oxidation technologies), there is a lack of detailed information

and experimental data on the effect of steam on the flammability of gaseous mixtures at high

pressures. Instead there is a general knowledge on the possibility of reducing the flammability

limits in the presence of steam [11,12,13,14].

In the process safety literature, the inerting effect of steam lowers the likelihood of

explosion [15]. The inerting effect is the process of adding an inert mixture, like N2, CO2, or

steam to reduce the concentration of oxygen below the limiting oxygen concentration (LOC)

where there is not enough oxygen required to propagate a flame [16]. Also, previous studies

indicated that the steam serves as both an energy sink for absorbing heat generated by the gaseous

chemical reactions and as a chemical reactant or suppressant that can affect the chemistry of the

mixture [7].

The energy balance between the chemical and physical effects is shown in Eq. (1). It

includes a pressure storage term on the left hand side; while the terms on the right hand side

represent the change in energy due to changes in total enthalpy and the heat loss term [17].

(1)

This equation assumes invisicid flow, neglects Dufour effect, assumes no external heat addition,

and no body forces. In summary, this equation can be interpreted to say that any rate of increase

6

in pressure is caused by the increase in total enthalpy of the system due to chemical reactions,

which overcome the heat losses from the system. The addition of steam to a flammable mixture

can affect both the enthalpy and the heat conduction terms in this equation.

Seiser and Seshardi (7) refer to the effect of steam on enthalpy change as the chemical

and the effect on heat loss as the physical effect on the flammability of a mixture. The addition of

water vapor to a premixed or non-premixed flame lowers the temperature in the reaction zone,

which increases the heat loss by conduction. The chemical effect of steam on the flammability of

the mixture is associated in the way steam interacts with the combustible mixture changing the

enthalpy of the system. Steam is unique as it has a higher chaperon efficiency than other reactants

and products found in hydrocarbon combustion. The chaperon efficiency is a parameter that

indicates the effectiveness of third body reactions (M). Steam addition can increase the

effectiveness of 3rd

body reactions, which can lead to new reaction paths.

The presence of steam in a combustible mixture greatly increases third body reactions

which can increase or decrease the flammability depending on the composition of the mixture [7].

In the case of mixtures with adiabatic flame temperatures around 2000 K, the addition of steam

makes the mixture more difficult to extinguish [7]. In the case of mixtures with adiabatic flame

temperatures around 1350 K, the addition of steam makes the mixture easier to extinguish [7].

This was observed in calculations and experiments by Seiser and Seshardi [7]. They posed that as

the adiabatic flame temperature is increased from 1350 K to 2000 K, the influence of the chain-

terminating reaction (R1) diminishes. Additionally in this range of increasing adiabatic

temperatures, the influence of the chain branching reaction (R2) increases.

(R1)

(R2)

If the terminating or branching reaction is more dominant, the flame would be harder or easier to

ignite, respectively. Additionally, it is known that the adiabatic flame temperature is strongly

7

dependent on the equivalence ratio of the mixture [17]. This means that the chemical effect on

flammability is strongly dependent on the equivalence ratio.

2.3 Flammability Limit Dependence on Equivalence Ratio

The flammability of the mixtures should be sensitive to small changes in equivalence

ratio ( ), which determines the adiabatic flame temperature of a given mixture. A NASA-CEA2

calculation [18] was performed for each mixture, without steam, at 1.01 MPa and the adiabatic

flame temperatures were calculated as 1357 and 2021 K for Mixtures 1 and 2, respectively. From

the large differences in calculated flame temperatures, it is anticipated that these two mixtures

have drastically different flammability limits.

Many studies have been conducted to determine an adiabatic flame temperature limit,

below which the mixture is not flammable. Chen et al. [15] summarized that the lower

flammability limit (LFL) and the upper flammability limit (UFL) are defined as the lean and rich

flammability limits of the mixture below or above which a flame cannot be sustained. For this

research, Mixture 1 and Mixture 2 are both fuel-rich mixtures, so investigation into UFL is of

interest. In the literature, there is a range of acceptable UFL adiabatic flame temperatures from

1000 to 1600 K. This indicates that Mixture 1 (Tf =1357 K) at 1.01 MPa would lie in this range of

temperatures implying that Mixture 1 is very close to the UFL without steam addition.

As shown by Seiser and Seshardi (7), the adiabatic flame temperature, which is directly

related to the equivalence ratio, determines the major chemical reaction paths taken in the

combustion process. If the equivalence ratio is closer to the stoichiometric value, the chain

branching reactions will be most effective during combustion. If the equivalence ratio is closer to

the UFL, the chain terminating reactions will be more dominant and slow the combustion process.

8

The literature shows that the equivalence ratio is a major determining factor in reaction paths of

the combustible mixture.

2.4 Influence of Pressure on Flammability Limits

The flammability limits for hydrocarbon/air have been shown to broaden at higher

pressures and reaction paths are also dependent on pressure [14,17]. The influence of steam on

flammability limits at higher pressures (P ~ 30 atm) is not available in the literature. For the

purpose of the present work, it is relevant to summarize some points concerning the role of steam

addition and initial chamber pressure on the chemistry of the hydrocarbon/oxygen mixtures. A

first point is made by observing that the hydrogen oxidation mechanism [17,19] is at the core of

hydrocarbon combustion chemistry. The driving chain-branching reaction is:

(R3)

The hydrocarbons in the mixture break down into hydrocarbon radicals, which then

produce a pool of hydrogen radicals through hydrogen abstraction. These hydrogen radicals

combine with the oxygen in the mixture and create hydroxyl radicals and oxygen atoms (R3),

thus branching the reaction paths and releasing thermal energy. The reaction, which is chain

terminating at pressures lower than 1.32 MPa (13 atm), along a 450 K isotherm in the classical

hydrogen-oxygen flammability limit [17], is the R1 reaction.

R1 competes with R3 for hydrogen radicals as pressure is increased, as the 3rd body

reactions (M) occur more frequently at higher pressures. The terminating reaction (R1) produces

the hydroperoxy radical (HO2). This is a heavy radical that can diffuse to the wall and will not

propagate the reaction at pressures less than 1.32 MPa and temperatures at the 450 K isotherm. At

pressures greater than 1.32 MPa, still along the 450 K isotherm, the hydroperoxy radical will no

9

longer diffuse to the wall, but can propagate reactions in the mixture through the following

overall chain propagating path:

(R4)

(R5)

The hydroperoxy radical combines with the hydrogen radicals to produce hydrogen

peroxide (R4) which easily decomposes to hydroxyl radicals (R5). R4 and R5 form an overall

chain propagating reaction causing the mixture to be flammable above 1.32 MPa along the 450 K

isotherm. In summary, the classical hydrogen-oxygen flammability limit, which is at the core of

hydrocarbon chemistry, shows that mixtures are non-flammable between 0.25 kPa and 1.32 MPa

along a 450 K isotherm and flammable at pressures higher than 1.32 MPa along the same

isotherm. These classical flammability limits dependency on pressure is due to the reaction path

selection of the HO2 molecule as pressure increases.

Shebko et al. [20,21] also shows that hydrocarbon combustion systems involve two

competing reaction paths that involve the reactions of HO2. Shebko [20,21] shows that once the

HO2 molecule is generated in R1, it can take one of the following paths:

(R6)

(R7)

(R8)

(R9)

Reactions R6 and R9 are both chain propagating, while R7 and R8 are both chain

terminating. From both the classical mechanism study and the work of Shebko [20,21] it is

known that the reaction path selection of HO2 is important in determining the flammability of a

hydrocarbon mixture.

The generation of the HO2 molecule is controlled by R1 which depends on 3rd

-body

reactions (M). These 3rd

-body reactions are more effective with increasing pressure. They are also

10

increased with increasing steam, due to its high chaperon efficiency as mentioned earlier. So,

pressure affects the reaction path selection of the HO2 molecule and the generation rate of HO2

molecules through more effective 3rd

-body reactions at higher pressures.

2.5 Influence of Reaction Zone Temperature on Flammability Limits

The amount of heat release and thermal energy losses for a given test can change the

induction time for ignition, as well as the temperature of the reaction zone. Additionally, the

presence of steam can lower the temperature of the reaction zone [7]. Westbrook et al. [9] and

Simmie [22] give summaries of hydrocarbon kinetic mechanisms at different reaction zone

temperatures.

Westbrook et al. [9] reports that at reaction-zone temperatures above 1200K, alkyl

radicals (R), like CH3, are produced by the hydrocarbon fuel due to beta decomposition. The

complex sequence of reactions in this high-temperature regime, is then initiated by R10:

(R10)

The major chain branching reactions in this regime are R3 and R11:

(R11)

Additionally in this temperature regime, an important reaction that retards ignition is R1. There

are also many other reactions that inhibit the chain branching by competing for H radicals.

At lower reaction-zone temperatures, less than 1000 K, the chain branching reaction, R3,

is quite slow. This reaction has a relatively high activation energy (70.3 kJ/mol), so it is not as

reactive at lower temperatures. Instead, the reaction R1 is most important in this temperature

regime as it has almost no temperature dependence. This reaction can initiate a branch by

colliding with RH in the following way:

(R12)

11

The H2O2 generated in R12 can then react by R5, depending on the pressure. This is the dominant

branching reaction mechanism at lower temperatures.

Simmie [22] gives a more recent, very detailed, review of chemical kinetic models. In

this review, the work of Petersen et al. [1] is of interest to this research. This research was for

CH4/O2/diluent (N2, He, Ar) mixtures at high pressures (4-26 MPa), high equivalence ratios (0.4-

6.0), and intermediate temperatures (1040-1500K) [1]. They show that at lower temperatures,

reactions involving acetaldehyde are important. They also show that the most important reactions

in determining ignition delay times at high pressures are the following:

(R13)

(-R5)

(R14)

It is important to note that all of these reactions depend on M.

At higher temperatures (1400K), they found that the dominant promoters of ignition are

the following reactions:

(R15)

(R16)

(R17)

(R18)

The most dominant inhibitor to these reactions is R14. The controlling radicals of ignition in this

fuel-rich, high pressure, high-temperature regime are the slow and inhibiting CH3 and HO2

radicals and the fast, chain-branching H and OH radicals. Pre-ignition build up of chain branching

H radicals is mainly due to the following reaction:

(R19)

12

At lower reaction temperatures (1100K) the most influential reactions change [1]. At

higher pressures and lower temperatures the following reaction occurs:

(R20)

which then produces the dominant chain branching path in the lower temperature regime:

(R21)

CH3O then branches in R19. Other chain branching reactions that are important here are R5 and

R22:

(R22)

At even lower temperatures (<1100 K), the fuel-rich data exhibited negative temperature

exponent behavior. The chemical kinetics in this regime was not explained by their work. The

addition of acetaldehyde chemistry, which was to account for the lower temperature regime had

insignificant effect. Understanding this regime requires more measurements of reaction rates. The

reaction rate of R21 needs to be experimentally determined as it is the most important reaction for

fuel-rich mixtures in the lower temperature, higher pressure regime.

Petersen’s work shows that as temperature is lowered, the reaction mechanism depends

more on the concentration of the 3rd

-body M for the branching reactions. At 1400 K, the reactions

R15-R18 do not involve M, since the temperature is high enough to overcome the activation

energy of these reaction paths. In both temperature regimes, the concentration of M plays a chain

terminating role in R14. In the lower temperature regime (~1100K), R5 and R19 are chain

branching reactions that also depend on M concentration. The propagation of a reaction in the low

temperature regime depends on the concentration of M to act as a catalyst to initiate reactions.

13

2.6 Cool Flame Phenomenon

A useful summary of the cool flame phenomenon was given by Fawcett and Wood [23];

certain important observations are summarized in this section. Autoignition and “cool flames” are

two different manifestations of the similar process of oxidation. A cool flame is a visual

phenomenon associated with the low-temperature oxidation of a hydrocarbon mixture in oxygen

or air. It is accompanied by a very small rise in temperature and pressure as compared to normal

ignitions. It is often referred to as a partial oxidation reaction. The cool flame can be seen with the

proper setup, as it emits a pale blue luminescence caused by the formation of excited

formaldehyde (CH2O) molecules. The formaldehyde molecules result from the decomposition of

hydroperoxides (H2O2 => 2OH) combining with the CH from aldhydes (O=CH-) which have a

weak oxygen bond. This chemical reaction is summarized as OH + CH => CH2O => pale blue

luminescence.

The cool flame phenomenon has explained the causes of many unexpected fires and

explosions. In high-speed jet fighter planes, the leading edge of fuel tanks on aircraft wing

becomes quite hot at Mach 2. This aerodynamic heating can initiate a cool flame that can

transition to an explosion as the jet plane lowers its altitude, which raises the pressure of the tank.

The cool flame phenomenon extends the upper flammability of hydrocarbon mixtures at

higher pressures. As seen in Figure 1, for a pressure of 5 atm, the upper flammability limit of the

hexane/air mixture is extended from 13% to 28% in terms of vol. % of hexane in air to the cool

flame boundary marked by shaded curve.

14

Figure 1: Effect of temperature and pressure on the flammability of hexane in air [23].

The cool flame phenomenon can exist in higher pressures and rich hydrocarbon mixtures. It is

characterized by very small changes in temperature and pressure. This phenomenon is usually not

visible in normal lighting condition, but can be seen with the specialized optical setup.

Warnatz et al. [24] describes the ignition limits for hydrocarbons with a p-T explosion

diagram. This diagram, shown in Figure 2, is very similar to the classical H2-O2 p-T explosion

diagram except that it includes the cool flame regime that is only found in hydrocarbon mixtures.

15

Figure 2: Ignition limits for hydrocarbons [24]

Multistage or multiple ignition is a phenomenon found in the 3rd

explosion limit where ignition

takes place after the emission of short visible light pulses. The cool flame regime is where

combustion takes place at low temperatures.

In these phenomena at the 3rd

explosion limit, the reaction R20 controls the chain

reactions that govern the ignition process. If the temperatures are high enough, the CH3O2

molecule will decompose and terminate reaction paths associated with this molecule. If the

temperatures are low enough the following degenerate chain branching path can be taken, i.e.:

CH3O2 + CH4 => CH3OOH + CH3 then CH3OOH => CH3O + OH. A degenerate path is one that

fails to branch at higher temperatures.

16

From reviewing the topic of cool flames, it was found that in hydrocarbon mixtures odd

flame spreading can be observed at higher pressures, lower temperatures, and fuel-rich mixtures.

The degenerate chain branching path, described above, that can only function at lower

temperatures, causes odd flame spreading. The OH produced in this chain reaction can then

combine with a CH radical to generate formaldehyde, which emits pale blue luminescence. Also,

multiple pulses of light are indicative that this same chemical mechanism is involved.

