Non-Catalytic Production of Hydrogen Via Reforming of Diesel, Hexadecane and Bio-Diesel for Nitrogen Oxides Remediation

Home , Hexadecane

NON-CATALYTIC PRODUCTION OF HYDROGEN VIA REFORMING OF DIESEL, HEXADECANE AND BIO-DIESEL FOR NITROGEN OXIDES REMEDIATION

DISSERTATION

Presented in Partial Fulﬁllment of the Requirements for

the Degree Doctor of Philosophy in the

Graduate School of The Ohio State University

Sergio Manuel Hernandez Gonzalez, M.S., B.S.

*****

The Ohio State University

2008

Dissertation Committee: Approved by

Professor Yann Guezennec, Adviser Professor Vish Subramaniam, Co-adviser Adviser Professor Giorgio Rizzoni Graduate Program in Dr. Shawn W. Midlam-Mohler Mechanical Engineering Professor Junmin Wang © Copyright by

Sergio Manuel Hernandez Gonzalez

2008 ABSTRACT

After-treatment technologies are required for diesel engines to meet the current and

future stringent emissions regulations. Lean NOx traps and SCR catalysts represent

the major routes for after-treatment for NOx mitigations under lean exhaust conditions.

These technologies require active agents (H2, CO and several hydrocarbons) that ei-

ther participate in nitrogen oxide reduction reactions or regenerate the NOx storage sites. Nevertheless, these species have to be obtained from either the on-board fuel or additional sources such as urea. Hydrocarbons are seen thus far as the most reliable source for generation of hydrogen. This work focuses then on the generation of hydrogen through Partial Oxidation of heavy hydrocarbons.

The research is ﬁrst oriented to assess numerically, using a proprietary code, the feasibility of the non-catalytic reforming of hexadecane (C16H34 a heavy hydrocarbon molecule used as a proxy for diesel fuel) under diﬀerent conditions of equivalence ratio, steam to carbon ratio, and inlet temperatures. The reforming process is then analyzed within the characteristic quantities ranges: 1 ≤ φ ≤ 1.9, 0 ≤ S/C ≤ 2, 1000 ≤ Tini ≤

1750 K. A novel contribution to the scientiﬁc community is the assessment of addition of water through secondary injection to improve the hydrogen production. Of special interest for automotive applications is the use of exhaust as a possible source for oxygen since diesel engines are operated under lean conditions. Such a concept is also studied

ii numerically in this dissertation. A discussion on the kinetic mechanisms producing

and consuming hydrogen in the partial oxidation of hexadecane is also included in this

dissertation.

Based upon the results obtained in the numerical simulation, a proof of concept

of the POx of hexadecane is experimentally performed. Diesel fuel POx reforming is

also experimentally studied since this concept is aimed to be utilized for diesel engine

after-treatment. Bio-diesel has become very popular due to its chemical conﬁguration

containing simple chains and its lower pollution characteristic and its renewability as a

bio-fuel. Hence, this research presents a novel study of partial oxidation of bio-diesel

(B-100) to generate hydrogen or syngas.

The results obtained using the 0D model strongly indicate that H2, CO and total hydrocarbons concentrations increase with increasing equivalence ratio (φ) for all temperatures and both sources of oxygen (air and lean exhaust). The three fuels tested experimentally showed an increasing H2 concentration with φ as well. Nevertheless,

H2 and CO saturate due to the decreasing adiabatic ﬂame temperature (Tad) with in-

creasing equivalence ratio. Vapor addition slightly increases the H2 % Vol for low φ

ratios, while it caused the H2 concentration to saturate at a faster rate as φ increased.

Increasing Tini yields higher H2 concentrations owing to the higher Tad for all φ and

S/C ratios. High H2 yields ( 25-30%) can be obtained for φ ≥ 1.6 for the low tem- ¬ perature case (1000 K), while even higher yields ( 30-40%) are seen for the high inlet ¬ temperature cases (1750 K). Addition of vapor along with the main feeds is beneﬁcial

at high inlet temperatures, while secondary injection of water showed a very slight

increase in H2 concentration even if injected at high temperatures. However, as S/C

is increased, a slower H2 saturation rate is seen for post-injection of vapor than for

iii its equivalent main injection of vapor. Lower product composition are obtained using

exhaust gas as O2 provider due to the lower ﬂame temperature caused by the amount of present diluent in the stream.

In summary, this study has demonstrated the feasibility of producing hydrogen-rich syngas from the partial oxidation of diesel fuel in a simple, practically implementable device which could be integrated into an on-board, after-treatment system for diesel engine vehicles. The system is capable of hydrogen yields consistent with the NOx

reduction needs of current engines and does not require any additional ﬂuids besides

the engine fuel. The detailed 0D kinetics simulations in this study have served to

understand the chemical kinetic mechanisms at play and, while not truly predictive,

to provide invaluable design guidelines for practical implementations, particularly with

respect to operating point equivalence ratio and the dominant role of temperature in

maximizing the hydrogen yield.

iv ACKNOWLEDGMENTS

I would like to extend my gratitude to my adviser Dr. Yann Guezennec, for his contributions throughout the research, to my co-adviser Dr. Vish Subramaniam for our discussions on chemical kinetics and everything referent to the numerical simulation, and specially to Dr. Shawn Midlam-Mohler for his constant interest, comments and revisions throughout this work. Furthermore, I would like to thank the professors of the Mechanical Engineering Department who took me under their guidance during these years. I am indelibly grateful to Ranjit Annamalai and Zhijun Zou for providing many of the tools to develop this research. Special appreciation is also given to Kenny Follen for his constant advice on the development of this work and availability for language corrections, and mostly on my daily life, thanks for being an excellent friend.

It is appropriate to thank Tenneco Inc. and the team that ﬁnancially sponsored this project for giving me the opportunity to enjoy and grow professionally through this experience. I also thank the National Council of Science and Technology of Mexico

(CONACyT, for its acronym in Spanish) for its sponsorship to carry out this research for the ﬁrst years of the program.

Special thanks are delivered to the staﬀ at CAR for providing me with a nice environment and the facilities to develop this work. Specially to Mr. Don Williams who is an excellent machinist who not only understood every single drawing I produced but

v also provided advice to enhance it. My fellow graduate and undergraduate students that contributed directly or indirectly on this project with ideas, thoughts, laughs and gatherings: Dr. Marcello Canova, Simone Bernasconi, Josh Supplee, Christopher Hoops and Orlando Inoa.

I would like to extend my deepest thanks to my friends who have contributed much towards the enjoyment of my time in Columbus. It would be unfair to mention a few and leave others out of the list, everyone has given me something very special. Special gratitude is extended to Joana Ferreti with whom I shared a big part of this time and who always motivated me to continue on the search of answers in every aspect of my life.

Finally and most important, I would like to strongly thank my family who has been there in every single moment of this journey for its patience and understanding.

Thanks to my parents who provided me with the means and curiosity to pursue education and embrace my professional career. Also, my brother and sister who listened to my ideas always. They know the journey has not been easy on the personal side but they were always accompanying me no matter the time of the day or night. Last but not least, special gratitude is extended to my uncle Abel Hernandez who always encouraged me to go beyond limitations and to pursue advanced education.

vi VITA

October 8, 1978 ...... Born - Salamanca, Guanajuato, Mexico 2001 ...... B.S. Mechanical Engineering Universidad de Guanajuato Guanajuato, Mexico 2003 ...... M.S. Mechanical Engineering University of Manchester Institute of Science and Technology Manchester, United Kingdom 2004 to present ...... Graduate Research Associate Mechanical Engineering The Ohio State University Center for Automotive Research United States of America

FIELDS OF STUDY

Major Field: Mechanical Engineering

vii TABLE OF CONTENTS

Page

Abstract ...... ii

Acknowledgments ...... v

Vita ...... vii

List of Figures ...... xiv

List of Tables ...... xxii

Chapters:

1. Introduction ...... 1

1.1 Introduction ...... 1 1.2 Motivation ...... 2 1.3 Thesis Outline ...... 4

2. Literature Review ...... 8

2.1 Introduction ...... 8 2.2 After-Treatment Technologies ...... 10

2.2.1 Lean NOx Traps, LNT ...... 13 2.2.2 Hydrocarbon-Selective Catalytic Reaction Systems, HC-SCR . 18

viii 2.2.3 Urea/Ammonia-Selective Catalytic Reaction, Urea/NH3-SCR 23

2.2.4 Hydrogen-Selective Catalytic Reaction, H2-SCR ...... 30

2.2.5 Plasma Assisted Catalytic Removal of NOx ...... 35 2.3 Fuel Reforming Technologies ...... 36 2.3.1 Catalytic Reforming ...... 38 2.3.1.1 Steam Reforming, SR ...... 39 2.3.1.2 Partial Oxidation Reforming, POx ...... 40 2.3.1.3 Autothermal Reforming, ATR ...... 43 2.3.2 Characterization of the Catalytic Reforming Process ..... 45 2.3.2.1 Reforming Temperature ...... 45

2.3.2.2 Steam-to and Oxygen-to Carbon Ratios, S/C & O2/C 48 2.3.3 Non-Catalytic Reforming Processes (Thermal POx) ...... 52 2.3.4 Characterization of Non-Catalytic Processes ...... 53 2.3.4.1 Temperature POx ...... 54 2.3.4.2 Equivalence Ratio and Steam to Carbon Ratio .... 57 2.3.5 Plasma Reforming ...... 59 2.3.6 Cleaning Up Processes ...... 62 2.3.6.1 Desulfurization ...... 63 2.3.6.2 Water Gas Shift Reaction ...... 65 2.3.6.3 Methanation ...... 66 2.3.6.4 Selective Membranes ...... 67 2.3.6.5 Selective or Preferential Oxidation ...... 68 2.3.6.6 Coke Formation ...... 70 2.4 Concluding Remarks ...... 71

3. Chemical Kinetics Modeling ...... 74

3.1 Introduction ...... 74 3.2 Problem Formulation ...... 76 3.2.1 Governing Equations ...... 78 3.2.2 Initial Temperature of the Feeds ...... 81

ix 3.2.3 Initial Concentrations when Using Pure Air ...... 85 3.2.4 Initial Concentrations when Using Exhaust Gas ...... 89 3.3 Deﬁnition of Hydrogen Yield ...... 94 3.4 Concluding Remarks ...... 96

4. Chemical Kinetics of Hydrogen Production/Consumption ...... 98

4.1 Introduction ...... 98 4.2 Methodology ...... 99 4.3 Eﬀect of Equivalence Ratio (φ) ...... 104 4.3.1 Pre-Combustion Zone ...... 105 4.3.2 Combustion Zone ...... 108 4.3.3 Cooling Zone ...... 114 4.4 Eﬀect of Steam to Carbon Ratio (S/C) ...... 117 4.4.1 Pre-Combustion Zone ...... 119 4.4.2 Combustion Zone ...... 121 4.4.3 Cooling Zone ...... 126

4.5 Eﬀect of Initial Temperature (Tini) ...... 128 4.5.1 Combustion Zone ...... 129 4.5.2 Cooling Zone ...... 138 4.6 Eﬀect of Initial Temperature and S/C ...... 139 4.6.1 Combustion Zone ...... 141 4.6.2 Cooling Zone ...... 146

4.7 Eﬀect of S/C Post-Injection at TPI = 2000 K ...... 147 4.7.1 Rapid Cooling ...... 149 4.7.2 Post-Injection Zone ...... 150 4.7.3 Cooling Zone ...... 151 4.8 Concluding Remarks and Comparisons ...... 152 4.8.1 Equivalence Ratio (φ = 1.6) and S/C = 1.0 Ratio Cases at Low

Tini ...... 153 4.8.1.1 Pre-Combustion ...... 153

x 4.8.1.2 Combustion ...... 154 4.8.1.3 Cooling ...... 155

4.8.2 Equivalence Ratio (φ = 1.6) cases at Low and High Tini ... 156 4.8.2.1 Combustion ...... 157 4.8.2.2 Cooling ...... 159 4.8.3 Equivalence Ratio (φ = 1.6) and S/C = 1.0 Ratio Cases at

High Tini ...... 159 4.8.3.1 Combustion ...... 160 4.8.3.2 Cooling ...... 162

4.8.4 S/C = 1.0 Ratio Cases at Low and High Tini ...... 162 4.8.4.1 Combustion ...... 163 4.8.4.2 Cooling ...... 165

4.8.5 S/C = 1.0 Ratio and Post-Injection (P.I.) Cases at Tini = 1000

K and TPI = 2000 K ...... 165

5. Chemical Equilibrium Compositions for Reforming of Hexadecane ..... 168

5.1 Introduction ...... 168 5.2 Effect of Equivalence Ratio at Low Temperature ...... 169 5.3 Effect of the Steam to Carbon Ratio at Low Temperature ...... 186 5.4 Effect of the Initial Temperature for φ and S/C ...... 191 5.5 Effect of Pressure for φ and S/C ...... 199 5.6 Effect of Post-Injection of Water Vapor for φ and S/C ...... 202 5.7 Effect of Exhaust as an Oxygen Provider ...... 209 5.8 Concluding Remarks ...... 215

6. Experimental Set up and Testing Procedures ...... 220

6.1 Introduction ...... 220 6.2 Experimental Apparatus & Reformer Designs ...... 222 6.2.1 Dilution and Measurement Systems ...... 227

xi 6.3 Testing Procedure ...... 230 6.4 Experimental Hydrogen Yield Deﬁnition ...... 235 6.5 Speciation for Hydrocarbon Composition ...... 236 6.6 Concluding Remarks ...... 238

7. Experimental Results ...... 240

7.1 Introduction ...... 240 7.2 Liquid Speciation ...... 242 7.3 Eﬀect of the Equivalence Ratio ...... 243 7.3.1 Hexadecane Fuel ...... 244 7.3.2 Diesel Fuel ...... 252 7.3.3 Bio-Diesel Fuel ...... 257 7.3.4 Experimental Results for Diesel, Hexadecane and Bio-Diesel . 263 7.4 Comparison of the Hydrocarbon Species obtained Experimentally and Those Calculated by the 0D Model ...... 266 7.5 Concluding Remarks ...... 270

8. Conclusions and Recommendations for Further Work ...... 272

8.1 Conclusions ...... 272 8.2 Further Work ...... 280

Appendices:

A. Chemical Reactions and Rates ...... 283

B. Speciation of Liquid Fuels ...... 303

C. Speciation of Reformate ...... 306

xii C.1 Hexadecane Cases ...... 306 C.2 Diesel Cases ...... 312 C.3 Bio-Diesel Cases ...... 317

Bibliography ...... 322

xiii LIST OF FIGURES

Figure Page

1.1 Schematic of a H2-SCR System with a Thermal Partial Oxidation Reformer 4

2.1 Product Composition as a Function of Temperature for IGO POx at λ = 0.4 - 0.46 [69] ...... 56

3.1 Fuel Decomposition and Major Species at 900 K and 1 atm ...... 83

3.2 Fuel Decomposition and Major Species at 1000 K and 1 atm ...... 83

3.3 Time History of the Temperature and Concentrations for φ = 1.0, S/C

= 0.0, 1 atm and Tini = 900 K ...... 84

4.1 Hydrogen Concentration and Temperature History for φ = 1.6, S/C =

0.0, P = 1 atm and Tini = 1000 K ...... 100

4.2 Hydrogen Number Density Gradients: Total, Reaction Term and Vol-

ume Term for φ = 1.6, S/C = 0.0, P = 1 atm and Tini = 1000 K at the Pre-Combustion Region ...... 108

4.3 Hydrogen Number Density and its Total Derivative for φ = 1.6, S/C =

0.0, P = 1 atm and Tini = 1000 K ...... 109

4.4 Hydrogen Number Density Gradients: Total, Reaction Term and Vol-

ume Term for φ = 1.6, S/C = 0.0, P = 1 atm and Tini = 1000 K at the Combustion Region ...... 110

4.5 Hydrogen Number Density Gradients: Total, Reaction Term and Vol-

ume Term for φ = 1.6, S/C = 0.0, P = 1 atm and Tini = 1000 K at the Cooling Region ...... 115

4.6 Hydrogen Concentration and Temperature History for φ = 1.6, S/C =

1.0, P = 1 atm and Tini = 1000 K ...... 118

xiv 4.7 Hydrogen Number Density Gradients: Total, Reaction Term and Vol-

ume Term for φ = 1.6, S/C = 1.0, P = 1 atm and Tini = 1000 K at the Pre-Combustion Region ...... 120

4.8 Hydrogen Number Density and its Total Derivative for φ = 1.6, S/C =

1.0, P = 1 atm and Tini = 1000 K ...... 122

4.9 Hydrogen Number Density Gradients: Total, Reaction Term and Vol-

ume Term for φ = 1.6, S/C = 1.0, P = 1 atm and Tini = 1000 K at the Combustion Region ...... 123

4.10 Hydrogen Number Density Gradients: Total, Reaction Term and Vol-

ume Term for φ = 1.6, S/C = 1.0, P = 1 atm and Tini = 1000 K at the Cooling Region ...... 127

4.11 Hydrogen Concentration and Temperature History for φ = 1.6, S/C =

1.0, P = 1 atm and Tini = 1750 K ...... 129

4.12 Hydrogen Concentration and Temperature History for φ = 1.6, S/C =

1.0, P = 1 atm and Tini = 1750 K ...... 130

4.13 Hydrogen Number Density and its Total Derivative for φ = 1.6, S/C =

0.0, P = 1 atm and Tini = 1750 K ...... 131

4.14 Hydrogen Number Density and its Total Derivative for φ = 1.6, S/C =

0.0, P = 1 atm and Tini = 1750 K ...... 131

4.15 Hydrogen Number Density Gradients: Total, Reaction Term and Vol-

ume Term for φ = 1.6, S/C = 0.0, P = 1 atm and Tini = 1750 K ... 133

4.16 Hydrogen Number Density Gradients: Total, Reaction Term and Vol-

ume Term for φ = 1.6, S/C = 0.0, P = 1 atm and Tini = 1750 K ... 134

4.17 Hydrogen Number Density Gradients: Total, Reaction Term and Vol-

ume Term for φ = 1.6, S/C = 0.0, P = 1 atm and Tini = 1750 K ... 139

4.18 Hydrogen Concentration and Temperature History for φ = 1.6, S/C =

1.0, P = 1 atm and Tini = 1000 K ...... 140

4.19 Hydrogen Concentration and Temperature History for φ = 1.6, S/C =

1.0, P = 1 atm and Tini = 1000 K ...... 141

xv 4.20 Hydrogen Number Density and its Total Derivative for φ = 1.6, S/C =

1.0, P = 1 atm and Tini = 1750 K ...... 142

4.21 Hydrogen Number Density and its Total Derivative for φ = 1.6, S/C =

1.0, P = 1 atm and Tini = 1750 K ...... 143

4.22 Hydrogen Number Density Gradients: Total, Reaction Term and Vol-

ume Term for φ = 1.6, S/C = 1.0, P = 1 atm and Tini = 1750 K ... 143

4.23 Hydrogen Number Density Gradients: Total, Reaction Term and Vol-

ume Term for φ = 1.6, S/C = 1.0, P = 1 atm and Tini = 1750 K ... 144

4.24 Hydrogen Number Density Gradients: Total, Reaction Term and Vol-

ume Term for φ = 1.6, S/C = 1.0, P = 1 atm and Tini = 1750 K at the Cooling Region ...... 147

4.25 Hydrogen Concentration and Temperature History for φ = 1.6, S/C =

1.0, P = 1 atm and Tini = 1000 K ...... 148

4.26 Hydrogen Number Density Gradients: Total, Reaction Term and Vol-

ume Term for φ = 1.6, S/C = 1.0, P = 1 atm, Tini = 1000 K and TPI = 2000 K at the Post-Injection Region ...... 150

4.27 Hydrogen Number Density Gradients: Total, Reaction Term and Vol-

ume Term for φ = 1.6, S/C = 1.0, P = 1 atm, Tini = 1000 K and TPI = 2000 K at the Cooling Region ...... 152

5.1 H2, CO, H2O and CO2 Equilibrium Concentrations for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 298 K based on [95, 96] 171

5.2 Equilibrium Adiabatic Flame Temperature for Hexadecane at Diﬀerent

φ and S/C Ratios at 1 atm and Tini = 298 K based on [95, 96] ..... 171

5.3 H2, CO, H2O and CO2 Equilibrium Concentrations for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1000 K based on [97] . 172

5.4 Equilibrium Adiabatic Flame Temperature for Hexadecane at Diﬀerent

φ and S/C Ratios at 1 atm and Tini = 1000 K based on [97] ...... 172

5.5 Adiabatic Flame Temperature for Hexadecane at Diﬀerent φ and S/C

Ratios at 1 atm and Tini = 1000 K ...... 173

xvi 5.6 H2 Concentration for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1000 K ...... 175

5.7 CO Concentration for Hexadecane at Diﬀerent φ and S/C Ratios at 1

atm and Tini = 1000 K ...... 178

5.8 CO2 Concentration for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1000 K ...... 178

5.9 THC Concentration for Hexadecane at Diﬀerent φ and S/C Ratios at 1

atm and Tini = 1000 K ...... 179

5.10 H2O Concentration for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1000 K ...... 179

5.11 H2 Yield for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1000 K ...... 181

5.12 Dry-N2-free H2 Concentration for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1000 K ...... 182

5.13 Ignition Delay for Hexadecane for Diﬀerent φ Ratios at S/C = 0, 1 atm

and Tini = 1000 K ...... 185

5.14 Ignition Delay for Hexadecane for Diﬀerent φ Ratios at S/C = 0, 1 atm

and Tini = 1000 K ...... 185

5.15 H2 Concentration for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1000 K ...... 188

5.16 Ignition Delay for Hexadecane for Diﬀerent S/C Ratios at φ = 1.6, 1

atm and Tini = 1000 K ...... 188

5.17 H2 Yield for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1000 K based on Fuel and Steam ...... 190

5.18 Adiabatic Flame Temperature for Hexadecane at Diﬀerent φ and S/C

Ratios at 1 atm and Tini = 1750 K ...... 193

5.19 H2 Concentration for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1500 K ...... 193

xvii 5.20 H2 Concentration for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1750 K ...... 194

5.21 H2 Yield for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1500 K ...... 194

5.22 H2 Yield for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1750 K ...... 195

5.23 H2 Yield for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1500 K based on Fuel and Steam ...... 195

5.24 H2 Yield for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1750 K based on Fuel and Steam ...... 196

5.25 Dry-N2-free H2 Concentration for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1750 K ...... 196

5.26 CO Concentration for Hexadecane at Diﬀerent φ and S/C Ratios at 1

atm and Tini = 1750 K ...... 198

5.27 CO2 Concentration for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1750 K ...... 198

5.28 THC Concentration for Hexadecane at Diﬀerent φ and S/C Ratios at 1

atm and Tini = 1750 K ...... 199

5.29 Equilibrium Adiabatic Flame Temperature for Hexadecane at Diﬀerent

φ and S/C Ratios at 2 atm and Tini = 1000 K based on [97] ...... 201

5.30 Kinetics Adiabatic Flame Temperature for Hexadecane at Diﬀerent φ at

2 atm and Tini = 1000 K ...... 201

5.31 H2 Concentration for Hexadecane at Diﬀerent φ and Post-Injection S/C Ratios at 1 atm and TPI = 2000 K ...... 206

5.32 H2 Yield for Hexadecane at Diﬀerent φ and Post-Injection S/C Ratios at 1 atm and TPI = 2000 K ...... 206

5.33 H2 Yield for Hexadecane at Diﬀerent φ and Post-Injection S/C Ratios at 1 atm and TPI = 2000 K based on Fuel and Vapor ...... 207

xviii 5.34 CO Concentration for Hexadecane at Diﬀerent φ and Post-Injection S/C

Ratios at 1 atm and TPI = 2000 K ...... 207

5.35 CO2 Concentration for Hexadecane at Diﬀerent φ and Post-Injection S/C Ratios at 1 atm and TPI = 2000 K ...... 208

5.36 THC Concentration for Hexadecane at Diﬀerent φ and Post-Injection

S/C Ratios at 1 atm and TPI = 2000 K ...... 208

5.37 Temperature History for Hexadecane at Diﬀerent O2% Exhaust for φ = 1.0 at 1 atm and Tini = 1000 K with no Cooling ...... 210

5.38 Temperature History for Hexadecane at Diﬀerent O2% Exhaust for φ = 1.6 at 1 atm and Tini = 1000 K with no Cooling ...... 211

5.39 H2 Concentration for Hexadecane at Diﬀerent O2% Exhaust at 1 atm and Tini = 1000 K with Cooling ...... 213

5.40 H2 Yield for Hexadecane at Diﬀerent O2% Exhaust at 1 atm and Tini = 1000 K with Cooling ...... 213

5.41 CO Concentration for Hexadecane at Diﬀerent O2% Exhaust at 1 atm and Tini = 1000 K with Cooling ...... 214

5.42 CO2 Concentration for Hexadecane at Diﬀerent O2% Exhaust at 1 atm and Tini = 1000 K with Cooling ...... 214

5.43 THC Concentration for Hexadecane at Diﬀerent O2% Exhaust at 1 atm and Tini = 1000 K with Cooling ...... 215

6.1 Schematic of the Reforming and Dilution System ...... 223

6.2 Thermal Partial Oxidation Reformer Version 1 ...... 224

6.3 Thermal Partial Oxidation Reformer Version 2 ...... 225

6.4 Thermal Partial Oxidation Reformer Version 3 ...... 226

6.5 Diluted Concentrations and Temperature History with Time [s] for Hex- adecane at φ = 1.5 ...... 234

xix 6.6 Non-Diluted Wet Concentrations and Temperature History with Time [s] for Hexadecane at φ = 1.5 ...... 234

6.7 CO2 Concentration for Hexadecane at Diﬀerent φ ...... 238

7.1 Temperature for Hexadecane at Diﬀerent Equivalence Ratios ...... 244

7.2 Adiabatic Flame Temperature for Hexadecane at Diﬀerent φ and S/C

Ratios at 1 atm and Tini = 1000 K ...... 245

7.3 H2 Concentration for Hexadecane at Diﬀerent Equivalence Ratios ... 249

7.4 H2 Yield for Hexadecane at Diﬀerent Equivalence Ratios ...... 249

7.5 CO Concentration for Hexadecane at Diﬀerent Equivalence Ratios ... 250

7.6 CO2 Concentration for Hexadecane at Diﬀerent Equivalence Ratios .. 250

7.7 CO Concentration for Hexadecane at Diﬀerent Equivalence Ratios ... 251

7.8 CO2 Concentration for Hexadecane at Diﬀerent Equivalence Ratios .. 251

7.9 THC Concentration for Hexadecane at Diﬀerent Equivalence Ratios .. 252

7.10 Temperature for Diesel at Diﬀerent Equivalence Ratios ...... 254

7.11 H2 Concentration for Diesel at Diﬀerent Equivalence Ratios ...... 254

7.12 H2 Yield for Diesel at Diﬀerent Equivalence Ratios ...... 255

7.13 CO Concentration for Diesel at Diﬀerent Equivalence Ratios ...... 255

7.14 CO2 Concentration for Diesel at Diﬀerent Equivalence Ratios ...... 256

7.15 THC Concentration for Diesel at Diﬀerent Equivalence Ratios ..... 256

7.16 Temperature for Bio-Diesel at Diﬀerent Equivalence Ratios ...... 257

7.17 H2 Concentration for Bio-Diesel at Diﬀerent Equivalence Ratios .... 258

7.18 H2, CO and CO2 Equilibrium Concentrations for Bio-diesel at Diﬀerent φ at 1 atm ...... 259

xx 7.19 H2 Yield for Bio-Diesel at Diﬀerent Equivalence Ratios ...... 260

7.20 CO Concentration for Bio-Diesel at Diﬀerent Equivalence Ratios .... 261

7.21 CO2 Concentration for Bio-Diesel at Diﬀerent Equivalence Ratios ... 262

7.22 THC Concentration for Bio-Diesel at Diﬀerent Equivalence Ratios ... 262

7.23 Temperature for Hexadecane, Diesel and Bio-Diesel at Diﬀerent Equiv- alence Ratios ...... 263

7.24 H2 Concentration for Hexadecane, Diesel and Bio-Diesel at Diﬀerent Equivalence Ratios ...... 264

7.25 H2 Yield for Hexadecane, Diesel and Bio-Diesel at Diﬀerent Equivalence Ratios ...... 264

7.26 CO Concentration for Hexadecane, Diesel and Bio-Diesel at Diﬀerent Equivalence Ratios ...... 265

7.27 CO2 Concentration for Hexadecane, Diesel and Bio-Diesel at Diﬀerent Equivalence Ratios ...... 265

7.28 THC Concentration for Hexadecane, Diesel and Bio-Diesel at Diﬀerent Equivalence Ratios ...... 266

B.1 Liquid Speciation: GC-MS Diesel Fuel BP ECD-1 ...... 304

B.2 Liquid Speciation: GC-MS Bio-Diesel Fuel B-100 ...... 305

xxi LIST OF TABLES

Table Page

3.1 Molar Coeﬃcients for the general Auto Thermal Equation (3.22) .... 86

3.2 Exhaust Conditions for 1.9 L Euro4 Diesel Engine ...... 90

3.3 Molar Coeﬃcients for Engine Combustion ...... 91

3.4 Reforming Inlet Concentrations (Vol%) for φ = 1.3 ...... 93

4.1 Hydrogen Producing Reactions in the Pre-Ignition Section ...... 107

4.2 Hydrogen Consuming Reactions in the Pre-Ignition Section ...... 107

4.3 Hydrogen Producing Reactions in the Combustion Section ...... 112

4.4 Hydrogen Consuming Reactions in the Combustion Section ...... 113

4.5 Hydrogen Producing Reactions in the Pre-Ignition Section ...... 121

4.6 Hydrogen Consuming Reactions in the Pre-Ignition Section ...... 121

4.7 Hydrogen Producing Reactions in the Combustion Section ...... 125

4.8 Hydrogen Consuming Reactions in the Combustion Section ...... 126

4.9 Hydrogen Producing Reactions in the Ignition Section ...... 137

4.10 Hydrogen Consuming Reactions in the Ignition Section ...... 137

4.11 H2 Producing Reactions in the Pre-Ignition Section for φ = 1.6 and S/C = 1.0 at Tini = 1000 K ...... 153

xxii 4.12 H2 Producing Reactions in the Combustion Section for φ = 1.6 and S/C = 1.0 at Tini = 1000 K ...... 154

4.13 H2 Consuming Reactions in the Pre-Ignition & Ignition Sections for φ = 1.6 and S/C = 1.0 at Tini = 1000 K ...... 155

4.14 H2 Producing Reactions in the Ignition Section for φ = 1.6 and S/C = 1.0 at Tini = 1750 K ...... 161

4.15 Hydrogen Consuming Reactions in the Ignition Section for φ = 1.6 and

S/C = 1.0 at Tini = 1750 K ...... 161

4.16 H2 Producing Reactions in the Ignition Section for φ = 1.6 and S/C = 1.0 at Tini = 1750 K ...... 163

4.17 Hydrogen Consuming Reactions in the Ignition Section for φ = 1.6 and

S/C = 1.0 at Tini = 1750 K ...... 164

5.1 H2 Producing Reactions ...... 176

5.2 H2 Consuming Reactions ...... 176

6.1 Fuels, Denomination and Hydrogen to Carbon Ratio ...... 221

7.1 Experimental Hydrocarbon Speciation for Combustion of Bio-diesel, Diesel

and Hexadecane in % Mass C1 basis ...... 269

7.2 Numerical Hydrocarbon Composition for Combustion of Hexadecane in

% Mass C1 basis ...... 270

8.1 H2 Producing Reactions in the Pre-Ignition Section for φ = 1.6 and S/C = 1.0 at Tini = 1000 K ...... 273

8.2 H2 Producing Reactions in the Combustion Section for φ = 1.6 and S/C = 1.0 at Tini = 1000 K ...... 274

8.3 H2 Consuming Reactions in the Pre-Ignition & Ignition Sections for φ = 1.6 and S/C = 1.0 at Tini = 1000 K ...... 274

8.4 H2 Producing Reactions in the Ignition Section for φ = 1.6 and S/C = 1.0 at Tini = 1750 K ...... 276

xxiii A.1 Chemical Reactions: Reaction Rates, Activation Energy and Coeﬃcient 283

C.1 Speciation for Hexadecane ...... 306

C.2 Speciation for Diesel ECD-1 ...... 312

C.3 Speciation for Bio-Diesel 100 ...... 317

xxiv CHAPTER 1

INTRODUCTION

1.1 Introduction

This dissertation comprises a study of non-catalytic reformation of hydrocarbons to

generate sufficient amounts of hydrogen for the remediation of nitrogen oxides, NOx, from the exhaust stream of diesel engines. Although different methods exist to ac- complish a reasonable production of hydrogen, the majority are mainly oriented to fuel cell applications. These technologies can be applied towards the reduction of nitrogen oxides, however, their lifetime is limited, require expensive materials or represent high energy costs. Thus, the optimal route to produce hydrogen is still debatable and strongly depends on its final application. Remediation of NOx does not require excessively large hydrogen yields and must be performed within an admissible fuel penalty. The present research presents an alternative method capable of satisfying these demands.

1 1.2 Motivation

Concern about air pollution and global warming has led to more stringent regula-

tions to control the amount of pollutants that industry and automobiles are allowed to

emit to the atmosphere. On this tenure, nitrogen oxides have been cataloged as part of

the green house gases provoking ground-level ozone, acid rain and global warming. As

a matter of fact, this group of highly reactive gases (NOx) are mainly produced by on

road vehicles followed by electricity generation plants and fossil fuel combustion in the

United States, according to the Environmental Protection Agency. Based upon this,

emissions regulations for automotive applications have become more restrictive during

the past decade. For instance, in the United States the emission regulations enforced

the NOx levels to be 5.36 g/kWh (4 g/bhp-hr) in 2001-2004, 1.35g/kWh (1 g/bhp-hr)

in 2004-2007. Currently, the standard is to decrease the levels to 0.27 g/kWh (0.2

g/bhp-hr) between 2007-2010, according to EPA.

The majority of passenger cars are propelled by gasoline engines which mainly oper-

ate at stoichiometric conditions generating a little amount of NOx which is alleviated by means of a three-way catalyst. Heavy duty vehicles, instead, are driven by diesel engines operated at lean conditions. The remediation of NOx for the latter engines can be performed via engine management i.e., Exhaust Gas Recirculation (EGR) which lowers the temperature in the cylinder below the activation temperature for NOx producing

chemical mechanisms. Nevertheless, complete nitrogen oxides remediation through

EGR is limited by engine eﬃciency. Therefore, diﬀerent after-treatment technologies

have been implemented (Lean NOx traps), while others are still undergoing extensive

research (Selective Catalytic Reduction) in order to satisfy the stringent regulations.

2 After-treatment technologies, however, require active agents that either participate

in nitrogen oxide reduction reactions or regenerate the catalytic activity after hours of

operation. Addition of hydrogen to these systems or as the main reductant has shown

interesting results, enhancing the NOx conversion eﬃciency. Unfortunately, hydrogen is not readily available and has to be obtained from diﬀerent sources. Hydrocarbons are seen thus far as the most reliable source for generation of hydrogen. Catalytic reforming of hydrocarbons has been employed in the petrochemical industry for several decades, however, adapting that technology for mobile applications is challenging.

Several studies on this reformation technique for automotive applications can be found in the literature and are introduced in the following chapter. Although high conversion eﬃciencies can be achieved, these catalyst require precious materials and have a limited lifetime due to poisoning.

Non-catalytic reforming of hydrocarbons, rich combustion, has also been performed particularly for methane gas. There are a few studies on the oxidation of heavy hydrocarbons (i.e., diesel, gasoline, kerosene). Nevertheless, the published literature is scarce. Plasma assisted catalytic reforming is another technique employed to extract hydrogen from fuel molecules; however, it implicates high electric power consumption. Catalytic and plasma reforming are principally targeted for fuel cell applications.

Combined thermal partial oxidation (TPOx) with catalytic steam reforming also tar- gets fuel cell applications. The amount of hydrogen required for after-treatment applications is lower than that for fuel cells, opening then a feasible window for thermal partial oxidation systems which do not require precious metals and are not sensitive to poisoning. Also, research has been performed on the remediation of nitrogen oxides via hydrocarbons, as well as regeneration of catalytic activity through CO and HC’s.

3 Hydrocarbons and CO along with H2, commonly named as syngas, are products of rich combustion systems (or TPOx). Based upon these facts, the present dissertation focuses on the thermal partial oxidation and possible scenarios to enhance its eﬃ- ciency towards its utilization for the remediation of nitrogen oxides (H2-SCR systems).

Figure 1.1 shows an schematic of a H2-SCR system with a thermal partial oxidation reformer as the hydrogen production method.

Figure 1.1. Schematic of a H2-SCR System with a Thermal Partial Oxidation Reformer

1.3 Thesis Outline

Based upon the motivation presented before, a brief review of the immense published literature on exhaust after-treatment and fuel reforming processes, including cleaning up processes, is introduced in Chapter 2. Lean NOx traps, SCR catalysts and plasma assisted NOx remediation are undergoing extensive research to satisfy the stringent regulations established by the EPA and the international treaties, such as the

Kyoto protocol. As mentioned above, addition of hydrogen boosts the NOx reducing mechanisms. Hydrogen is then obtained from hydrocarbons. Hence, explanation of

4 diﬀerent hydrocarbon reforming methods (catalytic and non-catalytic) and the inﬂu- ence of several parameters such as the oxygen to carbon ratio, equivalence ratio, steam to carbon ratio, and the reforming temperature on the production of hydrogen and other by-products is presented as well.

Although non-catalytic reforming of hydrocarbons shows a lower hydrogen yield than that for catalytic reforming, this concentration is suﬃcient to satisfy the hydrogen demand for the reduction of NOx in diesel engine exhaust. Based upon this, a detailed study of a non-catalytic reformer through numerical simulation of the chemical kinetics of the oxidation of rich mixtures of hexadecane (diesel surrogate) and experimental reforming of hexadecane, diesel, and bio-diesel fuels is presented in this thesis.

On this tenure, Chapter 3 presents the governing equations solved for a 0D model capable to simulate the combustion process through 794 reactions and 113 species. A series of calculations of the initial conditions to account for the inﬂuence of equivalence ratio, steam to carbon ratio, pressure and temperature (i.e., characteristic properties) on the hexadecane reforming process, as well as a new metric of the eﬃcacy of the process, are presented as the hydrogen yield. Secondary injection of water vapor in the reforming process and the utilization of exhaust gas as a possible oxygen provider are also exhibited in this chapter.

The results obtained for the numerical simulation are classiﬁed into two categories.

Chapter 4 presents a detailed explanation of the chemical kinetics on the generation of hydrogen through the reforming process under the inﬂuence of each characteristic properties for speciﬁc cases selected for comparison. Chapter 5 displays the chemical

5 product compositions at equilibrium obtained in the 0D model for the diﬀerent characteristic quantities. It also introduces analyses of the equilibrium product compositions acquired for the secondary injection of water vapor and the use of exhaust as an oxygen source.

Three diﬀerent designs of the experimental reformer, as well as the instrumenta- tion used for injection of the feeds and the measurement equipment, are presented in

Chapter 6. Hexadecane, diesel and bio-diesel are the fuels employed to investigate the reforming of hydrocarbons using this experimental reformer. Hexadecane is a pure substance and hence its composition is well known. However, diesel and bio-diesel are composed by a wide variety of hydrocarbons, therefore, their compositions have to be obtained through a GC-MS analysis before undergoing the reformation process. A brief introduction to the GC-MS process is then introduced in this chapter.

As expected, CO, CO2,H2 and H2O are the main products of rich combustion.

However, several hydrocarbons are also formed as byproducts, which can also be used for the remediation of NOx via HC-SCR systems. While the major compositions and an approximated total hydrocarbon concentration (C1 basis) contained in the reformate gas can be measured by the Horiba and H-sense analyzers available at the Center for

Automotive Research facilities, speciﬁc hydrocarbon compositions cannot be obtained.

Therefore, GC-FID analysis of the reformate to obtain a detailed hydrocarbon composition is performed at the Transportation Research Center which counts with a more advance measurement system (GC-FID, Varian 3600).

6 The results obtained for the experimental reformation of these three fuels, as well as the speciation of the liquid fuels and the reformate are presented in Chapter 7.

Comparisons between the results obtained in the experiments and those derived from the numerical 0D model are performed in this chapter.

Finally, a series of conclusions of the analyses introduced above and a section of further work that can be carried out to improve either the numerical simulation or the experimentation are included in Chapter 8. Three appendixes are added to this work. Appendix A comprises the chemical mechanism, their activation energies and coeﬃcients, employed for the 0D model. The spectrograms of the GC-MS of the liquid fuels are introduced in Appendix B, while the hydrocarbon speciation tables for the reformate gas for hexadecane, diesel and bio-diesel are given in Appendix C.

7 CHAPTER 2

LITERATURE REVIEW

2.1 Introduction

The increasing demand of energy, the scarcity of natural resources, the high prices of petroleum, and the high environmental standards for clean fuels have created an impetus to ﬁnd alternative energy sources and exhaust gas after-treatment technologies. On the energy side, the automotive and petrochemical industries are in reality contemplating hydrogen as an energy carrier for fuel cell applications and its key role in a long-term transition towards a clean and sustainable energy future [1]. On this tenure, tremendous amounts of research have been concentrated on hydrogen fuel cell systems as energy sources because of their high eﬃciency and pollution free characteristics [2, 3, 4]. On the after-treatment side, CO2 sequestration, catalytic remediation of nitrogen oxides (NOx) and particulate matter (PM) from engine exhaust [5] are the technologies being investigated research to eliminate or reduce these contaminants from the exhaust gases.

The present work focuses on the production of a reducing agent utilized for the remediation of nitrogen oxides from a diesel engine. According to the Environmental

8 Protection Agency (EPA), NOx is involved in the formation of ground-level ozone, ni-

trate particles, acid aerosols and nitrogen dioxide which cause respiratory problems [6].

Furthermore, it contributes to acid rain and global warming (nitrous oxide); nitrous oxide is considered to be the fourth largest contributor to the green house gases and over a

100-year period has about 310 times more impact trapping the heat than carbon dioxide [7]. Statistically, in the United States, automobiles generate approximately 49%

of the produced NOx, whereas commercial/industrial/residential applications produce

19% and 27% from electric utilities [6].

Control of nitrogen oxides is generally easier for stationary applications since they

are operated under more steady cycles, have less constrains on system size, and ready

access to reductants. Thus, the challenge for the automotive industry not only comes

from the rapid transient conditions but also from the weight and size constraints. Cat-

alytic conversion is the only practical solution to reduce or eliminate nitrogen oxides.

It is noteworthy that in the past years special interest has been focused on using non-

thermal plasma to reduce NOx and soot under rich conditions; however, their results

show very small conversion when using only the discharge. Catalytic NOx conversion

has been improved when a plasma discharge is used prior to the catalyst since the NO

oxidizes to NO2 [8, 9]. This leads to two main concerns are derived. First, the se-

lection of the catalytic material and second, the reducing agent. The present chapter

oﬀers a review of the current technical literature on the diﬀerent available technologies

used for the remediation of the nitrogen oxides generated by diesel engines, and then

focuses on the methods for hydrogen generation which is the reducing agent selected

for this research.

9 2.2 After-Treatment Technologies

As mentioned in the introduction, international treaties such as the Kyoto protocol

or national emissions regulations have become very stringent in the last decade. A brief

summary of the emissions from gasoline and diesel engines is necessary to understand

the diﬀerent ways that these regulations can be or have been achieved. Gasoline en-

gines operate very close to stoichiometric conditions, i.e., the equivalence ratio slightly oscillates between below and above unity. Several beneﬁts are achieved by operating under these conditions: 1) relatively high exhaust temperature which is beneﬁcial for catalytic emission reduction, 2) low production of particulate matter (PM) or soot, and

3) simultaneous reduction of NOx and oxidation of HC and CO (three way catalyst).

The latter is achieved using catalysts that contain precious metals (Pt, Pd and Rh) supported on alumina with cerium to provide oxygen storage.

Under rich conditions more reductants such as HC and CO are available to reduce

NOx; however, the oxidation of CO and HC is incomplete since the mixture is not reducing. While operating lean, the conversion of HC and CO signiﬁcantly improves due to the presence of excess oxygen in the form of O2, NO2 and NO2 which results in reduction of NOx. Perfect stoichiometric conditions provide the highest reduction of the three pollutants at steady state, unfortunately, it is not possible to achieve this under real driving conditions [10]. Oxygen storage under lean conditions then becomes

a key factor to eﬃciently reduce CO and HC in the rich phase, where the gas entering

the converter is net oxygen deﬁcient. Another factor that aﬀects the conversion of

these pollutants is the temperature of the catalyst, where good conversion is achieved

after 250 .

10 Emissions from gasoline engines are very signiﬁcant during cold starts due to the high equivalence ratio required for start as well as the cold catalyst temperature. One solution can be to implement hydrocarbon traps inside or before the three way catalysts to store the excessive hydrocarbons and release them once the catalyst achieves a reasonable temperature. Based upon these facts, gasoline engines show very good emission control although the new regulations require further reduction in the produced amount of carbon dioxide and also improvement in the vehicle fuel economy (mpg).

Diesel engines, which are generally more efficient than gasoline engines, operate under lean conditions. A new set of problems arises because of the non-homogeneous distribution of fuel/air in the cylinder which causes part of the fuel to be burnt by diffusion (i.e., oxygen and fuel must be present) creating thus high levels of nitrogen oxides and particulate matter. These two pollutants must be significantly diminished to satisfy the emissions levels. On the nitrogen oxides reduction, for instance, Japan set the NOx standard for diesel heavy duty vehicles to 2 g/kWh in 2005 while Europe is expected to meet the same standard by 2008-2009 [11, 12]. In the United States, the emission regulations have forced the NOx levels to be 5.36 g/kWh (4 g/bhp-hr) in 2001-

2004, 1.35 g/kWh (1 g/bhp-hr) in 2004-2007. Currently, the standard is changing to 0.27 g/kWh (0.2 g/bhp-hr) between 2007-2010, according to the Environmental

Protection Agency (EPA).

Possible solutions for these issues have been proposed during the last decade. Par- ticulate matter is usually reduced by a Diesel Particulate Filter (DPF), which is commonly a type of ceramic ﬁlter, which has to be regenerated by oxidizing the trapped

11 soot [13]. NOx, for instance, has been controlled by means of Exhaust Gas Recircu-

lation (EGR), where part of the exhaust is injected in the engine cylinder along with

fresh air. Unfortunately, combustion instability occurs at high EGR levels and also

soot levels increase with increasing EGR [14]. Exploratory investigation on multiple

injectors, swirling in-cylinder ﬂows and timing of the fuel injection has been carried out

to achieve low-temperature combustion for which it is known that NOx and PM can be reduced.

Presently, it is impossible to meet the environmental requirements by engine management only, hence, several after-treatment technologies have to be developed. One thing that has to be kept in mind is that the after-treatment system must satisfy these demands while maintaining a good or acceptable fuel economy. Three way catalysts, for instance, have been widely used for gasoline engines to control the NOx production but are not suitable for diesel engines due to the low or null NOx conversion under lean conditions [15, 16, 10]. Hence, a new challenge for diesel automotive engineers arises.

Direct injection of diesel fuel over a catalyst (i.e., Hydrocarbon SCR) has been studied to reduce NOx; however, it increases the amount of hydrocarbons in the exhaust with marginal NOx remediation [17]. Based upon this concept, new technologies have

emerged to control the production of nitrogen oxides, such as Lean NOx traps and

Selective Catalytic Reduction systems using ammonia, urea or mixtures of hydrogen,

hydrocarbons and carbon monoxide as reductant agents. A brief introduction of these

systems is given in the following sections.