2.7 Flammability Experiments

Bartknecht [25] highlights important parameters of explosion testing, which is directly

related to the flammability testing in this research. A change in pressure of the test chamber can

occur if a flammable mixture ignites. The rate of pressure change and maximum pressure

observed characterizes the amount of energy released from the exothermic reactions. The volume

of the reactor governs the rate of pressurization as well. Lower rates of pressure change (dP/dt)

are usually observed as the mixture’s composition approaches its flammability limits. The

flammability limits of a mixture can be broadened with greater energy release from the igniter

[25]. Also, a richer mixture usually has a higher minimum ignition energy required than a

stoichiometric mixture [26].

Information for these experiments performed in a tube reactor was found in reference to

explosions in pipelines [25]. In pipelines carrying potentially flammable mixtures, protective

measures should be taken to confine an explosion to a limited volume. This can be achieved by

using gaps or clearances, called flame barriers, for flames to expand into and then be quenched.

Two processes are important in causing a flame to quench as it is being ejected from a

pressurized enclosure into a flame barrier:

1. heat loss to the walls; and

17

2. reduction of temperature due to expansion of the hot combustion gases leaving the

enclosure.

The dimensions of the flame barriers are highly influenced by the concentration of the mixture. A

less flammable mixture could be quenched by a flame barrier that could not quench a more

flammable mixture. In addition, the velocity of the product gases and the residence time during

which the combustion gases are in contact with the reactant mixture affects the possibility of

onset of ignition. As contact time between the reactants and hot products increases, the mixture is

more flammable. This contact time decreases with increasing product gas velocity [25].

2.8 Summary of Processes Governing Flammability Limits

From this literature review, flammability of a mixture is governed by the rate of radical

growth. The accumulation of radicals occurs in the chemical induction time, , and the radical

concentration can either increase or decrease during the thermal induction time, .

Steam can affect the rate of radical accumulation in two main ways. Firstly, steam, in

high enough quantity, can act as an inerting agent, which lowers the available oxygen content of

the system. This reduces the productivity of chemical reactions as well as the amount of radicals

generated, and hence lowers the reaction-zone temperature. Secondly, steam has a high chaperon

efficiency, this makes 3rd

-body reactions much more effective and increase the rate of both chain-

branching and chain-terminating reactions involving the 3rd

body M.

The equivalence ratio of the mixture governs the rates of chemical reactions in generating

radicals. The closer the equivalence ratio is to the stoichiometric value, the higher the production

rate of radicals. Conversely, if the equivalence ratio is close to the UFL, the production rate of

radicals is low.

18

Initial chamber pressure governs the effect of 3rd

body reactions (M). The higher the

pressure, the faster the reaction rates of paths that involve M. In the classical H2/O2 reaction

mechanism, the HO2 and H2O2 molecules become much more reactive under higher pressures, as

their collision rates are increased. 3rd

body reactions are very important in this study as their rates

are greatly enhanced with increased initial pressure and steam content.

The chemical kinetic reaction mechanisms that lead to ignition are very sensitive to the

temperature of the reaction zone. The work of Simmie [22], Westbrook et al. [9], and Petersen

[1], showed that small changes in reaction-zone temperature can drastically change the chemical

kinetic mechanism. It was determined that for fuel-rich mixtures at high pressure, as the

temperature is decreased from 1400 to 1100K, the reaction mechanism transitions from M

playing a chain-terminating role to M playing a chain-branching role. As temperature is

decreased, reactions need the catalytic energy of the 3rd

-body M to initiate most reactions.

Conversely, at higher temperatures, 3rd

body reactions are not required for chemical reaction to

occur since the required activation energy is available due to the higher temperature.

Additionally, fuel-rich, high-pressure, low-temperature reactions can involve

acetaldehyde, which produces formaldehyde. This product emits a pale blue luminescence and is

the signature of the cool flame regime. Ignitions that occur in the cool flame regime are beyond

the upper flammability limit.

Ignition experiments emphasize the importance of total induction time ( ) in

flammability tests. Reactions that occur slowly are more likely to be near the flammability limits.

Also, the amount of time hot products are in contact with the unburnt reactants increases the

production of radicals in the chamber thus increasing the probability for onset of ignition.

Overall, the flammability of a mixture depends on the rate at which radicals accumulate.

This can be increased or decreased by steam depending on the vol.% of steam. The rate of

accumulation decreases with increasing equivalence ratio, decreasing pressure, and decreasing

19

temperature. There is also interdependency between these three variables. The main conclusion

from this literature review is that flammability limits can be viewed as the condition in which the

highly dynamic process of radical production by chemical reactions is just balanced by the heat

loss effect. These governing processes are summarized in Figure 3:

Figure 3: Schematic of chemical and physical processes governing flammability test

2.9 Chemical Process Safety Characterization of Flammability Limits

For characterizing the flammability limits of Mixture 1 and 2, a method found in the

chemical process safety literature can be used. This method uses calculations with experimental

data compiled from fuel vapor-air mixture explosion tests. Fuel vapor-air explosion tests are

executed by taking one type of fuel vapor and mixing it with air in a closed vessel and igniting it.

The amount of fuel is raised or lowered in these tests until the UFL and LFL are determined.

From Crowl and Louvar [16], flammability data is compiled from several sources. In their book,

flammability data for each of the constituents in the mixture are given. These data can be

combined using Le Chatelier’s equation, given as Eqs. (2) and (3), to calculate the values of UFL

and LFL of the mixture.

(2)

20

(3)

where Yi is the mass fraction of the ith species. In addition, correction factors can be added to

these limits that account for effects of pressure and temperature, as shown in Eqs. (4), (5), and

(6).

(4)

(5)

(6)

where T is in degrees Celsius and P is given in absolute scale of MPa. The LOC may also be

estimated by multiplying the LFL of the mixture by the stoichiometric coefficient. The values of

LFL, UFL, and LOC have been used in a “flammability triangle diagram” [16] to determine the

flammability of any mixture composed of nitrogen, oxygen, and combustibles. Figure 4 shows the

flammability triangle diagram for a mixture of methane vapor and air.

21

Figure 4: Flammability Triangle Diagram for a mixture of methane vapor in air [16]

The point A in Figure 4 is a non-flammable point as it lies outside the flammability

region bounded by the two dashed lines. These dashed lines are constructed from the connection

of two points. The UFL and LFL on the “Air Line” are connected to the intersection of the

stoichiometric line and the LOC line.

The closed vessel ignition experiments performed to gain this empirical data is quite

different from the flowing experimental test conditions of this study. The induction time in the

present study should be much longer as some intermediate combustion products are being

transported out of the reactor. For this reason any characterization of flammability limits with the

chemical process safety method will have wider flammability limits than what will be observed in

this study. Therefore, this method of characterization of flammability limits cannot be adopted,

22

unless a correlation between the induction times of the fuel-vapor/air experiments could relate the

test results from the present study.

2.10 Flammability Modeling

The following references were useful in determining the best way for utilizing the well-

established Chemkin code to define flammability boundaries. To achieve this goal it was

necessary to select the criteria for defining flammability limits. For this, the work of Law and

Egolfopoulos was referenced [27]. They define a flammability limit to exist when the major chain

branching reaction has a rate of production of radical species, which is equal to that of the major

chain terminating reaction. They refer to this condition as the “turning point” in which there is a

balance between the two reactions causing the energy generation rate to sharply decrease after

this point.

Westbrook [28] showed that a qualitative rule can be defined. Based upon experimental

observations, a premixed hydrocarbon/air flame with one-dimensional adiabatic laminar flame

speed lower than 50 mm/s can be considered non-flammable [28]. Womeldorf and Grosshandler

[29] used this rule, with Chemkin code, in their study of flammability limits of CH2F2 / air

mixtures.

Bui-Pham and Miller [30] studied the flammability limits of rich methane/air mixtures by

using the flammability limit definition of both Law and Egolfopoulos [27] and Westbrook [28].

They found very good agreement in the upper flammability limit between the two methods. When

using Law and Egolfopoulos’ method [27] they defined the chain branching reaction and chain

terminating reaction as the following:

(R3)

(R23)

23

Their work was very useful to determine the definition of the flammability limit in the Chemkin

code calculations performed in this study.

Liang and Zeng [11] used the Chemkin batch reactor coupled with GRI-Mech 3.0 to

show that water addition to a gas explosion would increase the time to induce an explosion. They

used a methane and air mixture and showed that 10% water addition decreased the explosion

pressure from 2.15 atm for the dry mixture to 0.15 atm with 10% water. From a sensitivity study,

they found that the dilution effect (or physical effect) water addition to reduce the reaction zone

temperature is the cause of the observed decrease in explosion pressure.

24

Chapter 3

Method of Approach

The influence of steam on the flammability of Premixed Natural Gas/Oxygen/Steam

mixtures was investigated experimentally for Mixture 2 and computationally for both Mixture 1

and Mixture 2.

3.1 Experimental Method of Approach

In this study, a tube reactor with a 40-mm diameter bore was used to study ignition, flame

propagation, and explosion characteristics of the two simulated NG mixtures as inputs to the

SCT-CPO reactor. The composition of NG delivered to the pilot plant can vary [31]; thus,

multiple compositions were investigated in the present study. Variation in the initial chamber

pressure and gaseous flow rates were also studied.

The reactor (shown in Figure 5) is equipped with multiple ports housing fast-response

photodetectors and dynamic pressure gauges to verify the onset of ignition and to measure flame

propagation rate. The initial pressure was measured with static diaphragm pressure gauges. The

test rig is also instrumented with multiple K-type thermocouples to verify the temperatures of all

the reactants individually, before mixing. The temperature of the gaseous mixture was measured

at several locations, including: the mixer section, the top, and bottom of the reactor. The reactor

was a stainless steel, thick-walled cylinder and the reactor portion of the chamber was lined with

a Pyrex glass tube to ensure a chemically inert test environment. Through two slit windows, the

flame propagation process was viewed and recorded with a high-speed camera. The process flow

diagram of the experimental setup is shown in Figure 6. The components of the mixture are

preheated through the steam generator heat exchanger and pre-heater to elevated temperatures

25

around 450 K. The gaseous flow rates were set by multiple flow meters and controlled by

computer actuated valves. After steady flow rates were established, an electric match was

triggered at the bottom of the test rig. The multi-channel data acquisition system was

synchronized with the ignition switch to activate and acquire signals from all instrumentation.

Figure 7 and Figure 8 show the experimental setup in more detail.

Figure 5: A 135o-sectional view of the windowed high-pressure tube reactor

26

Figure 6: Process flow diagram for flammability study of hydrocarbon/steam/oxygen mixtures

Figure 7: Liquid petroleum gas (LPG) reservoirs and stock tanks, and methane and hydrogen

manifolds

27

Figure 8: Test reactor experimental setup with flow controls, pre-heater, and steam generator

From a total of 76 tests performed, three different regimes of pressurization rate (rapid,

intermediate, and slow) were observed. They were defined according to the time required to reach

the maximum pressure recorded by the dynamic pressure transducers. When a mixture was

ignited, the pressure in the reactor rose significantly above the initial chamber pressure. Figure 9

shows three pressure-time (P-t) traces from Test No. 53, 54, and 55. These traces were recorded

by the dynamic pressure transducer at the bottom of the reactor, near the igniter.

For all tests with flammable mixtures, the time to peak pressure from the initiation time

was measured. The average time duration was determined to be = 0.051 s. This value was

used to normalize the time to peak pressure ( of the ith test. A dimensionless time (τi) was

defined as . If τ < 0.5, it is considered a rapid rate. If 0.5 <τ < 1, it is considered an

intermediate rate. If τ > 1, it is considered a slow rate.

28

Figure 9: Three representative P-t traces recorded from the tube reactor

Figure 10 shows the turbulent (ReD=20,000 – 40,000) premixed flame front propagating

up the reactor tube.

Figure 10: High-speed camera visualization of turbulent premixed flame front [35]

The flame spreading rate was also measured with 12 photodetectors (PD), spaced at 16.1 mm

between adjacent detectors. PD1 is located at the top and PD12 at the bottom of the reactor. An

example of the results is shown in Figure 11. Figure 12 shows the turbulent flame speed deduced

from the data in Figure 11.

29

Figure 11: Representative photodetector intensity-time traces [35]

Figure 12: Representative linear fit of flame-propagation speed [35]

Knowing the spacing between adjacent photodetectors, the relative turbulent flame speed

of the ignited mixture was determined from the slope of the line in Figure 12. The flow velocity

of the reactant mixture from the top region is added to the relative turbulent flame speed to

determine the absolute turbulent flame speed. The absolute turbulent flame speed from the reactor

bottom (ST1,abs) was determined to be 29.9 m/s, during the initial time interval. When the flame

approached the mixer exit, the flame speed decreased to 9.1 m/s. This was mainly due to the

entrance and jet mixing effect of the reactants passing through the multi-perforated discharge

plate at the inlet portion of the reactor. In this region, there is also a cooling effect due to the

30

presence of the discharge plate. This region is usually small compared to the overall length of the

reactor.

3.2 Computational Method of Approach

The homogenous batch reactor and the premix code in the Chemkin package [8] were

utilized to perform a heuristic study of the flammability trends. The chemical kinetics used in the

calculations was based upon a high-pressure butane chemical kinetic mechanism [32] with

thermodynamic data and transport properties corresponding to these high pressures. The kinetic

modeling [32] was compared with experimental data taken at =0.3 to 2.0, T = 1056 to 1598 K,

and pressure = 1 to 21 atm. Moreover, it is noted that the reference kinetic mechanism was

designed for use in power generation applications utilizing NG fuels. For this reason it was

selected with an awareness that the present experimental work was performed at >2.5 and initial

temperature of 450 K. Indeed, Donato et al. [32], found good agreement between model and

experiments at pressures as high as 30 atm and in the temperature range from 650 to 1400 K with

a stoichiometric mixture of n-Butane, Iso-Butane, and air. Their results diverge slightly from

experiments at T= 1075 K, =2.0 and P = 7.8 atm and 18.1 atm.

It is recognized that in the present study, the initial temperature of the reactant mixture is

lower than that of the test conditions conducted by Donato et al. [32] Also the equivalence ratios

of the mixtures in this study are more rich than theirs. Nevertheless, the adopted kinetic model

from Donato et al. [32] represents one of the most suitable hydrocarbon combustion mechanisms

for mixtures containing butane for elevated pressures. It is also recognized that the

comprehensive simulation of the exact experimental test condition is beyond the scope of this

investigation. The main reason for using Chemkin code is to study the trends of steam addition

effect on the flammability of the combustible mixture. Chemkin code can also be used to

31

perform sensitivity studies, which were beneficial in the investigation of reaction mechanisms

that could suppress flammability.

32

Chapter 4

Experimentation

4.1 Instrumentation

The data obtained from experimentation was organized into pressure, photodetector, and

flow rate data. Other data, such as temperature, was recorded immediately before triggering the

system with omega K-type thermocouples at various locations, given in Appendix A.