12 2.2.1 Lean NOx Traps, LNT

Nitrogen oxide traps are catalysts that contain alkaline-earth metals (e.g., barium) or alkali metals (e.g., potassium). Essentially, NO is oxidized to NO2 on the platinum sites during lean conditions. The nitrogen dioxide is then transformed into nitrates involving the alkali-metals and excess oxygen or the alkaline-metals. During the short period of rich conditions, the stored nitrates are decomposed and then the NOx is reduced by hydrogen, carbon monoxide and hydrocarbons over the precious metal to form molecular nitrogen [18, 19, 20, 21]. Such reductants can be generated on-board by catalytic or non-catalytic partial oxidation of diesel fuel or by engine under rich conditions. NOx conversions around 90% during a cycle (lean-rich operating conditions) have been reported by this mechanism. Unfortunately, NOx traps tend to be relatively large and require high loads of precious metals (Pt) which increase their cost.

These catalysts also suffer from deactivation due to the reactions between adsorbents and compounds within the wash-coat through/or particle growth of adsorbents, as well as sulfur poisoning. Moreover, exposure to high temperatures in very lean conditions can coarsen the Pt contained in the traps, whereas exposure to high temperatures in rich conditions can coarsen the storage materials. Both effects have detrimental effects on the conversion efficiency of the trap [18, 19].

Hence, one can conclude that the performance with temperature depends mainly on the storing materials, the loading of platinum, and the volume of the catalyst. Al- kaline metals generally oﬀer a good performance for low temperatures (250 to 450 ), whereas alkali metals present a better storage eﬃciency at high temperatures (up to

13 550 ). Some LNT formulations contain both alkaline metal and alkali metal to maxi-

mize the LNT temperature window and broaden the range of operation [20]. While a

high precious metal loading enhances the NOx conversion at low temperatures, it also catalyzes the decomposition of nitrates during lean conditions creating negative eﬀects on the NOx storage performance at high temperatures (> 450 ).

With respect to the volume of the LNT, low exhaust temperatures are normally accompanied by low ﬂow rates which require only a small LNT, while high exhaust

ﬂows and high temperatures require a larger volume to maintain reasonable space velocities to assure a good NOx reduction. Based upon all these facts, a catalyst with a low Pt load zone and then a high Pt load zone will provide a high NOx performance within a broad temperature range, while maintaining a reasonable volume. This also impacts the cost of the LNT since it utilizes less Pt than a similarly-sized LNT with uniform Pt distribution [19, 20].

Mahzoul et al.[18] studied the mechanism of the storage by using excess air. Ni-

trogen oxides as well as oxygen are adsorbed in two diﬀerent ways depending on the

location of the platinum sites related to the barium oxide crystallites. For those Pt

sites located close to the BaO crystallites, subindex ‘1’, the following reactions take

place

NO + (x)1 (NO∗)1 (2.1)

NO2 + (x)1 (NO2∗)1 (2.2)

O2 + 2(x)1 2(O∗)1 (2.3)

14 Then, these species react with the BaO to form Barium nitrate as follows,

BaO + 2(NO∗)1 + 3(O∗)1 −→ Ba(NO3)2 (2.4)

BaO + 2(NO2∗)1 + (O∗)1 −→ Ba(NO3)2 (2.5)

Thus, the barium nitrate decomposes according to

Ba(NO3)2 −→ BaO + 2NO + 1.5O2 (2.6)

For those Pt sites located too far from BaO crystallites, the adsorbed species (NO*)2,

(NO2*)2 and (O*)2 do not react with BaO in the same way. These species, subindex

‘2’, react to form barium nitrites Ba(NO2*)2 over these sites. The authors state that

NO oxidation to NO2 also takes place at the platinum sites as well as adsorption of both

nitrogen oxides and oxygen dissociation close to BaO before nitrate formation. The

storage process is driven by the nitrate formation and decomposition rates, Equations

(2.4), (2.5), (2.6). Once these rates are the same, the process stops. Increasing the oxygen concentration enhance the oxygen adsorption and also accelerates the rate of nitrate formation. Nevertheless, increasing the oxygen concentration more than 3% has no further inﬂuence on the NOx reduction [18].

Forzatti et al.[22] strongly agreed on the nitrate route towards the reduction of

NOx. However, they suggested that the nitrite route might play a very important role on the reduction of nitrogen oxides based on the fact that the adsorption of NO/O2 on the Ba/γ-Al2O3 occurs initially in the form of nitrites, which are then transformed into nitrates. The rate of nitrite formation and that of its oxidation to nitrates are higher when the catalyst contains Pt. Also, increasing the Ba loading in the catalyst

15 led to a larger NOx storage, which could be explained by the fact that the number

of Pt-Ba areas increases favoring thus the nitrite route. The adsorption of NO2 over

their catalysts was signiﬁcant, occurring with a simultaneous release of NO according

to the stoichiometry of the disproportionation reaction

2− O + 3NO2 −→ 2NO3 + NO (2.7)

Periodic regeneration of the catalyst trap has to be performed in order to recuperate the catalytic activity. Hydrogen, carbon dioxide and other reductant agents have been widely used for this purpose [21, 22]. Equation (2.8) shows the selective reaction

towards nitrogen, which is very fast; however, after prolonged exposition the reduction

process favors the formation of ammonia instead as illustrated in Equation (2.9)[22].

If the latter reaction occurs, an oxidation catalyst must be used to avoid ammonia slip.

Ba(NO3)2 + 5H2 −→ N2 + BaO + 5H2O (2.8)

Ba(NO3)2 + 8H2 −→ 2NH3 + BaO + 5H2O (2.9)

Sulfur poisoning (formation of stable barium sulphates) and Pt sintering (due to

high temperatures) are two phenomena that take place due to high or long exposure to

sulfur and high temperatures respectively. This indeed decreases the storage capacity

of the catalyst and in some cases totally inhibits it. Based on this, hydrogen and CO

have been widely used to regenerate the catalyst from poisoning due to their high activ-

ity within a broad temperature window. However, hydrogen is more eﬀective than CO

at low temperatures [21, 22]. Another problem seen in these traps is the degradation

by hydrocarbons which occupy the active Pt sites. Also, if hydrocarbons are used to

16 purge the LNT system, the temperature due to the oxidation of HC rapidly increases

(the reaction is more exothermic for hydrocarbons than the oxidation of hydrogen or

CO) achieving high values which impact the NOx conversion negatively. In order to alleviate these issues, a DOC (Diesel Oxidation Catalyst) is usually installed prior to the trap to eliminate most of the hydrocarbons. Furthermore, if an integrated system such as LNT/Urea-SCR would be used, the DOC would be utilized not only to reduce hydrocarbons but also to convert the NO to NO2 with a 50% ratio being preferred.

This could be explained by the fact that a 50/50 mix of NO to NO2 has improved the reduction of NOx due to the fast SCR reaction given by Equation (2.15)[20, 25].

Kong et al.[23] applied on-board hydrogen to regenerate a lean NOx trap system employed for a 8.3L diesel engine under six ESC (European Stationary Cycle) modes, while maintaining a reasonable fuel penalty. Hydrogen is generated through partial oxidation of low sulfur diesel fuel via assisted plasma reforming which provides about

21% H2 and 20% CO concentrations by using only 200 W of electric power. No detailed information on the plasma reformer is given in this publication; however, in a later publication [24] from the same company a very low H2 concentration of about 6-8% is

obtained by using only the plasma reformer (the power employed to generate the plasma

is not mentioned), while using a plasma-catalytic reformer the hydrogen concentration

is about 21-26%. Nevertheless, the relevance of this paper is on the regeneration of

the trap by hydrogen and carbon monoxide.

For all ESC modes, 85-90% NOx conversion is achieved while the fuel penalty is

about 2.99-4.74% [23]. Diesel fuel is employed as a regenerating agent as well showing

very low NOx conversions (50-79%) for the same ESC modes at the same fuel penalty

17 oﬀer by the reformate. In order to obtain the same NOx conversion for both regen-

erating agents at ESC 5 mode, the fuel penalties are 8.23% for diesel and 4.74% for

reformate.

2.2.2 Hydrocarbon-Selective Catalytic Reaction Systems, HC-SCR

Another technology has been employed hydrocarbons on metal based or zeolitic

catalysts as the reductants for NOx for SCR [15, 27, 28, 29]. Oleoﬁns, oxygenated

hydrocarbons, ethene, propene, among others present a high reactivity and activity

towards the NOx reduction. Furthermore, these products can be produced on-board

through the partial oxidation of diesel fuel. This clearly oﬀers a huge advantage over

urea catalysts, since HC storage and injection are not a problem. Unfortunately, the

main disadvantages of NO conversion via hydrocarbons are on the catalyst side.

Zeolitic catalysts have a high resistance to sulfation and coking; however, they are

not thermally stable and suﬀer from dealumination and degeneration of the active

components (e.g., formation of copper cations) [27, 28, 29]. Silver/platinum alumina-

supported catalysts are not suﬃciently active in the temperature range of real diesel

exhausts (200-400 ) in the presence of sulfur dioxide in the feed [27]. Moreover,

the NOx-removal below 200 has remained open as a common problem for both

catalyst types. Adsorbing catalysts showed sulfur poisoning and water elution of the

NOx-adsorbing agents, as well as time lags between the catalytic NOx-reduction and desorption rates [15]. It has been found that the activity of these catalysts is drasti-

cally enhanced by hydrogen addition using H2-HC-SCR [30].

18 Kubacka et al.[31] proposed two diﬀerent formulations of xIn/yCo-FER(II), where

x = 0.6, 1.2 and y = 0.25, 0.5, 0.75, employing methane and ethene. Ethene achieved

a good conversion 80-85% within a temperature range 325-400 , while methane ob-

tained a 90-100% over 350-450 . A 0.6In/0.5Co-ZSM-5 catalyst was studied for comparison with the FER(II). The catalyst containing FER(II) has a faster rate and higher conversions than the ZSM-5 within the 300-500 range. However, the ZSM-5 cat-

alyst presents better performance than the FER(II) at higher temperatures than 500

. He et al.[32] studied the eﬀect of C3H8 over diﬀerent CoOx/Al2O3 catalysts, while

varying the Co concentration (0.5 & 1 mol%). Both catalysts showed a 91 and 98.8%

conversion, respectively at 400 . A higher cobalt loading decreased the NO conversion but increased the conversion of propane. Their results showed a negligible formation of N2O since the selectivity of NO to N2 is higher than 98% in all experiments.

Komatsu et al.[15] obtained a 97-100% conversion at 160-200 over a catalyst with 2%Pt/MPs. Unfortunately, the conversion decreases with aging, shifting the temperature for the maximum conversion towards 170-240 . In order to recover the activity of the catalyst, a process of regeneration had to be carried out using additional small amounts of Fe or Pt, shifting back the conversion by 20 . This catalyst exhibited

100% de-NOx in the range 300-600 with excess amounts of hydrocarbons used as reductants (rich conditions).

A double layer catalyst for propene in the presence of oxygen was proposed by Kang et al.[28]. This catalyst is composed of two layers: 1) the lower layer made of Pt (or

Au)/Al2O3 (or SiO2), and 2) the upper layer of H-mordenite and HZSM-5 (Cu-zeolites).

The HM/M catalyst alone showed a NOx conversion of 40% at 450 , while the CuM/M

19 catalyst increases the conversion eﬃciency to 70% even at 400 . The concentration

of propene decreases with increasing reaction temperature until it is fully oxidized at

the maximum NOx conversion temperature. The conversion of Au/Al2O3/M catalyst at 400 was 55%, while the monolith catalyst coated with H-mordenite as the upper layer on Au/Al2O3/M was increased to 70% at same temperature. Propene was completely oxidized at 450 . When using CuM//Au/Al2O3/M the maximum conversion

was shifted to lower temperature 350 and with faster rate. The temperature window

was remarkably broadened comparing with that of Au/Al2O3/M catalyst. Moreover,

removal performance of NOx was enhanced when gold was substituted for platinum.

The maximum activity on CuM//Pt/Al2O3/M was obtained at 250 . The temper-

ature window of this catalyst was broadened more than that of CuM//Au/Al2O3/M catalyst and shifted to lower temperature, but the conversion was lower. These results implied that the double wash-coat catalyst was very eﬀective for the SCR of NOx with

propene.

2− Lin et al.[29] studied the performance of C3H6 on a 2 wt% Cu/SO4 /Al-Ce-PILC catalyst. Their results showed a maximum NO conversion of 56% at 350 , decreasing

slowly to a 22% at 700 . At low temperatures the conversion of NO increases as does

that of the reductant propene. As temperature increases propene oxidation begins

to compete with NO reduction reducing then the rate of NO conversion although the

NO conversion maximum is reached in this section. At high temperatures, propene

oxidation is favored and NO reduction decreases. The authors clearly stated that the

Cu loading in the catalyst plays an important role on the NO conversion, exhibiting

a maximum NO conversion for 2 wt%. On the contrary, the complete conversion

(oxidation) of propene was achieved at lower temperatures with increasing Cu content.

20 The addition of water in the gas stream caused the NOx reduction to shift to higher

2− temperatures for the Cu/SO4 /Al-Ce-PILC; however, its eﬀect was less pronounced at high temperatures. No damage or change to the structure of the catalyst occurred.

Sadykov et al.[27] stated that oxygen and water vapor are required for the high eﬃ- ciency of NOx conversion into N2 on Pt + Cu/ZrPILC catalysts. Nevertheless, oxygen

was seen to be more important since it accelerates the generation rate of reactive inter-

mediates such as nitrate-nitrite species. Decane conversion to CO2 is less sensitive to

the feed composition provided either water or oxygen are in a great excess, apparently

suppressing coking.

It is well known that sulfur poisoning of the catalysts inhibits the reduction of

NOx in the SCR process. Houel et al.[33] focused on this problem on Ag/Al2O3

catalysts using hydrocarbons (octane and diesel) as the reductant. At 350 the NOx

conversion is constant in the absence of sulfur dioxide. In the presence of 20 ppm

sulfur dioxide, the conversion increases initially; however, it rapidly decays with time

due to sulphur poisoning. This decrease is attributed to formation of sulphate around

the silver sites blocking them and causing mostly hydrocarbon combustion. At high

temperature (500 ), the NOx conversion is slightly increased and then stabilized.

Regeneration of the catalyst has to be done periodically to recover the reducing

activity. However, sulfur dioxide does not decompose below 800 under N2 com-

plicating the regeneration since exposing the catalyst to such high temperatures will

result in thermal deactivation. The authors [33] stated that regeneration of the cata-

lyst has to be done in the presence of a reductant, which could be either hydrocarbons

or hydrogen otherwise it would ineﬀective. Based on their results, most of the sulphur

21 could be removed by a high temperature treatment (500-650 ) in the reaction gas with hydrocarbons (n-octane) present due to sulphate reduction by the hydrocarbons.

This indeed recovered almost full activity of the catalyst. Diesel is also employed as a NOx reductant, unfortunately coking occurs at very low temperatures leading to deactivation of the catalyst. Also, only partial recovery of the SCR activity was noticed after desulphation. The authors believed that it is due to the fact that diesel fuel is a mix of diﬀerent hydrocarbons making it more diﬃcult to activate.

Klein et al.[34] tested four diﬀerent high temperature catalysts (3 metal type and 1

Pt type) using a 7.2L DI turbo charged diesel engine under Euro III 13 mode test. Two of these metal type catalysts showed a 77% conversion in mode 6, at a high temperature of 465 , while the platinum type did not show conversion due to the high temperature.

When low temperatures are employed (285 ) the Pt type catalyst shows the best conversion in mode 7 out of the 4 catalysts. Unfortunately, that conversion maximum is very low 20%, which is very impractical in real conditions. The metal type catalysts showed a better performance over the same Euro III cycle as the fuel penalty is increased

(more fuel is injected in the exhaust stream); the opposite occurs for the platinum type. Two catalysts are used in-line improving the conversion from 35% using only type A to 40% using type A-B. The authors also studied the inﬂuence of propene and hexadecane in the nitrogen oxides reduction. It is seen that the higher the hydrocarbon the higher the NOx reduction. A mixture of several hydrocarbons was tested obtaining a lower conversion peak than when using a high single hydrocarbon, and also lower HC conversion which will result in elevated hydrocarbon emissions if no oxidation catalyst is used after the SCR.

22 2.2.3 Urea/Ammonia-Selective Catalytic Reaction, Urea/NH3-SCR

Ammonia can be used as a reductant for nitrogen oxides in stationary plants since

keeping a supply of ammonia are minimal. Moreover, the combustion processes in

stationary applications are generally slowly varying. Hence, the amount required to

minimize or eliminate the nitrogen oxides is also constant. For mobile applications

this becomes challenging since the conditions of a diesel engine are mainly dynamic

(varying driving conditions). Managing the precise dose of NH3 is important since

injecting a lower amount will decrease the NOx conversion, but injecting a higher dose

will oxidize to molecular nitrogen and NO over the catalyst or oxidize directly to NO

in the atmosphere if it slips beyond the catalyst. Another drawback is the ammonia

storage for mobile applications, because it requires a complex injection system to dose

it to the exhaust due to the dynamic NOx behavior.

Liqueﬁed NH3 has been proposed to be used in large stationary plants, as well as for mobile applications. Nevertheless, its transportation to reﬁlling stations or onboard vehicles involves safety issues. Ammonium carbonate, NH4NH2COO, has also been

suggested as another ammonia carrier. Its advantage is that it is a solid reductant

and that it decomposes into urea ((NH2)2CO) and water when heated. However, it

does require a large injection system involving heating of the storage recipient, pumps,

etc. [16].

Based upon these issues, a 32.5% urea solution in water has been used to decompose

into ammonia in a suitable catalyst on board resolving the safety and storage prob-

lems [12, 17]. Urea storage could be a problem only at low temperature operating

conditions since it freezes at -11 [12]. The European Automobile Manufacturers

23 Association concludes that fuel optimized engines combined with a urea-SCR system

could reduced fuel consumption by 7% and running cost by 3% compared with today’s

engine technology [16].

Sullivan et al.[17, 35] and Baik et al.[36] explained the chemical process that the urea solution undergoes. First, the urea solution is injected into the hot exhaust stream provoking water evaporation. After the water has evaporated, thermal decomposition of urea into one molecule of ammonia and one molecule of isocyanic oxide is expected to happen along the length of the exhaust pipe according to Equation (2.10).

The isocyanic oxides undergoes a hydrolysis process on the catalyst surface, according

to Equation (2.11), to produce one more ammonia molecule and one of carbon dioxide.

(NH2)2CO −→ NH3 + HNCO (2.10)

HNCO + H2O −→ NH3 + CO2 (2.11)

These two steps can be summarized into one general equation, Equation (2.12),

given as

(NH2)2CO + H2O −→ 2NH3 + CO2 (2.12)

Based on this pathway, the stoichiometric molar ratio typical of a Urea-SCR system

is when Urea:NOx is 1:2. An overall equation for the urea-SCR process when oxygen

and nitrogen monoxide are present in the exhaust is given by Baik et al.[36] by Equation

(2.13)

2H2N − CO − NH2 + 4NO + O2 −→ 4N2 + 4H2O + 2CO2 (2.13)

24 Fang et al.[25], also suggested a mechanism for the decomposition of urea that

does not account for complete decomposition of urea into ammonia and cyanic acid,

but eﬀectively allows the formation of other species. This mechanisms is given by

Equations (2.14) and (2.11).

(x + 1)(NH2)2C = O + H2O −→ (HCNO)x + (x + 2)NH3 + CO2 (2.14)

If urea is decomposed into cyanuric acid (x = 3) rather than into isocyanic acid (x =

1), the production of ammonia will decrease since such a product is very stable and will

no longer decompose into ammonia. They also proposed a more detailed mechanism

for the selective catalytic reduction of nitrogen oxides. The authors strongly claimed

that in absence of oxygen the direct reactions with NO and NO2 are relatively slow, but when oxygen is present in the exhaust two diﬀerent routes are most likely to occur

4NH3 + 2NO2 + 2NO −→ 4N2 + 6H2O (2.15)

4NH3 + 4NO + O2 −→ 4N2 + 6H2O (2.16)

They reported two other patterns seen in their experiments involving pure urea and

isocyanic acid that reduce nitrogen oxides,

2(NH2)2CO + 6NO −→ 5N2 + 4H2O + 2CO2 (2.17)

4HNCO + 6NO −→ 5N2 + 2H2O + 4CO2 (2.18)

The authors suggest that two temperature regimes should be deﬁned for the decom-

position of urea. At low temperature ( 250 ) the presence of moisture accelerates ¬ the decomposition of cyanic acid into ammonia, whereas at high temperatures ( 350 ¬ 25 ) the decomposed products of Equations (2.10) and (2.11) form a water-insoluble

layer that does not allow further interaction between the water and the cyanic acid.

Hirata et al.[12] agrees with the thermolysis process undergone by urea. However,

it complements the SCR process by adding two more reactions to Equations (2.12) and

(2.16)

6NO2 + 8NH3 −→ 7N2 + 12H2O (2.19)

NO + NO2 + 2NH3 −→ 2N2 + 3H2O (2.20)

where it can be seen that a 1:1 ratio of NO:NO2 is needed. Equation (2.20) is indeed

very active within the whole range of temperature; however, its eﬀect on the NOx

conversion is stronger at low temperatures [12].

Special care has to be taken when sulfur dioxide coexists in the exhaust since it causes undesirable reactions to occur. Ammonia oxidation, which is promoted by sulfur dioxide, not only consumes ammonia but also forms nitrogen oxides. Furthermore, degradation of ammonia is seen by the contact of ammonia with sulfur dioxides in the presence of water at low temperature [25].

Gabrielsson et al.[16] studied the diﬀerence between using a full catalyst, which

has a channel wall made of catalytic material and one that has a ceramic base with

a metallic wash-coat or substrate. Indeed the substrate contains less active material

than the full catalyst, this impacts its activity at low temperatures. The activity

at low temperature has to be improved by the ratio between NO and NO2 (50/50).

Another way to improve the performance of the wash-coated monolith is by applying a

26 pre-oxidation catalyst upstream of the urea-SCR catalyst. The full catalyst presents the advantage of being able to store more ammonia than the wash-coated catalyst at low temperature, which is very important for dynamic response in mobile applications.

Hence, the precision to inject urea is diminished. However, the control of ammonia slip is more complicated for this type of catalysts if a rise in temperature occurs. At high temperatures a very similar activity was registered for both catalysts.

Two main catalyst groups can be distinguished from the technical literatures, the metal oxide catalysts and the zeolitic catalysts. Vanadium over titanium oxide catalysts have been widely used in industry to reduce nitrogen oxides; however, they are considered to be large in size for automotive applications [12]. Tungsten, molybde- num are used to increase the acidity of the catalyst, which has demonstrated good performance in a wide temperature window (larger than 200 ). For the low temperature range MnOx-CeO2 catalysts show good activity. Nevertheless, special care must be taken when using these for high temperature since they tend to oxidized ammonia rather than reducing NOx [16].

Zeolite catalysts are supported on ceramic substrates, and can easily be small in size.

Advantageously compared with the vanadium type, these catalysts do not decrease signiﬁcantly their NOx conversion when downsized [12]. Gabrielsson et al.[16] studied a zeolitic catalyst showing that their activity is low for the low temperature regime, although its activity shifted up four times when a POx catalyst was used in front of the zeolitic catalyst. They also noticed that sulfur poisoning is strongly diminished when the POx catalyst is used.

27 Improvements on the low temperature regime have been made since it is a very common range of operation for diesel engines. Baik et al.[36] tested 13 diﬀerent catalyst compositions with Cu, Pt, Fe, V2O5/TiO2, among others. The catalyst Cu-ZSM5

(metal-zeolitic) revealed the highest performance of NO removal activity ( 60-90%), ¬ particularly below 250 . A 90% conversion was maintained over a wide operating temperature window from 200 to 400 . The authors suggested that increasing the residence time of the feed gas stream in the reactor signiﬁcantly improves NO conversion particularly at low reaction temperature.

Gabrielsson et al.[16] and Sullivan et al.[35] agreed that water has a slight negative eﬀect on the conversion of nitrogen oxides for urea-SCR processes due to the fact that water competes with ammonia for adsorption sites on the surface of the catalyst.

The catalyst activity is then shifted to higher temperatures (about 15-20 ). Another explanation is that the urea does not hydrolyze on the catalyst surface as described in the previous mechanism when water is present in the exhaust stream, but directly oxidize to N2 thus not producing ammonia to reduce the NOx [35]. Another possibility is that the urea decomposes into polymeric melamine complexes on the catalyst surface causing surface passivation agreeing with Fang et al.[25], Equation (2.14).

(NH2)2CO + O2 −→ CO2 + H2O + N2 (2.21)

(x + 1)(NH2)2CO + H2O −→ (HNCO)x + (x + 2)NH3 + CO2 (2.22)

Combined technologies have also shown improvements in the NOx reduction. Sulli- van et al.[17] studied the selective NOx trapping on barium-oxide based catalysts with

Fe-Ba-ZSM5 and Fe-ZSM-5 to account for urea/NH3-SCR behavior. First, gaseous

28 NOx is trapped on the material as Ba(NO3)2 until the BaO is fully saturated. Then, urea is injected over the material to be hydrolyzed to ammonia and carbon dioxide.

Ammonia reduces the Ba(NO3)2 to BaO + N2 + H2O. Ammonia slip is prevented by

+ NH4 trapping on the acid sites. This buﬀer capacity compensates for any changes in NO under transient conditions. While testing both catalysts with direct urea or ammonia injection, a better performance was achieved for ammonia. This emphasizes that either the catalysts should possess urea-SCR activity or a separate catalyst to hydrolyze urea would be required. A pulse of hydrocarbons, CO and/or hydrogen can be used for regeneration of BaO instead of the NH3. The ﬁrst catalyst showed slight poisoning of the Fe sites on the zeolite by barium [17].

A urea system comprised of a pre-oxidation catalyst upstream of the SCR and a

NH3 oxidation catalyst downstream of the SCR to avoid ammonia slip is tested with a 2L DI TCI engine for the New European Driving Cycle (NEDC) [26]. The 32.5% urea solution is injected into the system at 120 . The average NOx conversion over the entire cycle is 63% (49% for the ECE part and 73% for the EUDC part) for a fresh catalyst. As known, all catalysts suﬀer from aging, thus the system is tested for 200 hours under the same driving cycle showing a very stable average conversion (63-65%).

It is remarkable that the conversion at EUDC part of the cycle showed an increase to

80% but the ECE part decreases with time to approximately 42%.

29 2.2.4 Hydrogen-Selective Catalytic Reaction, H2-SCR

Urea and HC selective catalytic reduction systems have been brieﬂy described in the

previous sections. A disadvantage of urea is that it requires a storage unit and frequent

replenishment. Moreover, HC-SCR requires to be operated at temperatures higher

than 200 and the addition of hydrocarbons contributes to polluting the atmospheric

air if not converted in the catalyst or as CO2 if burnt (oxidized) [37, 38].

Diﬀerent studies have demonstrated that hydrogen can be used as the reducing agent

for the selective reduction of nitrogen oxides in the absence and presence of oxygen

and other species in the exhaust stream. Even when using HC and Urea systems, the

addition of hydrogen has often increased their conversion eﬃciency. Shimizu et al.[30]

added (0.5%) pure hydrogen to the exhaust/urea stream passing through their urea-

silver/alumina system signiﬁcantly improving its activity achieving NO conversions

from 0% with no H2 addition at 200 to 84-90% within a 200-500 while no formation

of N2O was seen.

Shimizu et al.[39] studied the conversion of NO over a silver-alumina catalyst when using C3H8 as the main reductant. Addition of hydrogen boosted the NO conversion from 20-30% to 50-60% with increasing temperature (500-750 ), while increasing the

conversion of the hydrocarbon from 20-30 to 50-60%.

Tomita et al.[37] suggested that hydrogen is more active than other compounds

at low temperatures (≤ 300 ). The NO-H2 mechanism proposed by these authors is given as follows

2NO + 4H2 + O2 −→ N2 + 4H2O (2.23)

30 2NO + 3H2 + O2 −→ N2O + 3H2O (2.24)

It is also known that NO dissociates into N(ad) and O(ad) on the platinum sites,

promoted by the dissociated H(ad). Then, molecular nitrogen and nitrous oxide (N2O) are generated through combination of N(ad) with each other and with another molecular

NO, respectively. NH3 is seemed to be formed by the combination of H(ad) and N(ad) on this catalyst at high temperatures [37]. However, the preferred mechanism is that

of protons migrating via dissociation of hydrogen bonds with oxide ions in the P2O7

units by a hopping mechanism in the Sn0.9In0.1P2O7 catalyst

+ − H2 −→ 2H + 2e (2.25)

1 NO + 2H+ + 2e− −→ N + H O (2.26) 2 2 2

When hydrogen is not used, a negligible conversion of NO is seen, due to the fact

that the adsorption of oxygen atoms is favored on the platinum sites. The selectivity

was as high as 84% for all H2 concentrations evaluated [37]. ¬ Wen et al.[40] utilized a catalyst with a composition of Pd/Al = 0.33 and adding hydrogen showing high activity at 100 ; however, its activity rapidly decreased with

temperature (up to 30% at 300-400 ). Another catalyst with a composition of Pd/Al =

0.98 was tested over a hundred seconds at diﬀerent temperatures. At low temperature

(100-200 ) the conversion falls between 70-60%. However, at high temperature (300-

400 ) a high conversion 70% occurs at relatively beginning of the test, but decreases ¬ exponentially to about 30-40% with time. The mechanism is given as follows

aNO + bH2 −→ N2(N2O,NH3) + dH2O (2.27)

31 1 H + O −→ H O (2.28) 2 2 2 2

1 NO + O −→ NO (2.29) 2 2 2 where a, b and d depend on the nitrogen containing product. The authors stated that the ﬁrst two reactions happen only at low or high temperature.

Costa et al.[38] studied an H2-SCR of NO at 1 atm total pressure. The only products after the catalyst were molecular nitrogen, nitrous oxide with nitrogen selectivities of approximately 80%, 95% NO conversion, and 85% H2 conversion at 150 . The reactions considered based on the ﬁnal product compositions are those given by Equations

(2.23), (2.24) and (2.28).

A new type of catalyst, proton-conducting Sn0.9In0.1P2O7, evaluated in the present of NO, H2 and O2 shows a nitrogen selectivity greater than 80% within 50-350 range

[37]. It was seen that adding Pt or Rh the selectivity was enhanced due to the fact that these metals do not contribute to the oxidation of hydrogen. A combined Pt-

Rh catalyst showed a wider operating window by achieving good conversion at low temperature ( 100 ) and peaks about 200 , decreasing to 40% at 350 . The ¬ authors strongly suggested that the N(ad) formation is increased by the promotion of

Equation (2.26), improving then the selectivity [37].

Low temperature NO conversion (100 ) about 70% was achieved using a Pd/ZSM-

5 catalyst [40]. However, the formation of Pd-β hydride deactivates the catalyst for

NO reduction. The authors found that the rate of the reaction NO + H2 is fast or comparable to that of the H2 + O2 reaction.

32 Costa et al.[41] studied the Pt/MgO and Pt/CeO2 catalysts showing a high N2 yield

(80-85%) and a broad temperature window of operation. High NO conversions (95%) were obtained in the 100-200 range for the lowest 0.1 wt% Pt loading. A combined catalyst using both materials (50 wt% MgO-50 wt% CeO2) allowed the highest N2 yields

in the whole 100-400 range. The support catalyst alone presented a very low activity

at low temperatures; however, its conversion increased to a maximum of about 40% at

350 and its N2 selectivity was high 80%. The catalyst containing Pt showed an ¬ improved NO conversion at low temperatures, reaching a maximum of 94% at about

150 . After that the conversion decreases linearly from 50-40 within the temperature

range 250-400 . The N2 selectivity goes from 78% in the low temperatures to 92%

at the high regime. In a later publication Costa et al.[38] showed a 100% conversion

of NO and 85% N2-selectivity at 200 over a Pt/MgO-CeO2 catalyst when having

a feed stream containing NO, O2,H2O, CO2, CO, and H2. No further detail on the

composition of the catalyst is given.

Ag/alumina supported catalysts were doped with Zn and Mg to study their eﬀect

on the NO conversion when using a mixture of propane and hydrogen [39]. The Zn-

doped catalysts showed higher NO conversion than pure Ag/alumina at the same C3H8 conversion, indicating improvement of the selectivity by Zn-doping. The Mg-doping increased the C3H8 conversion, while the Zn-doped catalyst showed lower conversion than Ag/Al2O3. The NO conversion was largely enhanced by the doping of Zn, and slightly increased by the doping of Mg.

The eﬀect of CO on the NO conversion was studied by Costa et al.[38]. They found

that CO concentrations higher than 1 vol% have a signiﬁcant decrease in Pt/MgO-CeO2

33 catalyst’s activity when temperature was lower than 200 . This can be explained by two paths: 1) CO competes for H-adsorption sites on Pt which results in a signiﬁ- cant lowering of the surface coverage, and 2) CO oxidizes on the catalyst surface and increases the hydrogen combustion rate on those platinum sites. Full catalytic activity was recovered when CO was removed from the feed stream. A slight increase in activity with increasing CO concentration (0.3-1.0 vol%) was observed at high temperatures. This can be explained by the fact that the water-gas shift reaction is largely promoted at T ≥ 200 over Pt surfaces, and also a larger surface coverage of H occurs.

The fact that CO beneﬁts the NO conversion leads to think that these catalysts might be well suited for real operating conditions since CO levels for industrial applications are usually ≤ 0.5 vol%. While CO2 showed negligible variation to the NO conversion within the temperature range (100-600 ) over the same catalyst, C3H6 shows a slightly negative eﬀect (maximum conversion was 70% at 200 ). The N2 selectivity ¬ remained very similar to that obtained when this compound was absent.

Tomita et al.[37] and Costa et al.[41] agreed that water coexisting in the exhaust stream has a positive impact on the NO conversion for both catalysts. The NO conversion obtained at 150 was almost complete, whereas the selectivity increased by

5-12 percentage units at T ≤ 200 [41]. At 400 , the selectivity increased from

85 to 97% in the presence of 5 vol% H2O in the feed. [37] mentioned that increasing the concentration of oxygen in the exhaust decreases the conversion of NO2 since more hydrogen oxidizes; therefore, less hydrogen is available for conversion.

Costa et al.[38] studied the inﬂuence of SO2 on the NO conversion for their catalyst. For an aged catalyst, a very stable conversion was seen within a long period of

34 time, 50 h of continuous operation. A N2 selectivity and NO conversion of 85% were

obtained under certain conditions (200 , 0.1% NO/1% H2/5% O2/5% H2O/0.3%

CO/10% CO2/50 ppmv SO2/He). The authors claimed that this is the highest value

ever reported in the open literature for H2-SCR of NO under speciﬁc feed composition

and reaction conditions used in their work.

In another publication, Costa et al.[41] tested their catalyst in the presence of water and sulfur at 200 . It was seen that the catalyst (containing Pt) gradually decreased its activity from 80 to 0 % within a period of 24 h. After a given pre-nitration and pre-sulfation processes over the support, the activity appeared to be extremely stable during the same time, 92% N2 selectivity and a 80 % conversion.

2.2.5 Plasma Assisted Catalytic Removal of NOx

The reduction of nitrogen oxides and soot using a plasma in combination with the

catalytic systems described in the previous subsections, have attracted the attention of

the researchers. The literature on this topic, however, is very scarce. A very short

introduction to this ﬁeld is presented here due to its relevance on reducing NOx.

Plasma assisted catalytic removal has shown to eﬀectively reduce NOx when using

a small concentration of C3H6 (0.27 - 0.09 %) as the reducing agent, from a range between 42-54% when using only plasma to 68-72% when using plasma-assisted-HC-

SCR catalyst [8]. In the same study, the CO2 concentration was seen to increase with increasing temperature due to soot oxidation for the same hydrocarbon concentrations.

Nevertheless, there is little information on the reaction paths followed in this con- juncted system. Several authors [8, 9] agreed on the dissociation of oxygen by electron

35 impact. The atomic oxygen is then combined with NO to form NO2. When adding

hydrocarbons such as ethylene some radicals are formed (RO, RO2, OH, OH2, etc.)

which cause rapid formation of NO2 due to NO oxidation. The formed NO2 is con-

verted with the hydrocarbon to N2 in the catalyst.

Mok et al.[9] used this system with ammonia as the reducing agent. The NOx

conversion went from 40-50% using only the catalyst to 80-90% with the combined sys-

tem depending on the NOx initial concentration and the energy density of the plasma.

It is seen that the conversion increases with energy density for any given concentration

of NOx. The authors explained that the reactions in the catalysts are usually faster

for NO since the NO2 concentration in the exhaust is much lower than that for NO.

From the explanation of the ’plasma’ mechanism, the NO2 concentration is increased

which favors Equation (2.20) in the urea/NH3 catalyst accelerating the reduction of

nitrogen oxides.

2.3 Fuel Reforming Technologies

As discussed in Chapter 1, the purpose of this thesis is to provide hydrogen for the remediation of exhaust nitrogen oxides as the main reductant via H2-SCR or as

additional agent to enhance the NOx reduction in other SCR systems since hydrogen

is not readily available on board. Hydrogen can be generated from diﬀerent sources

such as biomass, alcohols, reforming of light and heavy hydrocarbons. If hydrogen

were to be generated at the chemical plant or reﬁnery, two problems would arise:

1) its storage would require special conditions (high pressure) and materials, and 2)

the implementation of hydrogen stations and/or networks would result in a radical

36 expensive investment [42]. This indeed creates a big challenge for the automotive

industry and, therefore, on-board reforming seems to be the most suitable option.

Methanol and ethanol reforming has been studied because of their ultra-clean and

sulfur-free nature, as well as their low reforming temperature [43, 44, 45, 46]. Never-

theless, its main drawbacks are that the energy density (19.7 and 30 MJ/kg, respec-

tively) is much lower than that of hydrocarbons, and the lack (very few plants) of

existing infrastructure of production and distribution. Based upon this and the fact

that gasoline and diesel networks are already available make this use of hydrocarbons

a more attractive solution [47]. However, special care must be taken since the fuel

reforming processes also produce pollutants such as CO and CO2.

In the current literature of fuel (alcohols and hydrocarbons) reforming technology

two methods can be recognized, the catalytic and the non-catalytic. Catalytic reform-

ing is commonly performed over two diﬀerent types of catalysts: the non-precious

metal based catalyst and the precious metal based catalyst. The selection of the cata-

lyst plays a very important role in the reforming process since it will drastically inﬂuence

the hydrogen production [46, 48, 49]. Although high hydrogen concentrations ( 60- ¬ 80% in dry conditions) are usually generated by catalytic reforming, it is expensive due to the materials involved and their tendency to age and be poisoned by sulfur and CO.

Another way to produce hydrogen that is often classiﬁed as catalytic because it accelerates the production of radicals, is the plasma technology. There are studies that show high hydrogen concentrations ( 14-30% in dry conditions) but with a large ¬ fuel penalty due to the electrical power needed to produce the plasma (larger than 1 kW

for [50, 51, 52]). On the contrary, non-catalytic reforming presents the advantages of

37 having no catalyst, being sulfur and CO tolerant, and has a low cost of electrical energy.

However, their use has a trade oﬀ in eﬃciency because their production of hydrogen is lower compared to catalytic reforming and also requires higher temperatures (above

1000 ).

Further classiﬁcation of the reforming processes are steam reforming, partial oxidation and autothermal reforming of any fuel, such as methanol, gasoline, iso-octane or diesel. “Hybrid” systems composed of either non-catalytic partial oxidation and catalytic steam reforming [69] or plasma assisted catalytic partial oxidation [8] are also found in the literature. These systems are used to either lower the cost or generate larger amounts of hydrogen depending on the application. If hydrogen were to be used for fuel cell systems, its concentration would need to be large, therefore, catalytic or “hybrid” reforming would be the preferred methods . On the contrary, lower hydrogen concentrations are often needed for reduction of nitrogen oxides which makes non-catalytic methods more attractive. A brief introduction to the catalytic, non-catalytic, and plasma methods is presented in the following subsections.

2.3.1 Catalytic Reforming

During the last decade, catalytic technology has undergone much progress to over- come the drawbacks of size, weight and cost. Within this technology, diﬀerent devices can be found such as alcohol reformers and hydrocarbon (natural gas, gasoline and diesel) reformers. Depending on the application, the fuel processor comprises the reformer(s), the desulphurization process, the water gas shift reactors (CO cleaning up), and in some processors, a membrane to separate the hydrogen from the reformate

[53, 54].

38 As mentioned previously, the reforming process can be performed over the non-

precious metal base catalyst and over the precious metal base catalyst. The non-

precious metal catalyst is typically made of nickel (Ni) supported on Al2O3, whereas

the precious metal catalyst is made of platinum (Pt). It has been seen that the Pt-base

catalyst increases the eﬃciency of the reformer but is also more expensive and prone to

be contaminated by CO and sulfur than the nickel-based catalyst. More studies have

been done on the catalyst materials [46, 48] but are not included in this review since the main interest of this study is to understand the behavior of the reformer.

Steam reforming, partial oxidation, and autothermal reforming processes are explained in the following subsections. As already stated, these systems provide a high conversion of the feeding fuel providing a high eﬃciency, which varies according to each reformer and therefore it is necessary to compare the advantages and disadvantages of each system. Moreover, a brief introduction to the desulphurization and the cleaning up processes is provided in further sections.

2.3.1.1 Steam Reforming, SR

Steam reforming is probably the most common, cheapest and more energy eﬃcient method for producing hydrogen, and is achieved by reaction over a catalyst at high temperature [42, 55]. In this process, steam reacts with the fuel (e.g., natural gas) in

the presence of a catalyst to produce hydrogen, carbon monoxide, and carbon dioxide

[42].

The overall reaction of the steam reaction is endothermic (occurring at around 800

) and therefore an external heat supplier must be used limiting the thermal design

39 and eﬃciency of the reactors. Commonly either a burner of a second fuel, a fraction of the primary fuel, or the residual fuel that remains in the reformate is used to provide this heat. The chemical reaction for steam reforming is given in Equation (2.30).

m C H + (S/C)nH O(l) −→ + 2n H + nCO + ((S/C) − 2)nH O(g) (2.30) n m 2 2 2 2 2 where n and m deﬁne the composition of the fuel, and S/C is the steam-to-carbon ratio for the mixture. According to Lutz et al.[55] the reaction is global, meaning that the net resulting reaction comes from a series of elementary reactions, including those catalytic interactions with surfaces.

Ahmed et al.[42] and Ming et al.[56] mentioned that these reformers are well suited for long periods of steady-state operation and can deliver relatively high concentrations of hydrogen (70-80% on a dry basis) in the crude reformate gas. Nevertheless, these reactors present a slow dynamic response, due to the use of indirect heat transfer that provides the heat for the endothermic reaction. Also, special care has to be taken in this process since coke can be formed depending on the temperature and the steam to fuel ratio. Furthermore, as seen in the reaction mechanism, carbon oxides are present in the reformate gas stream and must be removed according to their ﬁnal application.

CO2 can be mostly eliminated by absorption on amine solutions [42], whereas CO can be diminished by means of the water gas shift reaction or methanation processes

[56, 57].

2.3.1.2 Partial Oxidation Reforming, POx

This process involves the partial combustion of the fuel in an exothermic reaction with oxygen. The oxygen fed to the system is sub-stoichiometric so that both CO and

40 CO2 are formed in the system [42, 58]. The ideal reaction for the partial oxidation process is given in Equation (2.31) considering only hydrogen and carbon monoxide as the main products. It is noteworthy that other species may appear during the POx of the fuel; however, their ﬁnal compositions are presumably small.

n m C H + O −→ H + nCO (2.31) n m 2 2 2 2

It is noteworthy that partial oxidation systems can be characterized by the oxygen- to-carbon ratio, O2/C, which is 1/2 for an ideal product mixture of CO and H2. Fur- thermore, this reaction is exothermic (raising the reformate gas temperature over 1000

) which makes the POx more attractive than the steam reforming since no external heating is necessary to start the reaction. This is indeed reﬂected in a faster start-up of a vehicle [56]. As the process is a partial oxidation, the CO obtained from this reaction is combined with an appropriate amount of steam to complete the water gas shift reaction and convert the carbon monoxide into carbon dioxide, and consequently achieving a higher hydrogen concentration [42, 58].

Lutz et al.[58] proposed a general chemical reaction, Equation (2.32), that sum- marizes the POx reforming and the steam reforming step. In addition, this reaction allows the possibility of varying the steam to carbon ratio and the oxygen to carbon ratio.

m C H + AnO + SnH O(l) −→ nCO + + 2n(1 − A) H (2.32) n m 2 2 2 2 2

+ (S − 2(1 − A)) nH2O(g)

According to Lutz et al.[58] this global equation has a validation range for the oxygen to carbon ratio: 0 ≤ A ≤ 1 + (m/4n), where A corresponds to the oxygen

41 to carbon ratio. It is easily recognized that the steam reforming process takes place

when A = 0, whereas the stoichiometric condition for complete oxidation is given by the upper limit of A where, as known no H2 is produced. It can also be observed that

for any value of the oxygen to carbon ratio, the global balance requires a minimum

amount of steam given by S = S/C ≥ 2(1 - A).

The partial oxidation reforming of hydrocarbons produces a smaller concentration

of hydrogen than that reported for the steam reforming. This can be explained due

to the fact that in the steam reforming the steam, as well as the hydrocarbon, is split

apart, whereas in the POx reforming the amount of hydrogen that is split from the

steam is much smaller. As mentioned by Lutz et al.[58] another drawback is that the

nitrogen contained in air dilutes the product mixture obtained from POx reformers,

which reduces the hydrogen concentration in the output stream. This indeed will be

reﬂected on the eﬃciency of the fuel cell stack since the stream is directly used in the

fuel cell. However, hydrogen could be separated from the stream by diﬀerent systems

to increase the fuel cell stack eﬃciency. Nevertheless, this separation process requires

some work onto the system thus decreasing the eﬃciency of the power output of the

fuel cell system.

It is noteworthy that steam reforming and partial oxidation are often carried out

in the presence of catalysts to reduce the operating temperature and therefore reduce

losses. For this reason the selection of the catalyst will drastically inﬂuence the hy-

drogen production and thus the eﬃciency of the reformer [49].

42 2.3.1.3 Autothermal Reforming, ATR

Autothermal reforming emerges from the combination of the steam reformer and

the POx reformer. By this means, the rapid cold start-up capability and the transient

load of the POx reformer are mixed with the high concentration of hydrogen in the

reformate stream. This increases the eﬃciency of the reformer although it is smaller

than that for the steam reformer [49]. The autothermal reformers present the advantage of employing the heat provided by the POx reformer on the steam reformer [59].

Ahmed and Krumpelt [42] stated an idealized stoichiometric reaction for the conver-

sion of a hydrocarbon or oxygenate using the autothermal reforming process given by

Equation (2.33)

CnHmOp + x(O2 + 3.76N2) + (2n − 2x − p)H2O(l) −→ (2.33)

m nCO + 2n − 2x − p + H + 3.76xN 2 2 2 2 where x represents the oxygen-to-fuel molar ratio. This variable allows to obtain the number of moles necessary to convert the carbon in the fuel to CO2 (2n - 2x - p), the amount of hydrogen in the output stream (maximum 2n - 2x - p + m/2) as well as the maximum hydrogen concentration, and the heat of reaction. It can be seen that decreasing the oxygen-to-fuel ratio demands a higher consumption of water vapor, which therefore increases the steam-to-fuel ratio. As in the POx reformers, the steam reformer equation is obtained when x = 0, while increasing x = x c = [n - (p/2) +

(m/4)] gives the combustion reaction [42].

Bandi [49] mentioned that there are diﬀerent types of ATR’s such as the two-bed systems and the single bed systems. The two bed systems operate with the POX

43 reaction occurring in the ﬁrst bed. This delivers a hot stream that is discharged into

the second bed where the steam reforming takes place. For this 2-bed system however,

an external heating source is introduced in the reactor to balance for the losses in the

general system and also to enable the use of steam reforming only [49]. The single

bed system occupies less air than the 2-bed system, but has a higher risk of coke

formation, air starvation and the possibility of having explosive mixtures present in the

reforming bed. However, this conﬁguration has attained a better performance than

two-bed systems due to a better heat exchange in the single bed [60].

The temperature achieved by these reformers is lower than that of the POx which

favors the water gas shift reaction [42], therefore resulting in a higher selectivity for carbon dioxide and hydrogen. Carrete et al.[60] reported that a 60-65% methane

conversion could be attained with a selectivity of 80% towards hydrogen production.

Moreover, the insulation required for this reformer, as well as the thermal integration

of the system (between the cooler, the reactants and the hot eﬄuent) is smaller than

that for higher-temperature processes. Another advantage of the ATR’s is that the

carbon monoxide production is smaller than that for the steam reforming, as well as

the operating pressure [49].

44 2.3.2 Characterization of the Catalytic Reforming Process

Several studies [42, 58, 59, 61, 62] conclude that the control of the oxygen-to-carbon ratio, steam-to-carbon ratio, and the temperature of the reactor have a remarkable effect on the behavior of the reformer, the concentration of the products in the reformate stream, and the overall efficiency of the reformer. A brief introduction on their impact on the efficiency of the fuel processor and their management is then presented in the following sections.

2.3.2.1 Reforming Temperature

The temperature of the reactor plays a very important role in the reformation of hydrocarbons because it affects the rate of coke formation and the concentrations of hydrogen and carbon oxides, and therefore the efficiency of the reformer. The reactor temperature depends on the temperatures of the feeds (fuel, air and steam) and also the insulation or the heat flux applied on the walls of the reformer. High temperatures increase the production of hydrogen and carbon monoxide, and decrease the fractions of methane and carbon dioxide [61, 62]. For instance, Moon et al.[62] found a maximum value of hydrogen production at 700 for the autothermal reforming of iso-octane.

Docter et al.[61] mentioned that vaporization of the fuel, occurring prior to the reformer, increases the eﬃciency since the heat of vaporization would be compensated for by the combustion of a fraction of the fuel.

Liu et al.[63] suggested that in order to maintain a high temperature of the gases

(steam and fuel) prior to the reformer some points should be considered: i) the temperature of the fuel vaporizer must be always maintained 30-50 above the boiling

45 temperature of the fuel to assure complete vaporization; ii) water must be completely vaporized and superheated to a temperature close to that of the fuel vapor; iii) the fuel vapor and steam have to be mixed before being combined with the preheated air to form a uniform feed mixture; iv) the vapor mixing inlet lines must be heated to avoid condensation of the mixture prior to entering the reactor; and v) the heat losses to the environment can be prevented by thermally insulating the outside shell of the reactor.

It is noteworthy that the reactor temperature aﬀects all downstream processes such as the high/lower gas shift reactors [64].

The behavior of the temperature inside an ATR reformer presents a significant axial temperature gradient due to the fact that POx is completed at the first stages creating a large exotherm. Thereafter, the temperature decreases because of the steam reforming which is strongly endothermic (O-H bonds require a large activation energy to break up) [63]. The oxygen-to-carbon ratio increases the temperature of the reactor along the axial direction. This can be simply explained since increasing the mass flow rate of oxygen will generate or carry more energy inside the reformer (than at lower oxygen-to-carbon ratio) while the heat loss from the reactor remains essentially the same, therefore the temperature increases. In order to achieve the highest reforming efficiency, the O2/C ratio must be maintained above 0.4 [63] which agrees with the value reported by Ahmed et al.[42]. Increasing the S/C ratio from 1.1 to 2.8 produces negligible variations of the reactor temperature and the efficiency while maintaining the

O2/C ﬁxed at 0.34.

Liu et al.[63] found that the equilibrium values of the concentrations in the ATR reaction are achieved at 590 . However, it was seen that a better reforming eﬃciency

46 can be achieved at Teq = 700 , where methane and amorphous carbon are reduced to negligible levels. Carbon formation was seen for T < 700 from thermodynamic equilibrium calculations. At oxygen-to-carbon ratio of 0.6, the steam-to-carbon of

2 and temperature larger than (about) 600 no carbon is formed in the monolith; however, some coke is seen on the tubing connecting the reactor and the condenser [63].

Ming et al.[56] studied steam reforming of iso-octane at diﬀerent reactor temperatures (650 < T < 800 ), concluding that carbon monoxide increases along with the

reacting temperature. It was seen that only about 25% of the fuel was converted to

C1 products (CO, CO2 and CH4) at the lowest temperature, whereas about 100% was

converted at the highest temperature. However, the equilibrium composition calcu-

lation of iso-octane/steam within the same temperature range, while maintaining the

same steam-to-carbon ratio and pressure on the system, showed its maximum value

at the lowest temperature. Hydrogen and carbon dioxide yields tended to decrease

as temperature increased, whereas carbon monoxide increased due to the reverse wa-

ter gas shift reaction. As expected, a small amount of methane is found due to the

methanation reaction at a lower temperature.

Lutz et al.[58] investigated the eﬀect of temperature on the equilibrium species

composition for the reforming of heptane under the following conditions, S/C = 2/5

and O2/C = 0.85. The equilibrium model predicted that at 500 the hydrogen

production is low because a signiﬁcant amount of methane coexists. As expected, the

methane concentration is diminished to negligible levels as the reactor temperature

increases, creating therefore a local maximum of the hydrogen concentration at about

600 . However, temperatures above this produce a negative eﬀect in the decreasing

47 concentration of hydrogen due to a reverse water-gas shift which favors the formation of carbon monoxide, as mentioned by [56]. They also concluded that the existence of this local maximum of hydrogen concentration suggests that the partial oxidation and the water gas shift reaction should be accomplished at diﬀerent stages in the process.

Special care must be taken whilst using diesel in the experiments due to their tendency to form more carbon monoxide than carbon dioxide, which clearly indicates that the shift reaction is not in equilibrium. Decane has shown a closer equilibrium since its production of CO is smaller than that for the diesel cases. Nevertheless, the fact that CO is above the equilibrium value while CO2 is below it is also consistent with incomplete reaction due to kinetic limitations [58].

2.3.2.2 Steam-to and Oxygen-to Carbon Ratios, S/C & O2/C

As mentioned before, numerical simulations and experiments have been performed to obtain the optimal values of oxygen-to-carbon ratio and the steam-to-carbon to produce a high hydrogen concentration and low CO concentration. Docter et al.[61] performed a theoretical-numerical analysis where the oxygen-to-carbon ratio ranged from 0 to 1 (cracking of the fuel and complete combustion conditions respectively) and the steam-to-carbon ratio ranged from 0 to 2 (zero recalls the POX reforming process). Air was injected at 400 whereas fuel was injected at 20 and water at

200 . They reported that increasing the S/C ratio enhances the production of H2 and

CO2, diminishes the production of CO and reduces the possibility of coke formation.

Methane seemed to increase with increasing S/C; however, it vanishes as the air-to-fuel ratio increases. Based on these data, the optimal values were O2/C = 0.3 for S/C

= 0.7 and O2/C = 0.35 for S/C = 0, where no coke is seen to be formed. Within

48 these limits, the reactor temperature is seen to decrease with steam-to-carbon ratio.

However, at O2/C = 0.31 around S/C = 1.2 the temperature stopped decreasing and

started to grow approximately 10 above the minimum temperature registered in a very slow pace. Finally, the temperature of the reformer decreases as the O2/C ratio

is decreased.

In other studies performed on the reforming of iso-octane [48, 62], the minimum

oxygen-to-carbon ratio to assure an exothermic reaction (heat of reaction) was found to

be larger than 0.6. The production of hydrogen decreases as the O2/C ratio increases

while maintaining a constant value of steam-to-carbon ratio of 3 and at 700 . This

indeed is opposite to the results reported by Docter et al.[61] in which in all the

S/C cases the increase of O2/C ratio increased the production of hydrogen up to a

certain limit and then started to decrease. The concentration of carbon dioxide for

these analyses is seen to increase with an increasing O2/C, in agreement with [61], whereas the concentration of carbon monoxide decreases. Increasing the steam-to- carbon ratio enhanced the hydrogen and carbon dioxide concentrations. As expected, the concentration of carbon monoxide decreased. Based upon these tendencies the maximum production of hydrogen occurs at a high steam-to-carbon ratio S/C = 3 where methane and carbon monoxide present their lowest value [62].

Ahmed et al.[42] stated that the ATR reaction, Equation (2.33), becomes ther-

moneutral (Hr,298 = 0) at x = 0.44, and above this point is increasingly exothermic.

However, the nominal value of the amount of oxygen is slightly smaller for [42] than

that for [48] and [62] (O2/C = 0.6). It is important to point out that the thermoneu- tral oxygen-to-fuel ratio is a function of the fuel because it depends on the number of

49 atoms of carbons, hydrogen and oxygen, and its enthalpy of formation. The maximum

eﬃciency of the reforming tends to increase with the steam to carbon ratio (S/C).

Lutz et al.[58] also investigated the eﬀect of the steam-to-carbon ratio (1 ≤ S/C

≤ 4) and the oxygen-to-carbon ratio (0.5 ≤ O2/C ≤ 0.8) on the efficiency of a POX processor. As the steam-to-carbon ratio varies, their model adjusts the oxygen-to- carbon to balance the energy required to make the steam with the exothermicity of the reaction. It was seen that at low S/C, methane is produced due to incomplete fuel conversion. Therefore, the steam-to-carbon ratio is increased raising the efficiency of the system to a maximum value of 57% at S/C = 2.25 and T = 800 . After achieving ¬ this condition, any increase on the steam-to-carbon ratio seems to decrease the efficiency due to the energy required to heat the excess steam. In addition, the molar fraction of hydrogen is seen to slightly decrease with increasing S/C due to dilution caused by the increase of water content in the reformate. However, the total number of hydrogen moles increases with the steam-to-carbon ratio.

The eﬀect of the oxygen to carbon ratio was studied at a ﬁxed steam-to-carbon ratio of 2 [58]. Contrary to the steam to carbon ratio, the oxygen-to-carbon ratio

has a negative impact on the eﬃciency of the system decreasing nearly linear as the

O2/C increases because more fuel is oxidized, as opposed to being reformed. It was

emphasized that the energy balance of this system requires at least an O2/C of 0.5,

which lies between those given by Ahmed et al.[42] and Moon et al.[62]. Furthermore,

hydrogen decreases markedly, much more than the carbon monoxide for increasing

oxygen-to-carbon ratio. As a result of the increase of oxygen in the system, equilibrium

shifts toward more steam and CO2 but as products of oxidation of the fuel (opposed

50 to reforming). The eﬃciency follows the behavior of the mole fraction of hydrogen.

Hence, a maximum efficiency of 46% was shown at O2/C = 0.85 and S/C = 2.5, which is lower than the expected efficiency predicted by their theoretical model (efficiency

52%) [58].

On the same type of reformer, POx, Moon et al.[48] found that the formation of coke is directly related to the reactor temperature (which is related to the S/C and

O2/C) which must be controlled. Three combinations of steam to carbon and oxygen to carbon ratio were tested to ﬁnd the ‘optimal’ ratios that could diminish the coke production. Low production of coke was found at S/C = 1, O2/C = 1, and T = 560

. No hydrocarbons such as C2H4,C2H6 and C3 were detected in the product stream.

Ahmed et al.[42] supported the theory that coking is reduced at high oxygen-to and steam-to carbon ratios. They deﬁned that less coking is formed in the order POx >

ATR > SR. In another publication, Ahmed et al.[65] mentioned that cooking could be avoided by employing an O2/C = 2 and by increasing the amount of vapor at temperatures above 575 .

Pasel et al.[59] showed that the O2/C ratio is the most dominant factor on their

ATR processor because the temperature of the catalyst is strongly increased as this ratio increases. Contrary to the results obtained for POX by Lutz et al.[58], the hydrocarbon (diesel) conversion strongly increases with increasing O2/C ratio and is further enhanced by using a higher S/C ratio. In their case the maximum production was found at S/C = 2.2.

51 2.3.3 Non-Catalytic Reforming Processes (Thermal POx)

Although catalytic reforming is a good approach to the generation of hydrogen, it is still undergoing extensive research. Moreover, these methods are still highly expensive since they require catalyst materials to drive the reactions, usually have inherently limited lifetimes, and ﬁnally are often very sensitive to impurities contained in the gas supply. Therefore, the optimum route to produce hydrogen from conventional fuels is still debatable.

Technical literature on non-catalytic reforming is scarce and it is mostly focused on partial oxidation of methane. This opens the possibility of exploring not only the partial oxidation of diesel but also the autothermal reforming of diesel using water vapor. Rostrup et al.[66] claimed that addition of vapor is not beneﬁcial because the temperature of the reformer is lower and results in soot formation. However, no further analysis is found to explain this behavior. Conversely, Roth et al.[67] mentioned that addition of steam increases the production of hydrogen. The production of hydrogen depends directly on the type of application, e.g., for reduction of NOx particles in engine exhaust, using H2-SCR catalysts a low ratio between NO and H2 is targeted in order to reduce the fuel penalty when using on-board reformation of diesel fuel.

For non-catalytic reforming, most of the eﬀorts have been directed to the partial oxidation of methane and only a few have focused on diesel fuels. The process of partial oxidation is based on rich combustion (high equivalence ratio) [68]. The fuel is ideally oxidized to produce mainly CO, CO2,H2 and H2O. Due to the nature of the process, non-catalytic reforming is sulfur poisoning-free while maintaining high temperatures.

This reaction is highly exothermic, hence, the heat generated can supply the energy

52 necessary for steam reforming. Hartmann et al.[69] studied the reforming of IGO

(fuel similar to diesel fuel) obtaining eﬃciencies of 70% for pure partial oxidation and

claimed that addition of water will improve it. The deﬁnition of the eﬃciency is given

in Equation (2.34).

(n ˙ +n ˙ ) H η = H2 CO m,H2 (2.34) m˙ fuelHu

wheren ˙ H 2 andn ˙ CO represent the molar ﬂux for hydrogen and carbon monoxide (mol/s),

respectively. Hm,H2 is the lower heating value of hydrogen (J/mol) and Hu is the lower

heating value of the fuel (J/kg), andm ˙ f is the mass ﬂow rate of the fuel (kg/mol). On

the contrary, Rostrup-Nielsen et al.[70] analyzed the reforming of methane and claimed

that the gas composition cannot be adjusted by addition of steam to the same degree

as for ATR, because the addition of steam to POx results in a lower temperature and

thus coke formation. Roth et al.[67] mentioned that addition of vapor has very little

eﬀect on the soot depletion, however, the hydrogen production rises with additional

vapor.

2.3.4 Characterization of Non-Catalytic Processes

The chemical composition and structure are inherent to the fuel, but the physical

properties such as the temperature of the reactor and feeds, equivalence ratio, pressure

and steam to carbon ratio inﬂuence the reactor’s behavior. According to Naidja et al.

[68] even the droplet size and distribution have to be controlled prior to reforming to

achieve a maximum conversion of the liquid fuel minimal soot and NOx formation [68].

This applies to both processes, catalytic and non-catalytic.

53 A brief introduction to the eﬀect of the temperature, steam-to-carbon ratio (S/C)

and the equivalence ratio (φ) on the hydrogen production for non-catalytic processes

is given in the following sections.

2.3.4.1 Temperature POx

Two temperature regimes have been identiﬁed in the oxidation of hydrocarbons.

The ﬁrst regime focuses on low temperatures, i.e., cool ﬂames. According to Naidja

et al.[68], it has been seen that some fuel-air mixtures react chemically at low tem-

peratures ( 393 K) and produce weak flames generating very little heat. This heat ¬ comes mainly from the breaking and reforming of the fuel chemical bonds. The authors strongly suggested that further study on this field is necessary (unstable flames) as this seems to be a source for higher hydrogen yields.

The second regime concentrates on the actual combustion of the hydrocarbon, reaching high temperatures. Lemke et al.[71] studied the combustion of extremely fuel-rich

(φ = 4) methane/air mixtures at elevated pressures by non-catalytic partial oxidation.

They claimed that this technique seemed like the obvious alternative to catalytic partial oxidation and steam reforming of methane, since they accomplished essentially the same process as in CPOx. Thus, this combustion reactor could be very compact and have low capital and operating costs. However, the ﬂame temperatures are fairly low

(1000 K), and thus kinetics are much slower than conventional ﬂame reactors. They stated that there are many items that have to be addressed for these reformers: a) it is unknown whether or not equilibrium compositions can be realized in reasonable residence times, b) the kinetics of very fuel-rich mixtures are less well-known than those of near stoichiometric mixtures, c) the formation of carbonaceous deposits in the reactor

54 and, finally, d) the combustion of fuel rich mixtures is often characterized by transients that complicate stable flame operation and by low flame speeds that make stabilization very difficult.

Lemke et al.[71] used numerical techniques to investigate a TPOx reformer for adiabatic conditions, as well as for isothermal ﬂows held at the inlet temperature, using the

Barbieri chemical mechanism [72]. It was identified that this mechanism has a potential weakness since it does not account for the prediction of higher order hydrocarbons and carbonaceous deposits. The authors argued that carbonaceous deposits can be neglected since they are typically formed relatively late in the combustion sequence, well after ignition, peak temperature, and the major oxidation reactions. Their results for the adiabatic case showed that the reaction proceeds quickly to autoignition and the temperature overshoots the adiabatic flame temperature before falling towards the equilibrium. Conversion of fuel to products is only around 60% at 4 s for the adiabatic case, whereas it is roughly 34% for the isothermal cases. Neither the adiabatic nor isothermal case reaches chemical equilibrium. Instead, the combustion event produces a mixture that has significant chemical potential, but the reactions required to get this mixture to equilibrium are very slow. Increases in pressure (above 10 atm) seem to positively impact the reaction times for some cases faster than 10s. However, the temperature effect on the reaction times is crucial since increasing the inlet temperature by only 50 K, results in a decrease of an order of magnitude ( 1 s). ¬ Hartmann et al.[69] utilized marketable IGO as fuel for TPOx. Their reformer consisted of two chambers, the first type is only a mixing chamber that allows cool

ﬂames, while the second is where the main oxidation reactions or combustion take place.

55 Figure 2.1. Product Composition as a Function of Temperature for IGO POx at λ = 0.4 - 0.46 [69]

They concluded that reforming temperatures below 750 lead to undesirable formation of hydrocarbons and soot. Pure partial oxidation reforming eﬃciencies of ηref = 0.75 can be expected for λ = 0.35 (φ = 2.85). They claimed that this value can be raised by supplying water to the system. For hydrogen production, the inﬂuence of the reaction temperature is seen to be more important than the air ratio. Temperatures above 1200

are necessary to achieve a satisfying hydrocarbon conversion, as shown in Figure 2.1.

Therefore, air to fuel ratios about λ = 0.45 (φ = 2.22) and higher have to be used.

Their best reforming eﬃciency ηref = 0.70 is reached at λ = 0.46 (φ = 2.17).

Roth et al.[67] studied the temperature eﬀect on soot formation for non-catalytic reforming of diesel fuel within 800-1300 and brieﬂy described the behavior of CO,

56 CO2,H2 and CH4. Soot is seen to increase strongly with temperature although a max-

imum is observed between 1200 and 1300 while maintaining φ = 3 and S/C = 0.6.

Hydrogen and carbon monoxide concentrations are seen to increase with increasing

temperature. Carbon dioxide is seen to decrease with temperature. Methane pre-

sented an increasing behavior with temperature reaching a maximum at 1000 and

totally vanishing at 1300 .

2.3.4.2 Equivalence Ratio and Steam to Carbon Ratio

Roth et al.[67] investigated the sooting behavior of diesel fuel for the main oper-

ating parameters temperature (T = 800-1300 ), pressure (p = 1-3 bar), equivalence

ratio (φ = 1-3), and steam ratio (S/C = 0.2-0.6) at constant residence time of 400

ms. Numerical simulations were performed using the diesel combustion mechanism

of Dagaut [73] consisting of 298 species up to C16 and 2350 reactions and the mech-

anism of Marinov [74, 75] for soot modeling by PAH formation and growth utilizing

149 species and 671 equations. The ﬁnal mechanism consists of 356 species and 2542

reactions. As the equivalence ratio (φ) is varied, under the conditions of 1 atm, 1300

and S/C = 0.6, the rate of formation of soot increases. For the same conditions, the

hydrogen and carbon dioxide concentrations increase. Their model predicts that car-

bon monoxide (CO) will saturate faster than the hydrogen concentration. Fairly good

results are shown between the simulation and the experiments. However, the model

seems to overestimate the production of hydrogen and carbon monoxide. Carbon diox-

ide (CO2) shows a decreasing behavior with increasing equivalence ratio agreeing very well with their numerical prediction. Methane (CH4) formation is almost negligible

57 reaching a maximum of 0.9% at equivalence ratio φ = 2.0 for their experiments but ¬ their simulation shows a slow increasing behavior with increasing equivalence ratio.

Steam addition can be characterized based on the temperature [67]. At low temperature steam acts rather as a diluent, whereas at high temperatures it decomposes to H, O and OH radicals, ensuring a smoother reaction process. A minor eﬀect on the soot formation is seen by increasing the steam to carbon ratio from 0.2 to 0.6, under the conditions of 1 atm, 1300 and φ = 3. Nevertheless, increasing the steam ratio results in a soot abatement of nearly 30% experimentally and 20% numerically. The hydrogen and CO concentrations experience a slight increase ( 5% units) with increas- ¬ ing steam ratio. The predicted CO concentrations show the opposite trend than the experimental. As expected, carbon dioxide increases with the steam reforming due to the water gas shift reaction. Hartmann et al.[69] observed that the reforming temperatures can be controlled by supplying water vapor because it provides endothermic reactions. Their hydrogen and carbon dioxide yields increase with higher water vapor supply agreeing with [67], whereas the concentration of carbon monoxide decreases.

Lemke et al.[71] performed an equilibrium analysis at 20 atm by using the Gor- don/McBride [76] code for methane and air mixtures in order to study the inﬂuence of pressure and equivalence ratio on the non-catalytic processes. The equilibrium calculations showed that hydrogen increased to a maximum value at φ = 3, where 75% of the available hydrogen from the methane molecule is produced; thereafter, it decreases drastically. The methane conversion achieved a 100% for equivalence ratios lower than

2.5; however, it decreases. Carbon monoxide followed a similar behavior to that of hydrogen, while carbon monoxide and water vapor were relatively constant. The authors

58 strongly emphasized that solid carbon is signiﬁcant for equivalence ratios larger than 3.

The experimental results showed that the hydrogen concentration decreased with equivalence ratio; however, their values are much lower than those obtained at equilibrium

(e.g., 0.8 instead of 1.2 at equilibrium at φ = 4).

Borup et al.[77] studied a gas-phase (thermal) partial oxidation reactor coupled with catalytic steam reforming, for gasoline, since the gas-phase partial oxidation satisfies the on-board reforming start-up target of < 30 seconds to maximum power demanded by the DOE (Department of Energy) since no pre-heating before light-off of the reactor is required. Since the fuel mixture is rich, incomplete oxygen conversion occurs and then the adiabatic temperature rise is reduced. Significant conversion is found only for fuel-rich cases. The results show that the steam content lowers slightly the oxidation temperature and has a negligible effect on the ability for the gas-phase oxidation to occur. Special care must be taken while adding aromatic compounds since they decrease the relative conversion and increase the need for additional noble metal loading. The authors mentioned that the addition of aromatics will increase the amount of catalyst required for the catalytic partial oxidation reactor by almost a factor of 2.

In addition, aromatic hydrocarbons cause signiﬁcant carbon formation mainly during the start-up, when water may not be available.

2.3.5 Plasma Reforming

Plasma reforming is often referred as a catalytic process due to the eﬀect produced by non-equilibrium plasmas. However, in this thesis the term “catalytic reforming” refers to the process in which there is a physical catalyst, and “plasma reforming” refers to a reformer that uses only any type of plasma to generate hydrogen. Plasma

59 reforming has attracted attention because it eliminates some of the problems encoun-

tered in catalytic reformers, i.e., deactivation/poisoning of the active surface, while achieving relatively high hydrogen yields. Therefore, the main task of these reformers is to provide the energy and to create free radical species that enhance the hydrogen production reaction mechanism. Plasmas can be classiﬁed into two distinctive categories: thermal (equilibrium) and non-thermal (non-equilibrium).

Thermal plasmas are generated by low voltage ( 100 V) and high currents (> 20 ¬ A) and, therefore, the power injected in the discharge is high (higher than 1 kW) [52,

87]. The principal characteristic of these plasmas is that electrons and neutral species

are at the same high temperature ( 5000-10000 K) [52]. The electrons gain energy ¬ not only from the electric ﬁeld but also from collisions with heavy particles (neutral

species). These reformers tend to be compact and low weight and oﬀer high conversion

eﬃciency and fast responses. A disadvantage, however, is that the temperature in the

reactor is too high and aggressive cooling of the electrodes is performed to decrease

erosion. Furthermore, the energy consumption is too large which make their use on-

board unsuitable. Arc discharges are one of the most typical thermal plasmas.

Non-thermal plasmas are generated by very low electrical energy (50-300 W). The

temperature of the electrons is very high (up to 5000 K), while that of the neutral

species is low (near room temperatures) [52, 87]. Opposite to the thermal plasma, electrons in the non-equilibrium plasma only receive energy directly from the external electric ﬁeld and not from collisions. Hence, the main goal of these reformers is to generate chemical excited species and radicals through collisions with the excited electrons to initiate and enhance the chemical mechanisms. Due to their low-temperature,

60 low-electrode erosion and low-power consumption characteristics, non-equilibrium plasmas are very attractive for on-board reforming. Unfortunately, their production of hydrogen is lower than that for thermal plasmas. The most common types of non- equilibrium plasma reformers are dielectric barrier discharges, gliding arc discharges

(G.A.) and microwave discharges (M.W.). Examples of this type of plasma reformers are presented in this section because these reformers can be used for NOx remediation.

Partial oxidation of kerosene and methane using a microwave discharge at atmospheric pressure was studied by Babaritskii et al.[50]. A 2.45 GHz microwave frequency was employed and the M.W. power was varied in the range of 500 - 3000 W.

Methane reformation showed a dry gas composition of 31.1% H2, 17.0% CO, 47.8% N2,

1.2% CO2, 1.9% CH4, 0.8% C2H2 and 0.2% C2H4, while kerosene reformation at 3 kW results in 14.3% H2, 20.2% CO, 58.4% N2, 2.6% CO2, 2.9% CH4, 1.4% C2H2 and 2.6%

C2H4. The hydrogen concentrations for both fuels are relatively low compared with those obtained for catalytic reforming, although they could satisfy the hydrogen demand for NO2 reduction technologies. Nevertheless, the amount of energy used by this system is extremely high and would result in an elevated fuel penalty. From this work the main contributions are that the ﬂame propagation time for a methane-air mixture is observed to be less than 0.1 s, achieving half of the equilibrium amount of synthesis gas. However, the other half is obtained very slowly ( 105) in the remaining of the ¬ chamber. Kerosene presents a similar behavior, however, the chemical equilibrium is achieved faster ( 103-104) than that for methane in the reaction chamber. ¬ Sekiguchi et al.[51] employed a M.W. with a frequency of 2.45 GHz and maximum discharge power of 2.8 kW for steam reforming using n-hexadecane as a model of

61 gasoline. H, OH and O radicals as well as high energy electrons are generated, allowing

reduction and oxidation processes to occur in the plasma. The discharge power was

varied from 1.5 to 2.5 kW to study its eﬀect on the hydrogen production, which shows

an increasing behavior with increasing power. CO behaves similarly to H2, while water

content decreases with discharge power. It is important to notice that neither water

vapor nor hexane are completely converted even at the high discharge power. The

authors believe this is due to distribution and mixing of vapor and fuel.

Paulmier et al.[52] used a gliding arc to reform gasoline (California gasoline -

Syntroleum) by ATR or SR under air ratios between 0.1-0.4 and steam ratios between

1-3 for ATR and 1-5 for SR, pressure of 1-3 x 105 and feeds temperature of 773 K. Their results show that the maximum H2 yield for ATR is 7% for an air ratio of 0.1, S/C = 2.5

and P = 2 x 105. The hydrogen concentration decreases as the air ratio increases, while

CO increases. Steam reforming shows, instead, a much better performance achieving

a maximum hydrogen generation of 31% using an air ratio of 0, S/C = 4-5 and P = 1

atm. The hydrogen concentration increases with increasing S/C ratio.

2.3.6 Cleaning Up Processes

Depending on the ﬁnal application of the reformate, several clean up processes must

be carried out upstream and downstream of the processor to avoid deactivation of either

the SCR or LNT catalysts, or membranes to separate hydrogen for fuel cells. Within

these processes are: i) desulfurization, ii) water gas shift (WGS) reaction (CO elim-

ination), methanation (CO elimination), selective membranes (highly selectivity for

62 hydrogen), preferential oxidation (PrOx to eliminate CO), and avoiding coke forma-

tion. A brief review of the technical literature concerning these cleaning up systems

is introduced in the following sections.

2.3.6.1 Desulfurization

As mentioned before, sulfur is one of the most poisonous elements present in current

heavy hydrocarbons; in some of them the content of sulfur is excessive (50-300 for gaso-

line and 0.5% higher for diesels) [56]. Even for the ultra-low sulfur fuels, which only

contain < 30 ppmw in gasoline and < 15 ppmw in diesel, that satisfy the EPA speciﬁ-

cations, the sulfur contents are still too high [4]. Thiophene alkyl- or aryl-substituted

thiophenes, with lesser concentrations of sulﬁdes, disulﬁdes, and mercaptans are com-

mon sulfur constituents [78]. If diesel were directly employed in a processor, the life

of such a processor would be short (with low eﬃciency) due to poisoning of the catalyst

as sulfur chemisorbs on any metal surface [60, 79, 80]. Besides, sulfur reacts with hy-

drogen at low temperatures (300-400 ) producing H2S in a non-oxidizing atmosphere,

and with oxygen producing sulfur oxides (SOx) at higher temperatures (> 600 ) in an oxidizing atmosphere, polluting the air and decreasing the eﬃciency of the reactor

[53]. For these reasons, the sulfur content of fuels must be regulated through diﬀerent

techniques.

One of the methods to reduce the amount of sulfur in the hydrocarbons is the

hydrodesulfurization, commonly employed in the reﬁneries; however, in order to enable

the reaction of sulfur with hydrogen to yield H2S, it employs elevated temperatures and

high pressures, as well as a high H2-to-feedstock ratio [79]. This, unfortunately, would

not be applicable to the fuel processors mainly because of the energy cost of compressing

63 hydrogen. L¨oﬄer et al.[79] recommended the use of absorbents beds upstream of the processor to capture the sulfur species. These beds are generally metal-impregnated high-surface area supports such as active carbon, zeolites, or alumina. They claimed that this technology can be applied to the processors since the size of the beds depends only on the amount of sulfur contained in the hydrocarbon. Fukunaga et al.[81] agreed with this theory and developed a Ni-based catalyst that was implemented in the desulfurization process of kerosene. Such a catalyst reduced the content of sulfur to 1 ppm in a 4000 h test (which corresponds to 1 year of operation of a fuel cell system during daytime only). Commercial Ni catalysts were also tested in this analysis for comparison purposes showing that the reduction of sulfur content with these catalysts was only 15-42 ppm.

King et al.[78] proposed an adsorbent-ﬁlled cannister as the sulfur removal technology due to its easy replacement for automotive applications. This liquid phase desulfurizer (ZSM5) was submitted to several cycles to determine its degree of reversibility by measuring the diﬀerence in adsorption capacity between several cycles. A cycle ranges from time zero until the adsorbent is completely saturated. Once saturation is acquired, the desulfurizer is replaced by another one, and then cleaned with nitrogen for reuse. For testing purposes a network of desulfurizers is used in order to assure continuity of the testing and simulate a real application. These adsorbents showed a very slight decrease of reversibility after two cycles at atmospheric pressure. Besides, the adsorption capacity of these adsorbers shows a capacity of 0.012 g sulfur (0.032 g thiophene) per gram of adsorbent.

64 2.3.6.2 Water Gas Shift Reaction

Catalysts are very susceptible to deactivation due to the presence of CO and other

contaminants in the reformate stream. As seen in several studies [54, 82, 83], the

tolerance of CO contamination in the stream must be smaller than 10 ppm for suc-

cessful operation for fuel cells. Carpenter et al.[83] cited that the usual amount of

CO produced in most of the reformers is about 0.5-3% during steady state operation,

and approximately 5-10% during start-up or transients. Based upon these facts, the

stream arriving to the fuel cell stack must be cleaned up so that only CO2 and H2 are contained in the reformate.

Hagan et al.[82] illustrate several CO clean-up mechanisms such as i) the pressure

swing adsorption systems (very complex to control and low product recovery), ii) the

separation membranes (highly successful but limited by cost and the need for a high

pressure feed), iii) the selective methanation (being the simplest option but not useful

for high CO concentrations and loss of hydrogen concentration), iv) the water gas shift

reaction (increases the concentration of H2) and, ﬁnally, v) the preferential oxidation

(which has reported a good eﬃciency reducing the CO concentration to less than 10

ppm).

The reduction of the CO concentration is achieved by the addition of the water gas

shift reaction in the processor [64]. This reaction is given in Equation (2.35).

CO + H2O −→ CO2 + H2 (2.35)

By analyzing this equation it can be seen that additional H2 is generated even though

the main objective of using this reaction is to eliminate or reduce the amount of CO in

65 the reformate gas [54]. Doss et al.[64] observed that the reaction equilibrium favors the conversion of CO to CO2 at lower temperatures than the temperature dictated by the kinetics of this reaction. The water gas shift reaction is carried out in two stages: i) the high temperature shift reactor occurring at approximately 700 K and over Fe-

Cr-oxides catalysts) to convert part of the CO contained in the gas stream followed by, ii) the low temperature shift reactor ( 480 K) to provide further reduction of the CO ¬ concentration [64, 84]. Wieland et al.[84] mentioned that these reactors were ﬁrstly developed for industrial applications as pelletized ﬁxed bed catalysts, which made the

WGS reaction not suitable for mobile applications. However, there are now several non-pyrophoric shift catalysts (not self-igniting catalysts upon exposure to air), such as coated monoliths systems, foams or metallic substrates that achieve high space-time yields.

2.3.6.3 Methanation

As mentioned in the previous section, the selective methanation mechanism is a good option to remove the CO in the reformate only if the CO concentration is low (preferably less than 0.5%) [83]. However, while this reaction takes place in the fuel processor, the hydrogen concentration decreases because hydrogen tends to react with the CO as shown in Equation (2.36) producing methane and water [60, 64, 83], decreasing therefore the eﬃciency of the reforming process.

CO + 3H2 −→ CH4 + H2O (2.36)

It has been pointed out that special care must be taken while having CO2 in the reformate stream, because i) carbon dioxide can react with the produced water in a

66 reverse water gas shift reaction producing CO [60]; ii) carbon dioxide can react with

hydrogen producing methane and resulting in the generation of only methane and water

(run-away reactions), as shown in Equation (2.37)[83]. Qi et al.[85] emphasized the

fact that the methanation process is highly exothermic, which can cause the run-away

reactions, and operates in the temperature range of 200 to 400 .

CO2 + 4H2 −→ CH4 + 2H2O (2.37)

The methanation reaction has been found to occur due to the POx reforming tem-

perature where the amount of methane is reduced as the temperature in the reformate

increases [64]. Moreover, the increase in the steam to carbon ratio drastically reduces

the production of methane. While combining the eﬀects of the POx temperature and

the S/C ratio, the CH4 production seems to be reduced to less than 1%. Doss et al.

[64] reported that a 1000 K temperature favored the formation of CH4 thermodynami-

cally. Thus far, it could be claimed that the creation of methane could be considered

as a drawback because it does not undergo any reaction in the fuel cell system. How-

ever, they suggested that this CH4 can be burned in an external burner to heat up the

reactants of the fuel processor. Hence, the generation of methane through the fuel

processor may or may not result in an eﬃciency penalty (obviously depending on how

eﬀectively the burner heat can be used).

2.3.6.4 Selective Membranes

The Pd (Palladium) and Pd-alloy membranes can be very eﬀective in the removal

of carbon monoxide from the reformate stream. These membranes selectively allow

the hydrogen among the other constituents of the stream gas to permeate [60, 83].

67 The permeation of hydrogen through the membrane occurs in a series of processes:

“dissociative adsorption on the reformate face, diﬀusion of the hydrogen atoms through

the bulk metal, and recombination of the hydrogen atoms on the permeate face” [79].

In these CO-removal mechanisms the partial pressure difference over the membranes defines the driving force, which depends then on high inlet pressure of the stream. In addition, the rate of permeation is proportional to the difference in concentration of hydrogen atoms on both sides of the membrane and to the difference of the partial pressure of hydrogen [79].

Carpenter et al.[83] illustrate that at least 10 bar pressure diﬀerence is required in

recovering an 80% of the hydrogen in a 40% (wet) mixture. In order to obtain such a

big pressure diﬀerence, a multistage compressor would be required to compress the air

to the reformate, which would decrease the eﬃciency of the fuel cell system and would

increase its size. Another disadvantage of this technology is that the temperature

required is high, which negatively impacts the eﬃciency of the system [60].

2.3.6.5 Selective or Preferential Oxidation

In the selective or preferential oxidation reactor, the reformate undergoes a reaction

with a controlled amount of air over a suitable catalyst, such as alumina (Al2O3) supported structures (Ru and Rh are among the most active catalysts) [60]. Preferential

oxidation could be used either by itself as the CO removal process or following the

water gas shift (WGS) reaction. When used after the WGS, oxygen is injected into

the stream as air to react with the remaining CO following Equations (2.38) and (2.39).

1 CO + O ←→ CO (2.38) 2 2 2

68 1 H + O ←→ H O (2.39) 2 2 2 2

Pukrushpan et al.[54] cited that the amount of air injected into the PrOx must be twice the amount of CO at the inlet of the reactor, i.e., twice the amount needed to

maintain the stoichiometric reactions in Equations (2.38) and (2.39) (100% excess O2).

This can be explained by the description given by Doss et al.[64] and Qi et al.[85], in

which the eﬀective factor for the catalyst is assumed to be 50%, i.e., only half of the

amount of oxygen introduced reacts with the remaining CO and the other half reacts

with very small portions of hydrogen in the fuel gas. The gas leaving the PrOx is

then assumed to have no carbon monoxide in it [54, 64]. Besides, selective oxidation

prevents the loss of hydrogen as opposed to the methanation process described in the

previous section [60]. Qi et al.[85] introduced the concept of the catalyst selectivity as the ratio between the number of moles of CO removed and the total number of moles of hydrogen and carbon monoxide removed. The selectivity depends greatly not only on the type of catalyst, but also on temperature, pressure, steam content, and hydrogen and carbon monoxide concentrations in the reformate stream [85].

The operating temperature of the PrOx reaction must be low because the catalyst

employed to separate CO favors the reaction of hydrogen with carbon dioxide at high

temperatures which will result in unwanted CO [86]. This low temperature regime

is considered advantageous for the overall fuel processor eﬃciency since the PrOx can

therefore be operated between the LTS ( 200 ) and the PEM Fuel Cell Stack ( 80 ¬ ¬ ) without external heating, and also from low to high pressure [85]. Mitchell et al.

[86] indicate that PrOx is a suitable application under steady state conditions and that

thus far it is more diﬃcult to achieve good transient response. Conversely, Qi et al.

69 [85] emphasize that PrOx technology requires short start up times which ﬁt the transient response of automotive applications well. Mitchell et al.[86] performed several tests to corroborate their theory under steady state conditions with a 45 kW thermal input fuel processor. The attained CO concentration was less than 10 ppm and the hydrogen losses were about 0.1 to 2% in all cases. In addition, the same processor was used for the transient analysis operating on gasoline, natural gas and methanol. This showed that this processor was capable of performing 17.5 kWthermal/sec transients and was able to maintain a CO concentration smaller than 6 ppm. These transient tests comprised step increases and decreases in thermal input (from 10-45 kW and 45-10 kW, respectively) for up to 2 seconds, and were adjusted by controllers to maintain the

CO concentration less than 10 ppm.

2.3.6.6 Coke Formation

Coke formation has been considered a pervasive problem in steam reformers, because decomposition of hydrocarbons can be promoted at high temperatures (typically over 300 ). Carrete et al.[60] pointed out that this decomposition depends on the saturation of the hydrocarbons fed into the reformer (i.e., the amount of double bonds in the system), which can favor the dehydrogenating instead of the steam reforming, forming thus solid species [79]. Those species known as coke, can deactivate the catalyst in diﬀerent ways: i) coke diﬀusing through the metal crystal causing an aging of the catalysts, and ii) depositing on the active sites, causing a progressive deactivation

(possible polymerization of CnHm radicals) [60].

The accumulation of those solid particles limits the catalytic activity by blocking the catalyst pores and can cause the metal to separate from the support [60, 79]. It

70 has been seen that coke may grow in the interparticle spaces which will increase the

pressure drop in the catalyst bed leading to an undesirable blocking of the gas ﬂow

[79].

In order to avoid coke formation in the steam reformers, an optimum fuel to steam

ratio must be maintained throughout the entire system. Steam can act as a coke

inhibitor, and the produced hydrogen is also a inhibitor of coke formation [53, 79, 61].

Furthermore, the risk of coke formation can be reduced by contacting the fuel and the steam with the catalyst at relatively low temperatures (lower than the carbon limit temperature 500 ) and allowing the reaction to proceed to equilibrium [79]. ¬ This pre-reformer stream can be heated up by means of the heat produced in the burner in the steam reforming, and then be fed into the high temperature reformer, improving then the eﬃciency of the reformer. Finally, other techniques to reduce the coke formation are smaller metal crystallites and promotion of the catalysts with alkaline material [60].

2.4 Concluding Remarks

Diﬀerent technologies for the remediation of nitrogen oxides have been presented

in this Chapter. Lean NOx traps work by storing the nitrogen oxides and have to be

regenerated over a period of time by reducing agents such as hydrocarbons, carbon

monoxide or hydrogen. HC-SCR utilizes additional hydrocarbons (diesel, propane,

ethene, methane, among others) as the reductants that have to be administered up-

stream of the catalyst (Cu or Pt). These catalysts have shown conversions of about

71 22-77% [29, 34]. Urea-SCR systems have achieved a conversion eﬃciency about 60-

90% over Cu-ZSM5 catalyst in laboratory tests [36], and about 63-65% over a the

driving cycle for real applications. H2-SCR systems have also shown high conversions

about 60-70% for low and high temperatures [40] over a Pd-Alumina catalyst and about

95% for very low temperatures for [38] over a Pt/MgO-CeO2 catalyst. Finally, plasma assisted catalytic NOx removal presented a conversion 70%. ¬ In this literature review, for non-hydrogen after-treatment technologies, several studies have showed an increase on the NOx conversion eﬃciencies when hydrogen was added along with the other reductants. For instance, hydrogen showed to be a better reductant than CO at low temperatures for the regeneration of Lean NOx traps.

Furthermore, when 0.5% hydrogen is used in combination with urea and the exhaust stream for a Urea-SCR system, the NOx conversion over a silver-alumina catalyst increased from very low values to 84-90% within a large temperature range. An HC-SCR system using propane as the main reductant showed a 23-30% conversion for a high temperature range, while using hydrogen as a complementary reductant the conversion increased to 50-60% within the same temperature range.

Based upon these improvements, the promising reductions of nitrogen oxides using

H2-SCR systems, and the feasibility of on-board reforming rather than using an extra storage tank (Urea and selected HC’s), lead to the development of a technology capable of producing hydrogen while maintaining a reasonable fuel penalty.

The most common mechanisms to generate hydrogen have also been presented in this chapter. While catalytic reforming is a good approach to the generation of hydrogen, it is still undergoing extensive research, it is highly expensive and lifetime limited

72 and, ﬁnally, very sensitive to certain impurities that co-exist in the exhaust gas. Al- though non-catalytic reforming provides lower hydrogen concentrations than catalytic reforming even at very high temperature, it is suitable for NOx reduction applications since the required amount according to the studies mentioned above is about 0.5% Vol in the exhaust stream, while maintaining a low fuel penalty.

73 CHAPTER 3

CHEMICAL KINETICS MODELING

3.1 Introduction

Prediction of the variation of concentrations of major species is important in reveal-

ing the signiﬁcant reaction pathways responsible for hydrogen production by thermal

partial oxidation (TPOx). Typically, equilibrium calculations [76] are used to predict

compositions of major species but these are not useful in elucidating mechanisms as

they do not consider the individual elementary processes. It is therefore necessary to

consider the chemistry as being out of thermodynamic equilibrium and allowing the

reacting mixture to approach equilibrium. This requires modeling reacting systems

with ﬁnite rate chemical kinetics.

Methane and small hydrocarbons have been the targeted fuels to be studied during

the past decades as far as ﬁnite chemical kinetics is concerned [57, 73, 74, 75]. However,

most of the fuels used for automotive applications, diesel, kerosene and gasolines, are

composed of a large mixture of low and high hydrocarbons. Unfortunately, literature

on the chemical kinetics of heavy hydrocarbons is very scarce. Ristori et al.[88] have been pioneers in investigating the chemical reactions that take place in the oxidation

74 of hexadecane, C16H34, which is a high carbon content molecule typically used as a surrogate for diesel fuel. The ﬁnal chemical reaction set contained 242 species and

1801 elementary processes to describe combustion at atmospheric pressure.

In general, the form of the chemical reactions for hydrocarbon combustion include those for oxidation (exothermic), those for decomposition or pyrolysis (endothermic), those reacting with small radicals and those for the decomposition of alkyl radicals.

CxH2x+2 + O2 −→ iCxH2x+1 + HO2 (3.1)

CxH2x+2 −→ Cx−iH2(x−i)+1 + CiH2i+1 (3.2)

CxH2x+2 −→ iCxH2x+1 + H (3.3)

CxH2x+2 + M −→ CxH2x+1 + HM (3.4)

CxH2x+1 −→ CjH2j+1 + Cx−jH2(x−j) (3.5) where x is the carbon content in the hydrocarbon. For instance, for hexadecane these reactions have the following values: x ≤ 16, 1 ≤ i ≤ 8, 1 ≤ j and M could be H,

OH, O, HO2, CH3,C2H5,C2H3,C3H5,C4H7,C6H5, etc. As known, the oxidation of a hydrocarbon leads to the formation of hydrogen, carbon monoxide, carbon dioxide, water vapor and methane as the major species. Thus, the mechanisms to obtain these species are: i) hydrogen is obtained by the reaction of small radicals with H atoms, ii) water comes from small radicals (such as H) and OH radicals, iii) CO is mainly produced by HCO and CH2 with O reactions and it is consumed by OH and HO2 radicals and iv) CO2 is produced by CH2O and CO or oxidation of CO. All these reactions are temperature dependent.