4.1.1 Pressurization Data

Static and dynamic pressure transducers recorded the pressure data. The following figure

shows pressure time traces from these experiments included here for discussion.

Figure 13: Recorded pressure traces of Test #55

33

The output from setra model # 206 static pressure transducers contained a 60-Hz noise as

in Test # 55, Figure 13. These oscillations in the static gauges could be smoothed graphically, but

as they were used mainly for initial chamber pressure readings corrections were performed by

averaging over 0.1 seconds before ignition. The dynamic pressure gauge p-t traces show a period

of slow rise followed by a flame propagation interval, and then by a sharp rise in pressure. At a

time of 0 seconds the igniter is triggered, the chemical reactions occur in a time scale on the order

of . The initial chemical reactions, that occur in , build a pool of radicals. The thermal

induction period, , then increases the growth rate of this pool until an ignition occurs. As

mentioned in the literature review, induction time is a key parameter in understanding any

ignition observed. Also, the rate of pressure rise, dP/dt, after the induction period indicates the

extent of the heat release of the reactions.

Five types of dynamic P-t traces were observed during the experiments. As mentioned

earlier, some occurred in a rapid, intermediate, or slow rate of pressurization. Figure 13 is an

example of a slow rate of pressurization. Figure 14 is an example of a rapid rate of pressurization.

Figure 14: Recorded dynamic pressure traces of Test #53

34

Notice that Figure 13 has an induction time about 5 times longer than that of Figure 14. The two

mixtures have close to the same equivalence ratio and are at similar initial pressures. The main

difference is that Test # 55 had 13% more steam in the mixture. The higher amount of steam in

the mixture lowers the amount of oxygen available and therefore increases the induction time and

lowers the maximum rate of pressurization.

The “slow” rate of pressurization can be further divided as some of the these traces did

not have a period of rapid pressurization, but only an induction period after which no rapid

pressurization occur. In these P-t traces, the maximum pressure obtained was very low. Figure 15

shows this slower rate of pressurization in an expanded pressure scale from Test #73, which had

60% Vol. of steam.

Figure 15: Recorded dynamic pressure traces of Test #73

The fifth type of P-t trace that was observed is the multiple ignition P-t trace. It did not

show the typical rate of pressurization characteristics. Figure 16 shows the multiple ignition

phenomena of the ignition of a mixture that contained 41% steam by Vol. The reason for the

35

multiple pressure spikes is due to the fact that this test condition is near the flammability

boundary.

Figure 16: Recorded dynamic pressure traces of Test #56

The dynamic pressure transducers were very useful in characterizing the flammability of

Mixture 2 with increasing steam content. As steam is increased the mixtures become less

flammable. This is due to two different effects the steam has on the mixture. The steam dilutes

the amount of oxygen available for reaction and decreases the concentration of the initial radical

pool; this led to longer induction times. Also, after a critical amount of steam is added, the

reaction zone temperature was reduced leading to a different reaction mechanism governing the

radical pool growth.

Transitions between different reaction mechanisms can be seen in the above P-t traces. It

is suggested that typical rapid, intermediate, and slow rates of pressurization are governed by the

same reaction mechanism, like the higher temperature mechanism of Petersen et al. [1]. The

slower pressurizations and the multiple ignition p-t traces are controlled by a different reaction

36

mechanism, like the intermediate and low-temperature mechanisms described by Petersen et al.

[1]

4.1.2 Photodetector Flame Spreading Data

The following figure represents a typical set of recorded photodetector intensity-time

traces from the flammability experiments using Perkin-Elmer photodiodes (Model #VTP3310LA)

with a detectable wavelength range 400-1150 nm.

Figure 17: Photodetector intensity-time traces of Test #43

Figure 17 shows the flame spreading up the reactor tube, based upon the advancement of the first

discernible signal above the noise level from the photodetectors closer to the igniter to the reactor

top. The voltage signals first rise at PD 12 as the flame passes that point. This flame spreading

continues up the tube as can be seen by the voltage signals rising at PD 11 then PD 10 and so on

37

until reaching PD 6. Some of the photodetector traces did not show abrupt rise due to the slow or

marginal flame spreading process. Figure 18 is an example of a test that showed unusual flame

spreading.

Figure 18: Photodetector intensity-time traces of Test #55

Figure 18 shows flame spreading down the tube, as opposed to up the tube, which was the typical

direction in most of the tests. This test had one of the highest Re numbers of all the experiments

conducted. The reverse flame spreading phenomenon was likely due to the transition between

reaction mechanisms. Regardless of the cause, unusual flame spreading, or no flame spreading,

was observed in all tests with a steam mixture mole fraction above 24%. Another type of unusual

flame spreading seen in the experimentation was found in Test #53.

38

Figure 19: Photodetector intensity-time traces of Test #53

Figure 19 shows that PD 11 receives a signal before PD 12 and then the flame spreading travels

up the tube in an unsteady manner. At about 0.008 seconds, there is a spike in all signals which is

indicative of noise interference. Figure 19 shows that there is a good deal of unsteadiness in the

flame spreading. This same unsteadiness is seen in Test # 54, 56, and 58. For Test # 59-76 any

flame spreading seen is small in magnitude, occurs over a long time, and does not register in all

PDs. Test # 39-52 show typical flame spreading up the reactor as in Figure 17. The higher test

numbers had higher mole fraction of steam. It is suggested that the steam damps the flame

spreading in Test # 59-76 and causes unsteadiness and irregularities in Test # 53-58. This resulted

in accurate flame speed measurements only for Test # 39-52. It is proposed that this observed

transition in PD data represents the transition between chemical reaction mechanisms.

A typical flame speed profile for normal flame spreading, from the bottom of the reactor

to the top, can be seen in Test# 43, shown in Figure 20.

39

Figure 20: Linear fit of flame-propagation speed for Test #43

The time from ignition was determined by the PD voltage rise above a threshold value. Figure 20

shows the flame spreading up the tube occurs at a relative speed of 82.59 m/s which corresponds

to an absolute flame propagation speed of 84.72 m/s for Test # 43 with the reactant supply rate of

2.13 m/s. Note that the PDs at the top of the reactor did not register a high enough signal and the

deceleration process to a lower flame propagation rate near the injector plate was not observed.

4.1.3 Flow Rate Data

Mixture 2 experiments required the flow of oxygen, methane, ethane, steam, propane,

and butane. Figure 21 represents typical flow rate profiles of these experiments.

40

Figure 21: Time variations of flow rates of reactants of Test #43

Figure 21 shows that prior to ignition at time t= 0.0 seconds, the flow rates are constant. The

steam, propane, and butane all showed small degree of oscillations before triggering the igniter at

t = 0. To account for oscillations in any of the flow rates, the average flow rate over 0.1 seconds

prior to ignition was used as the flow rate of each component for data analysis. The propane and

butane flows had to travel much further from the metering point to the reactor than all other

flows. To account for this, the propane and butane flows were averaged over 1.0 second prior to

triggering. Figure 21 also shows a rapid flow drop phenomenon observed in many of the

experiments. This rapid drop is due to rapid chamber pressurization due to ignition. This shows

that the pressure wave created in the chamber at ignition can interact with the oxygen and

methane flow meters. This interference is not observed in the other supply lines their metering

point was further upstream.

41

4.1.4 Ignition Source

One major point of interest that was not fully investigated in these experiments was the

effect of the igniter location on flammability of the mixture. In Mixture 1 testing, the igniter (a J-

tek electric match with its head covered with a thin layer of RTV as protective coating) was

inserted from the bottom plate, vertically into the lower portion of the reactor. In Mixture 2 test

series, the igniter was inserted from the side of the bottom plate, horizontally into the reactor. The

head of the electric match was recessed into the bottom plate and also covered by a protective

layer of RTV, so that the hot steam would not heat the electric match to initiate ignition. It is

thought that the configuration in Mixture 2 would create a more even plane of ignition. While,

both igniters release the same amount of minimum energy (4 mJ minimum energy release for J-

tek electric match) it was not shown that the two configurations yielded the same results. In fact,

two of the tests performed for Mixture 1 with the horizontal igniter configuration resulted in a

flammable mixture observation, which was seen as nonflammable with the vertical igniter

configuration. This difference could be due to the hot particles, generated from the electric match,

interacting with a large volume of the relatively cooler mixture in the reactor. For the horizontal

igniter configuration, the hot particles are interacting with a smaller amount of gaseous mixture.

Determining which igniter orientation is more effective at igniting the mixture would be a useful

study for future research. As the literature [23] shows, the flammability limits can be expanded

with greater ignition energy. Ideally, the experiments would incorporate a planar sheet of ignition

energy that would release at least 1 joule of energy [23], which is much higher than the ignition

energy used in this study.

42

4.2 Summary of Uncertainties

Since the flammability experiments involves multiple control variables, the dependency

of the test data accuracy on the initial condition is of interest. Potential errors in the flow rate

measurements of different gaseous components are presented in Table 2. The uncertainties in the

initial conditions of the experiments are due to the instrumentation errors of pressure gauges and

flow meters. The inherent uncertainties of the static diaphragm pressure gauges are ±0.13% of

full scale. The inherent uncertainties of the flow meters vary for different groups of tests, as

shown in Table 2. See appendix E for detailed error analysis.

Table 2: Percent uncertainty in various components of the mixture for different tests

Test 1-10 11-38 39-48 49-56 57-59 60-76

O2 1.67% 1.67% 1.58% 1.58% 3.20% 6.30%

CH4 1.12 % 1.12 % 3.90 % 3.90 % 8.00% 16.0%

C2H6 1.20 % 1.20 % 5.51% 5.51 % 11.0% 20.0%

H2 1.20 % 1.20% - - - -

C3H8 - - 1.00 % 1.00 % 1.00 % 1.00 %

C4H10 - - 1.00 % 1.00 % 1.00 % 1.00 %

H2O 1.60 % 2.70 % 2.70 % 1.60 % 1.60 % 1.60 %

0.07 % 0.07 % 0.25 % 0.25 % 1.05 % 3.87 %

Mixture 1 (shown in Table 1) was used in Test # 1-38. In these tests, the orifice of steam

flow meter was changed, beginning with Test # 11, to allow for less steam flow. This change

increased the uncertainty of the steam flow measurement. Mixture 2 was used in Test # 39-76.

During these tests, the steam flow meter orifice was changed back to its original configuration to

allow for more steam flow. At Test # 56, the steam generator reached its maximum steam

production limit, but the Mixture 2 experiments had not exhibited a non-flammable data point. To

reach a non-flammable condition, all of the Mixture 2 flows, except steam, were reduced by half.

After Test # 59, the Mixture 2 flows were again reduced by half. In Test # 60-76, the majority of

the non-ignitions were observed. As the Mixture 2 flows were reduced, the error increased as the

flow meters were designed to flow, more accurately, at higher rates. The flow rate of the ethane

43

and methane was reduced to a level that lies in the lower limit of the flow meter’s control

capability. This resulted in relatively high error of the methane and ethane flow rates in Test # 57-

76. The error in the measurement of ethane has a small effect on the flammability of the mixture

as it constitutes a very small fraction of the mixture. This resulted in a maximum error on the

equivalence ratio of the mixture, which was found to be 3.87% in Test # 60-76.

Other error found in the experiments is due to human error in controlling multiple

variables. As seen in Appendix D, an efficient checklist for controlling these experiments was

developed. Following this checklist ensured success by overcoming difficulties in the control of

the Liquid Petroleum Gas (LPG) system, eliminating potential auto-ignitions, and maintaining

run conditions at initiation. Accurate control of the LPG system was one of the most challenging

issues in this research. The saturation properties of Propane and Butane complicate the

vaporization of the liquid fuel.

Figure 22: Vaporization curves for propane and butane

As seen in Figure 22, at an ambient temperature of -0.5oC the pressure of liquid butane and

propane is at 1 and 3.8 atm, respectively. The vaporization pressure of the fuels governed how the

44

run tanks, seen in Figure 7, were filled. If the vaporization pressure is too low, around 1 atm, the

piston in the run tank cannot be compressed by the fuel vapor. Therefore, the run tank cannot be

filled by the liquid supply tank. If the vapor pressure is too high (e.g., warm propane during the

summer time), the run tank would contain mostly the vapor rather than liquid. This situation

causes the run time to be shorter than required test duration. To solve these problems, in relatively

cold ambient conditions (<10 oC ), the butane tank was wrapped in a hot water heat exchanger

and the piston in the run tank was lowered by pressurized N2 to allow the tank to be completely

filled. In warmer conditions (>20 oC), the propane piston was forced to the top of the run tank

with pressurized N2. As the propane was filling the run tank, the nitrogen vent was slightly

cracked open allowing the propane to fill with more liquid than vapor. Both of these methods

were incorporated in all test preparation to ensure proper filling of the run tanks.

In addition to this issue with the LPG system, the LPG control valves have integrated

controls that interfere with the Proportional Integral Derivative (PID) Labview controller, causing

oscillations in the flows. The PID controller was then tuned to damp out the oscillations within 1

minute of flow initiation.

In the first few tests, the oxygen flow rate was set to the specified value in just one step.

In one of the earlier tests, a rapid increase in oxygen flow caused an autoignition. This resulted in

changing the procedure to increasing the oxygen flow in two steps and also flowing nitrogen

while the oxygen flow was being initiated. This procedure prevented the occurrence of

autoignition in any subsequent tests.

Another major issue with these experiments was the difficulty in controlling the steam

flow rate. During Mixture 2 testing, it was found that the required steam flow rates were

significantly higher than those necessary to prevent ignition in the Mixture 1 study. In the

Mixture 1 configuration, the steam flow control system was capable of accurately flowing and

measuring up to ~1.4 Nm3/hr. By moving the Rosemount fixed orifice flow meter to a location

45

upstream of the control valve, the higher-density state of the steam allowed the flow to be

accurately measured up to values of approximately 40 Nm3/hr. The work performed in changing

the maximum flow rate of the steam flow meter was assisted by Mr. Robert Hutchinson of the

Office of the Physical Plant at Penn State University. His familiarity with the Rosemount steam

flow meters was very helpful and he allowed us to borrow his laptop which contained software

that enlightened the team on the odd problems that were occurring with the steam flow meter. It

was found that if the pressure drop across the orifice was higher than a critical value, the steam

flow rate could jump to the full value, which is not the actual metered value. This problem was

solved by using adequate orifice size to limit the magnitude of the pressure drop.