75 In this chapter, a chemical kinetics model of the TPOx of hexadecane based on the

mechanism developed by Ristori et al.[88] is described in Section 3.2. The formula-

tion of the constant pressure combustion (or reaction) of a given species is described

in Subsection 3.2.1. Also described in this section is the criteria used to select the initial temperature of the feeds and the calculation of the reactants concentrations for the two oxygen providers: fresh air and exhaust gas. Finally, Section 7.5 introduces

the numerical parameter used in this thesis to monitor the percentage of hydrogen ob-

tained by the TPOx under diﬀerent conditions to the maximum obtainable from the

hexadecane molecule.

The compositions of the major components obtained for this model are described

in Chapter 5 and are compared with those obtained experimentally in Chapter 7.

3.2 Problem Formulation

As the primary goal of this study is to predict the chemical compositions of the major

constituents (H2, CO, CO2,H2O, CH4) of the thermal partial oxidation of hexadecane,

as well as the correct amount of exothermic heat release, and aid in interpretation of

chemical pathways, the kinetic set of reactions found in [88] is simpliﬁed incorporating

only 113 species and 764 elementary processes, excluding the isomers of hydrocarbon

species and associated isomerization reactions. The complete set of reactions, rates,

activation energy, temperature exponent for each reaction used in this thesis is given

in Appendix A.

76 Each elementary process (forward and reverse) used is written separately. For

instance, the rth elementary process is written as:

N N X kr X αirSi −→ βirSi (3.6) i=1 i=1

where Si represents species i, kr is the temperature-dependent rate coeﬃcient for reac-

tion r, αir is the stoichiometric coeﬃcient for species i in reaction r on the reactants

side, and βir is the stoichiometric coeﬃcient for species i in reaction r on the products

side.

The model assumes that the reactor is adiabatic, under constant pressure and that

the fuel and air are homogeneously mixed. As the primary objective is to understand

the behavior of the chemical reactions and to predict the ﬁnal concentrations of the 113

selected species, the model is reduced to a 0D model, which allows only variation with

time. Should the model be extended to account for axial ﬂow, the 2-D momentum

equation accounting for mass ﬂow through the boundaries, diﬀusion, turbulence and

viscous terms would have to be implemented. Based on these assumptions, the gov-

erning equations for this system are derived in the following section. Further details

on the solution scheme for the governing equations is given in Section 3.2.1.

In order to study the behavior of an actual thermal partial oxidation reformer, the

characterizing quantities given in Chapter 2, i.e., equivalence ratio, steam to carbon ratio, temperature and pressure have to be implemented in the calculation of the initial concentrations of the feeds. Oxygen and fuel are provided as part of the initial reactants to study the inﬂuence of the equivalence ratio. Fresh air is used as the oxygen provider for this work. Water vapor can be introduced along with fuel and air as part

77 of the reactants as well. In this thesis, two different configurations for the injection of water vapor are studied to understand their effect on hydrogen generation: 1) as part of the initial reactants, and 2) at a secondary location (time) downstream of the combustion of pure air and fuel. The calculation of these initial concentrations is derived in Section 3.2.3.

Furthermore, special interest on using exhaust gas (from the diesel engine) as the oxygen provider for the reformation of fuels (syngas) to be used for the after-treatment technologies presented in Chapter 2 has emerged in the last several years. The purpose of that is to avoid the implementation of electric compressors or pumps. Therefore, a section of this chapter is dedicated to the study of exhaust gas as an oxygen carrier.

The calculation of the initial concentrations using exhaust is derived in Section 3.2.4.

3.2.1 Governing Equations

The governing equations to be solved in this model are the rate of change of each species (production and consumption) and the energy equation to calculate the temperature due to the combustion process. Starting thus from the deﬁnition of number density, ni,

N n = i (3.7) i V where Ni represents the number of molecules of the i species and V is the total volume.

Diﬀerentiating then this equation with time, t, to obtain the rate of change of the species,

dn 1 dN N dV i = i − i (3.8) dt V dt V 2 dt V Ni

78 As the pressure is maintained constant throughout the process, the volume changes.

Hence, the variation of the volume can be obtained from the ideal gas law by diﬀeren- tiating it with respect to time

PT V = NT kB T (3.9)

dP dV dN dT V T + P = k T T + k N (3.10) dt T dt B dt B T dt where NT represents the total number of molecules in the system, kB refers to Boltz- mann’s constant (1.38054 × 10−23J/K) and T is absolute temperature in Kelvin. This equation can then be reduced for the constant pressure case using the ideal gas law,

Equation (3.9), as

dV 1 dN 1 dT = V T + (3.11) dt NT dt T dt

Substituting Equation (3.11) into Equation (3.8), the governing equation for the number density is given as,

dn 1 dN N 1 dN 1 dT i = i − i T + (3.12) dt V dt V V NT dt T dt

The term dNi/dt represents the variation of the concentrations or molecules with time due to chemical reactions at constant volume,

ER " Ns # dNi X Y α = (β − α )k N jr (3.13) dt ir ir r j V r=1 j=1 where ER represents all elementary chemical processes, r is the number of reaction and

Ns is the total number of species. The coeﬃcient kr is deﬁned in the Arrhenius form

E k = AT n exp a (3.14) r RT

79 where A represents the pre-exponential factor (m3/molecules)nr−1/s), nr is the reaction

order, n is the dimensionless empirical power for the temperature, Ea is the activation

energy (J/mol), and R is the ideal gas constant (8.314 J· K−1· mol−1). The coeﬃcients

A, n and Ea are expressed for every reaction in Appendix A.

Substituting Equation (3.13) into Equation (3.12)

 dNT  ER " Ns # dni 1 X Y α Ni 1 dT = (β − α )k N jr −  dt +  (3.15) dt V ir ir r j V  N T dt  r=1 j=1 T

Rearranging then in terms of all species and number densities

 ns  X dnj ER " ns #  dt  dni X Y αjr  j=1 1 dT  = (βir − αir)kr n − ni  +  (3.16) dt j  ns T dt  r=1 j=1 X  nj  j=1

The ﬁrst term on the right hand side represents the rate of change of the concentration of species i due to chemical reactions, and the second term represents the rate of change due to a change in volume for species i at constant pressure. The temperature is then found by solving the energy equation which can be written in terms of the speciﬁc enthalpy (enthalpy per unit mass), hi,

N X Q = Nihi (3.17) i=1

Then diﬀerentiating with respect to time and as the process is assumed to be adiabatic, the change in the total enthalpy is zero,

N N X dNi X dhi 0 = h + N (3.18) i dt i dt i=1 i=1

80 Assuming ideal gas at constant pressure, hi = CpiT for the diﬀerential of the enthalpy. Thus,

N N X dNi X dT 0 = h + N C (3.19) i dt i pi dt i=1 i=1

Rearranging,

Ns X dNi hi dT dt = − i=1 (3.20) N dt Xs NiCpi i=1 Equations (3.16) and (3.20) represent the governing equations for the 0D model in

the (Ns+1) unknowns ni and T. This system of coupled equations is highly non-linear, and very stiﬀ due to the slow and rapid variations of the concentration and temperature gradients. Matlab is selected to be the platform to solve the governing equations and a built-in solver ODE15s is employed to account for the stiﬀness of the system. This is a multi-step variable-order solver.

Ignition is indirectly modeled so that it must be specified in the form of an initial condition for the temperature. The initial concentrations of the feeds are determined for the two different oxygen providers and the water vapor configurations. Further details on the initial temperature selection and the calculation of the initial feeds are given in the following subsections.

3.2.2 Initial Temperature of the Feeds

The initial temperature used in the simulation must be high enough that the mixture will ignite in a reasonable amount of time but within reasonable temperature limits

81 since the final temperature of the reacted products depends on this initial temperature in the adiabatic calculation. Ignition is defined here as the rapid or abrupt increase of the temperature. In order to identify the initial temperature, the conversion (dissociation) of hexadecane to small hydrocarbons (CH4,C2H4,C3H6,C4H6,C6H12, etc.) and hydrogen is analyzed for four different temperatures (900, 1000 1100 and 1200 K), while no air is present. This process is simulated over a period of time of 10 seconds, although the residence time of the mixture in the experimental reformer is about 1 second as discussed in Chapter 6.

The conversion rate of fuel increases with increasing temperature is shown in Figures

3.1 and 3.2. Figure 3.1 illustrates the conversion of hexadecane at 900 K. It can be seen that at the end of 1 second, only 30% of the fuel is converted while producing a

7% hydrogen concentration. Increasing the initial temperature by only 100 speeds up the fuel conversion rate so as to achieve 70% (see Figure 3.2) while the hydrogen concentration increases to 15%. The other initial temperatures (1100 and 1200 K) show a larger hexadecane conversion (85 and 93%) and larger hydrogen production

(20% and 23%) at 1s. Based upon the conversion rate, the simulation-ignition-time will strongly depend on the temperature of the feeds which is directly translated to a longer computing time.

An oxidation case at stoichiometry and initial temperature T = 900 K is shown in

Figure 3.3 to illustrate the ignition delay due to the value of the initial temperature.

It can be observed that ignition takes place slightly after 7 seconds and the ﬁnal temperature is close to 2800 K. If the same case is run for an initial temperature T = 1000

K, ignition occurs after only 0.2 seconds and the ﬁnal temperature is about 2850 K.

82 Figure 3.1. Fuel Decomposition and Major Species at 900 K and 1 atm

Figure 3.2. Fuel Decomposition and Major Species at 1000 K and 1 atm

83 Figure 3.3. Time History of the Temperature and Concentrations for φ = 1.0, S/C = 0.0, 1 atm and Tini = 900 K

A more detailed analysis on the inﬂuence of the initial temperature on the ﬁnal

product composition is given in Chapter 5, from which it is seen that starting at a higher temperature yields immediate ignition. Therefore, the selection of T = 1000 K as the initial temperature not only guarantees ignition in a reasonable computational time, but also allows for further insight into the species and reactions behavior prior to ignition, during combustion (fast temperature gradient) and thereafter (cooling).

Furthermore, in Chapter 7 the ‘t’ coordinate is transformed into an axial coordinate

‘x’ to correlate the numerical simulation of the partial oxidation of hexadecane and the

experimental reformer. The initial temperature of 1000 K gives reasonable ignition

lengths as observed in the experiments reported here. The experimental reformer used

in the experiments consists of a mixing chamber and a combustion chamber that is

84 directly linked to a reaction chamber where the products presumably reach equilibrium.

These are completely consistent with the case with initial temperature of 1000 K.

3.2.3 Initial Concentrations when Using Pure Air

In order to understand the behavior of the partial oxidation and autothermal reform-

ing processes when using pure air as the oxygen supplier, the characterizing parameters

described in Chapter 2, equivalence ratio, steam to carbon ratio, pressure and temperature are varied over 1.0 ≤ φ ≤ 1.9, 0 ≤ S/C ≤ 2.0, 1.0 ≤ P ≤ 3 atm and 1000

≤ T ≤ 1750 K, respectively. For a given equivalence ratio and initial temperature

Tinit, while maintaining the pressure, the number densities of C16H34,O2,N2 and H2O

are determined from the deﬁnition of the equivalence ratio, the steam to carbon ratio,

the ideal gas equation of state and the molar balance of the single elements (C, H, O

& N). Starting then from the ideal gas equation,

P = nT kT (3.21)

3 where P represents the pressure in Pa, nT is the total number density (molecules/m ),

k is the Boltzmann constant (1.38×10−23 J· K−1) and T is the temperature (K). For

instance, the number density for the calculations where P = 1 atm and T = 1000 K

assuming V = 1 m3 is 7.3423×1024 (molecules/m3). The individual initial composition

of the reactants (O2,N2,H2O and C16H34) can be calculated based on the reaction,

Equation (3.22).

aC16H34 + b(O2 + 3.7755N2) + cH2O −→ dCO2 + eH2O + fN2 (3.22)

85 where a, b, c, d, e and f are the reaction coeﬃcients which can be calculated from a molar balance, while using the deﬁnition of the equivalence ratio and the steam to carbon ratio given as

(F/A) φ = actual (3.23) (F/A)stoic

n n S/C = H2O = H2O (3.24) nC NC · nfuel

X X NT = Ni = Avo ni (3.25) where F/A represents the fuel to air ratio in mass basis, the subindex actual contains the mass of fuel and air used whereas stoic indicates the F/A ratio at stoichiometric

conditions, S/C represents the steam to carbon ratio, nH2O, nC and nfuel represent the water, carbon and fuel number of moles, respectively, and NC is the number of carbon atoms in the molecule. The expressions to calculate the coeﬃcients in their ﬁnal form are given in Table 3.1.

Coeﬃcient Equation Coeﬀ Equation N /Avo a = T d =16a (M /M ) 1 + 16(S/C) + f air φ(F/A)stoic M b = f a e =16a(S/C) + 17a 4.7755φMair(F/A)stoic 3.7755M c =16(S/C)a f = f a 4.7755φMair(F/A)stoic

Table 3.1. Molar Coeﬃcients for the general Auto Thermal Equation (3.22)

86 These formulations are valid for any steam to carbon ratio and equivalence ratio if injected at the same injection point. The extent of this research includes not only the injection of water vapor along with the other reactants, but also the post-injection or secondary injection of water vapor after the partial oxidation process to investigate diﬀerent routes to improve the production of hydrogen. Hence, the post-injection process is more complicated to simulate due to the fact that the volume is allowed to vary with temperature. At this point it is important to describe how the simulations are performed in order to complete a case. In order to mimic the experimental process with this numerical code, several simulations have to be performed to complete one case in a time frame of 0.7-0.9 s which is the residence time of the actual reformer.

As known, heat losses are inevitable in real experiments and, as it will be explained in the experimental section, the products have to be cooled to 500 K. For the cases where the feeds are injected at the same location, the ﬁrst step is calculate the number densities for each compound entering the reformer at the initial pressure and temperature. The mixture is then combusted in order to reach the adiabatic ﬂame temperature ( 0.2 s). ¬ Once this has been achieved, the products are cooled to 500 K for fair comparison with experimental data.

For the secondary water vapor injection cases, the process is more complex. In order to imitate the experimental process, the number densities of air and fuel with S/C

= 0 at any equivalence ratio are calculated. The mixture is ignited up to the point of maximum temperature (adiabatic temperature)and then cooled down to a certain temperature (1500 or 2000 K). As the process in the numerical code is at constant pressure, the volume expands while igniting and then contracts when cooling. These changes in

87 volume must be accounted for with water vapor injection. Hence, the amount of water that should be used for main injection at a particular S/C ratio is transformed to a new basis or volume. Adding water to the reactor will increase the volume, which has to also be taken into account. The post-injection section lasts about 0.4 s and then the products are cooled down to 500 K for comparison with the experimental data. The temperature of 500 K arises from the fact that the gases entering the gas analyzer are at this temperature. In order to determine the appropriate amount of water vapor to post-inject, a series of calculations has been performed. Starting from the state where the secondary injection will take place by employing the ideal gas equation,

PVcool = NT,coolkTcool (3.26)

PVnew = (NT,cool + NH2O,P I )kTcool (3.27)

where NT,cool is the number of particles after cooling the mixture and NH2O,P I is the amount of vapor molecules post-injected at the new volume. Here, the new volume can be calculated in terms of the post-injected water as

Vcool NH2O,P I Vnew = (NT,cool + NH2O,P I ) = Vcool + (3.28) NT,cool nT,cool

Based on this, the amount of water to be post-injected has to be determined in a diﬀerent stage, which is the main injection. However, special care must be taken since the code starts every simulation at V = 1 m3. On this basis, the amount of water calculated for main injection must be transported or changed in volume basis to the after-combustion volume:

Vini nH2O,comb = nH2O,ini (3.29) Vcomb

88 In the same way, the amount of water from the combustion is transported to the

cooling region where post-injection takes place

∗ Vcomb nH2O,P I = nH2O,comb (3.30) Vcool

∗ 3 The Vcomb, which represents the assumed combusted volume, is set to 1 m in the code assuming thus that the total number density on that volume basis occupies 1 m3

3 instead of the combusted volume, Vcomb (larger than 1 m ). The water number density

to be post-injected after cooling is ﬁnally calculated as

∗ ViniVcomb nH2O,P I = nH2O,ini (3.31) VcombVnew

1 NH2O,P I = nH2O,ini (3.32) Vcomb

hence the Vnew is calculated as

nH2O,in Vnew = Vcool + (3.33) nT,coolVcomb

The other species number densities must be recalculated since the volume increases

when water is added. This calculation is done by ni,new = ni,coolVcool/Vnew.

3.2.4 Initial Concentrations when Using Exhaust Gas

Diesel engines are operated under lean conditions (oxygen rich) for the majority of the engine map. Therefore, there is oxygen available in the exhaust which could be used for the fuel reformer in order to decrease the fuel penalty and eliminate the use of external devices (pumps). Based on this, the current section focuses in deriving the chemical compositions for the reactants: oxygen, fuel, carbon dioxide and water

89 to be used to simulate the feeds for the reformer at diﬀerent engine conditions. The

data for the exhaust was obtained for a 1.9L Euro4 Engine at 2000 rpm, illustrated

in Table 3.2. It is noteworthy that the concentrations of CO and other hydrocarbons

are negligible (< 1.0 %). Also, the reformer conditions are: atmospheric pressure,

no addition of water (S/C = 0.0), initial temperature at 1000 K and 1.0 ≤ φ ≤ 1.9.

Torque [Nm] A/F Ratio CO2 Vol% O2 Vol% H2OVol% N2Vol% 10.6 52.096 3.87 15.13 4.11 76.89 40.4 30.538 6.81 11.16 7.24 74.79 60.0 24.462 8.62 8.76 9.16 73.46 250.1 20.952 10.19 6.66 10.83 72.32 326.2 16.669 12.95 2.93 13.76 70.36

Table 3.2. Exhaust Conditions for 1.9 L Euro4 Diesel Engine

For the engine, from the deﬁnition of the air to fuel ratio the mass of fuel injected in the engine can be calculated as

m˙ air,eng m˙ f,eng = (3.34) (A/F )eng

As known, the air burned,m ˙ air,eng,b, in the engine is that combusted at stoichiometric conditions. Therefore

m˙ air,eng,b =m ˙ f,eng(A/F )eng,stoic (3.35)

Since the total amount of air is that burned plus that unburned, the amount of unburned air can be derived as

90 (A/F )stoic m˙ air,eng,ub =m ˙ air,eng − m˙ air,eng,b =m ˙ air,eng 1 − (3.36) (A/F )eng

From the engine, the global chemical reaction is given in Equation (3.37).

gC16H34 + h(O2 + 3.7755N2) −→ iCO2 + jH2O + kN2 (3.37)

The number of moles for the fuel (g) and burned air (h) are known, and from the molar balance i, j and k can be calculated from Table 3.3.

Coeﬀ Equation Coeﬀ Equation m˙ 17m ˙ g = air,eng j = air,eng (A/F )eng Mf (A/F )eng Mf

(A/F )stoic (A/F )stoic h =m ˙ air,eng k =3.7755m ˙ air,eng (A/F )eng (A/F )eng

i =16(S/C)a

Table 3.3. Molar Coeﬃcients for Engine Combustion

The oxygen content of unburned air can be easily derived as

n˙ O2 = yO2 nair,eng,ub (3.38)

m˙ air,eng,ub = yO2 (3.39) Mair m˙ air,eng (A/F )stoic = yO2 1 − (3.40) Mair (A/F )eng

91 where yO2 is that for atmospheric air (0.21), nitrogen can be obtained from the original amount of air. Hence, the total amounts of products coming from the engine are

m˙ air,eng (A/F )stoic n˙ O2 = yO2 1 − (3.41) Mair (A/F )eng

m˙ air,eng n˙ N2 = yN2 (3.42) Mair

16m ˙ air,eng n˙ CO2 = (3.43) (A/F )eng Mf

17m ˙ air,eng n˙ H2O = (3.44) (A/F )eng Mf

In order to account for all possible carbon containing compounds in the system including those coming from the reformer, so that we can calculate the real φ, the reformer analysis allows us to calculate

m˙ air,eng (A/F )ref = (3.45) m˙ f,ref +m ˙ f,eng

m˙ air,eng m˙ air,eng m˙ air,eng m˙ f,ref = − m˙ f,eng = − (3.46) (A/F )ref (A/F )ref (A/F )eng 1 1 =m ˙ air,eng − (3.47) (A/F )ref (A/F )eng

The number of moles of fuel in the reformer is given as

m˙ f,ref n˙ f,ref = (3.48) Mf,ref m˙ (A/F ) = air,eng 1 − ref (3.49) (A/F )ref Mf,ref (A/F )eng m˙ φ 1 = air,eng − (3.50) Mf,ref (A/F )stoic (A/F )eng

92 The total number of moles, nT , entering the reformer is then expressed as

n˙ T =n ˙ O2 +n ˙ N2 +n ˙ CO2 +n ˙ H2O +n ˙ C16H34 (3.51) yO2 (A/F )stoic yN2 16 =m ˙ air,eng 1 − + + + (3.52) Mair (A/F )eng Mair (A/F )eng Mf 17 1 φ 1 +m ˙ air,eng + − (3.53) (A/F )eng Mf Mf,ref (A/F )stoic (A/F )eng

Obtaining then the molar fractions

n˙ i Yi = (3.54) n˙ T where i corresponds to the reactant species. The molar fraction of each component is then multiplied by the total number density obtained at 1 atm, 1000 K and 1 m3 to obtain the number density of each species.

ni = Yi· nT (3.55)

Table 3.4 presents the initial concentrations for φ = 1.3 entering the reformer.

These concentrations change with equivalence ratio.

C16H34 O2 N2 CO2 H2O 0.8415 14.5549 76.7504 3.8076 4.0456 0.6652 10.2987 75.8042 6.4155 6.8165 0.5614 7.7939 75.2473 7.9502 8.4471 0.4753 5.7133 74.7848 9.2251 9.8016 0.3236 2.0513 73.9707 11.4688 12.1856

Table 3.4. Reforming Inlet Concentrations (Vol%) for φ = 1.3

93 3.3 Deﬁnition of Hydrogen Yield

The numerical model calculates the number densities of each species i. The con-

centrations or the volume percentages are calculated by dividing the number density of

that particular species by the sum of the number densities of all species. In this thesis

two diﬀerent forms of presenting the results are given in order to avoid discrepancies

with the technical literature since catalytic reformers generally present results disre-

garding the water vapor and nitrogen content in their ﬁnal concentrations, while results

on non-catalytic reformers include every product. In order to calculate the eﬃcacy of

the reforming process or the real conversion of hexadecane to hydrogen, a new variable

is defined as the hydrogen yield. This definition differs from what other authors refer

to as hydrogen yield or hydrogen concentration. Here, it is deﬁned as the number of

H2 molecules produced per molecule of diesel fuel (or C16H34) consumed, divided by the maximum number of H2 molecules that could be produced from a single diesel (or

C16H34) molecule.

Using C16H34 as a model molecule to represent diesel, the maximum number of H2

molecules that could be produced from one C16H34 molecule is 17. Furthermore, the

hydrogen yield formulation depends on the nature of the process, i.e., whether post-

injection is involved or only main injection of the feeds takes place. The numerical

process followed by the main injection only takes combustion and one cooling leading

to the hydrogen yield deﬁnition expressed in Equation (3.56), where as the process fol-

lowed by post-injection takes combustion-cooling-postinjection-cooling, so four changes

in volume must be taken into account rather than only two. One must remember that

94 for the post-injection process, the change in volume of the ﬁrst cooling process is al-

ready included in the deﬁnition of the Vnew basis and therefore does not need to be

included in the hydrogen yield. Thus, the ﬁnal deﬁnition of hydrogen yield for the

latter process is given by Equation (3.57).

1 NH2 1 Vcomb Vcool nH2,cool ΨH2,MI = = ∗ (3.56) 17 NC16H34 17 Vinitial Vcomb nC16H34,initial

1 NH2 1 Vcomb Vnew VafterP I Vcool,2 nH2,cool,2 ΨH2,P I = = ∗ ∗ ∗ (3.57) 17 NC16H34 17 Vinitial Vcomb Vnew VafterP I nC16H34,initial

∗ ∗ ∗ where Vinitial, Vcomb, Vnew and VafterP I are the initial volumes for each process in the simulation and are equal to 1 m3. This deﬁnition is only based on hexadecane as the

hydrogen provider and is used for comparison for all cases including those where water

vapor is added at any location (i.e., main or secondary injection). However, in case the additional water would be considered as a hydrogen provider, the maximum amount of hydrogen moles is given as

nH2 = 17 nC16H34 + nH2O (3.58)

The ideal amount of steam can be recast on terms of the amount of fuel by the S/C

deﬁnition given in Equation (3.24) as

nH2O + nC16H34 = nC16H34 (1 + S/C· NC ) (3.59)

95 Based upon these deﬁnitions, the hydrogen yield for main and secondary injections is expressed in Equations (3.61) and (3.63), respectively.

(1 + S/C· NC ) NH2 ΨH2,MI = (3.60) (17 + S/C· NC ) NC16H34 + NH2O

(1 + S/C· NC ) Vcomb Vcool nH2,cool = ∗ (3.61) (17 + S/C· NC ) Vinitial Vcomb (nC16H34,initial + nH2O,initial)

(1 + S/C· NC ) NH2 ΨH2,P I = (3.62) (17 + S/C· NC ) NC16H34 + NH2O

(1 + S/C· NC ) Vcomb Vnew VafterP I Vcool,2 nH2,cool,2 = ∗ ∗ ∗ (17 + S/C· NC ) Vinitial Vcomb Vnew VafterP I (nC16H34,initial + nH2O,initial) (3.63)

3.4 Concluding Remarks

In this chapter, the kinetics modeling used in this thesis to simulate the partial oxidation reforming of hexadecane (C16H34) is introduced. Two diﬀerent oxygen providers are considered, pure air (O2 and N2) and diesel engine lean exhaust gas (O2,

CO2,H2 and N2). Based upon this, two sets of initial concentrations are calculated to account for the different species present at this stage. The effect of steam addition at two different locations is also investigated. Hence, a new set of initial concentrations

96 using pure air and water vapor is calculated. Finally, the hydrogen yield used as metric to measure the eﬃcacy of the reforming process in terms of the hydrogen providers

(fuel or fuel+steam) in the following chapters is detailed.

97 CHAPTER 4

CHEMICAL KINETICS OF HYDROGEN PRODUCTION/CONSUMPTION

4.1 Introduction

The problem and methodology employed in this research to simulate the thermal partial oxidation of hexadecane towards hydrogen generation was mentioned in Chap- ter 3. It has to be kept in mind that the addition of water either at the main injection location or through a secondary injection point can be referred to as autothermal reforming based on the deﬁnition of ATR (i.e., the combination of SR and CPOx/TPOx).

Several simulations are performed to study the inﬂuence of the characterizing parameters mentioned in Section 2.3.4 (i.e., equivalence ratio, steam to carbon ratio, and inlet/initial temperature of the feeds) on the ﬁnal product composition. The results obtained for those simulations are discussed in Chapter 4 and Chapter 5.

Chapter 4 oﬀers a very detailed explanation of the chemical species and reaction pathways that take place during the reforming process of hexadecane/air/steam mixtures for the generation and consumption of hydrogen under the inﬂuence of each

98 characteristic property. Identifying these chemical mechanisms allows not only to understand the advantages and disadvantages of each process for each characterizing quantity, but also to make comparisons between these parameters. One case that captures most of the features observed per characteristic property is selected for this analysis, and is set as a baseline for the remaining cases that study the same quantity. Another foreseeable application for the chemical analysis is that this selection can be used to generate a small set of global reactions for model based control of the reforming process.

Chapter 5 presents the ﬁnal compositions (molar fractions) for all the cases analyzed in this research after chemical equilibrium is attained.

This chapter is divided into seven sections. First, Section 4.2 describes the methodology employed to identify the chemical pathways that lead to the production of hydrogen. Then, the analysis performed for each characterizing property is given in Sections

4.3 - 4.7. Finally, Section 4.8 provides a summary and comparison between the cases presented in the previous sections.

4.2 Methodology

Identifying the reactions that produce or consume hydrogen during the reforming process requires close examination of the number density over small periods of time.

These periods of time describe the events of pre-ignition, combustion and cooling that take place in the simulation. Figure 4.1 illustrates the percentage volume of hydrogen and the temperature history for a rich case with no additional water and an initial temperature of 1000 K. Here, it is important to mention that the term “hydrogen” in this thesis refers to molecular hydrogen, H2, and not to atomic hydrogen (or H radical). It

99 can be seen that hydrogen is mostly produced within the small period during combus-

tion. However, due to the high initial temperature at which hexadecane is introduced,

its rapid decomposition into several hydrocarbons and free radicals (H, OH, C16H33, among others) that lead to H2 production within the pre-ignition region is observed.

Furthermore, the cooling portion also reveals diﬀerent species that contribute to the further formation of hydrogen.

Figure 4.1. Hydrogen Concentration and Temperature History for φ = 1.6, S/C = 0.0, P = 1 atm and Tini = 1000 K

This clearly indicates that some reactions are important within a certain period

of time while others do not contribute signiﬁcantly. As time evolves, some of those

initial reactions perish (mostly those involving large hydrocarbons), as the species are

100 consumed and not formed in later times. There are other species, mainly small hydrocarbons, that are formed and consumed at very early stages, produced in large quantities during combustion and generated after combustion due to recombination of atomic H and reaction of small radicals with hydrogen atoms (H atom abstraction).

This study starts from the mathematical deﬁnition in diﬀerential form given by

Equation (3.8) for the temporal variation of number density. That equation reveals two components: the first term refers to the rate of change of the species due to chemical reactions, while the second term is linked to the change in volume. Based upon this, the first term allows to identify and categorize the reactions that have a significant influence on the hydrogen production. The second term of the equation is also very important due to the rapid increase of the volume within the combustion region and its slow decrease in the cooling section.

The hydrogen number density gradient can be approximated using a ﬁrst order

Taylor’s series expansion with the number densities generated in the simulation without losing signiﬁcant information. Hence, a numerical gradient is approximated within two consecutive time steps deﬁned in the original simulation. The same approximation scheme can be performed with the volume and reaction terms without compromising the computing time. Although the general reaction gradient does not provide information on the hydrogen-related reactions, it is used to derive a new time step for the calculation of the individual hydrogen-related reaction gradient. This gradient could be recorded while implementing the algorithm at every time step, however, the computing time and the storage would be excessively increased due to the size of the time step (10−4

- 10−6 s) and the number of reactions that participate. Thus, the reactions term can

101 be calculated using the number densities obtained in the simulation and Equation 3.13

in terms of number density (i.e., divided by the volume).

As the general reaction gradient does not change very rapidly in the pre-combustion

and cooling portions the time step used for the individual reaction gradient can be in-

creased so as to reduce the number of time points to be analyzed and avoid large

computing times and excessive storage. Extreme caution has to be taken when se-

lecting the time step within the combustion region due to the rapid changes for both

terms of Equation (3.8). An optimal time spacing for the individual reactions gradient

is then determined by comparing the sum of all individual gradients and the volume

term, according to Equation (3.8), with the global number density calculated using the

Taylosr’s expansion at each new time step. This optimal time grid is diﬀerent for each case analyzed in this chapter and therefore it has to be determined individually.

Once the gradient for each reaction is calculated, the reactions are sorted by their signiﬁcance for the production or consumption of hydrogen. In order to do this, a new parameter capable of quantitatively categorizing the contribution of each reaction is deﬁned as dnH2 dt %C = i (4.1) i R H2 X dnH2

dt i=1 i where the actual gradient for each reaction is divided by the summation of the abso-

lute contribution of all reactions that involve hydrogen. This percentage captures the

direction of each individual reaction gradient and clearly points out the reactions that

102 aﬀect the production of hydrogen. However, it does not quantify the actual percentage at which hydrogen is produced within the selected period of time, and it must not be confused with the hydrogen percentage volume (mole fraction).

With the aid of this parameter, the reactions can be identified for a given period of time. A set of five reactions for production and five for consumption are selected according to their percentage contribution. These reactions comprise more than 70% of the absolute gradient for each time period. As these sets of reactions vary with time, a sampling time is defined as the time in which a set of reactions behave with a defined trend, i.e., there is no sign change of the gradient of any important reaction or there is no substitution of any reaction by another reaction not already included in the set. If any of these events occurs, the times at which the reaction set started and changed are recorded forming then the sampling time for which the set of reactions is valid.

In the following sections, special cases that capture the inﬂuence of the characterizing parameters are selected for their complete dissection to illustrate the important chemical reactions taking place. The rich case for φ = 1.6 is an appropriate case to analyze for the inﬂuence of the equivalence ratio since it leads to a high hydrogen concentration and hydrogen yield when no additional water is introduced, S/C = 0.0.

Opposite to that, the introduction of water as part of the feeds shows a very slight increase of hydrogen when a low S/C ratio is added along with fuel for some equivalence ratios; however, it negatively inﬂuences the hydrogen (H2) concentration as the steam to carbon ratio (S/C) increases. Based on this, the inﬂuence of water is analyzed for

103 a case that shows a slight increase: φ = 1.6 and S/C = 1.0. Both cases present the initial feed concentrations at 1 atm and 1000 K initial temperature.

The initial temperature inﬂuence is analyzed by using φ = 1.6, S/C = 0.0 and Tini

= 1750 K. This allows the comparison between the reaction sets that take place in the

φ eﬀect case and this case. A fourth case for φ = 1.6 and S/C = 1.0 is selected to

study the inﬂuence of S/C when the feeds are injected at higher temperature (1750

K). Finally, a post-injection case (φ = 1.6 and S/C = 1.0) is employed to illustrate

the possibility of injecting water through secondary injection once the combustion has

taken place. Special attention must be taken in this case since it assumes the water

to be injected at 2000 K. Two diﬀerent post-injection temperatures (1500 & 2000 K)

are analyzed in this research since the location of the secondary injection of water in an

experimental reactor could be allocated right at the ﬂame or slightly downstream of the

ﬂame. Although both cases at steady state show a decreasing behavior, the decreasing

slope for the lower temperature case is larger than that for the high temperature case.

Further details on these cases are given in Chapter 5.

4.3 Eﬀect of Equivalence Ratio (φ)

The case φ = 1.6, S/C = 0.0, P = 1 atm, and Tini = 1000 K is thoroughly analyzed

in this section. As mentioned above, three main sections can be distinguished for all

cases, with only the times being unique for this particular case: 1) before combustion,

t < 0.164 s (about 1000-1300 K); 2) during combustion within t = 0.164-0.1663 s, where

the adiabatic temperature is attained, most of the reactions take place, and hydrogen

is produced in abundance; and 3) the cooling section that starts right after combustion

104 (temperature decreases linearly according to the imposed heat transfer introduced in

Chapter 3). The chemical reactions and the number density gradients for these zones

are explained in the following subsections.

4.3.1 Pre-Combustion Zone

Due to the high initial temperature, 1000 K, hexadecane rapidly decomposes into

several hydrocarbons and radicals (H, OH, HO2,C4H7,C16H33 among others). There-

after, the highly reactive hydrogen atom preferably recombines with small alkenes and

aldehydes to form hydrogen. The hydrogen abstraction reaction between hexadecane

and hydrogen radical, reaction 529 in Appendix A, represents the maximum contribu-

tion towards hydrogen production over t = 0.0 s - 0.12 s. Although its contribution decreases exponentially from above 50% to 8% (t = 0.1597 s), this reaction continues

until no more hexadecane is present because of its oxidation and further decomposition

in the combustion zone.

Another main reaction in this region is that of ethane decomposition (C2H6), re-

action 223. This reaction presents a decreasing contribution from 36% to 15.525%

within the time range 0.008-0.12 s. Thereafter the reaction increases to 36.5% right

before combustion at about t = 0.1639 s. This reaction becomes the biggest producer

of H2 at the time its increasing behavior starts. Conversely, reaction 661, heptene

going to hydrogen and heptanyl radical shows an increasing participation within the

range t = 0.008-0.104 s from 2.5% to 11.723%. Thereafter, its contribution slowly

vanishes (3.92% at combustion).

105 Reaction 136, formaldehyde, rapidly becomes important starting from about 1.95%

to 10.22% at t = 0.163 s, and then slowly decreases to 9.55% right before combus-

tion. This reaction is important for most of the combustion duration. Acetaldehyde,

CH3HCO, through reaction 169 produces hydrogen along with an acetyl radical. This

reaction increases to 13.9% at t = 0.16366 s, and slightly ﬂuctuates ( 1%) before com- ¬ bustion. Finally, propylene reacts with atomic hydrogen producing H2 and propenyl

radical (C3H5) via reaction 376. Its contribution ranges from 1.97-8.9% within 0.008-

0.162021 s, then starts to decay slowly within the combustion region.

The hydrogen consuming reactions comprise a slight 5% of absolute number density

gradient at t = 0.01 s and continue to increase approximately to 9% right before combustion. Reactions 59 and 153, hydrogen with OH radical and hydrogen with methyl are the main reducing reactions. While reaction 153 increases from 0.35% to

5.09% before combustion, reaction 59’s contribution reaches a maximum of 4.45% at t

= 0.16298 and then decreases to 3.22% at the beginning of combustion. This reaction

is the responsible of the water production during the pre-combustion region. These

reactions play a very important role in the combustion duration. Reactions 54, 103

and 28 represent less than 0.5% and are therefore neglected.

Summarizing, the most important hydrogen producing reactions during the pre-

ignition period are 529, 223, 661, 169, 136: these reactions are given in Table 4.1 for

easy access. The consuming reactions are 59 and 153, while 54, 103, 28 are negligible.

These reactions are illustrated in Table 4.2. As can be concluded from the previous

discussion, the reaction term of the number density gradient is strongly dominated by

the producing reactions within the pre-combustion limits. It is also noticeable that the

106 initial hydrocarbon (hexadecane) disappears with little energy release in this section.

However, as time evolves, temperature and volume increase, enhancing therefore the negative inﬂuence of the volume term on the total number density gradient according to Equation 3.8. This can clearly be seen in Figure 4.2, where the total gradient is lower than that for the reaction term.

No. Chemical Reaction 136 H + CH2O −→ H2 + HCO 169 H + CH3HCO −→ H2 + CH3CO 223 C2H6 −→ H2 + C2H4 376 H + C3H6 −→ H2 + NC3H5 529 NC16H34 + H −→ H2 + NC16H33 661 H + NC7H14 −→ H2 + NC7H13

Table 4.1. Hydrogen Producing Reactions in the Pre-Ignition Section

No. Chemical Reaction 59 H2 + OH −→ H + H2O 153 H2 + CH3 −→ H + CH4

Table 4.2. Hydrogen Consuming Reactions in the Pre-Ignition Section

107 Figure 4.2. Hydrogen Number Density Gradients: Total, Reaction Term and Volume Term for φ = 1.6, S/C = 0.0, P = 1 atm and Tini = 1000 K at the Pre- Combustion Region

4.3.2 Combustion Zone

The combustion zone is a very narrow section in the time domain, where the number density and its total gradient suﬀer very rapid changes, as shown in Figure 4.3. Figure

4.4 illustrates the variation of the number density gradient and its components. The total gradient follows the behavior of the reaction term; however, the volume term becomes slightly larger than the reaction term (still large) at about t = 0.16448 s causing the total gradient to become negative and decrease the hydrogen number density as can be seen in Figure 4.3.

108 Figure 4.3. Hydrogen Number Density and its Total Derivative for φ = 1.6, S/C = 0.0, P = 1 atm and Tini = 1000 K

After this minimum, the global gradient increases because of the volume term, which is still positive, decreases to a value smaller than that for the reaction term which is also decreasing (but always positive), see Figure 4.4. The reaction term starts to rise at about t = 0.1646 s presenting an inﬂection point and later a maximum around t

= 0.1654 s. Thereafter, the three gradients decrease which indicate that the number density begins to saturate, as seen in Figure 4.3 and Figure 4.4.

Many reactions such as 148, 451 and 75 to mention a few, appear and disappear suddenly. However, their contribution is low compared to that for other reactions and, therefore, their inﬂuence can be neglected (< 4% each). Some other reactions,

223, 376, 136 and 169, present at the pre-ignition zone transcend to this region with signiﬁcant contribution. Here, the combustion region can be divided into two sections.

109 Figure 4.4. Hydrogen Number Density Gradients: Total, Reaction Term and Volume Term for φ = 1.6, S/C = 0.0, P = 1 atm and Tini = 1000 K at the Com- bustion Region

The first section is highly influenced by these transcendent reactions (t = 0.1639-0.1645 s), while the second part is influenced by reactions 24, 194 and 244.

The first section of the combustion is then analyzed. Reaction 223 is very important at the beginning of the combustion where its contribution increases from 36.5% to 39.3% in only 0.6 ms and then slightly fluctuates 1% until t = 0.164261 s where it ¬ starts to decrease. Its contribution reaches 1.8% after only 0.5 ms and, thereafter its contribution is negligible. Reaction 376 continues to decrease from 7.95% at the beginning of the combustion to 0.16% only 0.4 ms after (t = 0.164441 s). Reaction 136’s percentage contribution presents a fluctuating behavior with the following sequence 10.034-

9.98-13.5-13.4-19% for the time sequence t = 0.164121-0.164181-0.164401-0.164421 s.

110 This ﬂuctuations correspond to the peak-valley-peak-valley-inﬂection points depicted

in Figure 4.4. Reaction 169 starts with a contribution of 13.9% and decreases to 2.26% at t = 0.1643 s.

The hydrogen-related reactions that comprise the second section of the combus-

tion (24, 194 and 244) correspond to 1) water-hydrogen radical producing H2 and OH

radical (elementary process of the water gas shift reaction), 2) methane-hydrogen rad-

ical generating H2 and methyl radical, and 3) ethene (or ethylene)-molecular nitrogen

(liberated during combustion from the air molecule) forming hydrogen, molecular ni-

trogen and ethyne (acetylene), respectively. From the last reaction it can be seen that

nitrogen acts as an inert molecule (third body).

Reaction 24 begins with 2% at combustion at t = 0.16456 s and rapidly increases to 30.96% in 1.9 ms where the cooling section is about to start. Reaction 194 also increases from 3.6% at the lowest value of the total gradient, Figure 4.4. Its contri-

bution increases to 24.18% right at inﬂection point (t = 0.16476 s) of the same ﬁgure.

Although a very small decrease is seen for this reaction at 0.165261 s, it continues

to increase 32.6313% at 0.165261 and starts to decay (right before the maximum of

the second section of the combustion) to 16.71% at 0.166441 s, where the cooling sec-

tion begins. A similar trend is seen for reaction 244, which increases from 1.3072%

to a 22.28% at the third maximum of the ﬁrst section of the combustion, then its

contribution decays to 5.9% at 0.16532 s before the maximum of the second section of

the combustion. Its concentration keeps decreasing to 1.5% at the combustion-cooling

limit. It can be inferred that the decrease of these two last reactions, completely linked

to the increasing consuming reactions, causes the number density to saturate.

111 The consuming reactions play a very important role in the combustion region. Re- action 153 increases from 5.6% at the beginning of the combustion to 21.14% at 0.16444 s (third maximum of the ﬁrst combustion section). Its contribution decreases down to

15.8% at 0.1646 s (right after the global gradient minimum), and increases to 44% right before the cooling zone starts. This is the biggest hydrogen consuming reaction for the second combustion section. Reaction 59 increases to 8.01% at 0.1644 s which is close to the instant of time of the third maximum of the total gradient. The contribution of this reaction fluctuates within 1% and suddenly increases to 10.22% where the first inflection point of the second combustion region takes place (the total gradient is barely positive). After that it decreases to 4.24% right before the cooling zone. Reaction

54 reaches its maximum contribution percentage 6.78% at 0.16444 s (right before the minimum of the global gradient), then it decreases rapidly to a percentage of 1% before cooling. Reactions 103, 28, 27, 291 and 65 appear in this region; however, their contribution is very small.

No. Chemical Reaction 24 H + H2O −→ H2 + OH 136 H + CH2O −→ H2 + HCO 169 H + CH3HCO −→ H2 + CH3CO 194 H + CH4 −→ H2 + CH3 223 C2H6 −→ H2 + C2H4 244 N2 + C2H4 −→ N2 + H2 + C2H2 376 H + C3H6 −→ H2 + NC3H5

Table 4.3. Hydrogen Producing Reactions in the Combustion Section

Table 4.3 and Table 4.4 illustrate the ﬁve reactions that produce and consume the most hydrogen, respectively. As a summary, the combustion region can be divided into

112 No. Chemical Reaction 54 H2 + O −→ H + OH 59 H2 + OH −→ H + H2O 153 H2 + CH3 −→ H + CH4

Table 4.4. Hydrogen Consuming Reactions in the Combustion Section

two sections. The first section is dominated by some reactions that were prominent in the pre-ignition zone (223, 376, 136 and 169), while the second section is driven by new reactions (24, 194 and 244). The consuming reactions (153, 59 and 54) become important in the combustion part. This indeed provokes the reaction term to decrease as combustion evolves. The volume term plays a very important role as well, since it increases rapidly to comparable or even larger values than for the reaction term during the first subregion of the combustion section. This causes the global gradient to become negative. Thereafter, the volume term decreases (still positive but starts to saturate) and stays lower (still decreasing) than the reaction term over the second part of combustion allowing the total gradient to be driven again by the reaction term. The reaction term increases during the first part of the second subregion of the combustion zone to then decrease reaching a value close to zero. The behavior of these two terms causes the hydrogen number density to saturate. It is also noticeable that reaction

59 is the reverse reaction of reaction 24. As deﬁned above, reaction 24 presents a large and fast increase of percentage contribution from 2% (right after the number density minimum) to 30% at the end of the second part of the combustion, while its ¬ counterpart decreases from 10% to 4% within the same subregion. This highlights ¬ the importance of water in the gas mixture and opens the possibility of using water with the feeds for further production of hydrogen.

113 4.3.3 Cooling Zone

This section takes place within the time interval t = 0.1665-0.658 seconds, starting

right after combustion. From ﬁgure 4.3 a small discontinuity in the total gradient can

be observed. This is due to the simulation procedure explained in Chapter 3. There

it was stated that the mixture of air and fuel is ignited and allowed to reach the adi-

abatic ﬂame temperature. Thereafter, the gases are linearly cooled down to 500 K

so as to emulate the experimental process explained in Chapter 6 and to allow com-

parison between the results obtained in the numerical simulation with those obtained

experimentally given in Chapter 7. The numerical code is written to start at a unitary

volume basis (1 m3). Thus, the volume calculated at the adiabatic ﬂame temperature

is reinitialized to a 1 m3 basis when cooling starts. Furthermore, the number densi-

ties calculated at that same location are set as the initial number densities for cooling.

Hence the magnitude of all gradients are drastically shifted down at the maximum

temperature location. Implementation of the post-processing time explained in the

methodology section, Section 4.2, introduces a slightly larger discontinuity/gap seen in

the combustion gradient ﬁgures throughout this chapter.

Figure 4.5 presents the number density gradients as calculated numerically for the

cooling section, i.e., the total or global gradient, the reaction term and the volume term. It is clearly seen that at the start of the cooling region the reaction term de- termines the behavior of the global term. The number density decreases slightly after combustion and this is captured by the reaction gradient within t = 0.1665-0.22 s.

The volume starts to decrease right after combustion which is reﬂected on the volume

term which becomes negative. However, its contribution is very small in that time

114 window. Thereafter, as time evolves the volume term continues to grow negatively, becoming then larger than the reaction term. This term dominates the behavior of the total gradient during most of the cooling time. From Figure 4.5 it can be seen that correcting for the volume step imposed in the cooling section mainly aﬀects the volume term.

Figure 4.5. Hydrogen Number Density Gradients: Total, Reaction Term and Volume Term for φ = 1.6, S/C = 0.0, P = 1 atm and Tini = 1000 K at the Cooling Region

Here it is important to remind the reader about the deﬁnition of number density, number of particles and mole fraction. The number density is the number of atoms or molecules per unit volume. Therefore, a decreasing volume results in an increasing

115 number density for a ﬁxed number or particles. The number of particles can be calculated by multiplying the number density by the corresponding volume, accounting already for the volume step imposed at the combustion-cooling border. The percentage volume of hydrogen is calculated at every time step as the ratio of the hydrogen number density to the sum of the number density of all species (including hydrogen).