4.3 Safety Analysis

From a safety operation point-of-view, a study was conducted to predict the maximum

chamber pressure if the reactant mixture is detonated upon the initiation of the igniter. In this

study, the baseline gas mixtures were given in Table 1 in Chapter 1. NASA’s Chemical

Equilibrium with Applications (CEA2) Code was used to predict the theoretical maximum

chamber pressures that could be reached if all reactants burned instantaneously to equilibrium

products. The constant-volume (uv) option was used to model the system. Results of these

calculations for the two different baseline gas mixtures containing different proportions of

propane and butane are shown in

Figure 23. In this figure the red line corresponds to Mixture 2 and the blue line corresponds to

another proposed mixture, which has a greater volumetric supply rate of butane than propane in

the Mixture 2. As shown in this figure, if there is no heat loss or mass loss, the reactor pressure

could reach very high levels (P ~400 atm) for an initial pressure of 30 atm. Real-world maximum

pressures will always be lower due to the realistic conditions of finite-rate chemistry, non-

46

equilibrium products, heat loss, and the vented steady-flow chamber condition. However, an

efficiency factor can be developed at lower initial pressures based on measured pressure jumps

compared to the calculated values. For example, the Test# 42 with initial pressure of 9.4 atm

resulted in peak pressure of about 41 atm. This efficiency factor is about 40%. However, earlier

tests with less steam resulted in a peak pressure of over 54 atm, yielding an efficiency of over

50%. The efficiency factor is expected to increase with the initial chamber pressure, as the

reaction rates become faster at higher pressures. Conservatively applied, this factor can then be

used to estimate a safe upper operating limit for the system. By controlling the initial chamber

pressure to be below 18 atm, the maximum chamber pressure will be safely below 100 atm,

which corresponds to maximum operating pressure of the test facility by using the dry mixture.

With the addition of steam higher initial pressures could be utilized since the steam significantly

lowered the efficiency factor. The highest initial chamber pressure tested for Mixture 2 was 21

atm, while Mixture 1 tests had a maximum initial pressure of 30 atm, since it is less energetic

than Mixture 2.

47

Figure 23: Theoretically calculated maximum pressures for equilibrium reaction of

NG+LPG/steam/O2 as a function of initial chamber pressure with two different fuel ratios of LPG

48

Chapter 5

Chemkin Code Calculations

5.1 Overview of Calculations

Two separate methods of calculation were attempted with Chemkin Code to show general

flammability trends of the mixtures. The premix code in the Chemkin package was used to look

at the chemical reactions that occur in the experiment and how they could affect flammability of

the mixture. The homogenous batch reactor model was used to study flammability trends and also

give a simple method for performing sensitivity studies.

5.2 Premixed Laminar Flame Speed Calculations

The following is a conceptual diagram illustrating the steps and options involved in the

development of the premixed laminar flame speed reactor model in Chemkin:

49

Figure 24: Conceptual diagram of the Chemkin premixed laminar flame speed model

Once the model was specified, a temperature, axial velocity, and species solution to the

conservation equations was solved. The initial guess of the temperature profile and flame

thickness was very important to obtain a converged solution. The profile guess has to be close to

the actual solution for convergence. The Chemkin premix code was used to find the point in

which the steam content of the mixture was high enough to reduce the axial velocity (laminar

flame speed) to below 50 mm/s in order to suppress ignition.

As seen in Figure 24, Chemkin gives several viable options for the premixed laminar

flame speed model. To determine the solution options and grid criteria that provide a reliable

solution, the following studies were conducted. These studies were performed at a pressure of 10

atm, which was the average pressure used in the reaction mechanism. A grid convergence study

was conducted on the sensitivity of the flame-front temperature and axial velocity to the number

Flame Speed Calculation Model

Solution Model Options

•Use Intermediate Fixed Temperature solution

•Use Mixture average or Multicomponent Diffusion

•Use Correction Velocity Formalism or Trace Species Approximation

Reactor Physical Properties

• Initial Temperature

• Initial Pressure

•Guess Temperature Profile

Initial Grid Properties

•Grid Points

•Gradient and Curvature

• Flame Thickness Descriptioin

Species Specific Properies

• Product Species (Calculated from CEA)

• Intermediate Species Assumptions

Inlet Description

Inlet Velocity

Inlet Reactants

Solver

Specify Solution

Convergence Criteria

Windward or Central

Differenceing

Output Control

Option used for Sensitivity

Studies

50

of initial grid points, solution curvature (on the temperature and velocity profile), and solution

gradient.

Figure 25: Solution sensitivity to initial number of grids

51

Figure 26: Solution sensitivity to value of curvature and gradient

From these results it was determined that the number of initial grid points to be used was

20 and the gradient and curvature was set to a value of 0.25.

The solution to the problem can be found using mixture averaged transport or multi-

component diffusion. Mixture average transport simplifies the molecular transport problem

compared to considering the multi-component diffusion, where the transport of each species is

considered. The following figure compares the solution of the two different transport methods.

52

Figure 27: Comparison of multi-component diffusion and mixture average transport

It is clear from Figure 27, that for this study, a mixture average gave a solution nearly

identical to the more detailed multi-component diffusion. The mixture average solution

computational time was 44 minutes while the multi-component diffusion solution computational

time was 510 minutes. These points make the mixture average transport the method of choice.

Another numerical method option is to choose between windward and central

differencing. Both options were run for the same case and the windward differencing converged

without issue, while the central differencing did not converge to a solution. For this reason, the

windward differencing method was adopted.

0

500

1000

1500

2000

2500

0 0.02 0.04 0.06 0.08 0.1

Data 2 4:35:43 PM 10/23/2009

Multicomponent Diffusion Temperature SolutionMulticomponent Diffusion Velocity SolutionMixture Average Temperature SolutionMixture Average Velocity Solution

Te

mp

era

ture

(K

)

Axial Distance (cm)

Ve

locity

(cm

/s)

53

CEA2 Code was run as an “hp” problem to determine the products of the reactants at

equilibrium. The product compositions were used as an initial guess as described in Figure 24.

The type of intermediated species were approximated based on typical intermediate species in

hydrocarbon combustion [33]. The mole fraction of each intermediate species was set to a small

value, less than 0.1% by volume. The intermediate species specification is not essential in

converging to a solution, but can help the computations converge.

The calculation process takes multiple iterations. As the amount of steam was increased,

the calculated temperature profile became lower and less steep. The following is a figure showing

the evolution of the temperature profile as the steam content of the system is increased (from

mole fractions of 0.34 to 0.65) for Mixture 2 at 10 atm. Figure 28 indicates that as steam is added

to the mixture it becomes less energetic and less flammable.

Figure 28: Chemkin Code calculated evolution of temperature profile with different amounts of

steam addition

54

The centerline velocity also decreases as the steam content in increased. The following is

a plot of the decreasing velocity with increasing steam mole fraction, which indicates the non-

flammable mixture condition at points below 50 mm/s, as suggested by Westbrook [28].

Figure 29: Chemkin calculated evolution of centerline velocity with increasing steam

Figure 29 shows that the flammability limit is non-linear with respect to pressure. The

non-flammable data points require higher steam mole fractions for the 15 atm case than the 20

atm case. The calculated Mixture 2 results are to the right of those of Mixture 1. This is due to the

more energetic nature of Mixture 2, requires greater amounts of steam to suppress ignition.

5.3 Homogenous Batch Reactor Calculations

The homogenous batch reactor model, allows the user to specify initial temperature,

pressure, and mixture composition. Chemkin also allows specification of other parameters such as

volume of reactor and heat loss. The purpose of the model is to determine how long it will take a

Non-Flammable Region

55

specified mixture to reach its ignition (t= tign) at the given initial temperature and pressure

condition. This model is much simpler than the flame speed model and it takes much less

computational time to obtain solutions. Because of the shorter computation time, sensitivity

studies can be conducted much more efficiently, than with the flame speed model.

5.3.1 Sensitivity Studies

Sensitivity studies were conducted on the solution sensitivity to the A-factor and

temperature of the Arrhenius equation. In the chemical kinetic mechanism, each chemical

reaction is given an A-factor defined experimentally [32]. The A-Factor is the coefficient A in the

Arrhenius equation:

(7)

where k is the rate constant, Ea is the activation energy, Ru is the universal gas constant,

and T is the temperature. For sensitivity studies, small perturbations in A or T are made to see

how sensitive the solution is to changes in the rate constant (k) of each chemical reaction.

From the literature review, it was determined that the adiabatic flame temperature at

suppression conditions was in the range from 1000-1600 K. Egolfopoulos and Law [27] defined

this temperature as 1375 K, so this temperature was used in the adiabatic sensitivity studies. The

following A-factor sensitivity studies for Mixture 2 were calculated using the Chemkin batch

reactor at a flame temperature of 1375K under adiabatic conditions:

56

Figure 30: A-Factor sensitivity studies for Mixture 2

Figure 30 shows the top A-factor sensitivities of individual reactions for Mixture 2 at the

indicated pressure and mole fraction of steam. This gives a general idea of how the reaction paths

change with increasing pressure. It is interesting to note that only the high-pressure (30 atm) case

is sensitive to H2O2 reactions. This is important because it demonstrates that R1 goes to R4 and

R5 at higher pressures making the mixture more flammable. In other words, at 30 atm the HO2

radical produces a chain-propagating path when H2O2 decomposes to 2OH. Figure 30 also shows

-15 -10 -5 0 5 10 15

Mixture 2 at 1.0 atm X H2O = 0.67

Mixture 2 at 10.0 atm X H2O = 0.67

Mixture 2 at 10.0 atm X H2O = 0.61

Mixture 2 at 30.0 atm X H2O = 0.67

A Factor Sensitivity

H+O2 <=> O + OH

H+O2 + M <=> HO

2 + M

CH3+HO

2<=>CH

4+O

2

CH3+CH

3 + M<=> C

2H

6 + M

C2H

5+O

2<=>C

2H

4+HO

2

C2H

4+CH

3<=>C

2H

3+CH

4

C2H

3+O

2<=>CH

2CHO+O

C3H

8 + M <=> CH

3+C

2H

5 + M

H+C3H

8<=>H

2+I-C

3H

7

C4H

10+H<=>C

4H

9+H

2

CH3+HO

2<=>CH

3O+OH

H+C3H

8<=>H

2+N-C

3H

7

C3H

8+HO

2<=>I-C

3H

7+H

2O

2

H2O

2+O

2<=>HO

2+HO

2

(R3)

(R1)

(R14)

57

multiple chain-branching reactions that could be most dominant depending on the pressure. This

means that the “turning point” could be defined by multiple chain branching reactions.

This A-factor sensitivity study shows that the HO2 molecule goes from a chain-

terminating path to a chain-propagating path as pressure is increased to 30 atm. It also shows that

the “turning point” dominant reactions could potentially change with pressure and steam mole

fraction. These calculated results agree with the information obtained from the literature survey.

Temperature sensitivity studies were also performed on Mixture 1 and 2. These results

can be seen in the figure below.

Figure 31: Temperature sensitivity study for Mixture 1 and 2 at 15 atm

-1 -0.5 0 0.5 1

Mixture 2 at 15 atm , X H20 = 0.86

Mixture 1 at 15 atm , X H20 = 0.012

Mixture 1 at 15 atm , X H20 = 0.197

Temperature Sensitivity

H+O2<=>O+OH

H+O2 + M <=> HO

2 + M

C2H

6+CH

3CO

3<=>C

2H

5+CH

3CO

3H

CH3+HO

2<=>CH

3O+OH

CH3+HO

2<=>CH

4+O

2

CH3+CH

3 + M<=>C

2H

6 + M

C2H

6+CH

3<=>C

2H

5+CH

4

CH3+O

2<=>CH

2O+OH

CH4+H<=>CH

3+H

2

(R3)

(R1)

(R24)

(R16) (R14)

(R26)

(R18)

(R23)

)

58

Figure 31 shows the sensitivity of elementary reactions to the flame temperature at the

two different steam concentrations for Mixture 1. It also compares the results for Mixture 2 at the

same pressure. The Mixture 2 data in this figure are for an experimentally non-flammable

mixture, as the mole fraction of steam is 0.86. The temperature sensitivity can be summarized to

say that Mixture 2 at 15 atm has R1, R3, and R16 as the main chain branching reactions. It also

indicates the following reaction as an important chain branching reaction:

(R24)

The main chain terminating reactions are R14 and the following reaction:

(R16)

These reactions show that at 15 atm the HO2 molecule becomes reactive. M effects one of the

main chain terminating reactions R14 and one of the main chain branching reactions R1. The

branching reactions R24 and R16 show that formaldehyde would be a product in these Mixture 2

reactions, which is indicative of the cool flame regime.

Data for Mixture 1 are for an experimentally determined flammable mixture at a steam

mole fraction of 0.012 and for an experimentally determined non-flammable mixture at a steam

mole fraction of 0.197. The temperature sensitivities are opposite in reaction direction for the

following shared reactions of the two mixture 1 cases:

(R3)

(R16)

(R14)

(R26)

(R23)

As the non-flammable mixture temperature sensitivity is opposite in sign of the flammable

mixture temperature sensitivity for the above reactions, there must exist a point, dependent on

steam mole fraction, in which the sensitivities are equal. This means that as the mole fraction of

59

steam is increased, a small amount, reactions could change from chain-branching to chain-

terminating.

Overall, both sensitivity studies show that there are many reactions that control the

flammability of the mixtures. Many of these reactions involve the HO2 molecule and third body

reactions (M). The reactions, which control flammability, are very dependent on steam content in

Mixture 1. Reactions typical of the cool flame regime were found to be important in Mixture 2

with high mole fraction of steam. Further reaction path analysis and more complex rate of

production studies could further improve understanding of the chemistry involved.

5.3.2 Flammability Limits Using Homogenous Batch Reactor Model

It was determined that flammability studies can be conducted with the Chemkin batch

reactor by comparing the rate of production of the dominant chain-branching reaction (R3) and

the dominant chain terminating reaction (R1). If they are found to be equal the chemical

processes involved are at what Egofopoulos and Law refer to as a “turning point” that defines the

flammability boundary of the mixture [27]. Using the Chemkin code batch reactor, at adiabatic

conditions, with a reaction temperature of 1375K, the rates of production (ROP) for these two

reactions were calculated. Figure 32 is an example of this ROP method using Mixture 2. It shows

the amount of steam required for ROP of R1 to equal ROP of R3. Figure 32 also compares the

flame speed model to the ROP model.

60

Figure 32: Chemkin Code Flammability Limit study for Mixture 2

Figure 32 clearly shows the difference between the two methods of calculation using

Chemkin code. The ROP method estimates a greater amount of steam than the flame speed model

and does not show the non-linearity expected due to the change in reaction path above 13 atm, as

discussed in the literature survey. This shows that the ROP method is not valid using only

equations R1 and R3. The chemical equations used in this ROP method do change with pressure

and steam content, which makes this method depend strongly on sensitivity studies. For these

reasons, the flame speed model was preferred as this calculation inherently accounted for the

change in the dominant reaction mechanisms at the specified pressure and initial mixture

composition.

ROP Model

Flame Speed Model

61

Chapter 6

Discussion of Flammability Results

6.1 Flammability Results Compared with Chemkin Model

Figure 33 shows the Mixture 1 flammability boundary on the plot of pressure vs. mole

fraction of steam. The data points that define the flammability boundary are described by their

regime of combustion defined in Chapter 4. The experimentally determined flammability limit for

Mixture 1 shows that a very small amount of steam (less than 2%) was needed to create a non-

flammable mixture for any of the initial pressures. The limit shows highly non-linear behavior

with increasing pressure. This is similar to the non-linear trends found in the flammability limits

of hydrocarbon combustion discussed in Chapter 2.

Figure 33: Flammability Limits of Mixture 1

62

Figure 34 shows the Mixture 2 flammability limits on the plot of pressure vs. steam mole

fraction. The data points defining the flammability boundary were described by their regime of

combustion as indicated by the symbols in the legend. If a data point has a symbol that is of a

different color than that found in the legend, an abnormality in that result was observed. Details

regarding the abnormality can be seen by the tables in Appendix A. Figure 34 had a flammable

region that was divided into lower and higher reaction temperature regions. The division was

defined at the point in which abnormal ignitions began to be observed.

63

Figure 34: Flammability Limits of Mixture 2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

0.00 0.10 0.20 0.30 0.40 0.50 0.60 0.70 0.80

Intermediate Reaction

Rapid Reaction

Slow Reaction

No Reaction

Chemkin Model: No Reaction

Cool Flame Reaction

Multiple Ignition

Norm

ali

zed

In

itia

l C

ham

ber

Pre

ssu

re

[Pc,

i / P

ref]

Mole Fraction of Steam

Non-Flammable Region

Flammable

Region

TR>1400K

Lower Temperature

Flammable Region

TR<1400K

64

As indicated in the Introduction, it was anticipated that Mixture 1 would require less

steam to be non-flammable than Mixture 2. This is verified experimentally, by comparing the

scale of steam mole fraction between Figures 33 and 34. The required steam mole fraction for

Mixture 2 to be non-flammable is about 35 times higher than that of Mixture 1. No ignitions were

observed below Pc,i /Pref =0.25. This could be due to the dependency of the rich limit on pressure.

The Chemkin code results in Figure 33 show a flammability boundary similar to the

experimentally determined profile, except this boundary is shifted to the right to much higher

(~102) mole fractions of steam. The Chemkin code results in Figure 34 show a similar

flammability boundary as experimental data except for the Pc,i /Pref = 1.5 data point. Comparing

the two Chemkin calculations, it was found that more steam is needed for Mixture 2 to be non-

flammable than Mixture 1. The Chemkin flame speed model overestimates the flammability

boundary for both mixtures below 20 atm. This overestimation is primarily due to the fact that the

experiments sweep away reactive radicals due to the open flow through the chamber. The

Chemkin flame speed model does not account for this loss so it overestimates the flammability

limit. The Chemkin flame speed model is a better estimate in Mixture 2 as the reactive radicals

are not swept away as fast as in Mixture 1 tests. The residence time for the experiments

performed near the flammability boundaries are 4 times shorter in Mixture 1 than in Mixture 2

tests.

These calculations overestimated the experimental trends, which showed, as expected,

that the laminar premixed flame model did not accurately represent the physical processes

involved in the experiments. The model does not accurately estimate the losses in the experiment.

Additionally, it is assumed that more complicated processes, such as turbulent mixing and short-

time scale ignition phenomena are of vital importance in defining the flammability limits.

Both Chemkin calculations show a non-linearity in the flammability boundary, which

matches the proposed theory that the limits will be non-linear due to the non-linear flammability

65

limits of the classical hydrogen-oxygen mechanism that is at the core of hydrocarbon-oxygen

flammability limits. The experimental trend in the Mixture 1 boundary also shows this non-

linearity.

6.2 Summary of Flammability Results

It was shown, from the literature review, that the steam controls the flammability of any

hydrocarbon mixture with two mechanisms: 1) the steam affects the reactions chemically, by

increasing the 3rd

-body collisions that increases the number of chain-terminating or chain

branching reaction paths; and 2) physically, by lowering the reaction-zone temperature through

dilution. The flammability limits are also controlled by initial chamber pressure. The increase of

initial chamber pressure can change the dominant chemical reaction mechanisms involved. While

the steam and initial chamber pressure determine what reaction paths are available, the

equivalence ratio determines which path is taken to complete the reaction. The reaction paths

taken determine the flammability limits of the mixture, and which steam suppression mechanism

is more dominant.

In the Mixture 1 flammability boundary, shown in Figure 33, the non-linearity has to be

due to the change in the chemical reaction mechanisms with increasing initial chamber pressure.

The mixture is within its empirically determined flammability limits and there is not enough

steam to allow for the physical effect to suppress ignition. The Mixture 2 boundary is mainly

controlled by the physical suppression mechanism and does not seem to be suppressed by chain

terminating reactions as the indicative non-linearity trend is not found in the boundary in Figure

34. Further evidence that Mixture 2 is suppressed by the physical mechanism lies in the fact that

the temperature of the reaction zone is clearly lowered with increasing steam. With above 25%

66

Vol. steam, Mixture 2 transitions to different chemical reaction mechanism as seen by the change

in flame spreading and pressurization discussed in Chapter 4.

The two mixtures are suppressed by different mechanisms because the amount of

available oxygen per volume for chain branching and propagation is higher in Mixture 2 than in

Mixture 1 as seen in the difference in equivalence ratio. This means that the amount of chain

branching and propagation of reactions is too high in Mixture 2 to be effected by the chain

terminating reactions, so Mixture 2 was mainly suppressed by the physical effect of steam. At

higher pressures, Mixture 2 is more difficult to completely suppress as the cool flame regime

exists well above the upper flammability limits.

In Mixture 1 the chain terminating reactions, increased by steam content, are able to

compete with the smaller amount of chain propagating and branching reactions due to the smaller

mole fraction of oxygen in the mixture. This competition between chain-branching and chain-

terminating reactions creates a non-flammable condition depending on the steam content of the

mixture, the initial chamber pressure, and the thermal induction time.

For both Mixtures 1 and 2 the test results indicated that steam can be used to control the

flammability. However, Mixture 2 requires such a large mole fraction of steam that using it to

control the flammability limits would be inefficient for industrial processes, such as SCT-CPO.

Newson and Truong [34] show that for safety in industrial CPO processes the oxygen content in

the system should remain below 10% Vol.

To determine if steam can effectively be used to control the flammability of a mixture

used in an industrial process, the equivalence ratio of the mixture without steam is the most

important parameter. If the equivalence ratio of Mixture 2 was increased it would require less

steam to make a non-flammable mixture. In addition, there must be some point in between the

two equivalence ratios that the chain-terminating suppression mechanism and the physical

suppression mechanism have an equal effect on the mixture’s flammability. Further investigation

67

of mixtures with equivalence ratios between these two mixtures is needed to support this

hypothesis.

68

Chapter 7

Conclusions

The highly non-linear flammability boundaries for two different

hydrocarbon/oxygen/steam mixtures were determined through a series of tests performed at

operating conditions close to those adopted in a SCT-CPO reactor. The tests demonstrated that

steam can be used to control and suppress the flammability of hydrocarbon/oxygen mixtures. For

mixtures with lower adiabatic flame temperature (Mixture 1), less steam is required to achieve a

non-flammable condition. In the Mixture 2 tests, unusual flame spreading was observed with

steam mole fraction greater than 24%, indicating that steam is physically affecting the flame

spreading processes by lowering the reaction zone temperature. Also in these tests, the pressure

threshold was observed from these flammability tests, below which a self sustained ignition was

not possible due to insufficient energy release at lower initial pressures.

Similar to the experimentally observed phenomena, the Chemkin calculations also show

that more steam is required to suppress the flammability of Mixture 2 than Mixture 1. The

sensitivity studies performed with the homogenous batch reactor showed that the dominant

chemical reaction mechanisms change with increasing pressure and steam content. Because of

these dependencies, flammability studies using the homogenous batch reactor rate of production

(ROP) flammability limit method needed to be coupled with sensitivity studies for every

condition tested in order to determine the dominant chain-branching and chain terminating

reactions. The ROP flammability limit method using only R1 and R3 overestimated the steam

content required for suppression and therefore could not be used to illustrate the non-linear

dependency on pressure.

The flame speed model was more accurate in calculating the flammability limits as it did

not rely on assuming certain reaction mechanisms are most dominant, as in the ROP method. The

69

flame speed model calculations overestimated the trends seen in the experiments, but they did

account for the expected non-linearity in the trend. The flame speed model was found to be

useful, but its flammability limit solution is an overestimate of the experiments as the solution

does not account for the steady flow of reactant through the reactor. Turbulent combustion and

small time scale ignition phenomena were not considered due to the fact that it is beyond the

scope of the current program.

The results of the experiments and Chemkin calculations showed that the flammability of

these mixtures is a strong function of the equivalence ratio and its effect on the sensitivity of the

mixture to the two possible suppression mechanisms: the chain-terminating chemical suppression

mechanism and the dilution or physical mechanism. Mixture 1 was suppressed by the chemical

suppression mechanism, where small amounts of steam greatly increased the rates of chain-

terminating reactions by increasing 3rd

-body reaction effectiveness. In particular R14 is increased

by small increase in steam and absorbs available CH3 radicals causing overall chain termination.

Mixture 2 was mainly suppressed by the physical suppression mechanism, where the heat release

from the chemical reactions was diluted by the addition of a significant amount of steam causing

the reaction zone temperature to decrease and thereby creating a non-flammable condition.

Further research into mixtures of similar composition but different equivalence ratios is

required to validate the existence of these mechanisms and their dependence on equivalence ratio.

It is proposed that there exists a mixture with a certain equivalence ratio that will allow it to be

suppressed equally by both mechanisms. Further research could incorporate laser diagnostics to

measure radical concentrations in the flame front that would be indicative of certain chemical

suppression mechanisms. Incorporating a spectrometer in the experimental setup could be highly

beneficial in identifying formaldehyde species, which is indicative of the lower temperature

regime. Using different inert agents, other than steam, like N2, could be useful for comparison to

gain a better understanding of the chemical effect steam has on a mixture. In addition, more

70

accurate modeling could be investigated to include effects of turbulence, thermal induction time,

and ignition phenomena.

71

References

[1] Petersen, E., Davidson, D., Hanson, R., Combustion and Flame 117 (1999) 272-290.

[2] Horn, R., Williams, K., Degenstein, N., Schmidt, L., Journal of Catalysis 242 (2006) 92-

102.

[3] L. Basini, Catalysis Today 117 (2006) 384–393

[4] L. Basini, Catalysis Today 106 (2005) 34–40

[5] D.A. Hickman, L.D. Schmidt, Science 259 (1993) 343

[6] L.D. Schmidt, JR, Salge, B.J. Dreyer, P.J. Dauenhauer, Science 314(5800) (2006) 801

[7] R. Seiser, K. Seshadri, Proc. Combust. Inst. 30 (2005) 407-414.

[8] R. J. Kee, F. M. Rupley, J. A. Miller, M. E. Coltrin, J. F. Grcar, E. Meeks, H. K. Moffat,

A. E. Lutz, G. Dixon-Lewis, M. D. Smooke, J. Warnatz, G. H. Evans, R. S. Larson, R. E.

Mitchell, L. R. Petzold, W. C. Reynolds, M. Caracotsios, W. E. Stewart, P. Glarborg, C.

Wang, O. Adigun, W. G. Houf, C. P. Chou, S. F. Miller, P. Ho, and D. J. Young,

CHEMKIN Release 4.0.1, Reaction Design, Inc., San Diego, CA (2004).

[9] Westbrook C.K., Pitz W.J., Curran H.J., Gaffuri P., Marinov N.M.. Chemical kinetic

modeling of hydrocarbon ignition. In: Wolfrum J., Volpp H.R., Rannacher R., Warnatz

J., editors. Gas phase chemical reaction systems. Springer Series in Chemical Physics,

vol. 61; 1996. 279-290.

[10] Glassman, I., Combustion, Academic Press, Orlando, U.S.A., 1987.

[11] Liang, Y., & Zeng, W., Journal of Hazardous Materials 174 (1-3) (2010) 386-92.

[12] Marshall, J. B., Hydrogen:Air:Steam Flammability Limits and Combustion

Charachteristics in the FITS Vessel, Nuclear Regulatory Commision, Albuquerque:

Sandia National Lab, 1986.

72

[13] Sapko, M. J., Furno, A. L., & Kuchta, J. M.,Quenching Methane-Air Ignitions with

Water Sprays. United States Department of the Interior, Bureau of Mines, 1977.

[14] Yamamoto, T. K., Kobayashi, N., Arai, N., & Tanaka, T., Energy Conversion

Management 38 (1997) 1093-1100.

[15] Chen, C., Wang, T., Liaw, H., Chen, H., Journal of Hazardous Materials 166 (2009) 880-

890.

[16] D.A. Crowl, J.F. Louvar, Chemical Process Safety-Fundamentals with Applications, 2nd

ed., Prentice Hall, New Jersey, 2002.

[17] Kuo, K. K., Principles of Combustion, John Wiley and Sons, Hoboken, U.S.A., 2005.

[18] NASA Computer Program CEA (Chemical Equilibrium with Applications).

<http://www.grc.nasa.gov/WWW/CEAWeb/ceaHome.htm> (version 05/21/04).

[19] C. K. Westbrook, F.L Dryer, Proc. Combust. Inst. 18 (1981) 749-767.

[20] Shebeko, Y. et al., Journal of Combustion, Explostion, and Shock Waves 30 (1994) 183-

188.

[21] Shebko, Y. et al. Fire safety Journal 37 (2002) 549-568.

[22] Simmie, J., Progress in Energy and Combustion Science 29 (2006) 599-634.

[23] Fawcett, H. and Wood, S., Safety and Accident Prevention in Chemical Operations,

Wiley, New York, U.S.A., 1982.

[24] Warnatz, J., Maas, U., Dibble, R.W., Combustion, 2nd

Edition Springer, Berlin, Germany.

1999.

[25] Bartknecht, W., Explosions, Springer, Berlin, Germany., 1981.

[26] Lewis, B. and Elbe, G., Combustion, Flames, and Explosions of Gases, Academic Press,

New York, U.S.A., 1961

[27] C. Law, F. Egolfopoulos, Proc. Combust. Inst. 24 (1992) 137-144.

[28] Westbrook, C. K., Proc. Combust. Inst. 30 (1982) 127-141.

73

[29] Womeldorf, C., Grosshandler, W., Combustion and Flame 118 (1999) 25-36.

[30] Bui-Pham, M. N., Lutz, A. E., Miller, J. A., Desjardin, M., O’Shaugnessey, D. M., and

Zondlak, R. J., Combust. Sci. Technol. 109:71–91 (1995).

[31] Petersen, E.L., Kalitan, D.M., Simmons, S., Bourque, G., Curran, H.J., Simmie, J.M.,

Proc. Combust. Inst. 31 (2007) 447-454.