As all the number densities are calculated at the same time step and on the same volume basis, no volume adjustment is required. Based upon these deﬁnitions, the increase in number density must not be directly taken as the behavior of the percentage volume or concentration. The percentage volume is fairly constant throughout the cooling section, as shown in Figure 4.1.

Some of the important chemical reactions from combustion are also present in the cooling section, i.e., 24, 194 and 244. Reaction 24 increases from 30.9% to 45.99% within the time window t = 0.166441-0.338241 s. Then, its contribution slowly decays to 0.0214% at the end of the cooling region. This reaction has the largest contribution for most of the cooling process. Reaction 194 decreases to 4.46% by t = 0.22 s and then slowly increases to 18.7% at 0.5582 s. Then, its contribution diminishes to 0.3337% in the following 0.1 s. Reaction 244 exhibits a decreasing behavior at the end of the combustion section (1.5%) and completely decays to 0.0008% about 0.2 s after combustion is ﬁnalized.

Other reactions such as 122, 148, 491, 512 and 694 appear during the cooling at diﬀerent times; however, their contribution at all times is very small compared to the

116 previous contributors. Interesting behavior is observed for reaction 223 which disap-

peared during combustion and reappears at the end of the cooling section. Its contri-

bution starts at about 0.008% at 0.338241 s and exponentially increases to 16.1185%

by 0.558241 and reaches 99.228% at the end of the cooling region. This reaction is

the largest contributor at the end of the cooling section.

The consuming reactions are the same as in the previous subsections. Reaction

153 dominates the beginning of the cooling while reaction 59 takes over after t = 0.26

s. Reaction 153 shows a slow decrease from 44% to 0.38% at the end of the cooling

zone. Reaction 59 experiences further decrease from 4.24% to 1.87% in about 0.9

ms from the combustion-cooling limit. Then, its percentage contribution starts to

increase exponentially as time evolves reaching a maximum contribution of 37.5732%

at t = 0.3982 s, and decreasing thereafter to 0.0054% at the end of the cooling zone.

Reactions 291, 27 and 65 also appear in this zone but once again their concentrations

are smaller than 1% throughout the process.

Summarizing, the producing hydrogen reactions are those seen in the combustion

region: 24, 194 and 244, with the sudden reappearance of reaction 223, while the

consuming hydrogen reactions are 153 and 59. These reactions are given in Table 4.3

and Table 4.4 presented for the combustion subsection.

4.4 Eﬀect of Steam to Carbon Ratio (S/C)

The presence of water in the equivalence ratio suggests that adding steam in the mixture can result in a larger hydrogen production. In order to study the eﬀect of steam as part of the initial reactants, the case φ = 1.6, S/C = 1.0, P = 1 atm, and

117 Tini = 1000 K is thoroughly analyzed in this section. As for the previous case where

the equivalence ratio eﬀect was analyzed, this case also comprises the three noticeable

zones: 1) pre-ignition, t < 0.2138 s (about 1000-1300 K); 2) combustion, within t =

0.2138-0.2272 s; and 3) the cooling section.

Figure 4.6. Hydrogen Concentration and Temperature History for φ = 1.6, S/C = 1.0, P = 1 atm and Tini = 1000 K

The ﬁrst remarks between the S/C and φ cases are that the combustion is delayed

about 0.05 seconds, the adiabatic ﬂame temperature is lower by about 200 K (Tad

2500 K for the φ case), and ﬁnally the H2 concentration is lower for the S/C case as ¬ seen when comparing Figure 4.6 and Figure 4.1. Furthermore, the concentration of

hydrogen and the temperature present a smoother behavior at the highest temperature

location when steam is used as part of the feeds. In the following subsections, the

118 reactions responsible for the hydrogen generation within the three zones mentioned above are explained.

4.4.1 Pre-Combustion Zone

As explained in Section 4.2, hexadecane is decomposed into smaller hydrocarbons and radicals due to the initial temperature. The set of producing hydrogen reactions during the pre-ignition zone is the same as those identiﬁed for the previous case. As for the previous case, reaction 529 is again one of the main contributors in this zone; however, its contribution starts at 46.2% ( 5% lower than for the previous case) and ¬ decreases exponentially to 8.45% at 0.2056 s and disappears at the beginning of the combustion zone. Reaction 223, ethane decomposition, also presents a decreasing behavior from 46.3% at t = 0.01 to 18.27% at 0.1554 s. Thereafter, its contribution rises to 26.37% at t = 0.2138 s right after combustion starts. Contrary to the previous case, reaction 661 shows a decreasing behavior from 12.09% at 0.01 s to 9.4548% at the combustion limit.

Reactions 376 and 136 show an increasing behavior from 1.6% at 0.03 s to about

8.8% at the combustion limit. An increasing trend is also observed for reaction 169, which appears a few milliseconds later contributing with a 5.1925% at 0.1122 s and reaching a 12.96% at 0.2153 s.

Once again, the consuming reactions again conform a very small percentage contribution from about 1% at 0.01 s to 9% right before combustion. Reaction 59 increases as time evolves from about 0.3% to 3.30% within 0.01-0.2036 s and then decreases to

3.07% at the combustion limit. Reaction 153 instead shows an increasing behavior

119 over the entire section from 0.34% to 5.23%. Reactions 54, 103 and 28 represent less than 0.5% and are thus neglected.

It can be concluded that the trends of reactions 529, 223, 136, 169, 59 and 153 agree with those for the equivalence ratio case. Nevertheless, the behavior of reaction

661 is opposite to that of the previous case. The hydrogen producing and consuming reaction sets are given in Table 4.5 and Table 4.6. From the percentage contributions exhibited by these reactions, the number density gradient due to chemical reactions is positive. As for the previous case, the volume term increases as time evolves causing the total number density gradient to decrease as can be seen in Figure 4.7.

Figure 4.7. Hydrogen Number Density Gradients: Total, Reaction Term and Volume Term for φ = 1.6, S/C = 1.0, P = 1 atm and Tini = 1000 K at the Pre- Combustion Region

120 No. Chemical Reaction 136 H + CH2O −→ H2 + HCO 169 H + CH3HCO −→ H2 + CH3CO 223 C2H6 −→ H2 + C2H4 376 H + C3H6 −→ H2 + NC3H5 529 NC16H34 + H −→ H2 + NC16H33 661 H + NC7H14 −→ H2 + NC7H13

Table 4.5. Hydrogen Producing Reactions in the Pre-Ignition Section

No. Chemical Reaction 59 H2 + OH −→ H + H2O 153 H2 + CH3 −→ H + CH4

Table 4.6. Hydrogen Consuming Reactions in the Pre-Ignition Section

4.4.2 Combustion Zone

Although the combustion zone is narrow as shown in Figure 4.8, its duration (12 ms)

is about four times larger than that reported for the equivalence ratio case previously

presented (3.5 ms). This is due to the temperature behavior, taking then longer to

reach its adiabatic value. The combustion is again divided into two sections as done for

the previous case. Figure 4.9 illustrates the variation of the number density gradient

and its components. The ﬁrst part of the combustion occurs in a very small time

window (as for the φ case), but the second part is elongated which indicates that

the slope of the number density increases at a slower rate than that for the φ case.

In agreement with the chemistry of the previous case, the ﬁrst part is dominated by reactions that transcended from the pre-combustion section, while the second part

121 comprises a new set of reactions. From the latter figure it must be noticed that the order of magnitude of the gradients for both cases is the same (except for the first spike); however, the values are about 5-6 times apart. This difference and that reported for the pre-ignition could explain the lower number density value of hydrogen for this case.

Figure 4.8. Hydrogen Number Density and its Total Derivative for φ = 1.6, S/C = 1.0, P = 1 atm and Tini = 1000 K

122 Figure 4.9. Hydrogen Number Density Gradients: Total, Reaction Term and Volume Term for φ = 1.6, S/C = 1.0, P = 1 atm and Tini = 1000 K at the Com- bustion Region

Reviewing then the chemistry for the ﬁrst part of the combustion, it is found that reaction 223 represents the largest contributor for H2 production from the start of combustion to t = 0.219603 s. Its percentage contribution is seen to increase from 26.37% to 41.85% in 4.97 ms and then decrease to 0.898% at t = 0.2272 s (combustion-cooling limit). An intermediate increase from 24% to 25% corresponding to a third spike is observed for this reaction within its decreasing behavior. Reaction 661 percentage contribution continues to decrease from the start of combustion until it reaches a 3% at t = 0.2180 s (at the ﬁrst spike of the reaction’s gradient).

A decreasing behavior is seen for reaction 376 which disappears after 0.2194 seconds. Reaction 136 shows small ﬂuctuations during its increasing behavior reaching

123 a maximum of 18.123% at 0.2197 s, decreasing rapidly to 1.59% after 3 ms. Finally,

reaction 169 increases from 12.96% to 13.18% at 0.2174 s and then decrease with a very

fast rate to 3.2% at 0.2189 s.

For the second part of the combustion, reaction 194 starts oﬀ with a contribution

of 3.45% at t = 0.2180 s and increases to 25% in about 1.8 ms. It is seen to decrease

to 23.8% at 0.2207 s and once again increase to 30.91% over the next 2.2 ms. It de-

creases to 21.06% right at the combustion-cooling limit. Reaction 24 instead increases

through the second combustion region from 2.5405% to 27.52% within the time win-

dow t = 0.219438-0.227201 s. Finally, reaction 244 increases from 4.64% at 0.2197 s

to 21.10% at 0.2207 s, and then decreases to 1.7734% after only 3.5 ms. Reactions

75 and 148, among others, appear and disappear within very small time windows and

their contributions are not that signiﬁcant compared with that for the main reactions.

On the H2 consuming side, reaction 59 presents a ﬂuctuating behavior that starts with 3.08% at the beginning of combustion and then reaches a maximum of 9.44% at

0.2197 s. Thereafter its contribution decreases with a similar behavior to 4.2% at the combustion-cooling limit. Reaction 153, instead, presents a large increase form 5.23% at the pre-ignition/combustion limit to 23.86% at t = 0.2196 s, decreasing then to

15.1% 0.3 ms later, to ﬁnally increase to 42.44% at 0.2272 s. Reaction 54 starts with a 0.5% and increases to 5.35% at 0.219 s and then decreases to 0.32% 6 ms after. The percentage contribution of Reaction 103 decreases from 3.7% to 0.91% within the time range 0.2202-0.2217 s. Reactions 28 and 291 also appear during combustion, however, their contributions are negligible.

124 As a summary, the set of reactions describing the combustion section agrees with

that for the equivalence case. The reactions in the ﬁrst subregion behave in a similar

manner to those for the previous case with the addition of reaction 661, while the

second part presents an increased time window. The consuming reactions also behave

as those reported for the previous case. Nevertheless, reactions 103 and 54 are seen

to have a slightly larger percentage contribution. It is noteworthy that reaction 24

presents a similar behavior for this case as for the equivalence ratio, however, with a

lower percentage contribution ( 3%). Its counterpart also presents a slightly decrease ¬ in its contribution 0.5%. The decrease for the depleting reaction is smaller than ¬ the decrease for the production reaction, which suggests that steam has a slightly lower global contribution towards hydrogen. Table 4.7 and Table 4.8 illustrate the

ﬁve reactions that produce and consume the most hydrogen, respectively. Finally,

the discontinuity at the combustion-cooling section is again seen for this case. As

explained in Section 4.3.3, this discontinuity is due to the computational volume change

from the expanded volume at combustion to the unitary volume imposed in the code

for the new section.

Table 4.7. Hydrogen Producing Reactions in the Combustion Section

125 No. Chemical Reaction 54 H2 + O −→ H + OH 59 H2 + OH −→ H + H2O 103 H2 + CH2 −→ H + CH3 153 H2 + CH3 −→ H + CH4

Table 4.8. Hydrogen Consuming Reactions in the Combustion Section

4.4.3 Cooling Zone

This section comprises the reactions taken place within the time interval t = 0.2272-

0.727 seconds. Figure 4.10 illustrates the number density gradients for this region, i.e., the total or global, the reaction term and the volume term. As for the previous case, the behavior of the global term is directly linked to that of the reaction term for the start of the cooling section. After t = 0.3 s, the behavior of the global gradient is then driven by the volume term.

The producing reactions that take place in this region are 24, 194, 244 and 223.

Reaction 24 increases from the combustion limit reaching a maximum 38.5% at t =

0.29 s. Thereafter, the contribution decreases with an increasing rate until it completely vanishes at the end of the cooling section. Reaction 244 continues to decrease from the combustion limit percentage contribution until it disappears at t = 0.457 s.

Reaction 194 shows a decreasing behavior from 21% at the combustion-cooling limit to 11.53% at t = 0.28 s, then it reaches a maximum of 27.94% after 0.3 s. It ﬁnally decreases to 0.12% at the end of the cooling region. Reaction 223 disappears at 0.417 s to then reappear at 0.477 s with a negligible percentage contribution. Thereafter, its

126 contribution increases exponentially to 99.6% at the end of the section. Other reactions, such as, 122, 148, 491 and 694 again appear during the cooling section; however, their contribution is negligible compared with the large contributors.

Figure 4.10. Hydrogen Number Density Gradients: Total, Reaction Term and Volume Term for φ = 1.6, S/C = 1.0, P = 1 atm and Tini = 1000 K at the Cooling Region

The H2 consuming reactions are the same as in the previous subsections. Reaction

153 is seen to decrease from its ﬁnal value at combustion to 18.5% after only 0.23 s.

Then, its percentage contribution increases to 27.42% 0.12 s later and ﬁnally decreases to 0.2% at the end of the cooling section. Reaction 59 instead increases to 29.5% after

0.21 s after the combustion is complete. Thereafter, its contribution disappears at the end of the cooling section. Finally, reactions 291, 27 and 65% also appear in this zone but once again their concentrations are smaller than 1% throughout the process.

127 Summarizing, the producing H2 reactions are once again those seen in the combus-

tion region: 24, 194 and 244, with the sudden reappearance of reaction 223, while the

consuming reactions are 153 and 59. These reactions are given in Table 4.7 and Table

4.8 presented for the combustion region in the previous subsection.

4.5 Eﬀect of Initial Temperature (Tini)

As mentioned in Chapter 3, the initial temperature of the feed plays a very impor-

tant role in determining the ﬁnal product composition, since the combustion is almost

instantaneous and hexadecane is decomposed into smaller hydrocarbons and radicals

with a much faster rate. Thus, the eﬀect of the initial temperature is analyzed for a

rich fuel case with no additional water. This case presents the following characteriz-

ing variables φ = 1.6, S/C = 0.0, P = 1 atm, and Tinitial = 1750 K. The evolution of

temperature and hydrogen concentration are depicted in Figure 4.11.

As mentioned before, the combustion is immediate, therefore no pre-ignition zone exists. Within the combustion duration, hydrogen presents a peak value once the temperature is close to the adiabatic ﬂame temperature, then decreases slightly ( 0.8 ¬ in % Vol] and after a few milliseconds slowly increases to its ﬁnal composition. This behavior is very similar to that presented by the equivalence ratio case cited above, although the peak duration is much smaller for this case. The cooling section extends for approximately 0.65 seconds. A more detailed explanation on these zones is given in the following subsections. Another remark from this case is that the hydrogen concentration, 6% Vol, is larger than that for the previous two cases where the H2 concentration falls to 4% Vol.

128 Figure 4.11. Hydrogen Concentration and Temperature History for φ = 1.6, S/C = 1.0, P = 1 atm and Tini = 1750 K

4.5.1 Combustion Zone

The combustion zone as deﬁned for the previous cases captures the rapid tempera-

ture and hydrogen concentration increase. For this case, however, the largest changes

for the number density gradient occur within 0.4 ms where the temperature maximum

has not yet been reached as shown in Figure 4.12. The maximum value is seen to occur

about t = 0.02 s. This can be explained by the fact that some small hydrocarbons are still undergoing oxidation. For consistency, the deﬁnition of the combustion duration is held as before, i.e., where the temperature reaches its maximum. Nevertheless, the drastic change in duration is about 0.25 ms, which is similar to that for the equivalence ratio at lower initial temperature case. The combustion is divided into two regions:

129 1) where the rapid increases occur over t = 0 - 1 × 10−5 s as shown in Figure 4.13, and 2) where the hydrogen concentration presents the increasing-decreasing behavior at about 1.2 x 10−4 s, as shown in Figure 4.14.

Figure 4.12. Hydrogen Concentration and Temperature History for φ = 1.6, S/C = 1.0, P = 1 atm and Tini = 1750 K

130 Figure 4.13. Hydrogen Number Density and its Total Derivative for φ = 1.6, S/C = 0.0, P = 1 atm and Tini = 1750 K

Figure 4.14. Hydrogen Number Density and its Total Derivative for φ = 1.6, S/C = 0.0, P = 1 atm and Tini = 1750 K

131 Figure 4.13 depicts the rapid increase of hydrogen number density that takes place in only the first microsecond. Thereafter the number density increases at a slower rate. This is captured by its derivative also plotted in the same figure. The gradient increases rapidly ( 3.2 × 1028) during the first microsecond and then decreases ¬ one order of magnitude in the following microsecond. Thereafter, the number density

decreases one more order of magnitude (5 × 1026) within 8 µs. For the period t =

0.1 × 10−4 - 2.5 × 10−4 s shown in Figure 4.14, the number density shows a slight

maximum at about 0.15 × 10−4 s and then continues to increase reaching its second

largest maximum at about t = 1.2 × 10−4 s. The gradient is seen to become slightly

negative after the ﬁrst maximum of the number density within this period of time, and

then largely increases to about 7.5 × 1026 before the second maximum. It becomes

strongly negative ( -4 × 1026) right after the second number density maximum, and ¬ slowly rises (although still negative) during the following 0.1 ms.

Figures 4.15 and 4.16 illustrate the behavior of the three gradients: total, reaction

term and volume term, over the same period of time given in the previous explanation.

It can be seen that the total gradient is strongly inﬂuenced by the reaction term over

the ﬁrst 5 µs. From there, the volume term and the reaction term start to level oﬀ

driving the total gradient to zero at about 11 microseconds as shown in Figure 4.16.

That time corresponds to the ﬁrst number density maximum in the second part of

the combustion shown in Figure 4.14. Then, both terms, volume and reaction, are

seen to increase over 80 µs, although the reaction term rises to a larger value than the

volume term. Furthermore, the reaction term is seen to saturate. At the point that

the volume term reaches its maximum, and both, reaction and volume, terms start

to decrease. However, the volume term decreases at a slower rate. This provokes

132 both terms to intersect, which represents the second maximum of the number density.

The reaction term becomes strongly negative after this maximum. This along with the positive value of the volume term drive the total gradient to become even more negative than the reaction term according to Equation (3.8). As the reaction term starts to increase and the volume term to decrease to zero, the global gradient increases as well dominated by the reaction term. The volume term and the reaction term are seen to achieve the zero value at about t = 0.29 ms. Thereafter, the volume term is negative and the reaction term is slightly positive increasing thus the total gradient.

Figure 4.15. Hydrogen Number Density Gradients: Total, Reaction Term and Volume Term for φ = 1.6, S/C = 0.0, P = 1 atm and Tini = 1750 K

133 Figure 4.16. Hydrogen Number Density Gradients: Total, Reaction Term and Volume Term for φ = 1.6, S/C = 0.0, P = 1 atm and Tini = 1750 K

The reaction term is then analyzed by finding the percentage contribution of each reaction over the combustion duration. Hexadecane is mainly decomposed to hydrocarbons and radicals, and no significant contribution towards hydrogen is seen from reaction 529. Reaction 223 instead is the main contributor during the first 16 microseconds. At t = 1 µs its percentage contribution is above 50% and increases to

72.4% at t = 2 µs. Thereafter, it decreases exponentially to 0.31% at t = 0.22 ms experiencing a small increase before that time, however, too small to be considered.

Reaction 232 provides a 9.48% at the ﬁrst microsecond and then decreases to 1.8% after 3 µs, to then disappear from the main producing reactions. Reaction 376 starts with a 6.97% at the ﬁrst microsecond and then decreases to 2.54% after 6 µs, while reaction 169 decreases from 5.88% to 1.47% within t = 1-3 µs. Reaction 148 presents

134 an increasing behavior from 2.33% to 4.54% from t = 2-11 µs. Then starts to decrease

to 2.8% 6 µs later. At t = 83 µs it reappears as part of the main contributors with an increasing percentage contributions 1.3% to 4.7% at t = 0.115 ms. Thereafter, its contribution decreases to nearly zero at t = 11.95 ms.

Continuing with the H2 producing reactions, 136 presents a ﬂuctuating behavior

increasing from 1.5% to 11.2% within t = 3-16 µs, then decreasing to 7.73% after 12

µs, ﬂuctuate ±0.5% during 68 µs, to ﬁnally decay to 0.52% after 69 µs (t = 0.16 ¬ ms). Reaction 194’s contribution increases from 1.94% to 14.45% within t = 4-17 µs.

Then, its contribution diminishes to 11.25% during 9 µs and again rises to 27.13% after

79 µs. Its percentage is seen to decrease to 0.842% at the combustion-cooling limit (t

= 0.0199 s). This reaction is dominant from t = 90-135 µs. Reaction 75 presents an increasing behavior from 2.4% to 4.55% during the time window t = 7-23 µs, then it deploys to 1.3% after 31 µs.

The percentage contribution for reaction 244 increases from 2.5% to 47.6% during t = 12-48 µs and then slowly decreases to a negligible value at t = 19.94 ms. This is the main reaction over t = 19-90 µs. Reaction 24 increases from 1.3% to 49.95% within t = 53 µs - 19.94 ms. It is the most important reaction over the period t =

0.135 - 19.94 ms. Reactions 6, 55, 144, 246 and 491 among others appear as part of the main reactions for very small periods but their contribution is less than 1%.

On the H2 depletion side, reaction 153 increases from 1.75% to 16.8% for the ﬁrst

19 µs, then decreases to 7.96% during the next 13 µs and increases to 37.225% over

168 µs (t = 0.2 ms). Thereafter, its contribution decreases slowly to 3.33% at the combustion-cooling limit (t = 0.01994s). This is the main depleting reaction over the

135 first 24 µs and again from t = 40 µs - 1 ms. Reaction 59 has a fluctuating behavior from the beginning of the combustion starting at 0.13% to 11.13% at t = 8 µs. Then its contribution depletes to 1.2% at t = 0.1 ms and then rises to 40.2% at the combustion- cooling limit. This is the main consuming reaction from t = 24-40 µs and also over t = 1.0-19.94 ms. Reaction 54 and 103 also present a fluctuating behavior; R-54:

0.5883-6.4-6.1-6.9-5-7.4-1.3% during the time interval 0-0.29 ms, while R-103: 0.49-

1.8-1-2.4-4.4-0.08 throughout the combustion duration. Reaction 291 reaches 10.87% during the ﬁrst 16 µs and then decreases to 1.28% after 0.11 ms. Reaction 65 appears after 0.1 ms with a percentage contribution of 1.2% and then increases to 3.7% at t

= 0.165 ms. Its contribution decreases after 0.2ms to 1.1%. Reaction 27 contributes with 1.4% to 6.3% from t = 0.29 ms to the combustion-cooling limit.

Summarizing, the H2 producing reactions are 223, 232, 376, 169, 136, 148, 194,

75, 244 and 24. Opposite to the case with similar characterizing quantities but lower temperature, equivalence ratio case, reactions 529 and 661 do not appear as part of the main reaction mechanism for the generation of hydrogen which suggest their rates are slower than those for these reactions. It could also be due to their fast depletion in the very early stages of the combustion. Furthermore, more reactions are seen to have a much larger contribution for this case than for the equivalence case. Table 4.9 comprises these producing reactions.

The H2 depleting reactions are not only 153 and 59, as for the equivalence case, but also reactions 54, 103, 291, 65 and 27, and their contribution is more signiﬁcant than that for the previous case. Although the depleting reactions play a very important role in several stages of the combustion section, the main production of hydrogen occurs

136 during the ﬁrst microseconds. Thereafter, the production presents a slower ﬂuctuating increasing rate. Table 4.10 illustrates the most important consuming reactions for this combustion section.

No. Chemical Reaction 24 H + H2O −→ H2 + OH 75 CHOCHO −→ H2 + 2CO 136 H + CH2O −→ H2 + HCO 148 2CH3 −→ H2 + C2H4 169 H + CH3HCO −→ H2 + CH3CO 194 H + CH4 −→ H2 + CH3 223 C2H6 −→ H2 + C2H4 232 H + C2H5 −→ H2 + C2H4 244 N2 + C2H4 −→ N2 + H2 + C2H2 376 H + C3H6 −→ H2 + NC3H5

Table 4.9. Hydrogen Producing Reactions in the Ignition Section

No. Chemical Reaction 27 N2 + H2 −→ 2H + N2 54 H2 + O −→ H + OH 59 H2 + OH −→ H + H2O 65 H2 + CH −→ H + CH2 103 H2 + CH2 −→ H + CH3 153 H2 + CH3 −→ H + CH4 291 H2 + C2H −→ H + C2H2

Table 4.10. Hydrogen Consuming Reactions in the Ignition Section

137 4.5.2 Cooling Zone

The cooling section for this case behaves in a similar manner to that described for

the previous cases as can be seen in Figure 4.17. The volume term increases as the volume decreases after combustion. This positively impacts the total gradient for the hydrogen number density. The number of molecules can be calculated by multiplying the number density by the appropriate volume that is already readjusted to account for the numerical volume imposition between combustion and cooling.

Reaction 24 is the main contributor during this section starting from 49.9% at the combustion-cooling limit increasing to 71.5% at the end of the cooling process.

Reactions 136, 194, 55, 145, 122, 232, 512, 254, 244 and 144 also contribute during this section; however, their percentage contribution is less than 2%.

On the depleting side, reaction 59 (the reverse reaction to 24) represents the main consuming reaction. Its contribution starts at 40% at the combustion limit, increases to 48% at t = 0.384 s, and then slowly decreases to 20.5% at the end of the cooling section. Reaction 54 shows a small increase from 6.28% to 6.3% during the ﬁrst 7 ms of cooling, then decreases to a negligible value by t = 0.6 s. Other reactions, such as

153, 54, 103, 291 and 65 appear during the cooling section; however, their contributions

are negligible.

138 Figure 4.17. Hydrogen Number Density Gradients: Total, Reaction Term and Volume Term for φ = 1.6, S/C = 0.0, P = 1 atm and Tini = 1750 K

4.6 Eﬀect of Initial Temperature and S/C

In order to study the diﬀerence between the addition of water and the previous case, a case with the following characterizing quantities is employed: φ = 1.6, S/C = 1.0,

P = 1 atm, and Tinitial = 1750 K. The global behavior of the hydrogen concentration and temperature depicted in Figure 4.18 resembles that for the previous case. Four main diﬀerences can be distinguished from comparing it with Figure 4.11: 1) the temperature, as expected, is lower for this case; 2) the hydrogen concentration also shows changes, see Figure 4.19, observed for the previous case for the combustion duration, however, the slopes are much smaller; 3) the combustion section is slightly

139 larger than that for the previous case; and ﬁnally, 4) the hydrogen percentage volume is larger for this case 6.8% Vol.

Figure 4.18. Hydrogen Concentration and Temperature History for φ = 1.6, S/C = 1.0, P = 1 atm and Tini = 1000 K

For the combustion region, the hydrogen number density for the high-temperature- no-water case exhibits not only larger maximum, but also larger drops (larger decreasing gradients), see Figure 4.12. Thus at the end of the combustion region, the number density decreases to 11 × 1022 (from 14 × 1022) for the previous case, while it only decreases to 12 × 1022 (from 12.8 × 1022). The combustion region is bounded at 0.0-

1.069 ms. The cooling section starts right after. Further details on the combustion and cooling regions are given in the following subsections.

140 Figure 4.19. Hydrogen Concentration and Temperature History for φ = 1.6, S/C = 1.0, P = 1 atm and Tini = 1000 K

4.6.1 Combustion Zone

The combustion zone, as for the previous case, is divided into two subregions as shown in Figures 4.20 and 4.21. The ﬁrst 10 microseconds deﬁne the largest maximum occurring at 1 µs, its rapid decrease over the second microsecond, and its continuing decreasing value. The maximum registered for this case is slightly lower than that for the previous case, while the remaining of this subregion achieves similar values.

The limits for the second subregion are 10 µs to 0.5 ms. This region exhibits similar behavior compared to the corresponding subregion for the previous case. The ﬁrst negative value is very similar for both cases. The maximum is much larger for the non-water case (7.5 × 1026) than that for the water case (3 × 1026). Here the number

141 density attained 14 and 12.8 × 1022, respectively. Thereafter, the gradients decrease; however, the negative value for the non-water case is much larger than that for the water case, -3.5 and -0.5 × 1026, respectively. This decrease clearly marks the diﬀerence in percentage volume seen for these cases. At the end-limit of the combustion, the gradients are 0.4 × 1026 for the water case and 0.85 × 1026 for the non-water case.

Figures 4.22 and 4.23 provide details on the gradients for the water case during the two combustion subsections. The trends deﬁned for the previous case apply to this case; however, the duration of this second combustion subregion is larger than that for the non-water case. No further explanation is needed for these plots.

Figure 4.20. Hydrogen Number Density and its Total Derivative for φ = 1.6, S/C = 1.0, P = 1 atm and Tini = 1750 K

142 Figure 4.21. Hydrogen Number Density and its Total Derivative for φ = 1.6, S/C = 1.0, P = 1 atm and Tini = 1750 K

Figure 4.22. Hydrogen Number Density Gradients: Total, Reaction Term and Volume Term for φ = 1.6, S/C = 1.0, P = 1 atm and Tini = 1750 K

143 Figure 4.23. Hydrogen Number Density Gradients: Total, Reaction Term and Volume Term for φ = 1.6, S/C = 1.0, P = 1 atm and Tini = 1750 K

The reactions taking place over the combustion interval are described as follows.

Reaction 223 is again the maximum contributor for the ﬁrst 3 µs, increasing from

51.55% to 78.8%. Thereafter, its percentage contribution decreases to 0.2051% at t =

0.78 ms. Reaction 232 presents a decreasing behavior, as for the previous case, from

9.4% to 0.9% after 5 µs. Reaction 376 also shows a decreasing pattern as well, from

6.6% to 1.26% within only 4 µs, while reaction 169 goes from 5% to 0.68% within the

ﬁrst 3 µs. Reaction 75 shows an increasing path from 1% to 3.84% at t = 3 µs and 14

µs, respectively. It slightly decreases but its increasing path continues to 6.23% at t =

36 µs. It ﬁnally decreases to 0.5% over 0.137 ms after its maximum. An oscillating behavior is seen for reaction 148 as follows: 2.3-4.15-3.6-1-3.2-0.36% during the time interval t = 2 µs - 1 ms.

144 Reaction 24 also exhibits a ﬂuctuating behavior during the ﬁrst 50 µs of the combustion from 3.27-4.44-2.63%. Thereafter, its percentage contribution largely increases to

47.5% right before cooling (t = 1.069 ms). Reaction 194 appear later than reaction 24 but with a faster increasing rate. Its contribution also ﬂuctuates 0.97-10.5-5.9-24.75-

2.13% for the following times 5-24-40-243-1069 µs. Reaction 136 varies as follows

1.4-8.7-4.3-5.3-6.4-1% for the times t = 7-23-41-98-215-351 µs, and then disappears.

Finally, reaction 244 starts with a very low contribution at t = 18 µs and rapidly in-

creases to 57.2% over a very small time interval of 64 µs (t = 82 µs). Thereafter, its

percentage contribution decreases to 5.1% at the end-combustion limit.

On the H2 consuming side, reaction 153 presents an oscillating behavior from 1.8-18-

19.2-6.24-39.6-25.5% over t = 1-18-28-50-461-1338 µs, right before cooling. Reaction

54 contribution increases from 0.6% to 6.3 over the ﬁrst 15 µs. Then, it decreases

through a ﬂuctuating path to 0.8% at t = 0.62 ms. Reaction 103 has a very small

contribution from 0.5% to a maximum of 3.5% and then decreases to 0.33% at t =

1.338 ms. An important contribution is seen from reaction 291 from 0.05% to 14.6%

within the ﬁrst 23 µs. Then decreases to 1.5% at t = 74 µs and starts to rise to 4.85%

at t = 0.24 ms. Thereafter, its contribution decreases even at the cooling section.

The other main reaction 59 has a very small contribution over the ﬁrst 17 µs then it

increases to 10.5% over the next 22 µs. Its contribution decreases to 1.5% at t = 0.253

ms and then starts to increase to 21.25% at the combustion-cooling limit. Reaction

65 and 27 also participate in this section; however, their contributions are smaller than

2% over the entire combustion region.

145 Hence, the most dominant reactions governing the reaction term during the following time intervals are on the producing side: 1) R-223 t = 0.0 - 27 µs, 2) R-244 t = 27 µs

- 0.215 ms, 3) R-194 t = 0.215 - 0.291 ms, and 5) thereafter R-24; on the consuming side: 1) R-153 t = 0 - 38 µs and t = 70 µs - 1.338 ms, and 2) R-59 t = 38 - 70 µ and t = 1.338 ms - 0.69 s. All these reactions are given in Table 4.9 and Table 4.10 of the previous section.

4.6.2 Cooling Zone

The cooling zone behaves as for the other cases, where the volume term dominates over the entire region as shown in Figure 4.24. Thus, the reactions that generate hydrogen and their contribution are as follows. Reaction 223, 0% at t = 0.493 s and increases to 8% throughout the cooling process. Reaction 24 is the most important reaction from t = 1.338 ms to throughout the entire process reaching a maximum 54.8% at t = 8.95 ms. Its percentage contribution decreases to 29.7% by the end of cooling.

Reaction 194 presents an increasing patter from 2.13% at the combustion-cooling limit to its ﬁnal contribution of 23.3%. Reactions 491, 244, 144 512 and 55 also appear in this region; however, their contribution is negligible.

The reducing reactions are: 153 with an decreasing behavior from 25.5% at the combustion-cooling limit to 0.67% at t = 0.4528 s. Thereafter, its contribution increases to its ﬁnal value 26.72%. Reaction 59 is the most important reaction over the entire cooling with an increase from 21.2% at the combustion-cooling limit to 47.6% at t = 0.393 s. Then, it starts to decay until its ﬁnal contribution 11.76%. Reactions

103, 54, 291, 65 and 17 also appear in this section but their percentage contributions are negligible as for the previous case.

146 Figure 4.24. Hydrogen Number Density Gradients: Total, Reaction Term and Volume Term for φ = 1.6, S/C = 1.0, P = 1 atm and Tini = 1750 K at the Cooling Region

4.7 Eﬀect of S/C Post-Injection at TPI = 2000 K

For the cases in Sections 4.4 and 4.6, the water addition as part of the main feeds

has been analyzed. However, the water injected at a secondary location is also stud-

ied in this thesis and it is worth analyzing a speciﬁc. The post-injection inﬂuence is

barely seen in the hydrogen concentration plots at equilibrium (steady state), shown in

Chapter 5, since the hydrogen concentration (% Vol) is seen to decrease with increas-

ing S/C for most of the cases. Equivalence ratio φ = 1.6 at equilibrium illustrates a slight increase for the post-injection at high temperature case, which agrees with that observed for the S/C case introduced in Section 4.4. Furthermore, diﬀerences between

the behavior of the ﬁnal hydrogen composition of the two post-injection temperatures

147 is observed, as well as diﬀerences between these two cases and that of water as part of the main injection. Based upon this, the post-injection case selected to be scrutinized in this section is φ = 1.6, S/C = 1.0, P = 1 atm, and Tinitial = 1000 K and PI at T =

2000K. This case displays the same characterizing quantities as the case in Section 4.4 allowing further comparison between the chemical mechanisms and their contributions.

Figure 4.25. Hydrogen Concentration and Temperature History for φ = 1.6, S/C = 1.0, P = 1 atm and Tini = 1000 K

Figure 4.25 outlines the temperature and hydrogen mole fraction history for the post-injection case. The post-injection case was divided into four diﬀerent stages, as follows: 1) Pre-ignition of the fuel/air mixture; 2) combustion of the feeds (rapid increase in temperature and hydrogen percentage volume); 3) rapid cooling to the desired temperature (1500 or 2000 K); 4) post-injection over 0.4 seconds; and, ﬁnally

148 5) cooling to 500 K over 0.1 seconds. Stages 1 and 2 are those already described in

Section 4.3.1 and 4.3.2 for the S/C eﬀect, therefore they are included in this section.

Stages 3, 4 and 5 are included in this section. A ﬁrst remark from this ﬁgure is that the percentage volume is slightly higher than that for the second case.

4.7.1 Rapid Cooling

In this section, the combusted mixture undergoes a rapid cooling over a very narrow time window (0.1 ms). Once more, the combusted volume is transformed at the beginning of this section to the unitary numerical volume explained in previous sections, which leads to discontinuities in the number density gradients. From Figure 4.4 it can be seen that the volume term is almost zero. The reaction term is therefore the driving term for the total gradient.

The reaction term is then analyzed to identify the reactions that contribute the most either for consuming or producing hydrogen. On the producing side, the contribution of reaction 24 decreases from 30.96% at the combustion limit to 28% at the end of this cooling section. Reaction 194 increases from 16.71% to 18.5%, while reaction 244 continues to decrease from 1.82% within the same time window. Reactions 223, 148 and 491 also appear in this section; however, their contributions are negligible. On the consuming side, reaction 153 is seen to increase from 44% at combustion limit to 46.5% over 1µs and then decrease to 37.6% at the limit of the ﬁrst cooling. The contribution of Reaction 59 rises from 4.24% to 11.5% within the same time period. Reactions 291,

103 and 27 also consume hydrogen; however, their contributions are very small to be considered.

149 4.7.2 Post-Injection Zone

Figure 4.26 displays the total gradient and its components during part of the post- injection period. It can be seen that the volume term is still very small making its inﬂuence negligible compared with that from the reaction term.

Figure 4.26. Hydrogen Number Density Gradients: Total, Reaction Term and Volume Term for φ = 1.6, S/C = 1.0, P = 1 atm, Tini = 1000 K and TPI = 2000 K at the Post-Injection Region

The reactions that produce hydrogen during this region are mainly 24 and 194.

Of these, reaction 24 is the most important contributor. Its percentage contribution starts increasing from 28% to 45% right after the combustion-cooling limit ( 1 µs), ¬ then decreases in an oscillating manner to its ﬁnal contribution of 38.4% at the end of the psrt-injection section. Secondly, reaction 194 shows a global decreasing behavior

150 from 18% to 12% within the same time period. Reactions 244, 148, 223, 491 and

30 also appear prominent but their contributions never reach more than 2.5%. The reducing reaction set is mainly constituted by 153 and 59. Reaction 153 continues to be the largest consumer during the entire period. Its contribution has a decreasing global behavior that ﬂuctuates as 37.6-28-35-28-31-18% within t = 0.166541-0.166542-

0.166543-0.166575-0.24784-0.600611 s. Reaction 59 shows an increasing behavior from

11.5% to 29.6%. All these reactions are already listed in the tables for the previous cases and therefore are not presented in this section.

4.7.3 Cooling Zone

As for the previous cases, the total number density gradient is dominated by the volume term which increases as time evolves in this section, see Figure 4.27. The reactions that contribute to the reaction term are: 24, 194 and 223 on the producing side, and reactions 153 and 59 on the depleting side. The contribution of reaction

24 contribution decreases rapidly in this region from 38.4% to 0%, while reaction 194 shows a fast increase from 12% to 24% within 0.06 s and then decreases to nearly zero at the end of this section. Reaction 223, as did in the other cases, is the most dominant reaction very close to the end of this region, increasing its contribution from nearly zero to 99.7%. Reactions 244, 148, 232, 694, 122, 136, 418 and 313 appear in this section but their contributions are very small. The consuming reactions: reaction

153 decreases from 18.5% to 0.12% at the limit of the second cooling, as does reaction

59 from 29.6% to 0% within the same period of time.

Comparison between all the previous cases related to this case are given in the following section, Section 4.8.

151 Figure 4.27. Hydrogen Number Density Gradients: Total, Reaction Term and Volume Term for φ = 1.6, S/C = 1.0, P = 1 atm, Tini = 1000 K and TPI = 2000 K at the Cooling Region

4.8 Concluding Remarks and Comparisons

This section compares the cases presented in this chapter by highlighting the main diﬀerences and emphasizing the importance of some reactions over a certain period of time. The equivalence ratio case is compared against the steam to carbon ratio case at the same initial temperature, and then against the equivalent case at higher temperature. The steam to carbon ratio case at low temperature is compared to the equivalent case at higher temperature. The high temperature equivalence case is compared to the high temperature S/C case. The post-injection case is compared with the low temperature S/C case.

152 4.8.1 Equivalence Ratio (φ = 1.6) and S/C = 1.0 Ratio Cases at Low Tini

In this subsection, the equivalence ratio and steam to carbon ratio cases at the same initial temperature, 1000 K, are compared. The general diﬀerences are that the equivalence ratio case provides a slightly larger hydrogen concentration, the adiabatic ﬂame temperature is lower for the S/C case by about 200 K, and the combustion is delayed about 0.05 s for the S/C case. Further comparison of the behavior of the chemical reaction sets for both cases is carried out for the three distinguished sections: pre-ignition, combustion and cooling regions, and it is presented in the following subsubsections.

4.8.1.1 Pre-Combustion

1. Both cases present a total number density gradient strongly dominated by the

reaction term in Equation (3.8).

2. The trends of H2 producing reactions 529, 223, 136 and 169, and H2 depleting

reactions 59 and 153, are the same as those for the φ case.

Table 4.11. H2 Producing Reactions in the Pre-Ignition Section for φ = 1.6 and S/C = 1.0 at Tini = 1000 K

153 3. Reaction 661 contributes more (12%) for the S/C case than that for the φ case

(2.5%) at the start of pre-combustion. Thereafter, it decreases to 9% for the S/C

case, while it increases to 11.7% and then decreases to 3.92% at the combustion

limit for the φ case.

4.8.1.2 Combustion

1. The combustion region is divided into two subregions. The ﬁrst subregion is

dominated by reactions that take place during the pre-ignition section, while the

second subregion is dominated by a new set of reactions that are carried to the

cooling section.

2. Both cases present the same sets of reaction mechanisms and similar trends for

the two combustion subregions.

3. The H2 producing reactions for the ﬁrst combustion subregion are 223, 661, 376,

136 and 169, while the H2 consuming reactions for the ﬁrst combustion subregion

are 153 and 59.

No. Chemical Reaction 24 H + H2O −→ H2 + OH 136 H + CH2O −→ H2 + HCO 169 H + CH3HCO −→ H2 + CH3CO 194 H + CH4 −→ H2 + CH3 223 C2H6 −→ H2 + C2H4 244 N2 + C2H4 −→ N2 + H2 + C2H2 376 H + C3H6 −→ H2 + NC3H5 661 H + NC7H14 −→ H2 + NC7H13

Table 4.12. H2 Producing Reactions in the Combustion Section for φ = 1.6 and S/C = 1.0 at Tini = 1000 K

154 No. Chemical Reaction 54 H2 + O −→ H + OH 59 H2 + OH −→ H + H2O 153 H2 + CH3 −→ H + CH4

Table 4.13. H2 Consuming Reactions in the Pre-Ignition & Ignition Sections for φ = 1.6 and S/C = 1.0 at Tini = 1000 K

4. The H2 producing reactions for the second combustion subregion are 24, 244 and

194, whereas the H2 consuming reactions for the second combustion subregion

are 153, 59 and 54.

5. The combustion region is longer for the S/C case ( 12 ms) than for the φ case ¬ ( 3.5 ms). The ﬁrst three spikes in the combustion region behave in a very ¬ similar manner. The next two spikes, for the second subregion, are slightly larger

for the S/C case than for the φ case.

6. The total, reaction and volume gradients are larger for the equivalence ratio case,

about 5-6 times greater (still on the same order of magnitude) than the S/C case.

7. The volume term has a strong inﬂuence on the behavior of the total number

density gradient for both cases.

8. The H2 producing and H2 consuming reaction sets are the same for both cases

and behave similarly.

4.8.1.3 Cooling

1. The H2 producing reaction mechanisms presented during this region are the same

for both cases, and behave similarly. These reactions are 24, 244, 194, and 223.

155 2. The H2 consuming reaction mechanisms presented during this region are the same

for both cases; however, their behaviors diﬀer. The sum of their contributions

to the depleting part of the reaction term is larger for the S/C case. The H2

consuming reactions are 153 and 59.

3. Reaction 153 shows a ﬂuctuating behavior for the S/C case throughout the cooling

to ﬁnally decrease to 0%, while it shows a smoother decreasing behavior for the

equivalence ratio case.

4. Reaction 59 shows an increasing behavior that reaches a maximum and then

decreases to a negligible value for the S/C ratio case. This reaction decreases at

the beginning of the cooling region for the equivalence ratio case, then it increases

to a maximum and then decreases to zero at the end of the region.

5. The reaction term is more important compared to the volume term at the begin-

ning of the cooling region.

6. The volume term becomes very important after a few milliseconds of the combustion-

cooling limit.

4.8.2 Equivalence Ratio (φ = 1.6) cases at Low and High Tini

In this subsection, the equivalence ratio at initial temperature, 1000 K, is com-

pared against its equivalent at higher initial temperature, 1750 K. The main diﬀerences

noticed for these two cases are: 1) virtually no pre-ignition is seen for the high temper-

ature case since combustion is immediate, while the lower case displays a pre-ignition

region; 2) the adiabatic ﬂame temperature is higher for the high temperature case as

expected, and ﬁnally, 3) the hydrogen concentration is larger for the high temperature

156 case. Both cases show a rapid increase of temperature and hydrogen concentration at the ﬁrst combustion subregion, as well as the oscillating behavior observed in the second combustion subregion. The duration of these regions is smaller for the high temperature case. Further comparison on the behavior of the chemical reactions for both cases is carried out within the two common zones: combustion and cooling, in the following subsubsections.

4.8.2.1 Combustion

3 1. The order of magnitude for the H2 number density gradients is larger about 10 -

101 for the high initial temperature case than for the low initial temperature

case; i.e., 1028-1026 for the high temperature case compared to 1025 for the low

temperature case.

2. The volume term is seen to be insigniﬁcant at the beginning of combustion for

the high initial temperature case, while it has more weight for the low initial

temperature case.

3. The volume term becomes important, a few microseconds later for the high initial

temperature case. In fact, it is larger than the reaction term and causes the total

number density gradient to become strongly negative. Similar behavior is seen

for the low initial temperature case; however, the negative gradient is shallower

than that for the high initial temperature case.

157 4. The H2 producing reaction set is very similar for both cases; however, the per-

centage contributions of the H2 producing reactions at the beginning of the com-

bustion are much larger than those for the low initial temperature case, as well

as the introduction of new reactions.

5. No signiﬁcant contribution from hexadecane or heptene towards hydrogen is seen

for the high initial temperature case as is seen for the low temperature case.

A possible cause is that both hydrocarbons present a faster rate of decomposi-

tion towards other hydrocarbons or radicals, rather than reacting with hydrogen

radicals to produce hydrogen.

6. Reactions 232, 148 and 75 present signiﬁcant contributions for the high initial

temperature case over diﬀerent periods of time. These reactions are not impor-

tant for the low initial temperature case.

7. The H2 depleting reactions become very dominant in the second combustion sub-

region, turning the reaction term negative for more than 0.16 ms for the high

temperature case, while the reaction term for the low temperature case is always

positive.