[32] N. Donato, C. Aul, E. Petersen, C. Zinner, H. Curran, & G. Bourque. (2009). Ignition and

Oxidation of 50/50 Butane Isomer Blends. Proceedings of ASME Turbo Expo 2009:

Power for Land, Sea, and Air (1-13). Orlando: ASME.

[33] Law, C. K., Combustion Physics, Cambridge University Press, New York, U.S.A., 2006.

[34] Newson, E., Troung, B., International Journal of Hydrogen Energy 28 (2002) 1379-1386.

[35] Kutzler, Patrick M. "Flammability Limits of a Premixed Gas with Steam Addition."

Thesis. The Pennsylvania State University, 2008.

[36] Kline, S.J. and McClintok, F.A., “Describing Uncertainties in Single-Sample

Experiments,” Mechanical Engineering, ASME, Vol. 75, No. 1, 1953, p. 3-8.

[37] Reference Fluid Thermodynamic and Transport Properties. NIST Standard Reference

Database 23, Version 8.0.

74

Appendix A

Experimental Test Matrix

The tables given in Appendix A show all the calculations and data collected for every test

executed in both Mixture 1 and Mixture 2 testing. These tables are useful as finalized records of

all observations made during testing. See Appendix E and F for how the calculations were made.

75

Table A.1: Mixture 1 test series initial flow rates and chamber pressure

* Test failure due to various reasons (e.g., igniter malfunction or depletion of reactants).

1 0.79 31.14 0.00 3.00 45.72 1.00 0.00 0.00

2 0.75 32.53 0.00 3.00 45.70 1.00 0.00 0.00

3 0.86 32.54 0.00 2.99 45.62 1.00 0.00 0.00

4 1.43 32.55 0.00 0.00 45.66 1.00 0.00 0.00

5 1.47 32.55 0.00 0.00 45.66 1.00 0.00 0.00

6 1.46 31.14 12.30 0.00 46.29 1.50 0.00 0.00

7 1.44 31.20 0.00 0.00 46.14 1.50 0.00 0.00

8 1.47 31.24 2.10 3.00 45.81 1.50 0.00 0.00

9 1.47 31.22 1.30 3.00 45.98 1.50 0.00 0.00

10 1.47 31.14 0.00 3.01 46.17 1.50 0.00 0.00

11 0.72 31.21 0.63 3.00 46.03 1.50 0.00 0.00

12 0.69 31.18 1.03 2.99 46.11 1.50 0.00 0.00

13 0.69 31.23 0.76 3.00 45.48 1.50 0.00 0.00

14 0.37 31.22 0.49 3.01 45.73 1.50 0.00 0.00

15 0.38 31.20 0.17 3.00 45.80 1.50 0.00 0.00

16 0.94 31.21 1.06 0.00 46.42 1.50 0.00 0.00

17 0.97 31.28 0.85 3.02 45.79 1.50 0.00 0.00

18 0.88 31.24 0.68 3.00 45.66 1.50 0.00 0.00

19 1.16 31.27 0.81 2.86 45.15 1.50 0.00 0.00

20 1.14 31.20 0.76 2.99 45.50 1.50 0.00 0.00

21 1.13 31.18 0.58 3.01 46.12 1.50 0.00 0.00

22 1.23 31.22 0.74 2.84 45.49 1.50 0.00 0.00

23 0.55 31.23 0.35 2.86 45.92 1.50 0.00 0.00

24 0.91 31.24 0.67 3.05 45.99 1.50 0.00 0.00

25 0.95 31.23 0.49 3.01 45.84 1.50 0.00 0.00

26 0.51 31.21 0.47 2.99 46.03 1.50 0.00 0.00

27 0.33 31.21 0.00 3.03 45.40 1.50 0.00 0.00

28 1.04 31.27 0.54 3.00 46.90 1.50 0.00 0.00

29 0.85 31.22 0.49 3.00 45.68 1.50 0.00 0.00

30 0.85 31.24 0.74 3.00 46.01 1.50 0.00 0.00

31 0.73 31.17 0.54 3.07 46.08 1.50 0.00 0.00

32 0.74 31.23 0.39 2.97 46.02 1.50 0.00 0.00

33 0.80 31.13 0.53 2.96 46.24 1.50 0.00 0.00

34 0.85 31.20 0.40 3.00 46.29 1.50 0.00 0.00

35*

36 0.64 31.18 0.51 3.05 46.32 1.50 0.00 0.00

37 0.61 31.14 0.54 2.95 46.16 1.50 0.00 0.00

38*

CH4

(Nm 3/hr)

H2

(Nm 3/hr)

C3H8

(Nm 3/hr)

C4H10

(Nm 3/hr)

O2

(Nm 3/hr)

Steam

(Nm 3/hr)

C2H6

(Nm 3/hr)Pc,i / PrefTest #

76

Table A.2: Mixture 1 test series individual reactant species mole fraction

* Test failure due to various reasons (e.g., igniter malfunction or depletion of reactants).

1 0.39 0.0000 0.04 0.57 0.0124 0.00 0.00

2 0.40 0.0000 0.04 0.56 0.0122 0.00 0.00

3 0.40 0.0000 0.04 0.56 0.0122 0.00 0.00

4 0.41 0.0000 0.00 0.58 0.0127 0.00 0.00

5 0.41 0.0000 0.00 0.58 0.0127 0.00 0.00

6 0.34 0.1348 0.00 0.51 0.0164 0.00 0.00

7 0.40 0.0000 0.00 0.59 0.0190 0.00 0.00

8 0.37 0.0251 0.04 0.55 0.0179 0.00 0.00

9 0.38 0.0157 0.04 0.55 0.0181 0.00 0.00

10 0.38 0.0000 0.04 0.56 0.0184 0.00 0.00

11 0.38 0.0076 0.04 0.56 0.0182 0.00 0.00

12 0.38 0.0124 0.04 0.56 0.0181 0.00 0.00

13 0.38 0.0093 0.04 0.55 0.0183 0.00 0.00

14 0.38 0.0060 0.04 0.56 0.0183 0.00 0.00

15 0.38 0.0021 0.04 0.56 0.0184 0.00 0.00

16 0.39 0.0132 0.00 0.58 0.0187 0.00 0.00

17 0.38 0.0103 0.04 0.56 0.0182 0.00 0.00

18 0.38 0.0083 0.04 0.56 0.0183 0.00 0.00

19 0.38 0.0099 0.04 0.55 0.0184 0.00 0.00

20 0.38 0.0093 0.04 0.56 0.0183 0.00 0.00

21 0.38 0.0070 0.04 0.56 0.0182 0.00 0.00

22 0.38 0.0090 0.03 0.56 0.0184 0.00 0.00

23 0.38 0.0043 0.03 0.56 0.0183 0.00 0.00

24 0.38 0.0081 0.04 0.56 0.0182 0.00 0.00

25 0.38 0.0060 0.04 0.56 0.0183 0.00 0.00

26 0.38 0.0057 0.04 0.56 0.0183 0.00 0.00

27 0.38 0.0000 0.04 0.56 0.0185 0.00 0.00

28 0.38 0.0065 0.04 0.56 0.0180 0.00 0.00

29 0.38 0.0060 0.04 0.56 0.0183 0.00 0.00

30 0.38 0.0090 0.04 0.56 0.0182 0.00 0.00

31 0.38 0.0066 0.04 0.56 0.0182 0.00 0.00

32 0.38 0.0047 0.04 0.56 0.0183 0.00 0.00

33 0.38 0.0065 0.04 0.56 0.0182 0.00 0.00

34 0.38 0.0049 0.04 0.56 0.0182 0.00 0.00

35*

36 0.38 0.0062 0.04 0.56 0.0182 0.00 0.00

37 0.38 0.0066 0.04 0.56 0.0182 0.00 0.00

38*

X Steam X C2H6 X CH4 X H2 X C3H8 X C4H10Test # X O2

77

Table A.3: Mixture 1 test series initial conditions: O/C , S/C , temperatures, and φ

* Test failure due to various reasons (e.g., igniter malfunction or depletion of reactants).

1 0.60 0.0000 177.00 175.00 Bot 3.29

2 0.63 0.0000 178.00 173.00 Bot 22.10 3.15

3 0.63 0.0000 167.00 160.00 Bot 18.60 3.14

4 0.63 0.0000 170.00 164.00 Bot 41.10 2.82

5 0.63 0.0000 185.00 175.00 Bot 2.82

6 0.60 0.2657 180.00 170.00 Bot 3.00

7 0.60 0.0000 180.00 175.00 Bot 2.98

8 0.60 0.0405 178.00 173.00 Bot 3.29

9 0.60 0.0250 178.00 173.00 Bot 3.31

10 0.60 0.0000 187.00 183.00 Bot 75.10 3.33

11 0.60 0.0121 189.00 174.00 Bot 33.00 3.31

12 0.60 0.0198 180.00 172.00 Bot 3.32

13 0.61 0.0148 181.00 175.00 Bot 3.27

14 0.60 0.0095 177.00 172.00 Bot 3.29

15 0.60 0.0033 180.00 178.00 Bot 3.30

16 0.60 0.0228 187.00 182.00 Bot 3.00

17 0.60 0.0164 185.00 178.00 Bot 3.29

18 0.60 0.0132 189.00 182.00 Bot 3.28

19 0.61 0.0159 179.00 170.00 Bot 3.23

20 0.61 0.0148 178.00 172.00 Bot 3.28

21 0.60 0.0111 178.00 172.00 Bot 3.32

22 0.61 0.0145 175.00 167.00 Bot 15.70 3.26

23 0.60 0.0068 160.00 160.00 Bot 28.30 3.29

24 0.60 0.0129 180.00 168.00 Bot 3.31

25 0.60 0.0094 187.00 176.00 Bot 3.30

26 0.60 0.0090 159.00 147.00 Bot 3.31

27 0.61 0.0000 170.00 160.00 Bot 13.50 3.27

28 0.59 0.0102 168.00 160.00 Bot 46.30 3.36

29 0.60 0.0095 170.00 158.00 Bot 3.29

30 0.60 0.0142 169.00 159.00 Bot 3.31

31 0.60 0.0103 169.00 160.00 Bot 3.32

32 0.60 0.0075 174.00 167.00 Bot 92.92 3.30

33 0.60 0.0102 167.00 163.00 Bot 3.33

34 0.60 0.0076 182.00 177.00 Bot 3.33

35*

36 0.60 0.0097 183.00 173.00 Bot 3.34

37 0.60 0.0104 177.00 170.00 Bot 3.32

38*

Igniter

Location

ST,abs

(m/s)φTest #

O/C

Vol.

Ratio

Steam/C

Vol.

Ratio

Tctop

(oC)

Tcbottom

(oC)

78

Table A.4: Mixture 1 test series flow parameters

* Test failure due to various reasons (e.g., igniter malfunction or depletion of reactants).

1 1.80 10.12 22.83 2.71E+04

2 1.94 9.62 23.39 2.78E+04

3 1.66 11.22 23.37 2.78E+04

4 0.97 18.18 22.26 2.64E+04

5 0.98 18.04 22.26 2.64E+04

6 1.08 18.06 24.61 2.92E+04

7 0.98 17.77 21.82 2.59E+04

8 1.02 18.33 23.38 2.78E+04

9 1.01 18.27 23.23 2.76E+04

10 1.02 17.95 22.94 2.72E+04

11 2.03 9.03 23.08 2.74E+04

12 2.08 8.85 23.17 2.75E+04

13 2.07 8.83 23.01 2.73E+04

14 3.62 5.06 23.00 2.73E+04

15 3.50 5.20 22.92 2.72E+04

16 1.51 11.66 22.13 2.63E+04

17 1.53 12.01 23.12 2.75E+04

18 1.68 10.91 23.03 2.73E+04

19 1.26 14.45 22.91 2.72E+04

20 1.28 14.30 23.00 2.73E+04

21 1.29 14.18 23.08 2.74E+04

22 1.18 15.43 22.94 2.72E+04

23 2.41 7.58 22.95 2.73E+04

24 1.60 11.49 23.11 2.74E+04

25 1.55 11.84 23.02 2.73E+04

26 2.61 7.03 23.04 2.74E+04

27 3.85 4.71 22.82 2.71E+04

28 1.39 13.35 23.26 2.76E+04

29 1.65 11.08 22.98 2.73E+04

30 1.66 11.10 23.11 2.74E+04

31 1.91 9.63 23.08 2.74E+04

32 1.92 9.52 23.02 2.73E+04

33 1.75 10.48 23.05 2.74E+04

34 1.70 10.78 23.08 2.74E+04

35*

36 2.24 8.20 23.12 2.75E+04

37 2.29 8.01 23.04 2.74E+04

38*

Test #V

(m/s)

ρ

(kg/m^3)

m dot

(g/s)ReD

79

Table A.5: Mixture 1 test series flammability and type of pressurization

* Test failure due to various reasons (e.g., igniter malfunction or depletion of reactants).

1 No

2 Yes Intermediate, I

3 Yes Intermediate, I

4 Yes Rapid, R

5 No

6 No

7 No

8 No

9 No

10 Yes Rapid, R

11 Yes, Low dP/dt Multi-Ignitions, M

12 No

13 No

14 No

15 No

16 No

17 No

18 No

19 No

20 No

21 No

22 Yes Slow, S

23 Yes Intermediate, I

24 No

25 No

26 No

27 Yes, Low dP/dt Cool Flame, C

28 Yes Rapid, R

29 No

30 No

31 No

32 Yes Rapid, R

33 No

34 Yes Slow, S

35*

36 Yes, Low dP/dt Cool Flame, C

37 Yes, Low dP/dt Cool Flame, C

38*

Flammable

Rate of

Pressurization

(R,I,S,C,M)

Test #

80

Table A.6: Mixture 1 test series steam parameters and orifice type

* Test failure due to various reasons (e.g., igniter malfunction or depletion of reactants).

1 0.16 3.00 13.44

2 0.16 3.00 13.44

3 0.16 3.00 13.44

4 0.16 3.00 13.44

5 0.16 3.00 13.44

6 0.16 3.00 13.44

7 0.16 3.00 13.44

8 0.16 3.00 13.44

9 0.16 3.00 13.44

10 0.16 3.00 13.44

11 0.03 0.50 2.24

12 0.03 0.50 2.24

13 0.03 0.50 2.24

14 0.03 0.50 2.24

15 0.03 0.50 2.24

16 0.03 0.50 2.24

17 0.03 0.50 2.24

18 0.03 0.50 2.24

19 0.03 0.50 2.24

20 0.03 0.50 2.24

21 0.03 0.50 2.24

22 0.03 0.50 2.24

23 0.03 0.50 2.24

24 0.03 0.50 2.24

25 0.03 0.50 2.24

26 0.03 0.50 2.24

27 0.03 0.50 2.24

28 0.03 0.50 2.24

29 0.03 0.50 2.24

30 0.03 0.50 2.24

31 0.03 0.50 2.24

32 0.03 0.50 2.24

33 0.03 0.50 2.24

34 0.03 0.50 2.24

35* 0.03 0.50 2.24

36 0.03 0.50 2.24

37 0.03 0.50 2.24

38* 0.03 0.50 2.24

Test #ORIFICE BORE ID

(in)

Max Steam

(g/s)

Max Steam

(Nm3/hr)

81

Table A.7: Mixture 2 test series initial flow rates and chamber pressure

* Test failure due to various reasons (e.g., igniter malfunction or depletion of reactants).

39 (LPG1) 0.48 33.55 0.20 0.22 13.24 0.00 5.78 3.58

40 (LPG 2) 0.54 33.54 0.55 0.65 13.30 0.00 6.98 4.08

41 (LPG 3) 0.46 33.51 1.37 0.65 13.20 0.00 7.22 3.79

42 (LPG 4)*

43 (LPG 5) 0.43 33.52 1.61 0.65 13.23 0.00 7.41 3.30

44 (LPG 6)*

45 (LPG 7) 0.42 33.54 2.25 0.65 13.20 0.00 7.48 3.33

46 (LPG 8) 0.27 33.58 2.25 0.66 13.06 0.00 7.55 3.64

47 (LPG9) 0.26 33.57 2.22 0.64 13.33 0.00 1.92 3.42

48 (LPG 10) 0.26 33.54 2.38 0.65 13.07 0.00 6.95 4.45

49 (LPG 11) 0.26 33.57 4.40 0.65 13.13 0.00 6.55 3.49

50 (LPG 12) 0.34 33.54 4.81 0.67 13.20 0.00 7.66 4.47

51 (LPG 13) 0.28 33.55 7.10 0.68 13.19 0.00 7.31 4.36

52 (LPG 14) 0.33 33.58 15.91 0.50 13.20 0.00 6.70 3.18

53 (LPG 15) 0.34 33.51 18.37 0.63 13.20 0.00 7.19 3.73

54 (LPG 17) 0.38 33.49 24.70 0.66 13.21 0.00 7.04 3.39

55 (LPG 18) 0.43 33.52 33.94 0.66 13.19 0.00 7.97 3.54

56 (LPG 19) 0.55 33.48 40.70 0.49 12.81 0.00 7.25 3.69

57 (LPG 20) 0.19 16.67 24.79 0.34 6.24 0.00 3.42 1.87

58 (LPG 21) 0.37 16.73 24.26 0.31 6.54 0.00 3.39 1.45

59 (LPG 22) 0.23 16.69 35.00 0.32 6.13 0.00 3.21 1.56

60 (LPG 23) 0.42 8.25 30.00 0.18 2.21 0.00 1.77 0.83

61 (LPG 24) 0.39 8.31 20.20 0.16 3.05 0.00 1.79 0.89

62 (LPG 25)*

63 (LPG 26) 0.48 8.30 29.30 0.18 2.74 0.00 1.32 0.41

64 (LPG 27) 0.49 8.39 30.12 0.16 3.47 0.00 1.82 0.66

65 (LPG 28)*

66 (LPG 29) 0.47 8.30 25.00 0.18 3.41 0.00 1.93 1.17

67 (LPG 30) 0.09 8.29 11.30 0.17 3.12 0.00 1.78 0.91

68 (LPG 31) 0.06 8.35 4.95 0.17 3.31 0.00 1.40 0.98

69 (LPG 32) 0.20 8.30 18.15 0.15 3.38 0.00 2.31 1.12

70 (LPG 33) 0.23 8.28 21.91 0.15 3.24 0.00 1.81 0.49

71 (LPG 34) 0.68 8.34 25.00 0.19 2.71 0.00 1.59 0.85

72 (LPG 35)*

73 (LPG 36) 0.72 8.29 19.90 0.17 2.96 0.00 1.41 0.68

74 (LPG 37) 0.97 8.40 26.05 0.17 3.10 0.00 1.70 1.01

75 (LPG 38) 1.02 8.35 30.12 0.20 3.19 0.00 1.64 0.89

76 (LPG 39) 1.00 8.31 34.16 0.17 3.69 0.00 1.45 0.88

Steam

(Nm 3/hr)

C2H6

(Nm 3/hr)

CH4

(Nm 3/hr)

H2

(Nm 3/hr)

C3H8

(Nm 3/hr)

C4H10

(Nm 3/hr)

O2

(Nm 3/hr)Pc,i / PrefTest #

82

Table A.8: Mixture 2 test series individual reactant species mole fraction

* Test failure due to various reasons (e.g., igniter malfunction or depletion of reactants).

39 (LPG1) 0.59 0.00 0.00 0.23 0.00 0.10 0.06

40 (LPG 2) 0.57 0.01 0.01 0.23 0.00 0.12 0.07

41 (LPG 3) 0.56 0.02 0.01 0.22 0.00 0.12 0.06

42 (LPG 4)*

43 (LPG 5) 0.56 0.03 0.01 0.22 0.00 0.12 0.06

44 (LPG 6)*

45 (LPG 7) 0.55 0.04 0.01 0.22 0.00 0.12 0.06

46 (LPG 8) 0.55 0.04 0.01 0.22 0.00 0.12 0.06

47 (LPG9) 0.61 0.04 0.01 0.24 0.00 0.03 0.06

48 (LPG 10) 0.55 0.04 0.01 0.21 0.00 0.11 0.07

49 (LPG 11) 0.54 0.07 0.01 0.21 0.00 0.11 0.06

50 (LPG 12) 0.52 0.07 0.01 0.21 0.00 0.12 0.07

51 (LPG 13) 0.51 0.11 0.01 0.20 0.00 0.11 0.07

52 (LPG 14) 0.46 0.22 0.01 0.18 0.00 0.09 0.04

53 (LPG 15) 0.44 0.24 0.01 0.17 0.00 0.09 0.05

54 (LPG 17) 0.41 0.30 0.01 0.16 0.00 0.09 0.04

55 (LPG 18) 0.36 0.37 0.01 0.14 0.00 0.09 0.04

56 (LPG 19) 0.34 0.41 0.00 0.13 0.00 0.07 0.04

57 (LPG 20) 0.31 0.46 0.01 0.12 0.00 0.06 0.04

58 (LPG 21) 0.32 0.46 0.01 0.12 0.00 0.06 0.03

59 (LPG 22) 0.27 0.56 0.01 0.10 0.00 0.05 0.02

60 (LPG 23) 0.19 0.69 0.00 0.05 0.00 0.04 0.02

61 (LPG 24) 0.24 0.59 0.00 0.09 0.00 0.05 0.03

62 (LPG 25)*

63 (LPG 26) 0.20 0.69 0.00 0.06 0.00 0.03 0.01

64 (LPG 27) 0.19 0.67 0.00 0.08 0.00 0.04 0.01

65 (LPG 28)*

66 (LPG 29) 0.21 0.63 0.00 0.09 0.00 0.05 0.03

67 (LPG 30) 0.32 0.44 0.01 0.12 0.00 0.07 0.04

68 (LPG 31) 0.44 0.26 0.01 0.17 0.00 0.07 0.05

69 (LPG 32) 0.25 0.54 0.00 0.10 0.00 0.07 0.03

70 (LPG 33) 0.23 0.61 0.00 0.09 0.00 0.05 0.01

71 (LPG 34) 0.22 0.65 0.00 0.07 0.00 0.04 0.02

72 (LPG 35)*

73 (LPG 36) 0.25 0.60 0.01 0.09 0.00 0.04 0.02

74 (LPG 37) 0.21 0.64 0.00 0.08 0.00 0.04 0.02

75 (LPG 38) 0.19 0.68 0.00 0.07 0.00 0.04 0.02

76 (LPG 39) 0.17 0.70 0.00 0.08 0.00 0.03 0.02

X C3H8 X C4H10Test # X O2 X Steam X C2H6 X CH4 X H2

83

Table A.9: Mixture 2 test series initial conditions: O/C , S/C , temperatures, and φ

* Test failure due to various reasons (e.g., igniter malfunction or depletion of reactants).

39 (LPG1) 0.74 0.00 166.00 165.00 Bot 404.90 2.37

40 (LPG 2) 0.65 0.01 145.00 145.00 Bot 2.69

41 (LPG 3) 0.65 0.03 138.00 127.00 Bot 36.92 2.67

42 (LPG 4)*

43 (LPG 5) 0.67 0.03 145.00 145.00 Bot 84.72 2.60

44 (LPG 6)*

45 (LPG 7) 0.67 0.04 172.00 150.00 Bot 56.85 2.62

46 (LPG 8) 0.65 0.04 133.00 152.00 Bot 24.71 2.68

47 (LPG9) 0.99 0.07 143.00 147.00 Bot 127.22 1.81

48 (LPG 10) 0.63 0.045 135.00 145.00 Bot 106.56 2.75

49 (LPG 11) 0.70 0.09 130.00 127.00 Bot 138.82 2.50

50 (LPG 12) 0.61 0.09 151.00 153.00 Bot 49.28 2.87

51 (LPG 13) 0.62 0.13 135.00 142.00 Bot 92.30 2.79

52 (LPG 14) 0.71 0.34 141.00 131.00 Bot 27.72 2.45

53 (LPG 15) 0.66 0.36 158.00 163.00 Bot 2.65

54 (LPG 17) 0.68 0.50 138.00 134.00 Bot 29.43 2.57

55 (LPG 18) 0.64 0.65 157.00 150.00 Bot 2.73

56 (LPG 19) 0.67 0.81 141.00 134.00 Bot 2.62

57 (LPG 20) 0.68 1.00 135.00 133.00 Bot 2.57

58 (LPG 21) 0.72 1.05 143.00 149.00 Bot 2.42

59 (LPG 22) 0.74 1.55 136.00 140.00 Bot 2.37

60 (LPG 23) 0.74 2.69 158.00 156.00 Bot 2.34

61 (LPG 24) 0.68 1.64 150.00 146.00 Bot 2.57

62 (LPG 25)*

63 (LPG 26) 0.95 3.37 150.00 180.00 Bot 1.85

64 (LPG 27) 0.70 2.53 170.00 167.00 Bot 2.49

65 (LPG 28)*

66 (LPG 29) 0.58 1.77 163.00 162.00 Bot 2.98

67 (LPG 30) 0.67 0.91 109.00 103.00 Bot 2.61

68 (LPG 31) 0.71 0.42 120.00 120.00 Bot 2.47

69 (LPG 32) 0.55 1.20 140.00 134.00 Bot 3.15

70 (LPG 33) 0.76 2.00 140.00 139.00 Bot 2.32

71 (LPG 34) 0.74 2.22 180.00 179.50 Bot 2.35

72 (LPG 35)*

73 (LPG 36) 0.81 1.94 166.00 156.00 Bot 2.17

74 (LPG 37) 0.68 2.07 194.00 194.00 Bot 2.60

75 (LPG 38) 0.69 2.49 195.00 192.50 Bot 2.52

76 (LPG 39) 0.69 2.49 201.00 197.00 Bot 2.52

Igniter

Location

ST,abs

(m/s)Test #

O/C

Vol.

Ratio

Steam/C

Vol.

Ratio

Tctop

(oC)

Tcbottom

(oC)φ

84

Table A.10: Mixture 2 test series flow parameters

* Test failure due to various reasons (e.g., igniter malfunction or depletion of reactants).

39 (LPG1) 1.92 9.14 22.05 2.62E+04

40 (LPG 2) 1.74 10.65 23.32 2.77E+04

41 (LPG 3) 1.99 9.37 23.40 2.78E+04

42 (LPG 4)*

43 (LPG 5) 2.15 8.60 23.22 2.76E+04

44 (LPG 6)*

45 (LPG 7) 2.34 7.95 23.42 2.78E+04

46 (LPG 8) 3.18 5.92 23.68 2.81E+04

47 (LPG9) 3.06 5.31 20.45 2.43E+04

48 (LPG 10) 3.37 5.65 23.95 2.84E+04

49 (LPG 11) 3.32 5.64 23.51 2.79E+04

50 (LPG 12) 2.93 6.78 24.93 2.96E+04

51 (LPG 13) 3.33 6.03 25.18 2.99E+04

52 (LPG 14) 3.23 6.38 25.93 3.08E+04

53 (LPG 15) 3.49 6.20 27.17 3.23E+04

54 (LPG 17) 3.27 6.88 28.28 3.36E+04

55 (LPG 18) 3.36 7.33 30.99 3.68E+04

56 (LPG 19) 2.83 9.02 32.08 3.81E+04

57 (LPG 20) 3.72 3.62 16.92 2.01E+04

58 (LPG 21) 2.15 6.12 16.55 1.97E+04

59 (LPG 22) 3.75 4.00 18.86 2.24E+04

60 (LPG 23) 1.61 6.03 12.18 1.45E+04

61 (LPG 24) 1.38 5.90 10.21 1.21E+04

62 (LPG 25)*

63 (LPG 26) 1.41 6.56 11.59 1.38E+04

64 (LPG 27) 1.45 6.79 12.41 1.47E+04

65 (LPG 28)*

66 (LPG 29) 1.35 6.87 11.65 1.38E+04

67 (LPG 30) 2.98 2.19 8.22 9.76E+03

68 (LPG 31) 2.01 2.65 6.68 7.94E+03

69 (LPG 32) 2.25 3.62 10.26 1.22E+04

70 (LPG 33) 2.20 3.75 10.34 1.23E+04

71 (LPG 34) 0.96 9.23 11.11 1.32E+04

72 (LPG 35)*

73 (LPG 36) 0.50 15.48 9.75 1.16E+04

74 (LPG 37) 0.73 12.65 11.62 1.38E+04

75 (LPG 38) 0.76 13.00 12.42 1.48E+04

76 (LPG 39) 0.81 13.00 13.30 1.58E+04

Test #V

(m/s)

ρ

(kg/m^3)

m dot

(g/s)ReD

85

Table A.11: Mixture 2 test series flammability and type of pressurization

* Test failure due to various reasons (e.g., igniter malfunction or depletion of reactants).