8. Reactions 54, 103, 291, 65 and 27 present a larger contribution to the depletion of

hydrogen for the high initial temperature case than for the low initial temperature

case.

9. The trends of the H2 producing and H2 consuming reactions between these two

cases cannot be directly compared due to the addition of other reactions to the

reaction sets in the high temperature case.

158 4.8.2.2 Cooling

1. The number density gradients for both cases behave in a similar manner.

2. The H2 producing reaction sets are diﬀerent for these cases. The high tempera-

ture case reports that only reaction 24 is the maximum hydrogen provider during

the entire cooling period (49-71.5%). Other H2 producing reactions exist but

their contributions are less than 2%. This is completely diﬀerent for the low

temperature case where reaction 144, 194 and 223 play an important role.

3. The H2 depleting reactions are mainly 59 and 54 for the high temperature case

while the rest of the reactions mentioned in the combustion comparison are neg-

ligible for this section. On the contrary, reaction 153 has a signiﬁcant inﬂuence

on the hydrogen depletion for the low temperature case.

4.8.3 Equivalence Ratio (φ = 1.6) and S/C = 1.0 Ratio Cases at High Tini

In this subsection, the equivalence ratio (φ = 1.6) and steam to carbon ratio (S/C =

1.0) cases at high initial temperature, 1750 K, are compared. The adiabatic temper-

ature is seen to be higher for the equivalence ratio (φ = 1.6) case (with no water vapor

addition), as expected from the discussion in Subsection 4.8.1. However, the S/C ra-

tio case (S?C = 1.0) presents a larger ﬁnal hydrogen concentration ( 6.8%) than that ¬ for the equivalence ratio (φ = 1.6) case ( 6.1%). This is explained in the following ¬ paragraph in terms of the number density gradients.

The rapid changes occurring in the combustion section take place in a very narrow

zone which is divided into two subregions. The behavior of the number density and

159 its gradient are fairly similar for both cases (equivalence ratio (φ = 1.6 and S/C = 0.0)

and that with water addition (φ = 1.6 and S/C = 1.0)) during the ﬁrst combustion

subregion, although the S/C ratio presents slightly lower values (same order of magni-

tude). The second subregion, however, occupies a longer time frame for the S/C case

than for the equivalence case (φ = 1.6). This agrees well with the behavior observed

between the same cases at lower initial temperature in Subsection 4.8.1. The trends

for the number density gradients in the second subregion are similar for these two high

temperature cases, although slightly higher values for the equivalence ratio case (φ =

1.6). This also applies to the negative portion of the reaction term which is much

larger for the equivalence ratio case (φ = 1.6) than for the S/C ratio case. Hence, it causes the ﬁnal value of the hydrogen percentage volume at the end of the second combustion region to be lower for the equivalence ratio case than for the S/C case at the combustion-cooling limit. This impacts the H2 value in the cooling process.

The trends for both cases have been already discussed in the previous subsections and therefore are not included here. The comparison of the chemical pathways followed for the production of hydrogen for these two cases is carried out within the two distinct sections: combustion and cooling regions in the following subsubsections.

4.8.3.1 Combustion

1. The H2 producing and H2 consuming reaction sets are the same for both cases,

and so are their behaviors.

2. Reactions 223, 232, 136, 194, 75, 24, 194, 244, 148 and 376 comprise the producing

reactions set. Reactions 24 and 376 are seen to be present since earlier times for

the S/C ratio case.

160 3. Reactions 153, 59, 54, 103, 291, 65 and 27 comprise the H2 consuming reactions

set. Reaction 27 is seen to have a stronger inﬂuence for the equivalence case (φ

= 1.6 and S/C = 0.0) at the end of the combustion than for the S/C case (φ =

1.6 and S/C = 1.0).

Table 4.14. H2 Producing Reactions in the Ignition Section for φ = 1.6 and S/C = 1.0 at Tini = 1750 K

Table 4.15. Hydrogen Consuming Reactions in the Ignition Section for φ = 1.6 and S/C = 1.0 at Tini = 1750 K

161 4.8.3.2 Cooling

1. The H2 producing reactions set is diﬀerent for these cases. The equivalence ratio

case (φ = 1.6 and S/C = 0.0) is dominated only by reaction 24, while the S/C

ratio case (φ = 1.6 and S/C = 1.0) presents reactions 24, 194 and 223 coexisting.

2. Reaction 24 is seen to be the most important overall (47-55-30%) for the S/C case

(φ = 1.6 and S/C = 1.0). However, reaction 194 slowly increases to a fairly large

23% contribution whereas reaction 223 appears later in this region increasing to

8% at the end of the section.

3. The H2 reducing reactions set also diﬀers. The equivalence ratio case (φ = 1.6

and S/C = 0.0) presents reactions 59 and 54 as the largest hydrogen depleting

mechanisms, while the S/C ratio case (φ = 1.6 and S/C = 1.0) introduces reactions

153 and 59.

4.8.4 S/C = 1.0 Ratio Cases at Low and High Tini

In this subsection, the steam to carbon ratio case (φ = 1.6 and S/C = 1.0) at low initial temperature, 1000 K, is compared against the case defined by the same characterizing properties but at higher initial temperature, 1750 K. The trends of the number density gradients defined for the equivalence ratio (φ = 1.6 and S/C = 0.0) at different temperatures in Subsection 4.8.2 agree with the trends for these two cases, and therefore no further discussion is required. It is noteworthy that the presence of steam elongates the second combustion subregion for both S/C cases. Comparison of the mechanisms producing hydrogen for these two S/C cases is performed over the

162 two common zones (combustion and cooling) in the following subsections since no pre- ignition is seen for the high initial temperature case.

4.8.4.1 Combustion

1. The adiabatic ﬂame temperature achieved for the high initial temperature case

is higher than that for the low temperature case as expected.

2. The production of hydrogen is larger due to the higher adiabatic temperature

achieved for the high initial temperature case.

3. The H2 producing mechanisms set for both cases contain reactions 223, 244, 24,

194, 136, 169 and 376.

4. The high case temperature does not present direct inﬂuence of hexadecane reac-

tion 529.

Table 4.16. H2 Producing Reactions in the Ignition Section for φ = 1.6 and S/C = 1.0 at Tini = 1750 K

163 5. Reaction 24 presents a larger percentage contribution for the high initial temper-

ature case, which demonstrates that the addition of steam is only beneﬁcial at

high initial temperatures. Reaction 244 also shows a larger contribution for the

high initial temperature case.

6. Reactions 232, 148 and 75 present signiﬁcant contributions for the high initial

temperature case over diﬀerent periods of time. These reactions are not impor-

tant for the low initial temperature case.

7. Reaction 153 shows a larger percentage contribution to the depletion of hydrogen

for the low initial temperature case.

8. Reactions 54, 103, 291, 65 and 27 present a larger contribution to the depletion of

hydrogen for the high initial temperature case than for the low initial temperature

case.

Table 4.17. Hydrogen Consuming Reactions in the Ignition Section for φ = 1.6 and S/C = 1.0 at Tini = 1750 K

164 9. The trends of the H2 producing and H2 consuming reactions between these two

initial temperature cases cannot be directly compared due to the prominence of

other reactions to the chemical reaction sets in the high initial temperature case.

4.8.4.2 Cooling

1. The number density gradients for both cases behaves in a similar manner.

2. The H2 producing mechanism set slightly diﬀers for these two cases. The high

initial temperature case shows the inﬂuence of only reactions 24, 194 and 223,

while the low initial temperature case introduces reaction 244 as well.

3. The inﬂuence of reaction 24, vapor and atomic hydrogen, is greater for the high

temperature case through the entire cooling period, whereas it is only important

for the beginning of the section for the low temperature case.

4. The H2 reducing mechanisms are the same for both cases. However, the high

initial temperature case shows reaction 153 as the most dominant at the tail of

the cooling and reaction 59 at the beginning. The low initial temperature case

shows an opposite behavior.

4.8.5 S/C = 1.0 Ratio and Post-Injection (P.I.) Cases at Tini = 1000 K and TPI = 2000 K

Finally, in this subsection the steam to carbon ratio case (φ = 1.6 and S/C =

1.0) and its equivalent for post-injection with same initial temperature, 1000 K, are

compared. The comparison is carried out within the cooling section of the S/C low

initial temperature case and over the post-injection and cooling periods of the P.I.

165 case. This is explained in the following subsubsections. The combustion section is

not compared since the P.I. case utilizes that for the equivalence ratio at low initial

temperature, which has already been discussed in Subsection 4.8.1.

The S/C case presents an increasing hydrogen concentration with slower rate than

that for the post-injection case, lower hydrogen concentration at equilibrium ( 4%). ¬ This can be explained by the fact that the total number density gradient and the reaction term decay at a slower rate during the post-injection section than for the cooling performed in the S/C case (φ = 1.6 and S/C = 1.0) within equivalent time

window. The volume term for this time period is nearly zero for both cases, therefore,

the total gradient is driven by the reaction term. The cooling period for the P.I.

behaves in a similar manner as that for the S/C, however, the duration is shorter for

the P.I. case.

During the post-injection section, reactions 24 and 194 drive the reaction term

contributing with an average of 35% and 15%, respectively. During the second cooling

section of the P.I. case, reaction 24 shows a decreasing behavior until it totally vanishes

at the end of the cooling. Reaction 194 increases for a small period and then decreases

to a negligible value. For the S/C ratio (φ = 1.6 and S/C = 1.0), reaction 24 behaves

in a similar manner to that for the P.I. case; however, reaction 194 strongly decreases

at the beginning of the cooling period to 4% and then rises. It can be seen that this

behavior agrees with that for the reaction term discussed in the previous paragraph,

which decreases faster for the S/C ratio case (φ = 1.6 and S/C = 1.0).

The H2 consuming reactions are 153 and 59 for both cases. Reaction 153 decreases

through a more ﬂuctuating process for the P.I. case than for the S/C case (φ = 1.6 and

166 S/C = 1.0); however, the global trend for this reaction is similar for both cases. The behavior of reaction 59 is also equivalent for both cases, S/C and P.I.

167 CHAPTER 5

CHEMICAL EQUILIBRIUM COMPOSITIONS FOR REFORMING OF HEXADECANE

5.1 Introduction

The mechanisms that inﬂuence the production of hydrogen under the diﬀerent characterizing quantities and their dynamics for selected cases were presented in Chapter 4.

This chapter introduces the concentrations not only of hydrogen but also that of the major species at equilibrium for all the studied cases. With the aid of the previous chapter, the eﬀect of some reactions on the formation of hydrogen can be highlighted.

Following the organization of the previous chapter, this chapter is divided into

Section 5.2 to Section 5.6 to differentiate the influence of each characteristic property on the production of hydrogen and the major species. Hence, the sections are organized as: 1) Effect of Equivalence Ratio at Low Temperature, 2) Effect of Steam to Carbon

Ratio at Low Temperature, 3) Eﬀect of Higher Initial Temperature for Equivalence and

Steam to Carbon Ratio, 4) Eﬀect of System Pressure, and 5) Eﬀect of Post-Injection

Temperature. The cases containing exhaust gas as the source of oxygen rather than pure air are discussed in Section 5.7. Finally, Section 5.8 highlights the main features

168 of each parameter (equivalence ratio, steam to carbon ratio, temperature, pressure and secondary injection of water at two diﬀerent temperatures) and presents a comparison among these quantities.

5.2 Eﬀect of Equivalence Ratio at Low Temperature

As stated in Chapter 3, the partial oxidation of hexadecane is modeled using a large set of chemical reactions with finite kinetic rates. The reacting mixture is allowed to achieve equilibrium as time evolves, approximately a few fractions of a second after combustion is completed. The pre-ignition, ignition and cooling regions identified in the previous chapter revealed interesting features of the different chemical pathways followed in the POx of hexadecane to produce hydrogen (mainly radicals and atomic hydrogen) under the influence of the characterizing properties. The equivalence ratio range studied in this thesis is 1 ≤ φ ≤ 1.9.

Thermodynamic equilibrium calculations are typically used to predict the compositions of the major species (H2, CO, CO2, NO, among others) using a global reaction rather than elementary mechanisms, and assuming equilibrium for the general process.

In order to compare the kinetic trends obtained with the 0D model at equilibrium, equilibrium calculations are discussed here. As an attempt to predict the trends of the major species as well as the adiabatic temperature at equilibrium, the combustion of the hexadecane molecule is simulated using the code provided by Turns [96] that utilizes the Olikara and Borman [95] equilibrium routines. Here, it has to be clariﬁed that this code accounts for an initial temperature of 298 K (STP) and not for 1000 K.

169 Hence, the equilibrium compositions and the adiabatic temperature within the equiva-

lence ratio limits for the constant-pressure POx of hexadecane are presented in Figure

5.1 and Figure 5.2.

One should expect the adiabatic ﬂame temperature to be higher as the temperature

of the feeds is increased. Hence, a second attempt to account for this factor was carried

out using the MatLab version of the Olikara and Borman [95] equilibrium routines de-

veloped by Buttsworth [97]. In this method, the air properties, i.e., enthalpy, speciﬁc

heat, entropy, etc., are obtained using the coeﬃcients provided by Gordon et al.[98]

based on the JANAF Thermochemical Tables [99]. This code is valid for several fuels such as diesel, methane, gasoline, among others, but does not contain the coeﬃcients for hexadecane to calculate its thermodynamic properties. In order to account for hexadecane, the 7-term NASA polynomials used in the actual kinetic model presented in this thesis obtained from [100] are employed. Thus, the equilibrium compositions and the adiabatic ﬂame temperature are calculated. Figure 5.3 and Figure 5.4 show the

equilibrium composition and adiabatic ﬂame temperature, respectively, for the higher

initial temperature of 1000 K.

As expected, the adiabatic ﬂame temperature decreases with equivalence ratio owing

to the fact that the heat of combustion (∆HR = Hprod -Hreac) decreases, which may be

explained as follows. On the one hand (as will be explained later), the equivalence ratio

causes the concentration of diatomic molecules, such as H2 and CO, to increase. These

molecules have a lower enthalpy of formation than triatomic molecules such as CO2 and

H2O, which have a large concentration at low equivalence ratios (near stoichiometry).

Thus, the product enthalpy term decreases. On the other hand, raising the amount of

170 Figure 5.1. H2, CO, H2O and CO2 Equilibrium Concentrations for Hexadecane at Dif- ferent φ and S/C Ratios at 1 atm and Tini = 298 K based on [95, 96]

Figure 5.2. Equilibrium Adiabatic Flame Temperature for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 298 K based on [95, 96]

171 Figure 5.3. H2, CO, H2O and CO2 Equilibrium Concentrations for Hexadecane at Dif- ferent φ and S/C Ratios at 1 atm and Tini = 1000 K based on [97]

Figure 5.4. Equilibrium Adiabatic Flame Temperature for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1000 K based on [97]

172 fuel increases the total enthalpy of the reactants. The combination of these two terms definitively affects the final adiabatic temperature. This can also be understood as the heat release is less as there is less oxygen available to burn the fuel, thus lowering the temperature [101].

Figure 5.5. Adiabatic Flame Temperature for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1000 K

The difference in the numerical value for two equilibrium compositions at the two different initial temperatures is completely due to the initial temperature, which affects the reactants’ sensible enthalpy. The adiabatic flame temperature calculated by the

0D-model, shown in Figure 5.5, also decreases with equivalence ratio. However, its numerical value is much higher for low equivalence ratios. This may be explained by the fact that the concentration of CO2 and H2O are slightly larger at those equivalence

173 ratios (as will be explained later), which increases the enthalpy of the products. Also, the equilibrium composition only considers hexadecane to be the only fuel undergoing oxidation, while the kinetic mechanism takes into account the oxidation of the small hydrocarbon species formed during the pre-combustion section which decreases the reactants’ enthalpy for the kinetics code.

For the equilibrium compositions, Figure 5.3, hydrogen and carbon monoxide are formed by the dissociation of water and carbon dioxide, respectively. Based upon this,

CO increases with increasing equivalence ratio, opposite to the carbon dioxide concentration. Formation of water vapor is preferred at low equivalence ratios, however, part of the steam undergoes dissociation and hydrogen is then produced. As the equivalence ratio increases, φ > 1.3, the hydrogen concentration increases at a faster rate while steam concentration decreases. This rate is controlled by the water gas shift reaction CO2 + H2 CO + H2O. Compositions for OH, H and O radicals are also obtained using the equilibrium codes, however, their values are very small in comparison with the major species.

The compositions of the major species calculated by means of the chemical kinetics code at equilibrium under the influence of the equivalence ratio (1 ≤ φ ≤ 1.9) and steam to carbon ratio (0 ≤ S/C ≤ 2) are outlined in Figures 5.6 - 5.9. In this section the main interest is the equivalence ratio effect, while the steam to carbon ratio effect is explained in the following section. As illustrated by the equilibrium compositions, the hydrogen concentration increases with increasing equivalence ratio. Based upon the results presented in the previous chapter, hydrogen is produced not only by the

174 recombination of atomic hydrogen and small hydrocarbons including aldehydes (R-

136, R-169 and R-194), but also by decomposition of hydrocarbons (R-529, R-223 and R-661). Furthermore, hydrogen is seen to be abundantly produced within the combustion region by an elementary step of the water gas shift reaction (R-24) and by the collision of molecular nitrogen with ethylene (R-244). These elementary processes, listed in Table 5.1, clearly provide more information on the formation of H2 than the simple dissociation of water used in the equilibrium codes.

Figure 5.6. H2 Concentration for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1000 K

175 No. Chemical Reaction 24 H + H2O −→ H2 + OH 136 H + CH2O −→ H2 + HCO 169 H + CH3HCO −→ H2 + CH3CO 194 H + CH4 −→ H2 + CH3 223 C2H6 −→ H2 + C2H4 244 N2 + C2H4 −→ N2 + H2 + C2H2 376 H + C3H6 −→ H2 + NC3H5 529 NC16H34 + H −→ H2 + NC16H33 661 H + NC7H14 −→ H2 + NC7H13

Table 5.1. H2 Producing Reactions

No. Chemical Reaction 28 H2 + O2 −→ 2OH 54 H2 + O −→ H + OH 59 H2 + OH −→ H + H2O 153 H2 + CH3 −→ H + CH4

Table 5.2. H2 Consuming Reactions

The hydrogen concentration predicted by the kinetics model is lower for than that predicted by the equilibrium model. This may be explained by the fact that the only hydrogen consuming mechanism employed in the equilibrium code is the dissociation of hydrogen into atomic hydrogen while the kinetics code considers other depleting mechanisms, i.e., R-59 and R-153, listed in Table 5.2. One can argue that the equilibrium compositions are evaluated at the adiabatic ﬂame temperature while the composition in the kinetics code is evaluated here at 500 K (after cooling). Nevertheless, Chap- ter 4 clearly stated that hydrogen can be considered in equilibrium a few fractions of

176 a second after the adiabatic ﬂame temperature is achieved. Furthermore, the equi-

librium hydrogen concentration for stoichiometric combustion suggests that hydrogen

is produced at combustion in small amountst. The transients of the stoichiometric

combustion studied with the kinetics model indicate that a small amount of hydrogen

exists at the combustion zone; however, its concentration rapidly vanishes through its

oxidation (R-28 and R-54: molecular and atomic oxygen, respectively) and through the

water gas shift reaction (R-153) given in Table 5.2 and Appendix A.

Figure 5.7 shows the predicted values of the CO volume fraction from the 0D chem-

ical kinetics model. As can be seen, the trend of CO is in good agreement with that

showed by the equilibrium calculations. In general, more CO is produced with increas-

ing φ. This is beneﬁcial in the case of reforming for some after-treatment applications

since both CO and H2 are found to increase with increasing equivalence ratio. The

numerical values are slightly diﬀerent between the kinetics calculation and the equi-

librium calculations. This is explained by the amount of CO2 reported in Figure 5.8

since they are directly related by the WGS reaction. As can be seen, the variation of

the volume fraction of CO2 versus equivalence ratio is opposite to that of the H2 and

CO volume fractions shown in Figures 5.6 and 5.7. As φ increases, the CO2 volume

fraction sharply decreases. Moreover, unlike the concave down shape of the curves for

H2 and CO, the volume fraction of CO2 versus φ is concave up. The trend obtained from the kinetics 0D model agrees with that obtained for the equilibrium calculations.

177 Figure 5.7. CO Concentration for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1000 K

Figure 5.8. CO2 Concentration for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1000 K

178 Figure 5.9. THC Concentration for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1000 K

Figure 5.10. H2O Concentration for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1000 K

179 As presented in Chapter 3 and the reaction mechanisms in Appendix A, the 0D ki-

netics model calculates the concentration of individual hydrocarbon species. A large

number of plots could be generated and presented here to account for each hydrocarbon

species; however, their individual concentrations are small. Furthermore, Chapter 6

and Chapter 7 present the concentration of the total amount of hydrocarbons present in the reformate on a C1 basis. Based upon these facts, the hydrocarbons calculated by the kinetics model are lumped into a quantity on C1 basis expressed as THC. It is

important to mention that for all equivalence ratios, the most important hydrocarbons

are methane, CH4, butadiynyl, C4H, ethylene, C2H4, and acetylene (ethyne), C2H2, among others. Figure 5.9 shows the variation of the volume fraction of total hydro-

carbons (THC) versus equivalence ratio. As expected, the volume fraction of THC

increases with φ sharply since less heat is released as the equivalence ratio increases,

leaving more hydrocarbons unburned.

As can be seen from Figure 5.6, the volume fraction of H2 in the ﬁnal reacted

mixture are found to increase with increasing φ. This would suggest that the POx of

hexadecane should be performed at as high an equivalent ratio as possible except that

excessive soot formation becomes limiting in actual experiments at values of φ > 2.2 and beyond. Although the kinetics mechanisms do not account for the formation of soot, butadiynil, a soot precursor, presents the highest hydrocarbon concentration indicating then that more soot can be formed as the equivalence ratio increases.

The concentration of vapor calculated by the 0D kinetics model is depicted in Figure

5.10. As can be observed, steam increases with increasing equivalence ratio reaching its

maximum between φ = 1.3-1.6. Thereafter, its concentration decreases with increasing

180 equivalence ratio. This trend agrees well with that obtained from the equilibrium

calculations. Again, the equilibrium calculations show a lower steam concentration

which may be justiﬁed by the slightly larger hydrogen concentration (WGS reaction).

The hydrogen yield deﬁned in Chapter 3 as a function of equivalence ratio consid-

ering fuel as the only source of hydrogen, since no additional water is administered for

this section, is depicted in Figure 5.11. It can be seen that this quantity exhibits a

saturation-like behavior with increasing φ. Although not shown here, increasing the equivalence ratio φ > 1.9 causes the hydrogen concentration to saturate as well.

Figure 5.11. H2 Yield for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1000 K

Based on the results of the 0D model, this trend may be caused by the decrease of

the adiabatic ﬂame temperature with increasing φ, as shown in Figure 5.5. Hence, this

181 decrease in temperature limits the kinetics of H2 generation. These opposing mechanisms, H2 increasing and temperature decreasing, lead to the saturation-like behavior evident in Figure 5.11. As mentioned before, the 0D model also reveals that there are other factors that tend to decrease H2 production at higher values of φ. These include the presence of smaller hydrocarbon radicals such as CH3, which tend to consume H2 via scavenger mechanisms such as R-153. This is corroborated by the fact that CH4,C2H2 and other small hydrocarbons are produced in increasing numbers as the equivalence ratio is increased. The same saturation-like behavior evident in the volume fraction of H2 and in the hydrogen yield in Figures 5.6 and 5.11 is also observed for CO. As in the case of H2, the decrease in the ﬂame temperature with increasing φ is responsible for this trend in the CO volume fraction.

Figure 5.12. Dry-N2-free H2 Concentration for Hexadecane at Diﬀerent φ and S/C Ra- tios at 1 atm and Tini = 1000 K

182 Figure 5.12 illustrates the concentration of hydrogen in dry-N2-free conditions.

This ﬁgure is included since existing work report results for most catalytic POx reform-

ers on a dry-nitrogen-free basis, showing then the high 70-90% hydrogen concentrations.

It can be seen that if the water vapor and the nitrogen were removed from the reformate

stream, the hydrogen concentration using non-catalytic partial oxidation reach values

close to 30% at φ = 1.9. This ﬁgure captures more clearly the saturation-like behavior

explained above.

From this section it can be seen that the 0D chemical kinetics model presented here

reveals some key details regarding thermal partial oxidation of diesel fuel (modeled

as C16H34). First, hydrogen and carbon monoxide concentrations increase (concave

down behavior) with increasing equivalence ratio. This is beneﬁcial for some of the

after-treatment technologies discussed in Chapter 2. Hence, the H2 yield as deﬁned by

Equation (3.56), a measure of the eﬃciency of the reforming process, also increases with increasing equivalence ratio. Furthermore, conversely to hydrogen, the concentration of carbon dioxide decreases with a concave up trend. These species show a saturation- like behavior for φ > 1.9 which can be explained by the decrease of the adiabatic

ﬂame temperature (heat release) and the increase of scavenger radical species that consume hydrogen. The total hydrocarbon concentration increases exponentially as the equivalence ratio is increased owing to the decrease in temperature. Thus, one of the main conclusions drawn from this discussion is that temperature plays a very important role in the production of hydrogen via non-catalytic reforming of hexadecane.

Finally, another interesting phenomenon caused by the equivalence ratio is the characteristic ignition time, which is larger as the equivalence ratio decreases, see Figure

183 5.13. According to Westbrook [102], the ignition time depends on the number of chain carriers, i.e., chain-branching mechanisms that can lead a system to completion in a very short period of time. Ignition of a reactive system was deﬁned by Westbrook

[102] as an exponential growth in temperature and number of chain carriers (radicals).

Hence, the ignition time depends on the amount of radicals present at a particular

temperature and pressure. Here, the sub-stoichiometric air/fuel mixture is adminis-

tered at the same temperature and pressure with the equivalence ratio being increased.

From the 0D model it can be seen that the amount of hydrogen radicals (H, OH, HO2) produced in the pre-ignition zone increases with φ, as does the concentration of hydrocarbon radicals (CH3, aldehydes (CH3HCO), and unsaturated HCs). Aldehydes, according to Glassman [101], inﬂuence the ignition time lag, either accelerating or slow-

ing it down. The production of these radicals, however, may increase their oxidation

rate (mainly with O2, O and OH) as φ increases. Also, the recombination of these radicals (chain termination) becomes important in this zone since the rate of production of the major species and stable hydrocarbons seems to be faster. These species also oxidize resulting in a larger exotherm, for instance hydrogen oxidizes with OH producing water vapor, also CO oxidizes with O2. The heat release, reﬂected as the

temperature, has a slightly faster increasing rate as φ is increased as can be seen in the

zoomed region illustrated in Figure 5.14.

184 Figure 5.13. Ignition Delay for Hexadecane for Diﬀerent φ Ratios at S/C = 0, 1 atm and Tini = 1000 K

Figure 5.14. Ignition Delay for Hexadecane for Diﬀerent φ Ratios at S/C = 0, 1 atm and Tini = 1000 K

185 5.3 Eﬀect of the Steam to Carbon Ratio at Low Temperature

This section focuses on the effect of the steam added as part of the feeds, or what is called in this thesis main injection. Figures 5.5 - 5.12, discussed in the previous section illustrate the influence of steam to carbon ratio within the limits 0 ≤ S/C ≤ 2, and for the range of equivalence ratios mentioned before. It is noteworthy to mention that a few studies related to the non-catalytic partial oxidation with the addition of water vapor are found in the literature. As mentioned in Chapter 2,[66] strongly states that the addition of water affects negatively the concentration of hydrogen by lowering the reaction temperature and could also increase the soot formation. Conversely, [67] presented a numerical and experimental analysis for the formation of soot in the oxidation of diesel fuels, in which adding water increased the percentage concentration of hydrogen. This discrepancy and the fact that high temperatures are observed for the partial oxidation of hexadecane, which could lead to the dissociation of water (via WGS step: H2O + M → H + OH + M, where M is a third body), provided the motivation to further investigate the effect of water. Furthermore, the availability of the reaction kinetics also facilitate this study.

For a given φ, the adiabatic flame temperature decreases as S/C increases owing to the fact that water has a high specific heat therefore some of the heat released from the combustion is absorbed by the water. As φ increases, the adiabatic flame temperature for all S/C ratios decreases. From figure 5.5 it can be seen that the temperature for the S/C = 0 case decreases at a faster rate than for the other S/C cases. This is related to the fact that after φ > 1.39, the product concentration (triatomic and higher

186 molecules) for the steam cases is slightly larger than that for the pure combustion case,

S/C = 0, increasing then the products total enthalpy for those S/C cases.

Figure 5.6 depicts the variation of hydrogen percent volume with S/C and φ ratios. From this ﬁgure it can be seen that a combination of low S/C and low φ provide a slightly larger concentration of hydrogen than the pure combustion case. As the equivalence ratio is increased, the hydrogen concentration starts to saturate as explained in the previous section. This is more noticeable as the S/C increases since the adiabatic temperature is lower as S/C increases, limiting therefore the kinetics of hydrogen production. Figure 5.15 is employed to aid with the interpretation of the hydrogen concentration, showing the slight increase in hydrogen concentration for low equivalence ratios and the negative eﬀect as S/C increases. The H2 producing reactions set, as explained in Chapter 4, is that of the equivalence ratio section presented above, with the variation of R-661: heptene reacting with atomic hydrogen (H +

NC7H14 −→ H2 + NC7H13). Another conclusion withdrawn from the previous chapter was that the combustion zone presents slower rates (lower gradients) for the S/C cases. It is also seen that the characteristic ignition time, already described for the equivalence ratio, increases as the S/C increases for a given φ ratio, as shown in Figure

5.16. This delay depends more on the speciﬁc heat of the water vapor, meaning that water vapor absorbs some of the heat release at the pre-ignition zone, slowing therefore the production of radicals species (i.e., CH3) and the oxidation of hydrocarbons.

187 Figure 5.15. H2 Concentration for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1000 K

Figure 5.16. Ignition Delay for Hexadecane for Diﬀerent S/C Ratios at φ = 1.6, 1 atm and Tini = 1000 K

188 The CO concentration increases with φ but the addition of steam limits the kinetics of CO production as temperature decreases with S/C ratio. Carbon dioxide follows a similar behavior as described in Section 5.2, decreasing with φ for all the S/C ratio cases.

However, the decreasing rate of CO2 is slower for the cases containing steam as part of the feeds. This causes the trends to converge at φ 1.39 with a concentration about ¬ 6.2%. The slow decreasing rate is completely related to the temperature decreasing rate which seems to be faster for the low S/C cases.

The total hydrocarbon concentration increases as the equivalence ratio increases and shows a slight increase as the S/C ratio increases. The vapor product concentration for any given equivalence ratio is seen to increase as the S/C ratio increases. This is expected since part of the water does not react or decompose during combustion part. It is noteworthy that the maximum steam concentration moves along with the equivalence ratio as S/C increases. This is a temperature eﬀect which decreases as

φ and S/C increase, therefore a large part of the initial vapor does not undergo any chemical interaction.

The hydrogen yield deﬁned in terms of only fuel as the hydrogen provider shows an increase ( 2% Vol) for all S/C ratios at low φ ratios. As the equivalence ratio ¬ increases, the yield starts to saturate with a higher eﬀect for the high S/C cases as depicted in Figure 5.11. The dry-nitrogen free hydrogen concentration outlined in

Figure 5.12 presents similar trends as those appointed by the hydrogen yield. One can argue that the deﬁnition of hydrogen yield presented previously does not consider vapor as a hydrogen provider. For this reason, Figure 5.17 illustrates the behavior of the yield deﬁned in Equation (3.61), which accounts for fuel and steam as hydrogen

189 providers. From Figure 5.17, it can be concluded that addition of steam lowers the hydrogen yield and emphasizes the saturation-like behavior as S/C increases.

Figure 5.17. H2 Yield for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1000 K based on Fuel and Steam

Based on the discussion presented in this section, it is concluded that the addition of steam slightly increases the hydrogen percent volume for low φ ratios agreeing with

[67]. As the S/C is relatively low, this implies that the amount of water required to achieve this increase in hydrogen percent is very small. For instance, for automotive applications (after-treatment), this amount of water may be extracted from the engine exhaust where it is generated as a byproduct of combustion. Should additional water be needed, a small storage tank would need to be installed on the vehicle. Nevertheless, it has to be kept in mind that the increase of hydrogen concentration and hydrogen

190 yield on fuel basis is not very signiﬁcant. The hydrogen yield based on fuel and water

vapor show an even lower percent, supporting the theory that for real applications the

increase of H2 [% Vol] is relatively small.

5.4 Eﬀect of the Initial Temperature for φ and S/C

As concluded from Section 5.2, temperature plays a very important role in the

reformation of hexadecane. Based upon this, an exploratory numerical analysis of the

initial temperature eﬀect is studied using Tini = 1500 K and Tini = 1750 K for the

equivalence and steam to carbon ratios deﬁned previously. These temperatures may

be diﬃcult to attain. Nevertheless, commercial electrical heaters capable of elevating

temperatures up to 1800 [105] can be used for this purpose. The results presented

in this section, for the discussion on the temperature eﬀect, are mainly those for 1750

K since the trends for some species are similar. Only the hydrogen related ﬁgures for

1500 K are also enclosed.

As expected, the adiabatic ﬂame temperature for a given combination of equivalence

ratio and steam to carbon ratio increases as the temperature of the feeds is increased.

This is shown in Figure 5.18. Furthermore, the adiabatic temperature decreases with

φ and S/C. However, it is noteworthy that the values of the adiabatic temperature are

much higher than those reported for Tini = 1000 K. This eﬀect is relevant to understand the behavior of the major species under the inﬂuence of φ and S/C ratios.

The hydrogen concentration further increases as the initial temperature is increased for a given φ and S/C as shown in Figures 5.19 and 5.20, for 1500 K and 1750 K

respectively. As explained in the previous sections, the hydrogen concentration tends

191 to saturate faster (with φ) as the S/C ratio is increased. This is also seen for the high

temperature cases, however, to a lesser extent owing to the much higher adiabatic ﬂame

temperature for a given φ and S/C as the initial temperature increases. Consequently,

the hydrogen yield in terms of fuel as the only source of hydrogen increases with initial

temperature for all φ and S/C ratios studied here as depicted in Figures 5.21 and

5.22. For instance, for φ = 1.6 and all S/C the yield increases from 20-26% at 1000

K to 30-37% at 1500 K and farther to 35-48% at 1750 K. It is seen that the addition

of water has a larger impact on the yield as the initial temperature increases due to

the higher adiabatic temperatures. This can be directly translated to the saturation

behavior present as φ and S/C increase. Then, the saturation is retarded as the initial

temperature increases.

The hydrogen yield based on fuel and water as sources of hydrogen is illustrated in

Figures 5.23 and 5.24 for 1500 k and 1750 K, respectively. As can be seen from these

ﬁgures, the yield increases for a given φ and S/C as initial temperature is increased.

However, the percentage that this yield increases ( 1-3%) is much smaller than that ¬ for the yield deﬁned only for fuel at a given φ. Figure 5.25 shows the hydrogen con-

centration in a dry-N2 free basis. For a given φ value, the increasing S/C ratio shows an enhancement on the hydrogen production for Tini = 1500 K and Tini = 1750 K while it is only seen for the low φ at Tini = 1000 K.

192 Figure 5.18. Adiabatic Flame Temperature for Hexadecane at Diﬀerent φ and S/C Ra- tios at 1 atm and Tini = 1750 K

Figure 5.19. H2 Concentration for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1500 K

193 Figure 5.20. H2 Concentration for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1750 K

Figure 5.21. H2 Yield for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1500 K

194 Figure 5.22. H2 Yield for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1750 K

Figure 5.23. H2 Yield for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1500 K based on Fuel and Steam

195 Figure 5.24. H2 Yield for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1750 K based on Fuel and Steam

Figure 5.25. Dry-N2-free H2 Concentration for Hexadecane at Diﬀerent φ and S/C Ra- tios at 1 atm and Tini = 1750 K

196 The concentration of CO depicted in Figure 5.26 is also enhanced by the increasing

initial temperature for a given φ and S/C ratios. As explained in the previous sections,

the S/C ratio is found to decrease the CO percent for all φ cases analyzed in this thesis

due to the decreasing adiabatic temperature. Carbon dioxide concentration slightly

increases with initial temperature for φ < 1.6 and S/C ratios as can be seen in Figure

5.27. As explained for the low temperature case in Section 5.3, the decreasing rate

of CO2 is more accentuated for the low S/C ratios forcing the trends to converge at

φ 1.3. Thereafter, the CO2 concentration is seen to increase slightly with S/C and ¬ initial temperature for a particular φ ratio. The emission of hydrocarbons, outlined

in Figure 5.28, decreases as the initial temperature is increased for all φ and S/C

ratios. This result is expected since higher temperatures lead to fast decomposition

of hydrocarbons as mentioned in Chapter 4, resulting hence in a larger production of

hydrogen and CO species as seen in this section.

197 Figure 5.26. CO Concentration for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1750 K

Figure 5.27. CO2 Concentration for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1750 K

198 Figure 5.28. THC Concentration for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1750 K

5.5 Eﬀect of Pressure for φ and S/C

The results exposed in the previous sections are calculated for atmospheric pressure

owing to fact that all the kinetics coeﬃcients for all the participant chemical equations

in the oxidation of hexadecane are calculated for 1 atm [88]. According to Glassman

[101], increasing the pressure results in an increase on the adiabatic ﬂame temperature

and a decrease on the amount of dissociation. It is claimed that the eﬀect is greater

as the equivalence ratio approaches stoichiometry, since the amount of dissociation is

greater for those φ ratios. Glassman [101] also highlights that recombination reactions involving a third-body (such as N2,H2O to mention a few) are pressure-sensitive, and

therefore their coeﬃcients must be calculated for the actual pressure. Based upon

199 these facts, any calculation for a system using a set of coefficients calculated for a different pressure would result in misleading conclusions. As mentioned before, the coefficients used in this thesis are expressed for 1 atm, and therefore, no further study can be performed.

The equilibrium code from [97] is employed to identify the trend of the adiabatic

ﬂame temperature at 2 atm for the equivalence ratios used in this thesis to verify

Glassman’s statement. Comparison between Figures 5.4 and 5.29 clearly shows an temperature increase for the low equivalence ratio, while the high equivalence ratio are less aﬀected. An exploratory analysis is then performed with the kinetics 0D model at

2 atm for the same equivalence ratio range. Figure 5.30 shows a decreasing adiabatic

ﬂame temperature for the low equivalence ratios and increasing Tad for the high φ when comparing with Figure 5.5; which completely disagrees with the expected trends deﬁned by Glassman [101] and those determined through the equilibrium calculation.

This indeed corroborates the fact that the coeﬃcients must be calculated for the correct pressure and, therefore, no further details on the major concentrations are presented here.

200 Figure 5.29. Equilibrium Adiabatic Flame Temperature for Hexadecane at Diﬀerent φ and S/C Ratios at 2 atm and Tini = 1000 K based on [97]

Figure 5.30. Kinetics Adiabatic Flame Temperature for Hexadecane at Diﬀerent φ at 2 atm and Tini = 1000 K

201 5.6 Eﬀect of Post-Injection of Water Vapor for φ and S/C

From Sections 5.3 and 5.4, it is clearly seen that main injection of water has a larger eﬀect at high inlet temperatures. Withdrawn from those sections, if water is not introduced as part of the main injection reactants, the mixture of fuel and air ignites and reaches higher adiabatic temperatures. One question of interest is to study the eﬀect of introducing water or additional fuel in the reactor to emulate a secondary injection.

The injection of additional fuel was studied on a separate project and therefore will not be discussed here. However, the main conclusion was that the decomposition of hexadecane produces highly reactive radicals (CH3) that tend to consume hydrogen.

Cases were compared based on an equivalent ratio deﬁned as φtotal = φMI + φPI , where

φPI represents the amount of fuel post-injected divided by the amount of air used at the main injection. This corresponds to a higher φMI depending on the amount of fuel utilized. Hence, a main injection φMI = 1.9 can be roughly emulated with post- injection as φ = 1.6 + 0.3. It was seen that the hydrogen concentration obtained using the secondary injection was lower than the correspondent employing a higher main injection equivalence ratio owing to the hydrogen consuming species. However, employing water at a secondary port will not reﬂect the same impact since no additional creation of those hydrogen consuming species will take place.

Based upon this conclusion and the fact that temperature highly impacts the reformation of hexadecane, two different post-injection temperatures are analyzed in this thesis, i.e., 1500 and 2000 K. This could be translated to an actual reformer as the location of the secondary port, which could be right at the flame or a few centimeters downstream where convective heat transfer has cooled down the flame a few degrees

202 Kelvin. As known from Section 5.2, the adiabatic ﬂame temperature of all φ with no

addition of water are above 2000 K. In order to account for the viable heat transfer

occurring in an actual reformer, the mixture of products (from air and fuel combustion)

is cooled down to the post-injection temperature. The amount of water post-injected

can be calculated using the equations given in Section 3.2.3.

The combustion of fuel and air provides the common contribution of hydrogen to

both cases. As shown in Figure 4.1, the hydrogen concentration decreases slightly

right after the cooling process has started. The post-injection cases, however, present

a much faster cooling rate over a period of 0.1 ms after combustion. During this

time interval the hydrogen percent decreases because the consuming reactions (R-153

and R-59) dominate the reaction term, as explained in Section 4.7. Over the post-

injection period the concentration of hydrogen is seen to be fairly constant for the low

post-injection temperature, while it increases about a 25% (from 3.5 to 4.5% Vol) for

the high temperature case at φ = 1.6. This again highlights the important role that temperature plays on the reformation of hexadecane.

The behavior of the species is similar for both cases, however, the equilibrium values are larger for the high temperature case. For simpliﬁcation, only the ﬁgures for 1750

K are presented in this thesis. Nevertheless, discussion on both temperature cases is carried out in this section. Hydrogen concentration is seen to be lower for the 1500

K cases in comparison with the higher post-injection temperature cases and also with the equivalent case at main injection. Figure 5.31 shows the percentage volume of

hydrogen. The saturation of the hydrogen concentration is also observed for this cases.

Nevertheless, the addition of water at the secondary injection does not increases the

203 saturation trend for the high S/C as it was observed for the equivalent main injection

cases. This is due to the fact that the main concentration of hydrogen was generated

through the partial oxidation with no water addition and its behavior through the

post-injection period is fairly constant. The 1750 K case shows very similar behavior

for all S/C for φ < 1.6. For higher equivalence ratios, the high S/C starts to saturate

faster. This case in comparison with the correspondent main injection case shows a

very slight enhancement for S/C = 0-0.5. This enhancement is more observable for

the S/C = 1-2 explained previously for the saturation of the concentration.

Based on the results for the hydrogen concentration, the hydrogen yield fuel-basis

for the low temperature case is fairly constant for all the S/C cases since the concen-

tration of hydrogen is mainly produced through the POx process rather than in the

post-injection. For the high temperature case, after the rapid cooling, the hydrogen

concentration increases within the post-injection section. Hence, the hydrogen yield

in Figure 5.32 shows larger values with S/C at a given φ. Once again, the yield for the high S/C cases is larger due to the production of hydrogen in the post injection zone for all S/C.

The hydrogen yield deﬁned on fuel-steam basis shown in Figure 5.33 shows an in-

crease for the high S/C when comparing them with the correspondent main injection

case in Figure 5.17. For the low temperature this is explained because the H2 concen-

tration is produced in the POx process, while for the high temperature cases H2 is also

produced through the post-injection section.

The CO concentrations, illustrated in ﬁgure 5.34, present an increasing behavior

with equivalence ratio as expected. CO is seen to decrease with S/C, however, the

204 eﬀect is less pronounced for the post-injection cases than that for the main injection case. In addition, the high post-injection temperature case shows a slightly higher CO concentration than the low temperature case for large φ due to the higher temperature.

This can be explained by means of the water gas shift reaction, since CO2 decreases farther with increasing S/C for a given φ. The CO2 concentration is depicted in Figure

5.35. The convergence φ for the high temperature cases occurs at about 1.6, due to the diﬀerent decreasing rates seen with S/C. This does not occur to the low temperature cases within the equivalence ratio range covered in this thesis. This is due to the fact that the CO trend is fairly equally spaced with S/C through the equivalence ratio φ domain.

Finally, the total hydrocarbon concentrations outlined in Figure 5.36, are seen to be smaller for the post-injection cases. This can be explained by the fact that the post-injection period is still at high temperature that some small hydrocarbons can continue to react to produce CO and water, or recombine into stable hydrocarbons.

205 Figure 5.31. H2 Concentration for Hexadecane at Diﬀerent φ and Post-Injection S/C Ratios at 1 atm and TPI = 2000 K

Figure 5.32. H2 Yield for Hexadecane at Diﬀerent φ and Post-Injection S/C Ratios at 1 atm and TPI = 2000 K

206 Figure 5.33. H2 Yield for Hexadecane at Diﬀerent φ and Post-Injection S/C Ratios at 1 atm and TPI = 2000 K based on Fuel and Vapor

Figure 5.34. CO Concentration for Hexadecane at Diﬀerent φ and Post-Injection S/C Ratios at 1 atm and TPI = 2000 K

207 Figure 5.35. CO2 Concentration for Hexadecane at Diﬀerent φ and Post-Injection S/C Ratios at 1 atm and TPI = 2000 K

Figure 5.36. THC Concentration for Hexadecane at Diﬀerent φ and Post-Injection S/C Ratios at 1 atm and TPI = 2000 K

208 5.7 Eﬀect of Exhaust as an Oxygen Provider

The analysis carried out in the previous sections employed fresh air as the oxygen supplier. For automotive applications, this can be done using an electrical or mechanical pump to provide the air for the reformer. Nevertheless, diesel engines are known to run under lean conditions suggesting therefore the attractive solution of using exhaust gas as the oxygen source, reducing the energy cost that a an electrical pump would introduce into the picture. Based upon this, ﬁve diﬀerent set points of exhaust gas are utilized in the numerical study to assess the feasibility of this possible solution within the equivalence ratio selected for the previous cases. The concentrations of the major gases contained in the exhaust, CO2,H2O, O2 and N2, are obtained from an

1.9L Euro4 Engine at 2000 rpm, illustrated in Table 3.2. The methodology employed to calculate the feeds for the POx reformer was introduced in Section 3.2.4.

The mixture of exhaust gas and fuel are allowed to ignite over a period of 1 s owing to the fact that the cases involving a low equivalence ratio and low oxygen concentration

(O2 > 8.76 showed a very negligible rise in temperature (less than 100 K) during the

ﬁrst 0.5 s. It is noteworthy that the low temperature rise is present for all equivalence ratios when using a very small oxygen concentration, as shown in Figure 5.37. This can also be seen in Figure 5.38, which outlines the temperature history for hexadecane for a φ = 1.6 at 1 atm and 1000 K case for the diﬀerent oxygen concentrations employed in this study. Although the necessary equivalence ratio to produce combustion is present, the exotherm generated through the combustion with such a small amount of oxygen is not capable of heating the remaining gas (excess hexadecane, CO2, etc.), therefore not allowing other important reactions to occur. This can be explained by the fact

209 that the amount of diluent (N2, CO2 and H2O) present in the initial gas increases as the oxygen concentration is decreased, which results in a initial mixture with a larger heat capacity.

Figure 5.37. Temperature History for Hexadecane at Diﬀerent O2% Exhaust for φ = 1.0 at 1 atm and Tini = 1000 K with no Cooling

For the numerical analysis, the sudden increase in temperature and concentrations is required to account for ignition. As can be seen in Figures 5.37 and 5.38,O2 <

8.76% present a small heat release, low temperature gradients and no signiﬁcant hydrogen production. Based upon these facts, only the largest oxygen concentrations are further analyzed, i.e., cases are cooled down to 500 K after they reach their maximum temperature for comparison with the results from previous sections.