39 (LPG1) Yes Intermediate, I

40 (LPG 2) Yes Rapid, R

41 (LPG 3) Yes Rapid, R

42 (LPG 4)*

43 (LPG 5) Yes Slow, S

44 (LPG 6)*

45 (LPG 7) Yes Rapid, R

46 (LPG 8) Yes Slow, S

47 (LPG9) Yes Rapid, R φ was Low

48 (LPG 10) Yes Rapid, R

49 (LPG 11) Yes Rapid, R

50 (LPG 12) Yes Rapid, R

51 (LPG 13) Yes Rapid, R

52 (LPG 14) Yes Rapid, R

53 (LPG 15) Yes Rapid, R

54 (LPG 17) Yes Multi-Ignitions, M

55 (LPG 18) Yes Slow, S Reverse Flame

56 (LPG 19) Yes, Low dP/dt Multi-Ignitions, M

57 (LPG 20) No

58 (LPG 21) Yes Rapid, R Low ST

59 (LPG 22) Yes, Low dP/dt Slow, S Lit in exhaust

60 (LPG 23) No

61 (LPG 24) Yes Intermediate, I Unusual dP/dt

62 (LPG 25)*

63 (LPG 26) Yes, Low dP/dt Cool Flame, C

64 (LPG 27) No

65 (LPG 28)*

66 (LPG 29) No

67 (LPG 30) No

68 (LPG 31) No

69 (LPG 32) No

70 (LPG 33) No

71 (LPG 34) No

72 (LPG 35)*

73 (LPG 36) Yes, Low dP/dt Cool Flame, C

74 (LPG 37) Yes Multi-Ignitions, M

75 (LPG 38) Cool Flame, C

76 (LPG 39) Yes, Low dP/dt Cool Flame, C

Flammable

Rate of

Pressurization

(R,I,S,C,M)

RemarksTest #

86

Table A.12: Mixture 2 test series steam parameters and orifice type

* Test failure due to various reasons (e.g., igniter malfunction or depletion of reactants).

39 (LPG1) 0.03 0.50 2.24

40 (LPG 2) 0.03 0.50 2.24

41 (LPG 3) 0.03 0.50 2.24

42 (LPG 4)* 0.03 0.50 2.24

43 (LPG 5) 0.03 0.50 2.24

44 (LPG 6)* 0.03 0.50 2.24

45 (LPG 7) 0.03 0.60 2.69

46 (LPG 8) 0.03 0.60 2.69

47 (LPG9) 0.03 0.60 2.69

48 (LPG 10) 0.03 0.60 2.69

49 (LPG 11) 0.16 4.00 17.92

50 (LPG 12) 0.16 4.00 17.92

51 (LPG 13) 0.16 4.00 17.92

52 (LPG 14) 0.16 4.00 17.92

53 (LPG 15) 0.16 8.00 35.84

54 (LPG 17) 0.16 8.00 35.84

55 (LPG 18) 0.16 8.00 35.84

56 (LPG 19) 0.16 10.00 44.80

57 (LPG 20) 0.16 10.00 44.80

58 (LPG 21) 0.16 10.00 44.80

59 (LPG 22) 0.16 10.00 44.80

60 (LPG 23) 0.16 10.00 44.80

61 (LPG 24) 0.16 10.00 44.80

62 (LPG 25)* 0.16 10.00 44.80

63 (LPG 26) 0.16 10.00 44.80

64 (LPG 27) 0.16 10.00 44.80

65 (LPG 28)* 0.16 10.00 44.80

66 (LPG 29) 0.16 10.00 44.80

67 (LPG 30) 0.16 10.00 44.80

68 (LPG 31) 0.16 10.00 44.80

69 (LPG 32) 0.16 10.00 44.80

70 (LPG 33) 0.16 10.00 44.80

71 (LPG 34) 0.16 10.00 44.80

72 (LPG 35)* 0.16 10.00 44.80

73 (LPG 36) 0.16 10.00 44.80

74 (LPG 37) 0.16 10.00 44.80

75 (LPG 38) 0.16 10.00 44.80

76 (LPG 39) 0.16 10.00 44.80

Test #ORIFICE BORE ID

(in)

MAX STEAM

(g/s)

Max Steam

(Nm3/hr)

Appendix B

Individual Test Summaries

B.1 Test Summaries with Mixture 1 Test Series

A premixed gas mixture, referred to as Mixture 1, composed of oxygen, methane,

hydrogen, ethane, and steam was investigated, in order to examine its flammability and flame-

propagation properties, by Pat Kutzler et al. at PSU in 2008 in tests 1 thru 29 [35]. Tests 30 thru

38 were run with Mixture 1 to allow the new research team to compare results to existing results

with the same overall experimental setup. The following results are for the recent test runs from

Test#34-38 with Mixture 1.

Test # 34: Ignition observed.

Test # 35: Test failed due to igniter malfunction.

Test # 36: Ignition observed.

Test # 37: Ignition observed. Pyrex tube partially sooted.

Figure 35: Glass tube liner post-test from Test #37

Test # 38: Test Failed as the Methane supply was depleted before the end of the test.

88

B.2 Test Summaries with Mixture 2 Test Series

A premixed gas mixture, referred to as Mixture 2, composed of oxygen, methane,

propane, butane, ethane, and steam was investigated in order to examine its flammability and

flame-propagation properties.

Test # 39: Ignition observed. There was a loud noise when ignition occurred that indicated high

chamber pressures were reached.

Test # 40: Ignition observed. PD5 was not aligned, therefore did not produce reliable results. The

flame propagation interval was not identified as the photodetectors did not register signals. This

failure in the photodetectors is due to windows being obstructed by debris generated from Test

#39. Even though the Chamber was cleaned after Test #39 the photodetectors were not cleaned

individually. Due to the lack of photodetector signals no flame speed was calculated for this test.

At ignition a loud noise was again associated with high pressurization of the chamber.

Test # 41: Ignition Observed. The typical rapid flow loss immediately after ignition is not

observed in this test.

Test # 42: Test failed due to erratic steam behavior.

Test # 43: Ignition observed.

Test # 44: Unplanned ignition observed. Possible autoignition. This incident led to the

development of a test procedure to slowly step up the oxygen flow.

Test # 45: Ignition observed. Loud noise heard.

Test # 46: Ignition observed. Had to clean exhaust line and exhaust valve of shattered glass from

Test # 45.

Test # 47: Ignition observed. Flame spreading was very unsteady.

89

Test # 48: Ignition observed. Very large noise oscillation seen in dynamic pressure transducer

data. Any oscillations of this magnitude were determined to be due to a loose connection at the

gauge.

Test # 49: Ignition observed.

Test # 50: Ignition observed.

Test # 51: Ignition observed. Changed to H trim on Steam "Badger" Valve to allow for more

steam flow.

Test # 52: Ignition observed. Mixer static pressure gauge failed in previous test.

Test # 53: Ignition observed. PD data indicated very turbulent, unsteady flame spreading.

Test # 54: Ignition observed.

Test # 55: Ignition observed.

Test # 56: Ignition observed. Highly oscillatory pressurization and flame spreading. This

indicates a different type of flame spreading, or combustion regime.

Test # 57: No ignition observed.

Test # 58: Ignition observed.

Test # 59: Ignition observed. Odd rate of pressurization and flame spreading was observed. As

the steam content increased, photodetector and dynamic pressure data appeared highly unsteady

as in this test.

Test # 60: No ignition observed.

Test # 61: Ignition observed.

Test # 62: No ignition observed by Vision DAQ.

Test # 63: Ignition observed.

Test # 64: No ignition observed.

Test # 65: Test failed due to igniter malfunction.

Test # 66: No ignition observed.

90

Test # 67: No ignition observed. Lower temperatures were observed during ignition due to lower

pressure decreasing the temperature of the steam which controls the mixture temperature.

Test # 68: No ignition observed. Lower temperatures during trigger initiation.

Test # 69: No ignition observed.

Test # 70: No ignition observed.

Test # 71: No ignition observed.

Test # 72: Ignition observed. Error with DAQ no recordings obtained.

Test # 73: Ignition observed. Nitrogen purge accidentally on during test.

Test # 74: Ignition observed. Burst disk failure occurred about 1 second after trigger, when flows

were shut off.

Test # 75: Ignition observed, dynamic pressure recordings failed due to operator error.

Test # 76: Ignition observed with very small pressure rise.

Appendix C

Test Data Sheet

ENI PREMIXED Test Data Sheet

Test No: Date: Participants:

TEST MATRIX EXPERIMENT NO.

PRESSURE: ________ atm(abs) __________ [psig]

FLOW CONDITION SET POINTS COMPOSITION

O2: ____[SLPM]______%mol___________Feed Press

C2H6: ____[SLPM]______%mol___________Feed Press

CH4: ____[SLPM]______%mol___________Feed Press

H2: ____[SLPM]______%mol___________Feed Press

H2O: ____[Nm3/H]_____ %mol___________Feed Press

C3H8 : ____[Nm3/H]______%mol___________Feed Press

C4H10: ____[Nm3/H]______%mol___________Feed Press

N2: ___________Purge Pressure

Location TC # Temperature [oC]

Pre-Heater 1

Steam Generator 2

Mixing Chamber 3

Chamber TOP 4

Chamber BOT 5

* Steam Generator Pressure: __________ [psig]

IGNITER INFORMATION

Type: Electric Match h

Match Resistance: ____________

Chamber Location: TOP or BOTTOM

NI DAQ DATA File :____________________

Nicolet Vision Instrumentation Table

Channel Description

1 PCB (Chamber Bottom)

2 N/A

3 Photodetector #1 (Not in Use)

4 Photodetector #2

5 Photodetector #3

6 Photodetector #4

7 Photodetector #5

8 Photodetector #6

9 Photodetector #7

10 Photodetector #8

11 Photodetector #9

12 Photodetector #10

13 Photodetector #11

14 Photodetector #12

15 PCB (Chamber Top)

16 Trigger

NI DAQ Computer System

Channel Description

0 Exhaust Line Pressure

1 Mixer Pressure

2 Chamber Top Pressure

3 Chamber Bottom Pressure

4 O2 Flow Rate

5 CH4 Flow Rate

6 C2H6 Flow Rate

7 H2 Flow Rate

8 C3H8 Flow Rate

9 C4H10 Flow Rate

10 STEAM Flow Rate

11 Trigger

12 STEAM PRESSURE

COMMENTS:

Appendix D

Premixed Gas Reactor Test Checklist

TEST NO.: DATE: OPERATORS:

Pre-Test Procedure

Photodetector box

Setra box

Validyne box

PCB signal conditioners

Teledyne-Hastings mass flow controller

Siemens mass flow meter power box

Load new Pyrex tube into chamber connecting the heat exchanger and exhaust line

install new igniter in bottom plate

Verify steam generator and pre-heater have sufficient water levels

Verify that the gas bottles have sufficient pressures to run the test

Verify the shop air pressure is around 90 psig and the air supply for the I/P controller and

Badger control valve is set to 22 psig

Make sure that igniter is electrically continuous and not grounded to the chamber

Make sure that the igniter is not connected to the igniter extension cord

Check that the transducer, photodetector, and TC probe wires are connected to the correct

channels in the data acquisition systems

Confirm Methane Plug Valve is open to Chamber

Prepare video camera systems for remotely monitoring the test facility and steam

pressure

Prepare Data Acquisition Systems (NI and Vision)

93

Test Procedure

Display Testing Signs on Room 128 doors

Turn on the steam generator well in advance of test time. Set Controller on Panel to 200

psig

Slightly Open the Vent Valve on the Steam Generator for Safe Heating Procedure

Open valve on water line to supply water to the exhaust heat exchanger

Main operator should keep the control panel arming key in his pocket

Check igniter resistance and continuity to ensure it is unfired

Turn on the Preheater when The Steam Generator Pressure is around 100 psig.

Preheat reactor with N2 flow and preheat steam line until suitable temperatures are

reached

Open N2 Cluster and confirm valves in Compressor Room are directing flow to lab

Run N2 flow to purge system of any condensation in the reactor from the steam flow

Once Chamber Temperatures are at 40 C and Steam Generator at 200 psig, Set steam

pressure controller to 450 psig

Open Steam Ball Valve to flow to chamber and Vent out water in lines

Set steam flow to 20-40% to get steam to start flowing through chamber

Prepare Fuel System

Open Ethane and set Regulator

Methane (& N2 for regulator) hold on setting regulator

Prepare LPG system

Close Plug Valves to Test Chamber

Close N2 fill valve to Butane

Open N2 fill valve to Propane

Open N2 Bottle, Set 2-way valve to fill

Pressurize Propane to 400 psig (Listen for piston to move to top)

Open Propane Bottle

94

Set 2-way valve to vent and slightly open vent gate valve as propane fills (5

mins)

Confirm Propane Piston is at bottom (no more venting)

Close N2 fill valve to Propane

Open N2 fill valve to Butane

Remove Butane plug above fill tank (Allow to fully vent)

Attach fitting and N2 fill line to push piston to bottom

Reassemble and fill Butane by Opening the Butane Bottle

Close N2 fill valve to Butane (Hold on setting regulator)

Do Final Visual Check of Test Cell

Turn off Pre-heater, Close N2 preheat bottle, let N2 flow out, close solenoid valve

Open oxygen bottles and keep regulator around 500 psig

Make announcement of the upcoming test using microphone and Public Address system

Test operators should go to test stations and assign the duties to all participants

Bring the flow rates of the gases and steam to the pre-selected values by setting the mass

flow controller channels and LabVIEW-controlled valves

Open Fuel Solenoid Valve

Set flow of Ethane

Set flow of Methane, coordinate with person setting methane regulator

Pressurize LPG flow system, open Valves to Test chamber

Set flow of Oxygen, coordinate with N2 Flow

Set flow of Steam

Set flow of LPG

Turn off N2 diluent flow

Bring the chamber to pre-selected pressure by slowly adjusting the exhaust valve on the

control area wall

Begin Recording with Vision DAQ System

Activate NI DAQ (Must hit trigger within 2 minutes of this step)

Record Temperatures

Turn Igniter safety key switch to ARMED (RED LIGHT) position

Countdown the test firing and activate the igniter switch

95

Post-test Procedure

Open the exhaust valve fully to depressurize the test chamber

Immediately set all flows to zero on the mass flow controller channels

Open N2 purge valve to discharge residual gases (allow N2 gas to flow until chamber is

cleared of exhaust gases)

Turn off power supply to low-pressure pre-heater and high-pressure steam

generator

Close all the gas bottle valves and turn off valve in water line to heat exchanger

Save the data files and copy to the appropriate HPCL network project folder

Disassemble chamber and prepare for next test run

Remove Test Signs on Room 128 doors

REMARKS:

96

Appendix E

Error and Uncertainty

The error analysis was defined as by Kline and McClintok’s [36] definition of

uncertainty:

Where Q is the dependent parameter, v is the independent parameter, and is the

uncertainty associated with the independent variable. If the uncertainties of the instruments are

known this formula can be used to calculate the uncertainty in the instrument. The equivalence

ratio error is of great importance as small variations in this parameter can cause a error in

determining the flammability. The following shows how the error of the equivalence ratio was

calculated for each test:

97

Where the following parameters are constant for all tests from manufacturer:

Using these parameters provided by the manufacturers and the flow data from each test a for

each test can be determined and an error is given by the following equation:

The uncertainty in each test was calculated and the uncertainty for each flow group was taken

from the average of those tests. This uncertainty analysis method could be used for any of the

data in this research. All uncertainties are accounted for in Chapter 3.

98

Appendix F

General Calculations

µ was approximated to be the value for oxygen at 450 K. This approximation gives an order of

magnitude but is not an accurate approximation as the mixture is not pure oxygen and the mixture

is not always at 450 K. The density of the mixture was calculated using Refprop software [37].