210 Figure 5.38. Temperature History for Hexadecane at Diﬀerent O2% Exhaust for φ = 1.6 at 1 atm and Tini = 1000 K with no Cooling

Based upon the little exotherm presented for this cases, implementation of an actual reformer becomes then more complicated due to mixing issues, turbulence, etc.

Studies on ﬂame propagation on methane and diesel fuel, Glassman [101] and Follen

[13], have shown that the amount of diluent present in the exhaust gas limits the ﬂame propagation. Glassman [101] showed that a mixture of CH4 in mixtures of CO2 or N2 required at least an oxygen concentration of 14.6% and 12.1% respectively. Follen [13] reported that for a gaseous mixture of diesel fuel and exhaust gas, the ﬂame can only propagate for O2 concentrations greater than 13.5%.

Hydrogen concentration in Figure 5.39, increases as the oxygen concentration because of the increase in temperature (greater heat release) observed in Figures 5.37 and

5.38. As for the previous cases, H2 concentration increases with increasing φ. It is

211 noticeable that the hydrogen percent for the two low O2 concentrations is lower than

for the larger O2 concentration. This can be explained by the fact that the amount of

diluent increases as the oxygen concentration is decreased. This eﬀect is clariﬁed by

the hydrogen yield, Figure 5.40, which is deﬁned only in terms of hydrogen and fuel.

The hydrogen yield shows an increasing trend as equivalence ratio increases. The hydrogen yield is very similar for these oxygen concentrations; however, they are much smaller than the hydrogen reported for fresh air at a given φ ratio.

CO concentration follows the deﬁned increasing trend with φ presented for the fresh

air cases, however, its is lower than that for the fresh air cases due to the lower temper-

ature achieved during combustion. As expected, the CO concentration for the lower

exhaust oxygen content is smaller than that for O2 = 15.13%. Conversely, CO2 con-

centration decreases as the equivalence ratio is increased and the concentration for O2

= 15.13% is lower than for smaller O2 concentrations. Furthermore, the concentration

of CO2 is larger than that produced by the air fresh cases. Finally, the total amount

of hydrocarbon (C1 basis) is lower as the oxygen concentration is increased due to the

higher temperature achieved during combustion. THC concentration is slightly larger

for exhaust cases than that for fresh air due to the temperature diﬀerence.

212 Figure 5.39. H2 Concentration for Hexadecane at Diﬀerent O2% Exhaust at 1 atm and Tini = 1000 K with Cooling

Figure 5.40. H2 Yield for Hexadecane at Diﬀerent O2% Exhaust at 1 atm and Tini = 1000 K with Cooling

213 Figure 5.41. CO Concentration for Hexadecane at Diﬀerent O2% Exhaust at 1 atm and Tini = 1000 K with Cooling

Figure 5.42. CO2 Concentration for Hexadecane at Diﬀerent O2% Exhaust at 1 atm and Tini = 1000 K with Cooling

214 Figure 5.43. THC Concentration for Hexadecane at Diﬀerent O2% Exhaust at 1 atm and Tini = 1000 K with Cooling

5.8 Concluding Remarks

The present section presents a brief summary of the results presented in this chapter and compares them to illustrate the advantages and disadvantages of each characterizing parameter over the other parameters. An increasing equivalence ratio enhances

H2, CO and THC concentrations for all temperatures and oxygen providers studied in this chapter. The hydrogen yield follows the trend deﬁned by the hydrogen concentration, increasing with a saturation behavior as φ is increased due to the lower adiabatic temperatures. It can be concluded that the trends deﬁned by the kinetics code are in good agreement with the predicted equilibrium concentrations performed using the

Matlab simulator from [97] with the addition of the fuel properties.

215 Addition of steam within 0.25 ≤ S/C ≤ 1 at 1000 K results in a very slight increase

( 0.3% Vol) in hydrogen concentration for φ < 1.6. Increasing further the equivalence ¬ ratio and/or the S/C ratio decreases the hydrogen concentration. The hydrogen yield deﬁned in terms on fuel as the only hydrogen provider also presents a slight increase within the same S/C range and decreases as S/C and φ over pass those limits. CO

concentration decays with increasing S/C, while the CO2 depleting slope with φ decreases as S/C is increased, forcing the CO2 concentration to be higher as S/C increases after φ = 1.39. THC concentration is slightly larger for steam cases than that for the case with no additional steam. This is explained by the lower adiabatic temperature achieved as the S/C increases. A dry-N2-free hydrogen concentration was introduced in Sections 5.2 and 5.3 for fair comparison with catalytic concentrations ( 70) found ¬ in the technical literature. A new deﬁnition of hydrogen yield is also introduced to

account for water vapor as a hydrogen provider. This new yield shows lower values as

the S/C is increased due to the lower adiabatic temperature as the S/C is increased. It

is noteworthy that water is not included in the deﬁnition of fuel penalty for automotive

applications. Furthermore, water can be recovered from the exhaust gas.

The temperature of the feeds plays an crucial role in the reformation of hydrocarbons

as was concluded in the previous sections. Administering the reactants at high tem-

peratures causes the combustion to be immediate and generate larger concentrations

of H2 and CO, while reducing those for CO2 and THC. This is explained by the fact

that higher initial temperatures induce higher adiabatic ﬂame temperatures, which en-

hances the conversion of hydrocarbons towards CO and H2. For these cases, addition

of water results in a larger contribution for hydrogen generation, which is quantiﬁed by

the hydrogen yield deﬁned for fuel as the only H2 provider. It is noteworthy that the

216 operation limits for the φ and S/C ratios can be increased as the initial temperature is

increased due to the increase in the adiabatic ﬂame temperature for those cases. The

hydrogen percent deﬁned in a dry-N2-free environment presents an increasing hydrogen

concentration with φ and S/C, (e.g., φ = 1.9, S/C = 2 and Tini = 1750 K results in H2

= 39%). The hydrogen yield deﬁned in terms of water vapor and fuel shows a similar

behavior for the three diﬀerent temperature, however, it increases with temperature for

a given S/C and φ ratios.

Although analysis of the pressure eﬀect on the hydrogen concentration was postu-

lated on the scope of this dissertation. Unfortunately, the set of kinetic rate coeﬃ-

cients is calculated for atmospheric pressure. Based upon this, the adiabatic ﬂame

temperature using the equilibrium code from [97] was calculated and compared with

that calculated by the kinetics code. According to the equilibrium temperature cal-

culations and Glassman [101], the temperature should increase with pressure for low equivalence ratios, while it does not present a signiﬁcant eﬀect for high equivalence ratios. The kinetics code show an opposite behavior for the low φ ratios, while the

temperature increases for the high φ ratios. Therefore, no further analysis on the concentrations is performed.

As temperature plays an important role in the reformation process and injection of vapor as part of the initial feeds is seen to decrease the adiabatic temperature of the process, secondary injection is studied to take the high temperature generated in the combustion of fuel/air mixtures. Two diﬀerent post-injection temperatures were introduced showing that the lower secondary injection temperature case does not increase the hydrogen percent already generated in the combustion of a rich mixture.

217 Instead, the concentration is seen to remain constant through the entire post-injection

process. The highest post-injection temperature show an increasing hydrogen percent

as time evolved resulting in a slightly larger value for selected cases (φ = 1.6). The

advantage of using the heat generated for the combustion is clearly seen for the high

S/C ratios, which show a much slower saturation behavior than that for main injection

of water. This causes the hydrogen yield (based on fuel) to be larger for all S/C ratios

at a given φ ratio, contrary to the main injection case at high φ ratios. The hydrogen

yield in terms of water vapor and fuel present a similar trend for both cases, main and

secondary injection. However, the numerical values are slightly larger for the post-

injection case. The CO concentration is seen to increase for the post-injection case,

while the CO2 concentration decreases. Furthermore, THC percent is slightly lower for the high-temperature post-injection case. Two main diﬀerences in the imposed heat transfer must be considered when comparing these cases, the main and secondary injections. The main injection case presents a linear cooling after the combustion, while the secondary injection case has a rapid cooling to the P.I. temperature, then the temperature slightly varies in the P.I. section (i.e., fairly constant through the process

0.4 s) and then second cooling to 500 K. This could raise the question about fair ¬ comparison between these cases. Nevertheless, the concentration reported for main injection at 2000 K, see Figure 4.25,( 0.285 s) is approximately 4.1% (close to the ¬ ¬ ﬁnal value). At that same time location, see Figure 4.25, the post-injection hydrogen

concentration is about 4%. If cooling were performed thereafter, the H2 concentration

would have continue to increase (as shown for the main injection).

Finally, studies on the reformation of hexadecane using exhaust as a the oxygen

provider were presented in Section 5.7. It is seen that the high content of diluting

218 species in the exhaust limits the amount of energy release from the exothermic oxidation reaction. This is observed for all the φ and oxygen concentrations, being more noticeable for the very lower O2 concentrations for which the maximum temperature show a negligible increase of less than 100 K. For the cases that presented ignition, the hydrogen and CO concentrations show a similar behavior than those for fresh air with increasing equivalence ratio; however, with lower numerical values due to the lower temperature. Furthermore, the oxygen concentration drives the available amount of energy CO2 thus presents slightly higher values as does the THC concentration. Al- though the hydrogen yields show an increasing trend with φ, their numerical values are lower than those reported by the correspondent fresh air case.

219 CHAPTER 6

EXPERIMENTAL SET UP AND TESTING PROCEDURES

6.1 Introduction

The proposed numerical model described in Chapter 3, although based on the reaction mechanisms provided by Ristori et al.[88], requires further validation since only a reduced reaction set is considered for the prediction of the main combustion products, including several hydrocarbons and heat release. For this reason a non-catalytic experimental reformer for diesel and hexadecane fuel, based on that developed at the

Ohio State University by Midlam-Mohler [10], is employed. Due to the importance of using alternative fuels to power up vehicles, bio-diesel (B-100) is also tested for the production of hydrogen and carbon monoxide in this work. It is noteworthy that no experimental work on bio-diesel has been found in published literature and therefore this will be a signiﬁcant breakthrough. Table 6.1 illustrates the three fuels used in this research as well as some characteristics.

The chemical composition of the bio-diesel and diesel fuels used in this experiment is obtained by mass spectrometry (MS) and gas chromatography (GC) at the Mass

Spectrometry & Proteomics (CCIC - Campus Chemical Instrument Center) of the

220 Fuel Denomination Chemical Formula H/C Ratio A/Fst Diesel BP-ECD1 CnH1.8n 1.8 14.5 Hexadecane Acros-Organics C16H34 2.125 14.96 Bio-diesel B-100 C19H36O2 1.89 14.24

Table 6.1. Fuels, Denomination and Hydrogen to Carbon Ratio

Ohio State University. Gas chromatography is basically a separation process which is achieved by distributing the substances to be separated between two phases, a moving phase and a stationary phase. The mobile phase could be either liquid or gas. If liquid is used the principle is called LC. On the contrary, when the mobile phase is a gas, the principle is called GC. The stationary phase could be either liquid or solid.

Hence, there are 4 sub-groups depending on the conditions of both phases. The sample is injected in the GC device where the sample vaporizes and separates into the various components which produce speciﬁc spectral peaks that are proportional to the quantity of the corresponding substances in the sample.

The mass spectrometry identiﬁes the substances by ionization (i.e., electrons are

separated by impact to give a positive ion), breaking the molecules into charge particles

which can be deﬂected by a magnetic ﬁeld according to their masses. The ions that

are not deflected pass through an amplifier. The size of the magnetic field employed to

bring each ion onto the detector is used to identify the mass of each ion being detected.

Hence, the compound can be identiﬁed by the GC retention time (injection-peak), the

parent ion and fragmentation pattern as searched by a database of known compounds

[89, 90, 91]. The retention time can help to diﬀerentiate between some compounds.

221 The chemical composition of bio-diesel and diesel fuel are reported in the Appendix

B. Hexadecane (ACROS Organics - [544-76-3]) is 99% pure, and therefore no mass spectrography analysis is required. Although the primary objective of this research is the production of hydrogen, the hydrocarbon composition leaving the reformer after the combustion is important since they can be used for reduction of NOx on HC-SCR or for regeneration of LNTs or DPF systems. Hence, speciation via gas chromatography of the reformate is performed for the three fuels in order to determine the ﬁnal hydrocarbon composition. This again represents a signiﬁcant contribution to the reforming of hydrocarbons by using bio-diesel as a hydrogen generator.

6.2 Experimental Apparatus & Reformer Designs

Figure 6.1 illustrates a diagram of the reforming-dilution system employed in this study. The liquid fuel enters the atomizer tubes via a positive displacement pump

(FMI Q) which is previously calibrated to account for approximated mass ﬂow rates.

Nevertheless, the fuel recipient is set on a scale and weighted before the test is initiated.

The mass of the recipient is recorded every minute for the duration of the test. This data is then collected and averaged to obtain an accurate fuel mass flow rate for each experimental case. The atomizers are heated up to approximately 400 to ensure total vaporization of the fuel. The vaporized fuel is then atomized axially through the atomizers in an extremely fine and uniform droplet form. Compressed air is injected tangentially, and in counterflow arrangement to the fuel atomizers, through four ports on the outer wall of the reformer at the same axial location, and metered using a mass

ﬂow controller (Omega FMA 5400/5500). This conﬁguration enables the administered fuel droplets and air to be well mixed before being ignited.

222 Figure 6.1. Schematic of the Reforming and Dilution System

Three versions of the TPOx reactor have been designed, built and tested to study the behavior of the partial oxidation process. The ﬁrst two versions are illustrated in

Figures 6.2 and 6.3. The first version consists of a 24 in. long, 2.5 in. OD stainless steel tube with three distinct sections. The first zone corresponds to the injection section which has three heated atomizers (AC heaters and capillary tubes) and three counterflow air tubes. The second section refers to the ignition section where a resistive heating element (Bosch glow plug 0250202 002) is located 2.5 in. downstream of the injection zone. This device provides high enough temperatures ( 800 ) to ¬ initiate the combustion of the mixture of fuel and air. Finally, the remainder of the tube is the reaction chamber located immediately after the ignition zone; wherein the combustion products are allowed to further react and attain equilibrium. As mentioned in Chapter 2, temperature plays a very important role in the reformation of

223 hydrocarbons; therefore, the entire reformer is insulated with high temperature insulation (Zircar Ceramics Inc. ASB-2600) so as to attain the highest possible temperature.

For this reactor type, thermocouples (type K) are located along the axis of the tube, 2 inches apart, to monitor the temperature along the centerline of the reformer as shown in Figure 6.2.

Figure 6.2. Thermal Partial Oxidation Reformer Version 1

The second version of the TPOx conserves the zone distribution and the counter-

ﬂow mixing. However, the injection system is constituted by four DC-heater(glow plug)-capillary tube systems instead. The reforming system is enhanced by using a concentric chamber, 4 in. OD, that encloses the reactor in order to transfer part of the heat released by the combustion to the incoming air. The igniter (GP) is replaced by a Chentronics exciter, which releases a spark to reliably initiate the combustion, due to the fact that long exposure to high temperatures damages the glow plug electrical resistance. The new igniter is tested under high soot formation, large equivalence ratios, and high temperature conditions showing a remarkably good performance at all times.

224 For this reactor type, only two thermocouples (type K) are used to monitor the axial thermal behavior inside the reformer. The ﬁrst thermocouple is located 4 in. downstream of the igniter, whereas the second thermocouple is 2.5 in. further downstream.

Figure 6.3. Thermal Partial Oxidation Reformer Version 2

The third version of the reformer captures the advantages of both reactors. The counterflow mixing concept is utilized as well as the injection system of the second reactor. The concentric arrangement employed for version 2 indeed increased the temperature in the reformer; however, the elevated temperatures and the continuous use of the reformer deformed and fractured the inner pipe. The flame is observed to move backwards from the igniter location to the air inlet as the inner temperature and the wall temperature increase with time exposure. A screen (conical reduction to acceler- ate the flow from the mixing chamber to the ignition region) was allocated prior to the igniter to anchor the flame for the second and third reactor versions. Nevertheless, the flame moved back and fractured the reformer. Emissions data was recorded for the second version of the reformer but could not be fairly compared with that obtained

225 for other reformers since it is unknown when the incoming air stopped being injected through the air inlets and started to filtered through the fracture which would definitely decrease the homogeneity of the mixing. The third version does not use the concentric pipe, instead it only uses the inner pipe, similar to the first reactor. The reformer is insulated downstream of the igniter and the anchorage screen is used prior to the igniter. Both concepts caused the flame to anchor right after the igniter as shown in

Figure 6.4 for all the cases performed in the reformer.

Figure 6.4. Thermal Partial Oxidation Reformer Version 3

226 6.2.1 Dilution and Measurement Systems

A gas sampling port, for all the reactors, is located beyond the outlet of the re-

former. The sample gas is cooled to 191 and diluted with molecular nitrogen (ratio of 1:4) before being analyzed by a Horiba exhaust gas analyzer (MEXA-7500DEGR) and a hydrogen analyzer (H-Sense). The sample is diluted in order to bring the gas concentrations within the range of measurement of the analyzers and to avoid condensation of hydrocarbons and water vapor. Hence, a sample of the reformate is extracted through a heated ﬁlter (Horiba DF-03) by a heated diaphragm pump (Horiba HTP306).

The gas ﬂow is restricted by a needle heated valve located upstream of the pump so as to satisfy the needed dilution ratio. The sample gas enters a heated mixing chamber downstream of the pump where it is mixed with molecular nitrogen heated to 191 .

The nitrogen ﬂow is metered by a mass ﬂow controller (Omega FMA 5400/5500). Part of the diluted gas is sent to the analyzers, while the rest is released into the exhaust.

The accurate value of the dilution ratio is determined by running ambient air through the reformer-dilution system. Hence, the dilution factor is directly induced by the reduction of the oxygen concentration to approximately 5%. The dilution factor is recorded prior and after every test in order to ensure that the reformate ﬂow through the sampling lines is fairly constant and that no signiﬁcant restriction due to soot is encountered. Then, both values have to show a good agreement within 10% in order for the results to be considered.

The Horiba and hydrogen analyzers are calibrated to measure the major species,

CO, CO2,H2,O2, NOx, as well as the total amount of hydrocarbons. The THC are

227 measured by using a hydrogen ﬂame ionization detection mechanism that uses the phe-

nomenon in which ions, generated by the heat energy (dissociation) when hydrocarbons

are introduced into a hydrogen ﬂame, migrate to the electrodes located on the side of

the ﬂame generating a current which is proportionally correlated to the number of

carbon atoms. The amount of hydrocarbons is then reported by the analyzers as the

conversion of all hydrocarbons present in the exhaust to C1 species. Water vapor is not directly measured but its concentration is inferred approximately from the measured, CO, CO2 and H2 concentrations by assuming that the water gas shift reaction,

Equation (6.1), attains equilibrium.

CO + H2O −→ H2 + CO2 (6.1)

KNCO2 NH2 NH2O = (6.2) NCO

where K is the dimensionless equilibrium constant with a value of 3.5 [92] and N is

the number of moles of each species. The measured concentrations, as well as the

axial temperatures, are recorded by a data acquisition system, Labview. The data

collected by the analyzers has to be post-processed because the values for hydrogen,

hydrocarbons and NOx are in wet conditions, while carbon oxides and oxygen are in dry

conditions. Therefore, the dry values must be converted to wet values and then to raw

values (undiluted) in the following manner. A dry mole fraction for any compound

would have the form,

X Y = i,wd (6.3) i,dd N−1 X Xj,wd j=1

228 where Y and X represent the mole fraction and the concentration of species i, the subscript dd and wd illustrate dry diluted and wet diluted conditions, respectively,

N is all measured species, and N-1 means water is not included. Hence, a wet mole fraction for any species has a similar form but includes water in the denominator

X Y = i,wd (6.4) i,wd N X Xj,wd j=1

Hence, diving Equation (6.3) by Equation (6.4), the wet diluted mol fraction of species i can be calculated as         Y  1  Y = i,dd = Y   (6.5) i,wd N i,dd  X  X 1 + H2O,wd  X  N−1  j,wd  X  j=1  Xj,wd  N−1 X j=1 Xj,wd j=1

  1 Y = Y   = Y [1 − Y ] (6.6) i,wd i,dd  Y  i,dd H2O,wd 1 + H2O,wd  1 − YH2O,wd

Finally, the wet diluted mole fractions are multiplied by the dilution factor in order to obtain the raw mole fraction.

229 6.3 Testing Procedure

A normal test is initiated by setting the air mass flow controller to a fixed value of 100 l/min. This mass flow rate is selected based on a set of experiments carried out to select the optimal value to achieve high temperatures. It was seen that by increasing the mass flow rate from 60 to 100 l/min, the temperature recorded at the

first thermocouple location increased about 200 K. This may be due to lengthening of the flame front and its approach toward the location of the thermocouple. Also, higher mass flow rates increase the thermal energy released per unit time in the reaction zone, so that the maximum flame temperature is expected to increase with increasing

ﬂow rate.

The dilution system (including pump, heaters and molecular nitrogen ﬂow (14.9 l/min)) is started after the volumetric ﬂow rate of air is adjusted to 100 l/min. The dilution factor is attained by the needle valve that controls the amount of air entering the dilution system coming from the reformer. The atomizer heaters are then set to 400 , while the ignition Chentronics system is energized using a power supply.

When these conditions are achieved, the positive displacement pump is turned on and the change in weight of the fuel recipient located on the scale is recorded every 30 to 1 minutes as fuel is injected into the reformer. The instantaneous air mass flow rate is recorded and averaged to obtain an average mass flow rate. The chemical composition of the exhaust is recorded and the final concentrations of the constituents are calculated in Matlab using the dilution ratio previously mentioned. Although the transients of the experiments are recorded to account for the dilution factor, the concentrations reported in this thesis are those at steady state or chemical equilibrium.

230 The experimental measurements reported here span a range of equivalence ratios

from 1.0 ≤ φ ≤ 2.2, which is similar to that deﬁned for the numerical simulation

1.0 ≤ φ ≤ 1.9. The upper limit of φ is selected so as to avoid problems that arise due to excessive soot formation that interferes with the functioning of the analyzers to measure the concentrations of the various species. The equivalence ratio is deﬁned on a mass ﬂow basis as

(m ˙ /m˙ ) φ = fuel air actual (6.7) (m ˙ fuel/m˙ air)stoic

wherem ˙ is the mass flow rate, the subscript actual refers to the mass flow rates measured at the syringe pump and by the mass flow controller, and the subscript stoic refers to the stoichiometric value of the fuel to air ratio, which depends on the fuel used. The stoichiometric air/fuel ratios for the fuels used in this thesis are reported in Table 6.1. Once the fuel and air flow rates are set to the appropriate amounts that

yield φ = 1, the mass ﬂow rate of the fuel is varied to attain the desired equivalence ra-

tio. Special attention has to be taken when calculating the stoichiometric equivalence

ratio for bio-diesel since it contains large concentrations of oxygenated hydrocarbons.

The molecule C19H36O2 is used in this study as the representative molecule for bio-

diesel. Hence the amount of oxygen contained in the fuel has to be considered part

of the oxidizer instead of part of the fuel. The stoichiometric equivalence ratio can be

calculated from a molar balance of Equation (6.8).

aC19H36 + aO2 + b(O2 + 3.7755N2) −→ cH2O + dCO2 + 3.7755eN2 (6.8)

where the coeﬃcients can be determined in terms of the amount of fuel a as: b = 27

a, c = 18 a, d = 19 a and c = 27 a. From the deﬁnition of air to fuel ratio using

231 oxidizer instead of air and the coeﬃcients previously deﬁned, the stoichiometric air to

fuel ratio can be found as

Oxidizer 28M + 3.7755(27)M = O2 N2 = 14.24057 (6.9) F uel stoic 19MC + 36MH

The mole fraction of the molecular oxygen contained in the gas can be inferred

from the coeﬃcients as yO2,f = nO2,f /nf = 1/56. As the actual masses of fuel and air are used in the experiment, the oxidizer to fuel ratio is also aﬀected by the oxygen contained in the fuel. Hence, that amount has to be subtracted from the fuel and added to the oxidizer side as M m˙ 1 − y O2 Oxidizer m˙ − m˙ f O2,f M = f O2,f = f (6.10) M F uel actual m˙ air +m ˙ O2,f O2 m˙ air + yO2,f m˙ f Mf

From Equation (6.10) it can be seen that the amount of oxygen contained in the

gas is small and can be neglected for the calculation of the actual oxidizer to fuel ratio.

An illustration of a typical measurement is shown in Figure 6.5, which displays the

temperature recorded by thermocouple T1 and the diluted concentrations of the major

species in percent of the total gas volume as recorded by the analyzers. Ignition occurs

approximately 10 seconds after the resistive ignition source is energized, after which the

temperature at station 1 rises sharply and proceeds to slowly approach a steady state.

Once steady state is reached, after variations on concentration are smaller than 5%,

the measured values are recorded and compared with those obtained by the numerical

calculation. Approximately 50 seconds after ignition all concentrations are found to

change by less than 5-10%, so that measurements after this time are taken to be the

steady state values. Thermal steady state requires much longer times on the order

232 of several minutes, so that steady state is taken here to be the point at which the

temperature increase is less than 35 over a period of 1 minute. At this point, the concentrations are already steady and vary by less than 5%.

All diluted concentrations are recorded and converted to wet non-diluted concentrations, as shown in Figure 6.6, following the procedure mentioned above. These

measurements are compared with the predictions of a 0D chemical kinetics model in

the following chapters. Comparisons between the chemical kinetics model described in

this section and the experimental data obtained require elaboration. The experiment

comprises an inhomogeneous, turbulent ﬂow with heat transfer whereas the 0D model

considers a perfectly mixed, homogeneous mixture, with no spatial variations. Thus,

the only common platform for comparison of course, is the initial equivalence ratio.

Beyond this parameter, no other parameters are comparable between the model and

experiment.

233 Figure 6.5. Diluted Concentrations and Temperature History with Time [s] for Hexade- cane at φ = 1.5

Figure 6.6. Non-Diluted Wet Concentrations and Temperature History with Time [s] for Hexadecane at φ = 1.5

234 6.4 Experimental Hydrogen Yield Deﬁnition

A previous deﬁnition has been given in the previous chapter for the numerical

calculation of the hydrogen yield. For the experimental results, the hydrogen yield,

based on the same concept as that for the numerical results, is calculated in diﬀerent

form depending on the fuel. For instance for diesel fuels, once the actual value of

hydrogen is calculated, the hydrogen yield is obtained as

(m ˙ H2 /m˙ f )actual ΨH2 = (6.11) (m ˙ H2 /m˙ f )ideal where the subscript actual represents the experimental value of H2 measured (after converted to mass of H2) and the mass ﬂow rate of fuel at the inlet. The subscript ideal represents the amount of H2 (gram) that can be obtained from a mol of fuel

(hydrocarbon CxHy) divided by the amount of fuel (gram) in one mol of fuel. This expression can be recast in terms of the H/C ratio of the fuel deﬁned as Y = y/x. The nominal value of the H/C ratio for light diesel is 1.8 according to Heywood [92], while

it is 2.125 for hexadecane (C16H34) and 1.8947 for bio-diesel B-100.

(m ˙ /m˙ ) Ψ = H2 f actual (6.12) H2 2 1/ 1 + MC MH2 Y ideal

m˙ 2 M = H2 1 + C (6.13) m˙ f actual Y MH2 ideal

235 6.5 Speciation for Hydrocarbon Composition

The experimental data collected by the Horiba and hydrogen measurement systems

allow comparison of the major products of the combustion between experimental and

numerical results. However, special interest exists in ﬁnding the hydrocarbon compo-

sition of the reformate due to the fact that not only hydrogen can be used as a reductant

for nitrogen oxides, but also diﬀerent hydrocarbons (e.g., ethylene) as well as CO. As

explained in Chapter 2, also LNT and Urea systems can improve their eﬃciency or recover/regenerate catalytic activity by using hydrogen. Therefore, speciation of the reformate from diesel, hexadecane and bio-diesel are included in this thesis.

The reformate speciation is performed at the Transportation Research Center (TRC) that counts with a Gas Chromatogram Detector-Flame Ionization (GC-FID, Varian

3600). This chromatogram has two columns (J&W Scientiﬁc 123-1063, 60m × 0.32mm

Id; DB1 phase; 1.0 micron ﬁlm thickness) and a two-head pump (Baldor MP-602 - Metal

Bellows Pump).

The speciation requires a 30:1 N2:exhaust ratio. The set up to collect the sample from the reformer and further diluted is illustrated in Figure 6.7. The original dilution

system is used to calculate the 4:1 N2:exhaust ratio to ensure that the reformer is

producing the major concentrations as expected (i.e., no speciation). Once hydrogen

reaches the expected concentration and steady state of all major species (as well as

temperature) is achieved, the valve that directs some of this mixture to the Horiba

analyzer is shut oﬀ and then the mixture is directed through a non-heated ﬁlter (Horiba

DF-03) for further cleaning of solid carbonaceous species to the chamber where the 4:1

clean-diluted mixture mixes with the remaining amount of nitrogen needed to achieve

236 the desired 30:1 ratio. Due to the restriction of the ﬁlter, the 4:1 clean-mixture is

allowed to pass through the mixing chamber for about 30 seconds. After this, the

nitrogen valve is opened. The new mixture is released to the environment for about

20 seconds to assure that the ﬂow going to the bags contains the 4:1 diluted sample

mixed with the second amount of nitrogen. Thereafter, the bag valve is opened while

the exhaust valve is partially closed to allow more ﬂow be driven to a black tedlar

sample bag.

The black tedlar sample bag with a shutoﬀ valve is delivered to TRC facilities for

speciation. The sample is connected to the GC-FID through a 2-head metal bellows

pump and pumped to the sample loop at the head of both columns. The ”A” side

(small Loop), receives 1/10 of the sample that the “B” side (large loop) receives. ¬ The component values from Methane to Propyne are taken from the “A” side because

the column cannot separate these components at the higher concentration. All com-

ponent values from “2-Methylpropane” on, are taken from the large loop side due to the

increased sensitivity. The software on the GC-FID is the Varion GC Star Workstation

- Software Version A2.

The GC Oven program takes the sample from -60 to 237 . The carrier gas is helium. The individual components are printed out as peaks on a baseline. The peak areas of the sample chromatogram are converted to ppmC (ppmCarbon) using coeﬃcients derived from a CRC #4 standard mix. These values can then be converted to ppm of each component by dividing by the number of carbons in the component, and to mass in the sample as the dilution of the sample is known. A summary of the

237 results of the GC-FID analysis is presented in Chapter 7, while the complete table of results for the reformate species is given in Appendix C.

Figure 6.7. CO2 Concentration for Hexadecane at Diﬀerent φ

6.6 Concluding Remarks

The experimental set up, the diﬀerent designs of the partial oxidation reformer and the procedure employed to validate the numerical results presented in the previous chapters is introduced in this chapter. GC-MS speciation of the liquid diesel and bio-diesel used in the experimentation is carried out to identify the main hydrocarbon composition of these fuels. Diesel fuel presents a very wide range of hydrocarbons,

238 while bio-diesel is mainly constituted by four diﬀerent oxygenated hydrocarbons. An deﬁnition of the hydrogen yield in terms of the hydrogen to carbon ratio that char- acterizes each fuel (diesel, hexadecane and bio-diesel) is introduced in this chapter to allow comparisons between the experimental results and those obtained numerically at equilibrium. Finally, the procedure and set up used to obtain a GC-FID speciation of the hydrocarbons contained in the reformate gas is also presented in this chapter.

239 CHAPTER 7

EXPERIMENTAL RESULTS

7.1 Introduction

Three diﬀerent designs of the partial oxidation reformer were described in the previous chapter. The three versions were tested following the methodology described in the previous chapter as well. The ﬁrst version, a single pipe with three injectors was used as a prototype to assess the feasibility of the concept. This design indicated that high temperatures can be achieved. Furthermore, hydrogen yields similar to those predicted from the kinetics code are observed. Nevertheless, the temperature of the

TPOx process as emphasized in Chapter 5 plays a very important role.

Based upon that conclusion, the second version of the reformer comprised concentric

ignition and reaction chambers. The inner pipe serves the same purpose as for the

ﬁrst version, while the channel formed between the two pipes utilized the heat transfer

from the combustion reaction to heat up the incoming air. This version achieved

high temperatures faster than the ﬁrst prototype. Unfortunately, the ﬂame always

anchored at the air injection ports. Hence, the diﬀerence in temperature between the

incoming air and the ﬂame temperature caused large thermal stresses in the inner pipe

240 which ﬁnally resulted in fracture. Furthermore, the atomizers (capillary tubing) suﬀer

of coking problems due to the near location of the ﬂame, which elevated the injectors

temperature to about 850 .

A combined version that used the single pipe reactor concept and the injection hard-

ware of the second version was employed in the third prototype. High temperature

insulation was utilized downstream of the igniter to avoid high temperatures in the

mixing temperatures. Moreover, a ﬂame screen/holder located right after the igniter

was implemented with successful results. The average temperature observed in the

mixing chamber was about 700-750 , which agrees with the injection temperature

used in the kinetics code. Three diﬀerent fuels (hexadecane, diesel and bio-diesel) are

reformed to produce hydrogen using this reactor for partial oxidation over the equiv-

alence ratio used for the kinetics model. It is noteworthy that no injection of water

(main or secondary) is implemented in this reactor due to the low enhancement showed

in the 0D model.

As mentioned in Chapter 6, liquid speciation using a GC-MS was performed to identify the main constituents of diesel and bio-diesel fuels. Hexadecane is a pure compound, therefore no further liquid analysis is required. A brief explanation is given in Section 7.2. The major species compositions recorded with the Horiba and

hydrogen analyzers for the three fuels used in this dissertation are discussed in Section

7.3. As presented in Chapter 3, the model is formulated under adiabatic and perfectly

homogeneous assumptions, and models the reforming process versus time (0D model).

The actual reformer involves a 3D, inhomogeneous and non-adiabatic environment.

These diﬀerences allow the equivalence ratio to be the only metric to compare the

241 kinetic modeling results with the experiments. Thus, these compositions are compared

to the predicted trends obtained from the kinetics model using the equivalence ratio.

Finally, a brief discussion on the hydrocarbon compositions predicted by the kinetics

code and those obtained for the experiments is given in Section 7.4.

7.2 Liquid Speciation

As mentioned in the previous chapter, GC-MS speciation of the diesel and bio-

diesel fuels is carried out to determine the initial set of compounds (hydrocarbons)

that constitute each fuel. Figures B.1 and B.2 present the spectrogram in terms of relative abundance of diesel and bio-diesel respectively. Abundance is understood as a quantiﬁcation based on the hydrocarbon with the largest composition. As expected, a wide variety of hydrocarbons is contained in diesel fuel; however, identiﬁcation of all the hydrocarbons requires a very extensive database search as well as more specialized

GC equipment, see [103], than that available at OSU. Nevertheless, the most abundant

hydrocarbons can be identiﬁed from this simple GC-MS analysis.

Benzene,1-butenyl (1-Phenyl-1-butene, C10H12, CAS # 824-90-8) is the most abun-

dant component of diesel fuel, therefore it represents 100%. Tetradecane (2,6,10,

trimethyl, C17H36, CAS# 14905-56-7) presents the largest second composition. As

mentioned in Chapter 3, hexadecane is commonly used as a surrogate for diesel fuel.

As can be seen from the speciation performed here, hexadecane presents about a 60%

abundance (retention time = 21.61 min). The high content of hexadecane allows fur-

ther comparisons between the experimental results for the reformation of diesel fuel

and those obtained numerically via the 0D model.

242 Bio-diesel (fatty acid alkyl esters) is an alternative diesel fuel produced from re-

newable sources such as vegetable oils (edible or non-edible oil) and animal fats, in

combination with a simple alcohol. The molecules that formed bio-diesel are simple

hydrocarbon chains having no sulfur or aromatics [104]. In the GC-MS analysis per-

formed on bio-diesel (B-100) at CCIC (Campus Chemical Instrument Center at OSU)

for this dissertation, 9,12-Octadecadienoic acid, C19H34O2 (Linoleic acid - methyl es-

ter, CAS # 112-63-0), is the most abundant component. The second most abundant

compound is 9-Octadecenoic acid, C19H36O2 (methyl 9-octadecenoate, CAS # 2462-

84-2), while the third one is hexadecanoic, acid methyl ester, C17H34O2 (Palmitic acid

- methyl ester, CAS # 112-39-0).

As can be inferred from this analysis, large hydrocarbon chains are found to consti-

tute a large percentage of diesel and bio-diesel fuels. This allows comparison for the

reforming results with hexadecane. Furthermore, the experimental reforming results

can be compared with those obtained from the 0D model discussed in Chapter 5.

7.3 Eﬀect of the Equivalence Ratio

The temperature measured at the ﬁrst thermocouple location (see Figure 6.3), the

hydrogen yield and concentrations of H2, CO, CO2, and total hydrocarbons (THC)

emanating from the experiment for all fuels are presented in this section. These results

are obtained over the range of equivalence ratios 0.8 ≤ φ ≤ 2.2 and plotted along with

the calculated values obtained from the 0D chemical kinetics model for comparison.

First, the results of hexadecane are examined, then those for diesel and bio-diesel are

discussed respectively, in the following subsections.

243 7.3.1 Hexadecane Fuel

Figure 7.1 shows the variation of temperature within the equivalence ratio range.

The adiabatic ﬂame temperature calculated with the numerical model versus φ, Figure

7.2, shows higher values than those observed in the experiments. This is explained by the fact that the apparatus is not adiabatic and therefore suﬀers heat loss from the igniter to the location of the thermocouple. The general trend of the calculated values is that of a decrease in the adiabatic ﬂame temperature with increasing φ, which is in good agreement with the experimental temperature.

Figure 7.1. Temperature for Hexadecane at Diﬀerent Equivalence Ratios

244 Figure 7.2. Adiabatic Flame Temperature for Hexadecane at Diﬀerent φ and S/C Ratios at 1 atm and Tini = 1000 K

The measured output of H2 from the experimental reformer as a function of the

equivalence ratio is shown in Figure 7.3. As predicted by the 0D model, the hydrogen percent by volume of the total gas mixture shows an increasing behavior with equivalence ratio. It can be seen that the predicted hydrogen concentration agrees well with the experimental concentration. Small diﬀerences are observed at equivalence ratios near stoichiometry, where hydrogen production is expected to be negligible. This can be explained by the fact that the mixture in the actual reformer is not completely homogeneous which could lead to isolated rich spots and lean spots. As expected, the hydrogen yield (only fuel basis) shown in Figure 7.4 for the reformation of hexadecane

increases with equivalence ratio and is in good agreement with the predicted values

from the 0D model. The saturation behavior shown by the numerical trends are also

245 seen for the hydrogen concentration and yield due to the decreasing temperature with

increasing equivalence ratio.

Although the increasing concentration of hydrogen and yield may suggest that the

reformer should be operated at as high an equivalent ratio as possible, it has to be

pointed out that excessive soot formation becomes limiting at values of φ > 2.2 and beyond. The data shows a tendency for the rates of increase of H2 yield and H2 volume fraction to decrease as φ increases. The same trend is displayed by the predictions of the 0D chemical kinetics model, shown by the solid curve.

Figure 7.5 shows measured values of the CO volume fraction, along with predictions

from the 0D chemical kinetics model. As can be seen, the trend of CO is predicted

quite well by the model showing an increasing behavior with increasing equivalence ra-

tio. However, the numerical values of the concentration are lower for the experimental

measurements. This can be explained by the fact that the numerical simulation is car-

ried out under the assumptions of perfect homogeneous mixture, and an adiabatic pro-

cess. The simulation also assumes that the residence time is long enough for chemical

equilibrium to be achieved. Instead, the experimental setup is not perfectly homo-

geneous and has signiﬁcant heat losses. The non-homogeneity leads to lean and rich

spots over the entire volume. Furthermore, the presence of the ﬂame screen increases

the speed of the ﬂame since the diameter is smaller than the actual pipe, provoking a

fast core with little residence time for the species to achieve complete equilibrium, as

well as recirculation zones where the species have a long residence time achieving then

equilibrium.

246 The CO2 volume fraction depicted in Figure 7.6 shows the trend calculated by the

numerical 0D model and that obtained experimentally. It can be seen that the pre-

dicted model reproduces the trend observed in the experimental results, decreasing as

the equivalence ratio increases. The fact that the variation of the volume fraction of

CO2 is underpredicted by the 0D model can be also explained by the diﬀerences be-

tween the 0D model and the actual reformer. Figures 7.7 and 7.8 show the CO and

CO2 concentrations based on equilibrium calculations, predictions from the 0D kinetics model, and experimental results. From these ﬁgures it is clearly seen that CO and

CO2 experimental concentrations are lower than those predicted by equilibrium calculations as well, concluding thus that the reactions forming these species (oxygenated hydrocarbons R-HCO) do not achieve equilibrium within the length of the experimental reformer due to the non-homogeneous mixing and insuﬃcient residence time.

Figure 7.9 shows the variation of the volume fraction of total hydrocarbons (THC)

versus equivalence ratio. As expected, the volume fraction of THC increases sharply

with increasing equivalence ratio since the temperature decreases. From the ﬁrst ver-

sion of the reformer, it was seen that increasing the mass ﬂow rates resulted in a higher

temperature, higher hydrogen yields and lower amount of hydrocarbons. Based on this

fact, the fact that THC concentration obtained experimentally is higher than that seen

for the numerical 0D model must be due to lower temperatures arising from heat losses.

It must be emphasized that the hydrocarbons that comprise the THC in the model are

likely diﬀerent from those in the experiments. For example, soot formation, which

clearly occurs in the experiments, is not included in the 0D chemical kinetics model.

The model underpredicts the experimental values for these reasons. Nevertheless, the

trends observed in the experiments are reproduced by the calculations since the fact

247 that certain common hydrocarbons (e.g., CH4,C2H2, etc. ) are produced in increasing amounts as φ is increased, is correctly represented in the model.

It can be concluded that the homogeneity of the mixture and the heat transfer in the system are different between the experimental reforming process and the numerical calculations. The fact that the reactions are temperature dependent suggests that the equilibrium compositions achieved by reaching the adiabatic flame temperature would be different if heat transfer were modeled for the entire process, reaching then lower temperatures. Although, the prediction of hydrogen agrees fairly well with that for the predictive model, it can be seen that hydrogen starts to decreases for the high equivalence ratio cases due to decreased temperatures. If the linear temperature profile imposed in the 0D model were to start after the temperature has achieved 2000 K, the equilibrium compositions for hydrogen would be different (slightly lower) as it is indicated by the experimental measurements. The CO composition would also be slightly larger since the CO consuming elementary process of the water gas shift reaction decreases as temperature decreases and φ increases. CO2, however, would continue to be underpredicted.

248 Figure 7.3. H2 Concentration for Hexadecane at Diﬀerent Equivalence Ratios

Figure 7.4. H2 Yield for Hexadecane at Diﬀerent Equivalence Ratios

249 Figure 7.5. CO Concentration for Hexadecane at Diﬀerent Equivalence Ratios

Figure 7.6. CO2 Concentration for Hexadecane at Diﬀerent Equivalence Ratios

250 Figure 7.7. CO Concentration for Hexadecane at Diﬀerent Equivalence Ratios

Figure 7.8. CO2 Concentration for Hexadecane at Diﬀerent Equivalence Ratios

251 Figure 7.9. THC Concentration for Hexadecane at Diﬀerent Equivalence Ratios

7.3.2 Diesel Fuel

The experimental results obtained for the reformation of diesel fuel are presented in

this section and plotted along with the predictive trends obtained by the 0D model for

hexadecane. As mentioned in Section 7.2, diesel fuel constitutes a wide mixture of hy-

drocarbons. Figure 7.10 illustrates the temperature recorded by the ﬁrst thermocouple

in the reformer. This temperature shows a maximum obtained at low equivalence ratio

which agrees well with the behavior of the oxidation of hydrocarbons expressed in [101].

Thereafter, the temperature decreases with increasing φ. It can be noticed that the temperature for diesel fuel is slightly larger than that for hexadecane, while most of the concentrations of products are very similar. Hence, a simple explanation is that the

252 diesel fuel contains small hydrocarbons apart from hexadecane. This indeed causes the enthalpy of the reactants to be slightly smaller for diesel fuel than for hexadecane

(Hr in the energy equation presented in Chapter 5).

The behavior of the experimentally measured hydrogen concentration shown in

Figure 7.11 for diesel is in good agreement with the predicted values for hexadecane using the 0D model and those obtained experimentally for hexadecane. As explained for hexadecane, the non-homogeneous nature of the actual reformer generates lean and rich pockets that lead to the varying production of hydrogen even for the stoichiometric equivalence ratio. The saturation behavior also takes place for diesel fuel due to the decreasing temperature with increasing equivalence ratio. The hydrogen yield shows slightly larger values for diesel fuel than that for hexadecane. This is owing to the fact that the hydrogen yield deﬁned in Equation (6.13), is given in terms of the hydrogen to carbon ratio, which is larger for hexadecane (2.125) than for diesel fuel (1.8).

Carbon monoxide and carbon dioxide concentrations depicted in Figures 7.13 and

7.6 show similar trends to those described earlier for hexadecane. The total volume percent of hydrocarbons shown in Figure 7.15, shows a sharp increase with increasing equivalence ratio, as it does for pure hexadecane fuel. Nevertheless, a very slight decrease while comparing diesel and hexadecane is seen owing to the slightly higher temperatures attained for diesel fuel.

253 Figure 7.10. Temperature for Diesel at Diﬀerent Equivalence Ratios

Figure 7.11. H2 Concentration for Diesel at Diﬀerent Equivalence Ratios

254 Figure 7.12. H2 Yield for Diesel at Diﬀerent Equivalence Ratios

Figure 7.13. CO Concentration for Diesel at Diﬀerent Equivalence Ratios

255 Figure 7.14. CO2 Concentration for Diesel at Diﬀerent Equivalence Ratios

Figure 7.15. THC Concentration for Diesel at Diﬀerent Equivalence Ratios

256 7.3.3 Bio-Diesel Fuel

The experimental results obtained for reforming of bio-diesel are presented in this section and shown along with the predictions from the 0D model for hexadecane. As mentioned in Section 7.2, this particular bio-diesel fuel constitutes large chains of oxygenated hydrocarbons. Figure 7.10 illustrates the temperature recorded at the ﬁrst thermocouple in the reformer. The temperature shows a slight increase for φ = 1.45, thereafter it decreases with increasing φ. It can be noticed that the temperature for bio-diesel fuel is slightly higher than that for hexadecane and lower than that for diesel.

The trends of the major species for bio-diesel follow those described for hexadecane and diesel fuels described in the previous sections. Nevertheless, slight diﬀerences are noticed.

Figure 7.16. Temperature for Bio-Diesel at Diﬀerent Equivalence Ratios

257 The experimentally measured hydrogen concentration increases as φ is increased

as shown in Figure 7.17. The hydrogen volume percent is, however, lower than that predicted by the 0D model for hexadecane. This can be explained by obtaining the equilibrium composition of the major species using C19H36O2 as the model molecule

using the equilibrium code found in [97]. The coeﬃcients to calculate the thermody-

namic properties for this molecule are obtained from [100]. The resulting CO, CO2,

H2O and H2 equilibrium compositions are shown in Figure 7.18.

Figure 7.17. H2 Concentration for Bio-Diesel at Diﬀerent Equivalence Ratios

258 Figure 7.18. H2, CO and CO2 Equilibrium Concentrations for Bio-diesel at Diﬀerent φ at 1 atm

Comparison between the equilibrium compositions for hexadecane shown in Fig-

ure 5.1 and those for bio-diesel illustrate that hydrogen production is expected to be

lower for bio-diesel. Based upon this, the experimental trends are in good agreement

with the equilibrium compositions trends. As explained in Chapter 5, the equilibrium

model accounts for chemical and thermodynamic equilibrium, as well as the adiabatic

assumption, which allows use of the equilibrium constants to calculate dissociation

products such as CO, H2, H and OH. Hence, the numerical values are not expected to completely agree with the experimental results. It is noteworthy that the equivalence ratio used to plot the results for the kinetics model already considers the oxygen present in the bio-diesel molecule. As explained for hexadecane, the saturation behavior also takes place for bio-diesel due to the decreasing temperature with increasing equivalence

259 ratio. The hydrogen yield shows slightly lower values for bio-diesel fuel than for the other two fuels. This is owing to the fact that the hydrogen concentration is lower for bio-diesel.

Figure 7.19. H2 Yield for Bio-Diesel at Diﬀerent Equivalence Ratios

Carbon monoxide and carbon dioxide concentrations shown in Figures 7.20 and

7.20 show similar trends to those described for hexadecane and diesel. CO shows slightly lower values than those for diesel and hexadecane, while CO2 presents slightly higher values. Once again the diﬀerence between the experimental results and those obtained from the 0D model can be explained by the series of assumptions inherent in the model, as discussed for hexadecane in Section 7.3.1. The total volume percent of hydrocarbons, shown in Figure 7.22, shows a similar increase as the equivalence ratio

260 increases, as for the hexadecane fuel. Nevertheless, a very slight decrease is observed when comparing diesel to hexadecane.

Figure 7.20. CO Concentration for Bio-Diesel at Diﬀerent Equivalence Ratios

261 Figure 7.21. CO2 Concentration for Bio-Diesel at Diﬀerent Equivalence Ratios

Figure 7.22. THC Concentration for Bio-Diesel at Diﬀerent Equivalence Ratios

262 7.3.4 Experimental Results for Diesel, Hexadecane and Bio-Diesel

Figures 7.23 to 7.28 show the temperature, the hydrogen yield and the volume percent of the major species H2, CO, CO2 and THC for the three fuels on the same plot in order to aid in visualizing the diﬀerences explained in the previous sections.

Figure 7.23. Temperature for Hexadecane, Diesel and Bio-Diesel at Diﬀerent Equiva- lence Ratios

263 Figure 7.24. H2 Concentration for Hexadecane, Diesel and Bio-Diesel at Diﬀerent Equivalence Ratios

Figure 7.25. H2 Yield for Hexadecane, Diesel and Bio-Diesel at Diﬀerent Equivalence Ratios

264 Figure 7.26. CO Concentration for Hexadecane, Diesel and Bio-Diesel at Diﬀerent Equivalence Ratios

Figure 7.27. CO2 Concentration for Hexadecane, Diesel and Bio-Diesel at Diﬀerent Equivalence Ratios

265 Figure 7.28. THC Concentration for Hexadecane, Diesel and Bio-Diesel at Diﬀerent Equivalence Ratios

7.4 Comparison of the Hydrocarbon Species obtained Exper- imentally and Those Calculated by the 0D Model

Table 7.1 presents the results obtained in the GC-FID speciation of diesel, bio-diesel and hexadecane in terms of mass (C1 basis) percent. For example, a hydrocarbon such as C4H10 would be counted as C1H2.5 in terms of C1 basis. Hence, the total number of moles of the C1 compound is given by the number of moles of the original compound multiplied by the number of carbon atoms. Three diﬀerent equivalence ratios for each fuel were run to analyze the hydrocarbon composition. Hence, the

B1, B2, and B3 cases correspond to bio-diesel at the equivalence ratios of 1.15, 1.5, and 1.9, respectively. D1, D2, D3 are the diesel cases at the equivalence ratios of

266 1.36, 1.15, and 1.8 respectively. Finally, the H1 and H2 cases correspond to those for

hexadecane reforming at equivalence ratios of 1.15 and 1.9, respectively. The third

case ran for hexadecane is not presented here due to a perforation found in the sample

bag compromising the actual composition of the reformate. For all cases, acetylene

exhibits the largest % mass C1 ( 45-65%) followed by methane ( 19.6-31.1%), benzene ¬ ¬ ( 0.6-10.35%) and ethylene ( 1.4-26%). Acetylene is known to be a soot precursor. ¬ ¬ As known, the 0D model predicts the number densities of several hydrocarbons for

the partial oxidation of hexadecane. In order to calculate the % mass in C1 basis of

the predicted hydrocarbon concentrations, Equation (7.1) is used.

Mi Xi(%massC1) = yi NC ∗ 100 (7.1) MT

where Xi is the mass fraction of the species i, yi represents the mole fraction of species

i calculated by the 0D model, NC is the number of carbon atoms in species i, Mi is

the molecular weight of species i while MT is the total molecular weight of all hy-

drocarbon species. Table 7.2 shows the predicted hydrocarbon compositions at three diﬀerent equivalence ratios, i.e., 1.3, 1.6, and 1.9 at 1 atm and Tini = 1000 K for hexadecane. A strong statement must be made here. The reaction set used for the numerical simulation does not account for soot formation mechanisms since the main objective is to predict the exothermicity and the composition of the major species.

Furthermore, some of the species that can be measured in the experimental speciation are not simulated in the 0D model. Nevertheless, from Table 7.2 and Table 7.1 it

can be seen that the numerical model predicts four species (methane, acetylene, ethy-

lene and benzene) that appear in the experimental set of species. From Table 7.2,

267 1-butadienol or 1-butadien-1-ol (HOC4H6) presents the largest % mass C1 ( 33-55%) ¬ followed by ethene, methoxy (C3H6O) ( 16-30%), methane (CH4)( 3.8-8.5%), and ¬ ¬ ethylene-aldehyde (CH2HCO) ( 2.3-8%). No further comparisons between the exper- ¬ imental concentrations and those obtained by the 0D model can be made due to the arguments presented in this discussion.

268 Compound B1 B2 B3 D1 D2 D3 H1 H2 φ 1.15 1.5 1.9 1.36 1.15 1.8 1.15 1.9 Methane 24.15 20.82 19.58 19.68 27.86 19.66 31.13 20.23 CH4 Ethylene 7.49 2.38 3.47 26.05 1.40 11.39 1.78 5.68 C2H4 Acetylene (Ethyne) 53.10 66.21 64.00 45.31 64.90 0.00 53.59 62.31 C2H2 Ethane 0.00 0.00 0.00 0.00 0.00 49.18 0.00 0.00 C2H6 Propene 1.00 0.09 0.00 0.04 0.00 0.00 0.00 0.05 C3H6 Allene (Propadiene) 0.92 0.53 0.02 0.27 0.36 0.30 0.24 0.63 C3H4 Propyne 0.00 0.00 0.53 0.59 0.00 0.67 0.00 0.00 C3H4 1,3-Butadiene 0.57 0.05 0.04 0.00 0.00 0.16 0.06 0.06 C4H6 2,2-Dimethylpropane 0.61 0.63 0.67 0.35 0.16 0.92 0.00 0.93 C5H12 Cyclopentadiene 0.66 0.13 0.16 0.16 0.04 0.32 0.04 0.22 C5H6 Benzene 6.58 3.36 7.99 4.55 2.29 10.35 0.62 5.98 C6H6 meta- & para-Xylenes 0.49 0.06 0.19 0.47 0.12 1.25 0.17 0.59 C8H10 Styrene 0.13 0.03 0.03 0.05 0.01 0.55 0.19 0.07 C8H8 Naphthalene 0.80 2.06 0.66 0.47 0.62 0.89 5.12 0.18 C10H8 Others 2.12 2.60 2.23 1.54 1.26 3.13 2.78 2.39 CH4 Methanol 0.14 0.26 0.12 0.16 0.18 0.06 0.29 0.19 CH3OH

Table 7.1. Experimental Hydrocarbon Speciation for Combustion of Bio-diesel, Diesel and Hexadecane in % Mass C1 basis

269 Compound Formula φ = 1.3 φ = 1.6 φ = 1.9 5 Ethyne C2H2 2.51058 4.24686 5.01193 Ethen-1-ol HCCOH 0.67021 0.09864 0.02684 Methane CH4 6.31761 3.82978 8.43229 Ethylene C2H4 0.00031 0.00018 0.00068 Ethylene-aldehyde CH2HCO 2.31431 7.62295 8.07330 Ethene, methoxy C3H6O 30.74563 24.55053 15.79910 Propionyl or propelene-aldehyde C3H5O 0.64816 0.79339 0.71433 Benzene C6H6 0.00635 0.44131 1.02426 1 Butadienol or 1 butadien-1-ol HOC4H6 40.38223 54.47224 33.49637 1,3 Butadien-1-ol C4H5O 1.54993 2.31324 1.54717 Phenoxy radical C6H5O 0.36697 0.00040 1.72014 2,4-Cyclopentadien-1-one C5H4O 4.05989 0.12231 23.70706

Table 7.2. Numerical Hydrocarbon Composition for Combustion of Hexadecane in % Mass C1 basis

7.5 Concluding Remarks

In this chapter, the experimental results obtained for the partial oxidation reform-

ing of hexadecane, diesel and bio-diesel fuels is introduced. The experimental major

species obtained for hexadecane are compared with those calculated by the 0D model

showing a very good agreement for hydrogen and hydrogen yield. The trend of CO

and CO2 concentrations exhibited a good agreement with that deﬁned by the 0D model.

However, the actual values showed a slight disagreement due to the diﬀerences between

the conditions of the experiment and the assumptions employed in the simulation, i.e., homogeneous mixing and adiabatic process. The species obtained for the reforming of diesel and bio-diesel are plotted along with the experimental and numerical hexadecane

270 concentrations for comparison. The model can be fairly used to predict the concentrations for diesel and hexadecane; however, bio-diesel shows slightly lower hydrogen concentration and yield.

A brief introduction of the liquid speciation of diesel and bio-diesel fuels is presented in this chapter. Some of the most abundant compounds for diesel fuel are benzene, tetradecane and hexadecane, while for bio-diesel are linoleic acid, methyl 9- octadecenoate and palmitic acid. The results of the hydrocarbon speciation of the reformate gas for three diﬀerent cases per fuel showed that the hydrocarbon species reported in the numerical calculation are diﬀerent from those obtained experimentally.

This is owing to the fact that some mechanisms (such as soot formation) are not included in the calculation and take place in the experimentation. Furthermore, some species that can be measured in the GC-FID analysis are not considered in the numerical analysis.

271 CHAPTER 8

CONCLUSIONS AND RECOMMENDATIONS FOR FURTHER WORK

8.1 Conclusions

The results presented in this dissertation strongly indicate that hydrogen can be

produced via non-catalytic partial oxidation controlled by a single parameter or com-

binations of various factors (φ, S/C, temperature, etc.) As mentioned in Chapter 1

and 2, the hydrogen concentrations required for exhaust after-treatment are relatively

low 0.5% Vol (exhaust) if used as an additive, or within a H2/NOx ratio of 10 for H2- ¬ SCR systems [38]. With further investigation on the H2-SCR formulations and slight

modiﬁcations of the reformer hardware the assessment of a TPOx-H2SCR system can

be performed. A brief summary of conclusions from this dissertation is given in the

following paragraphs.

A detailed analysis of the chemical mechanisms that produce hydrogen under the

inﬂuence of the characteristic properties (equivalence ratio, steam to carbon ratio and

temperature) was given in Chapter 4. Furthermore, a case on secondary injection of

272 vapor was also presented and analyzed. The main conclusions drawn for the mechanisms producing and consuming hydrogen are itemized below:

1. For a given equivalence ratio and steam to carbon ratio at low initial tempera-

tures, three regions are identiﬁed: the pre-ignition, combustion and cooling zones.

The set of hydrogen production reactions for both cases in the pre-ignition zone is

mainly dominated by hexadecane reacting with hydrogen radical and ethane de-

composition. Other reactions involving heptene, propylene, and aldehydes with

hydrogen atoms also take place. The latter reactions are also substantial for the

production of hydrogen (H2) during the ﬁrst combustion subregion.

Table 8.1. H2 Producing Reactions in the Pre-Ignition Section for φ = 1.6 and S/C = 1.0 at Tini = 1000 K

A second subregion of combustion is identiﬁed where water-hydrogen atom, CH4-

hydrogen atom, and ethylene-N2 reactions represent the largest contributors to

production of hydrogen.

273 No. Chemical Reaction 24 H + H2O −→ H2 + OH 136 H + CH2O −→ H2 + HCO 169 H + CH3HCO −→ H2 + CH3CO 194 H + CH4 −→ H2 + CH3 223 C2H6 −→ H2 + C2H4 244 N2 + C2H4 −→ N2 + H2 + C2H2 376 H + C3H6 −→ H2 + NC3H5

Table 8.2. H2 Producing Reactions in the Combustion Section for φ = 1.6 and S/C = 1.0 at Tini = 1000 K

Finally, in the cooling zone the reaction set consists of the same reactions as the

combustion subregion. Ignition delay increases with increasing S/C ratio due to

the speciﬁc heat of water. This also extends the combustion zone (in the time

domain).

2. For the same cases, equivalence ratio and steam to carbon ratio at low initial

temperatures, the depleting H2 reaction mechanisms include the reverse reac-

tions of water-atomic hydrogen and methane-atomic hydrogen (OH and methyl

consuming hydrogen, respectively). These two reactions are present throughout

the entire reforming process.

No. Chemical Reaction 59 H2 + OH −→ H + H2O 153 H2 + CH3 −→ H + CH4

Table 8.3. H2 Consuming Reactions in the Pre-Ignition & Ignition Sections for φ = 1.6 and S/C = 1.0 at Tini = 1000 K

274 3. For the high initial temperature cases selected to analyze the eﬀect of equivalence

ratio (φ = 1.6 and S/C = 0.0) and steam to carbon ratio (φ = 1.6 and S/C = 1.0)

on the hydrogen production, no ignition zone is seen. The combustion duration

is shorter as the temperature is increased. The equivalence ratio case (φ = 1.6

and S/C = 0.0) presented in Section 4.5 at this temperature shows slightly larger

number density gradients due to the H2 producing reaction mechanisms. How-

ever, the H2 depleting mechanisms become important in the second combustion

section, actually decreasing the hydrogen concentration further for the equiva-

lence ratio (φ = 1.6 and S/C = 0.0) than for the cases when water is added.

The high temperature cases do not show direct contribution from hexadecane

molecules. This is explained by the fact that the hexadecane molecules present

a higher rate for decomposition into small radicals and aldehydes. The S/C ra-

tio case (φ = 1.6 and S/C = 1.0) during the cooling section presents contribution

from reactions 24, 194 and 223, opposite to the equivalence ratio case (φ = 1.6

and S/C = 0.0) that shows only reaction 24 as the main contributor within the

same region.

4. The steam to carbon case (φ = 1.6 and S/C = 1.0) at high initial temperature

(1750 K) case presents contributions from other reactions (232, 148 & 75) while at

lower initial temperature for the same S/C ratio these mechanisms show negligible

production.

5. Secondary injection was simulated and thoroughly analyzed for a speciﬁc case (φ

= 1.6, S/C = 1.0 and TPI = 2000 K). Reactions 24 and 194 drive the reaction

contributions for the post-injection zone, while within the same time period the

275 No. Chemical Reaction 24 H + H2O −→ H2 + OH 75 CHOCHO −→ H2 + 2CO 136 H + CH2O −→ H2 + HCO 148 2CH3 −→ H2 + C2H4 169 H + CH3HCO −→ H2 + CH3CO 194 H + CH4 −→ H2 + CH3 223 C2H6 −→ H2 + C2H4 232 H + C2H5 −→ H2 + C2H4 244 N2 + C2H4 −→ N2 + H2 + C2H2 376 H + C3H6 −→ H2 + NC3H5

Table 8.4. H2 Producing Reactions in the Ignition Section for φ = 1.6 and S/C = 1.0 at Tini = 1750 K

S/C ratio case (φ = 1.6 and S/C = 1.0) at the same initial temperature presents

reaction 24 as the main contributor and reaction 194 decreases.

Equilibrium compositions for those conditions obtained using the kinetics 0D model were given in Chapter 5. Also, the cases of secondary injection of water vapor as well

as those using exhaust as the source of oxygen were discussed in the same chapter.

The main conclusions derived from this analysis are as follows:

1. H2, CO and THC concentrations increase with increasing equivalence ratio for all

temperatures and sources of oxygen (fresh air and exhaust gas). Nevertheless,

hydrogen and carbon monoxide concentrations saturate due to the decreasing

adiabatic ﬂame temperature with increasing equivalence ratio. High hydrogen

yields ( 25-30%) can be obtained for equivalence ratios above 1.6 for the low ¬ initial temperature case, while even higher yields ( 30-40%) are attained for the ¬

276 high initial temperature cases. Addition of water is not entirely beneﬁcial for

low initial temperatures, since it lowers the adiabatic ﬂame temperature.

2. The initial temperature of the feed gases plays a crucial role in the reformation

of hydrocarbons. Hence, if fuel and air are injected at high initial temperatures,

the hydrogen concentration increases as these three characteristic quantities are

increased. The eﬀect of steam addition on the hydrogen concentration is an

increase of about 1.5% Vol at low equivalence ratios (1.3) and about 0.3-1% Vol

for higher φ ratios (1.9 & 1.6 respectively). The hydrogen yield in terms of fuel as

the source of hydrogen showed a very large increase with temperature. The H2

yield ranges from 12% Vol for low equivalence ratios and low initial temperature

to about 20-24% Vol for the same equivalence and S/C ratio at a higher initial

temperature is increased. The yield continues to increase with a slower slope as

the equivalence ratio is further increased. The increase obtained by using water

vapor at these high initial temperatures and high equivalence ratios is 45-50% ¬ Vol while at low initial temperatures the H2 yield ranges between 40-42% Vol

with no water vapor addition. CO presents an increasing behavior with φ but

a decreasing behavior with S/C, while CO2 continues to decrease for all S/C,

however with a slower rate as S/C is increased. THC concentration is slightly

larger for the cases with steam addition than for the case with no additional steam

owing to the lower adiabatic temperature achieved as the S/C increases.

3. A new definition of H2 yield has been defined in this work. This is a definition in

terms of fuel and vapor as the sources of hydrogen. It was observed that addition

of water vapor decreases the numerical value of this yield for all temperatures.

277 4. Two diﬀerent post-injection temperatures were studied. The hydrogen volume

percent for the low initial temperature case shows no further increase than that al-

ready obtained at combustion, while the higher post-injection temperature (2000

K) slightly increases the hydrogen concentration only for φ = 1.6. The advan-

tage of using the heat generated for the combustion has a larger eﬀect for the high

S/C ratio cases, causing the hydrogen yield (based on fuel) to be larger for all

S/C ratios at a given φ ratio, contrary to the main injection case at high φ ratios.

The CO concentration is seen to increase for the post-injection case, while the

CO2 concentration decreases. Furthermore, THC percent is slightly lower for

the high-temperature post-injection cases than the normal S/C cases due to the

longer exposure to a high temperature (2000 K for 0.4 s).

5. Studies on the reformation of hexadecane using engine exhaust as the source of

oxygen were carried out. The high content of diluting species in the exhaust

limits the amount of energy release from the exothermic oxidation reaction for

all oxygen concentrations from exhaust. For the cases that resulted in ignition,

the hydrogen and CO concentrations show a similar behavior as those for fresh

air with increasing equivalence ratio; however, with lower numerical values due

to the lower temperature.

Finally, the steady state compositions obtained from the 0D model for various equivalence ratios are compared to the experimental results for hexadecane, diesel and bio- diesel fuels. The main conclusions are as follows:

278 1. The trends of the major species predicted by the 0D model are in good agreement

with the experimentally measured concentrations. The numerical values of hy-

drogen volume percent and hydrogen yield also agree between the model and the

experiments; however, the concentrations of CO are overpredicted while they are

underpredicted for CO2. This can be explained by the fact that the numerical

simulation is carried out under the assumptions of perfect homogeneous mix-

ture, and an adiabatic process. Also, long residence times resulting in chemical

equilibrium is assumed possible. Instead, the experimental setup is not per-

fectly homogeneous and presents signiﬁcant heat losses. The non-homogeneity

leads to lean and rich pockets over the entire volume as well as temperature

non-uniformities. Chemical equilibrium for the reactions forming these species

(oxygenated hydrocarbons R-HCO) is not achieved within the ﬁnite length of the

experimental reformer due to the non-homogeneous mixing. The total hydro-

carbon volume fraction increases sharply with increasing equivalence ratio since

the temperature decreases.

2. The trends for the product compositions obtained for diesel reforming are in

good agreement with those obtained for hexadecane. Bio-diesel compositions

also followed the same trend as hexadecane. Nevertheless, bio-diesel presents a

lower hydrogen concentration and yield than diesel and hexadecane.

3. The hydrocarbons that comprise the total hydrocarbon composition in the model

are diﬀerent from those in the experiments. This can be explained by the fact

that soot formation mechanisms, which clearly occurs in the experiments, are not

considered in the 0D chemical kinetics model. Furthermore, acetylene (C2H2),

a known soot precursor, was noted in large quantities.

279 4. It can be concluded that the homogeneity of the mixture and the heat transfer

in the system are relatively diﬀerent between the experimental reforming process

and the numerical calculations. The fact that the reactions are temperature

dependent suggests that the ﬁnal steady state compositions achieved by reaching

the adiabatic ﬂame temperature would be diﬀerent if heat transfer were modeled

for the entire process, reaching lower temperatures.

5. Despite the diﬀerences seen between the experimental and numerical results, the

trends of the compositions with the equivalence ratio are well predicted by the

0D model.

8.2 Further Work

The use of hydrogen or syngas as a reductant for remediation of nitrogen oxides is a promising solution to meet diesel emission regulations. Although the present work clearly describes a feasible concept for the on-board production of hydrogen, the following are some directions for future research:

1. The 0D model developed in this work is a powerful tool for the prediction of the

chemical product composition of the thermal oxidation process for heavy hydro-

carbons. Validation of the model was done using comparison with experimental

data and separately by comparing the model prediction with published results

on the combustion of pure hydrogen and pure methane. Nevertheless, a more

accurate model would require a heat transfer model that accounts for radiation

and convection heat losses. This would be an intermediate step towards a more

advanced tool for research purposes.

280 2. In order to accurately predict the hydrocarbon composition, other reactions such

as soot formation mechanisms would need to be added to the set of chemical

reactions.

3. The model could be used to predict the product composition at intermediate

equivalence ratios to provide a more detailed functionality map, which could be

used in control-oriented models used to model and control the performance of

H2-SCR or HC-SCR catalysts in the presence of syngas or hydrogen.

4. A CFD analysis to simulate the heat transfer and the distribution of the ﬂow

(including mixing) would be necessary if a high ﬁdelity model is needed. Such a

model would need to solve the 3-D energy, continuity and momentum equations

(Navier-Stokes) under turbulent regime. Nevertheless, the number of reactions

considered for the analysis is usually reduced to global reactions or the reaction

set has to be solved within a diﬀerent routine due to the stiﬀness of the system

(the ﬂow and chemical reactions present diﬀerent time scales).

5. While the 0D model, with its limitations and assumptions, is not fully capable

of precise quantitative predictions, the trends are correctly captured and the im-

portant effects, such as the favorable effect of temperature, are clearly identified,

hence providing a vary valuable design guiding tool.

6. Although the thermal partial oxidation reformer studied in this dissertation demon-

strated good performance, further research on the mixing of the feeds and the

ﬂame holder/sheet is needed.

7. It was seen that the temperature of the feeds plays an important role on the

amount of hydrogen generated through the reforming of hydrocarbons. High

281 temperatures were measured by the thermocouple used in the injection system.

These elevated temperatures caused soot formation inside the atomizers after sev-

eral hours of operation. A possible solution would be to increase the length of the

mixing chamber to guarantee that the atomizers are not exposed to temperature

higher than 500 to increase the durability of the system.

8. Although the igniter showed an impeccable performance in the experiments re-

ported here, its long exposure to high temperatures caused the connector between

the tip and the semiconductor to melt. Hence, to diminish that problem, metal

ﬁns could be added on the area close to the igniter to increase the convective

area.

9. Although the second version of the reformer had several drawbacks, the temper-

ature achieved by the feeds was higher (since the air was heated). It would be

worth exploring diﬀerent heat transfer mechanisms (such as long longitudinal ﬁns

from the base of the inner pipe to the outside of the outer pipe) to recuperate

the heat lost to the environment.

10. Selection of the pipe material is also essential since high temperatures are re-

quired to achieve even higher hydrogen concentrations. A reﬂective material or

steel with a reﬂective coating can be used for the same hardware conﬁguration.

As predicted from the simulation, the majority of hydrogen is produced in the

combustion zone. Hence, a ceramic pipe could be adapted in this zone to main-

tain the high temperature from the combustion.

282 APPENDIX A

CHEMICAL REACTIONS AND RATES

This appendix comprises the reduced set of chemical reactions, the pre-exponential

3 nr−1 factor A (m /molecules) /s)), activation energy Ea (J/mol), and the temperature power n for each individual reaction used to simulate the partial oxidation of hexadecane based on the reaction mechanism given by Ristori et al.[88] at 1 atm introduced in Chapter 3.

No. Reaction A Ea n 1 2H + N2 −→ N2 + H2 7.31e+017 0 -1 2 H + OH + H2O −→ 2H2O 1.39977e+023 0 0.2 3 H + N2 + OH −→ N2 + H2O 8.615e+021 0 -2 4 H + O2 −→ OH + O 1.85e+014 16812 0 5 H + N2 + O2 −→ N2 + HO2 8e+017 0 -0.8 6 H + HO2 −→ H2 + O2 4.28e+013 1411 0 7 H + HO2 −→ 2OH 1.69e+014 874 0 8 H + HO2 −→ H2O + O 3.01e+013 1721 0 9 N2 + HCCO −→ N2 + CH + CO 6e+015 58821 0 10 OH + HCCO −→ H + CO + HCO 1e+013 0 0 11 OH + HCCO −→ H2O + C2O 3e+013 0 0 12 O + HCCO −→ H + 2CO 1.93e+014 590 0 13 H + HCCO −→ CO + CH2 1.5e+014 0 0 14 O2 + HCCO −→ H + CO + CO2 1.4e+009 0 1 Continued on next page

Table A.1. Chemical Reactions: Reaction Rates, Activation Energy and Coeﬃcient

283 Table A.1 – Continued from previous page No. Reaction A Ea n 15 HCCO + CH2 −→ C2H + CH2O 1e+013 2000 0 16 HCCO + CH2 −→ CO + C2H3 3e+013 0 0 17 HCCO + C2H2 −→ CO + C3H3 1e+011 3000 0 18 HCCO + CH −→ CO + C2H2 5e+013 0 0 19 2HCCO −→ 2CO + C2H2 1e+013 0 0 20 H + HCCOH −→ H + CH2CO 1e+013 0 0 21 OH + HO2 −→ H2O + O2 2.89e+013 -497 0 22 O + HO2 −→ OH + O2 1.81e+013 -400 0 23 2HO2 −→ O2 + H2O2 407.5 1979 3.321 24 H + H2O −→ H2 + OH 4.5996e+008 18570.8 1.6 25 2H2O −→ H + OH + H2O 1.59986e+017 114245 0 26 H2O + O −→ 2OH 1.4999e+010 17256.2 1.14 27 N2 + H2 −→ 2H + N2 2.2e+014 96080 0 28 H2 + O2 −→ 2OH 1.7e+013 47780 0 29 OH + H2O2 −→ H2O + HO2 5.8e+014 9557 0 30 H + H2O2 −→ H2 + HO2 1.7e+012 3750 0 31 H + H2O2 −→ OH + H2O 1e+013 3590 0 32 O + H2O2 −→ OH + HO2 2.8e+013 6400 0 33 OH + H2C4O −→ HCCO + CH2CO 1e+007 2000 2 34 H + H2C4O −→ HCCO + C2H2 5e+013 3000 0 35 N2 + HCO −→ H + N2 + CO 1.85e+017 17000 -1 36 OH + HCO −→ H2O + CO 1e+014 0 0 37 O + HCO −→ OH + CO 3e+013 0 0 38 O + HCO −→ H + CO2 3e+013 0 0 39 H + HCO −→ H2 + CO 7.224e+013 0 0 40 O2 + HCO −→ HO2 + CO 2.11e+009 -620 1.12 41 HCO + CH3 −→ CO + CH4 1.2e+014 0 0 42 HO2 + HCO −→ H + OH + CO2 3e+013 0 0 43 HCO + C2H6 −→ CH2O + C2H5 47000 18235 2.72 44 2HCO −→ CO + CH2O 1.8e+013 0 0 45 2HCO −→ H2 + 2CO 3e+012 0 0 46 2N2 −→ N2 + 2N 5.55e+017 225000 -2.5 47 N2 + 2N −→ 2N2 4.53e+015 0 0 48 O2 + N −→ O + NO 9.0402e+012 7108 0 49 N2 + O2 −→ O2 + 2N 1.9e+017 225000 -0.5 50 N2 + O2 −→ N2 + 2O 1.2e+014 107900 0 51 N2 + 2O −→ N2 + O2 1e+016 0 -1 52 2O2 −→ O2 + 2O 3.6102e+018 118000 -1 53 N2 + O −→ N + NO 7e+013 75500 0 Continued on next page

284 Table A.1 – Continued from previous page No. Reaction A Ea n 54 H2 + O −→ H + OH 1.5e+007 7560 2 55 H + OH −→ H2 + O 4876.6 3876 2.8 56 H + N2 + O −→ N2 + OH 6.2e+016 0 -0.6 57 OH + O −→ H + O2 1.8e+013 0 0 58 H + O2 −→ OH + O 1.19988e+017 16515.3 -0.91 59 H2 + OH −→ H + H2O 9.99893e+007 3298.28 1.6 60 OH + N −→ H + NO 2.29017e+013 -169.694 0 61 2OH −→ H2O + O 1.49948e+008 0 1.14 62 CH + C2H2 −→ H + C3H2 1e+014 0 0 63 CH + CH4 −→ H + C2H4 6e+013 0 0 64 CH + CH3 −→ H + C2H3 3e+013 0 0 65 H2 + CH −→ H + CH2 3.31872e+008 1668.26 1.79 66 O + CH −→ H + CO 3.99963e+013 0 0 67 OH + CH −→ H + HCO 3e+013 0 0 68 O2 + CH −→ O + HCO 3.3e+013 0 0 69 O2 + CH −→ OH + CO 2e+013 0 0 70 CH + CO2 −→ CO + HCO 3.4e+012 690 0 71 CH + CH4 −→ H + C2H4 6e+013 0 0 72 CH + CH3 −→ H + C2H3 3e+013 0 0 73 CH + C2H2 −→ H + C3H2 1e+014 0 0 74 CHOCHO −→ CO + CH2O 1.176e+016 50937 -1.28 75 CHOCHO −→ H2 + 2CO 6.519e+039 367469 -7.703 76 OH + CHOCHO −→ H2O + CHOCO 1e+013 0 0 77 O + CHOCHO −→ OH + CHOCO 7.24e+012 1970 0 78 H + CHOCHO −→ HCO + CH2O 1e+012 0 0 79 HO2 + CHOCHO −→ H2O2 + CHOCO 1.7e+012 10700 0 80 CH3 + CHOCHO −→ CH4 + CHOCO 1.74e+012 8440 0 81 O2 + CHOCHO −→ HO2 + CO + HCO 6.3e+013 30000 0 82 CHOCO −→ CO + HCO 2e+007 0 0 83 O2 + CHOCO −→ HO2 + 2CO 3.98e+012 30000 0 84 2CH2 −→ H2 + C2H2 3.2e+013 0 0 85 2CH2 −→ 2H + C2H2 4e+013 0 0 86 CH2 + CH3 −→ H + C2H4 4e+013 0 0 87 H + C2H2 −→ CH + CH2 2571.27 104520 3 88 CH + CH2 −→ H + C2H2 4e+013 0 0 89 CH2 + CH4 −→ 2CH3 4e+013 0 0 90 OH + CH2 −→ H2O + CH 1.13e+007 3000 2 91 OH + CH2 −→ H + CH2O 2.5e+013 0 0 92 O + CH2 −→ 2H + CO 5e+013 0 0 93 O + CH2 −→ H2 + CO 6e+013 0 0 94 CH2 + CO2 −→ CO + CH2O 1.1e+011 1000 0 Continued on next page

285 Table A.1 – Continued from previous page No. Reaction A Ea n 95 2CH2 −→ H2 + C2H2 3.2e+013 0 0 96 2CH2 −→ 2H + C2H2 4e+013 0 0 97 CH2 + CH3 −→ H + C2H4 4e+013 0 0 98 CH + CH2 −→ H + C2H2 4e+013 0 0 99 CH2 + C2H2 −→ H + C3H3 1.2e+013 6620 0 100 CH2 + C2H4 −→ C3H6 4.3e+012 10038 0 101 CH2 + C2H6 −→ CH3 + C2H5 6.5e+012 7911 0 102 H + CH2 −→ H2 + CH 1e+018 0 -1.56 103 H2 + CH2 −→ H + CH3 7e+013 0 0 104 O2 + CH2 −→ 2H + CO2 1.58499e+012 1000 0 105 O2 + CH2 −→ OH + HCO 4.3e+010 -500 0 106 O2 + CH2 −→ H2 + CO2 3.45e+011 1000 0 107 O2 + CH2 −→ H2O + CO 1.87e+010 -1000 0 108 O2 + CH2 −→ H + OH + CO 8.64e+010 -500 0 109 O2 + CH2 −→ O + CH2O 5e+013 9000 0 110 CH2 + C3H8 −→ CH3 + NC3H7 2.19e+012 6405 0 111 H + CH2CO −→ CH2HCO 1e+014 2500 0 112 N2 + CH2CO −→ N2 + CO + CH2 3.6e+015 59270 0 113 O2 + CH2CO −→ CO2 + CH2O 2e+013 61500 0 114 HO2 + CH2CO −→ OH + CO + CH2O 6e+011 12738 0 115 O + CH2CO −→ CH2 + CO2 1.76e+012 1349 0 116 O + CH2CO −→ OH + HCCO 1e+013 8000 0 117 OH + CH2CO −→ CO + CH2OH 6.93e+012 0 0 118 OH + CH2CO −→ HCO + CH2O 2.04e+011 0 0 119 OH + CH2CO −→ H2O + HCCO 1.02e+011 0 0 120 OH + CH2CO −→ CO2 + CH3 3.1e+012 0 0 121 H + CH2CO −→ CO + CH3 15000 673 2.83 122 H + CH2CO −→ H2 + HCCO 1.8e+014 8600 0 123 CH2CO + CH3 −→ CO + C2H5 6e+010 0 0 124 CH2CO + CH3 −→ HCCO + CH4 7.5e+012 13000 0 125 CH2 + CH2CO −→ HCCO + CH3 1e+012 0 0 126 CH2 + CH2CO −→ CO + C2H4 3.6e+013 11000 0 127 O2 + CH2HCO −→ HO2 + CH2CO 1.58e+010 0 0 128 O2 + CH2HCO −→ OH + CO + CH2O 2.51e+010 0 0 129 O2 + CH2HCO −→ OH + CHOCHO 2.51e+011 14640 0 130 O + CH2HCO −→ HCO + CH2O 3.98e+013 0 0 131 HO2 + CH2HCO −→ OH + HCO + CH2O 1e+013 0 0 132 N2 + CH2O −→ H + N2 + HCO 1.26e+016 77898 0 133 HO2 + CH2O −→ HCO + H2O2 4e+012 11665 0 134 OH + CH2O −→ H2O + HCO 1.716e+009 -447 1.18 135 O + CH2O −→ OH + HCO 1.807e+013 3088 0 Continued on next page

286 Table A.1 – Continued from previous page No. Reaction A Ea n 136 H + CH2O −→ H2 + HCO 1.26e+008 2170 1.62 137 O2 + CH2O −→ HO2 + HCO 2.04e+013 39000 0 138 CH2O + CH3 −→ HCO + CH4 4.09e+012 8843 0 139 N2 + CH2OH −→ H + N2 + CH2O 1e+014 25100 0 140 H + CH2OH −→ H2 + CH2O 3e+013 0 0 141 O2 + CH2OH −→ HO2 + CH2O 2.168e+014 4690 0 142 H2 + C2H2 −→ 2CH2 660.914 104491 3 143 N2 + CH3 −→ H + N2 + CH2 1.9e+016 91600 0 144 N2 + CH3 −→ N2 + H2 + CH 6.9e+014 82460 0 145 H + CH3 −→ H2 + CH2 7e+013 15100 0 146 H + N2 + CH3 −→ N2 + CH4 7.99849e+026 0 -3 147 2CH3 −→ H + C2H5 3.011e+013 13513 0 148 2CH3 −→ H2 + C2H4 9.99893e+015 32026.8 0 149 2CH3 −→ C2H6 2.39977e+014 0 -0.4 150 2CH3 −→ H + C2H5 3.011e+013 13513 0 151 CH3 + CH3O −→ CH2O + CH4 2.409e+013 0 0 152 CH3 + CH2OH −→ CH2O + CH4 2.41e+012 0 0 153 H2 + CH3 −→ H + CH4 659.951 7743.79 3 154 H2O + CH3 −→ OH + CH4 482.958 14866.2 2.9 155 HO2 + CH3 −→ OH + CH3O 5e+012 0 0 156 O + CH3 −→ H + CH2O 6.99937e+013 0 0 157 OH + CH3 −→ H + CH2OH 2.64e+019 8068 -1.8 158 OH + CH3 −→ H + CH3O 5.74e+012 13931 -0.23 159 OH + CH3 −→ H2O + CH2 8.9e+018 8067 -1.8 160 OH + CH3 −→ H2 + CH2O 3.19e+012 10810 -0.53 161 O2 + CH3 −→ O + CH3O 1.32e+014 31398 0 162 O2 + CH3 −→ OH + CH2O 4.38e+011 14656 0 163 N2 + CH3 + C2H5 −→ N2 + C3H8 4.9e+014 0 -0.5 164 N2 + CH3CO −→ N2 + CO + CH3 8.64e+015 14400 0 165 CH3HCO −→ HCO + CH3 9.586e+014 74180 0 166 HO2 + CH3HCO −→ H2O2 + CH3CO 1.7e+012 10700 0 167 OH + CH3HCO −→ H2O + CH3CO 2.35e+010 -1113 0.73 168 O + CH3HCO −→ OH + CH3CO 5.85e+012 1808 0 169 H + CH3HCO −→ H2 + CH3CO 4.1e+009 2405 1.16 170 O2 + CH3HCO −→ HO2 + CH3CO 2e+013 42200 0.5 171 CH3 + CH3HCO −→ CH4 + CH3CO 1.7e+012 8440 0 172 HCO + CH3HCO −→ CH2O + CH3CO 7.8e+013 8440 0 173 N2 + CH3O −→ H + N2 + CH2O 4.88e+015 22773 0 174 HO2 + CH3O −→ CH2O + H2O2 3e+011 0 0 175 OH + CH3O −→ H2O + CH2O 1e+013 0 0 176 O + CH3O −→ OH + CH2O 1.3e+013 0 0 Continued on next page

287 Table A.1 – Continued from previous page No. Reaction A Ea n 177 H + CH3O −→ H2 + CH2O 2e+013 0 0 178 O2 + CH3O −→ HO2 + CH2O 2.349e+010 1788 0 179 CH2O + CH3O −→ HCO + CH3OH 1.15e+011 1280 0 180 CO + CH3O −→ CO2 + CH3 1.566e+013 11804 0 181 HCO + CH3O −→ CO + CH3OH 9e+013 0 0 182 C2H5 + CH3O −→ CH2O + C2H6 2.41e+013 0 0 183 C2H3 + CH3O −→ CH2O + C2H4 2.41e+013 0 0 184 C2H4 + CH3O −→ CH2O + C2H5 1.2e+011 7000 0 185 CH3OH −→ OH + CH3 1.565e+046 103522 -9.28 186 HO2 + CH3OH −→ H2O2 + CH2OH 6.3e+012 19360 0 187 OH + CH3OH −→ H2O + CH2OH 4.532e+011 1160 0.33 188 OH + CH3OH −→ H2O + CH3O 3.629e+011 5868 0.7 189 O + CH3OH −→ OH + CH2OH 1.63e+013 5030 0 190 H + CH3OH −→ H2 + CH2OH 4e+013 6100 0 191 CH2O + CH3OH −→ 2CH3O 1.549e+012 79570 0 192 CH3 + CH3OH −→ CH4 + CH2OH 3.566e+011 8663 0 193 CH3 + CH3OH −→ CH4 + CH3O 467700 12764 2.328 194 H + CH4 −→ H2 + CH3 22500 8756.6 3 195 N2 + CH4 −→ H + N2 + CH3 1.99985e+017 88432.1 0 196 O + CH4 −→ OH + CH3 1.1999e+007 7624.28 2.1 197 OH + CH4 −→ H2O + CH3 1.59986e+006 2461.76 2.1 198 HO2 + CH4 −→ H2O2 + CH3 1.122e+013 24641 0 199 O2 + CH4 −→ HO2 + CH3 4.04e+013 56913 0 200 H2 + C2 −→ H + C2H 400000 1000 2.4 201 O2 + C2 −→ 2CO 5e+013 0 0 202 OH + C2 −→ H + C2O 5e+013 0 0 203 C2H2OH −→ H + CH2CO 5e+015 28000 0 204 H + C2H2OH −→ H2 + CH2CO 2e+013 4000 0 205 O + C2H2OH −→ OH + CH2CO 2e+013 4000 0 206 OH + C2H2OH −→ H2O + CH2CO 1e+013 2000 0 207 O2 + C2H2OH −→ H2 + HCO + CO2 4e+012 -250 0 208 C2H3O −→ H + CH2CO 1.6e+013 35000 0 209 C2H3O −→ CH3CO 8.511e+014 14000 0 210 C2H4O −→ CO + CH4 1.21e+013 57200 0 211 C2H4O −→ CH3HCO 6e+013 57200 0 212 C2H4O −→ HCO + CH3 4.9e+013 57200 0 213 O2 + C2H4O −→ HO2 + C2H3O 4e+013 61500 0 214 H + C2H4O −→ H2 + C2H3O 2e+013 8300 0 215 H + C2H4O −→ H2O + C2H3 5e+009 5000 0 216 H + C2H4O −→ OH + C2H4 9.51e+010 5000 0 217 O + C2H4O −→ OH + C2H3O 1.91e+012 5250 0 Continued on next page

288 Table A.1 – Continued from previous page No. Reaction A Ea n 218 OH + C2H4O −→ H2O + C2H3O 4.79e+013 5955 0 219 HO2 + C2H4O −→ H2O2 + C2H3O 4e+012 17000 0 220 CH3 + C2H4O −→ CH4 + C2H3O 1.07e+012 11830 0 221 C2H4O2H −→ OH + C2H4O 3.16e+011 19500 0 222 C2H6 −→ 2CH3 1.54e+018 45700 -1.24 223 C2H6 −→ H2 + C2H4 2.29e+017 34039 0 224 C2H6 −→ H + C2H5 8.11e+017 51439 -1.23 225 H + C2H6 −→ H2 + C2H5 5.25e+014 12800 0 226 OH + C2H6 −→ H2O + C2H5 4.1e+006 860 2.06 227 O + C2H6 −→ OH + C2H5 9.99e+008 5803 1.5 228 O2 + C2H6 −→ HO2 + C2H5 1e+013 51000 0 229 C2H6 + CH3O −→ C2H5 + CH3OH 3.02e+011 7000 0 230 CH3 + C2H6 −→ CH4 + C2H5 7.536 9883 3.727 231 HO2 + C2H6 −→ H2O2 + C2H5 1.21e+012 17600 0 232 H + C2H5 −→ H2 + C2H4 1.25e+014 8000 0 233 HO2 + C2H5 −→ H2O2 + C2H4 3e+011 0 0 234 CH3 + C2H5 −→ CH4 + C2H4 1.144e+012 0 0 235 2C2H5 −→ C2H6 + C2H4 1.4e+012 0 0 236 OH + C2H5 −→ H2O + C2H4 2.409e+013 0 0 237 OH + C2H5 −→ H + CH2O + CH3 2.409e+013 0 0 238 O + C2H5 −→ CH2O + CH3 4.238e+013 0 0 239 O + C2H5 −→ H + CH3HCO 5.298e+013 0 0 240 O + C2H5 −→ OH + C2H4 3.046e+013 0 0 241 O2 + C2H5 −→ HO2 + C2H4 3e+020 6761 -2.86 242 O2 + C2H5 −→ C2H4O2H 1.25e+040 11030 -9.41 243 O2 + C2H5 −→ OH + CH3HCO 1.585e+014 10390 -1.17 244 N2 + C2H4 −→ N2 + H2 + C2H2 3e+017 79350 0 245 N2 + C2H4 −→ H + N2 + C2H3 2.97e+017 96560 0 246 H + C2H4 −→ H2 + C2H3 1e+014 15009 0 247 CH3 + C2H4 −→ C2H3 + CH4 6.63 9499 3.7 248 CH3 + C2H4 −→ C2H3 + CH4 4.16e+012 5600 0 249 HO2 + C2H4 −→ C2H4O2H 2e+011 8000 0 250 OH + C2H4 −→ H2O + C2H3 2.024e+013 5936 0 251 O + C2H4 −→ HCO + CH3 1.2e+008 530 1.44 252 O + C2H4 −→ H + CH2HCO 2e+008 530 1.44 253 O2 + C2H4 −→ HO2 + C2H3 4e+013 61500 0 254 H + C2H3 −→ H2 + C2H2 3e+013 0 0 255 HO2 + C2H3 −→ OH + CO + CH3 3e+013 0 0 256 HCO + C2H3 −→ CO + C2H4 9.034e+013 0 0 257 OH + C2H3 −→ H2O + C2H2 3e+013 0 0 258 O + C2H3 −→ CO + CH3 1.5e+013 0 0 Continued on next page

289 Table A.1 – Continued from previous page No. Reaction A Ea n 259 O + C2H3 −→ H + CH2CO 1.5e+013 0 0 260 O2 + C2H3 −→ HO2 + C2H2 1.975e+012 10000 0 261 CH + C2H3 −→ CH2 + C2H2 5e+013 0 0 262 CH2 + C2H3 −→ H + NC3H4 3e+013 0 0 263 CH2O + C2H3 −→ HCO + C2H4 5420 5862 2.81 264 C2H3 + CH3 −→ C2H2 + CH4 3.91e+011 0 0 265 C2H3 + C2H6 −→ C2H5 + C2H4 1.5e+013 10000 0 266 C2H + C2H3 −→ 2C2H2 9.64e+011 0 0 267 C2H3 + C2H2 −→ H + C4H4 2e+012 5000 0 268 2C2H3 −→ C2H2 + C2H4 1.084e+013 0 0 269 2C2H3 −→ NC4H6 5e+013 0 0 270 C2H3 + C2H4 −→ H + NC4H6 1.2e+012 4683 0 271 C2H2 + CH3 −→ C2H + CH4 1.8e+011 17290 0 272 C2H2 + CH3 −→ H + NC3H4 4785 12893 2.42 273 C2H2 + CH3 −→ NC3H5 13960 16502 2.21 274 C2H2 + CH3 −→ NC3H5 3.853e+056 727892 -13.6 275 2C2H2 −→ H + NC4H3 6.31e+013 41600 0 276 C2H + C2H2 −→ H + C4H2 1.82e+014 467 0 277 O2 + C2H2 −→ OH + HCCO 2e+008 30100 1.5 278 O2 + C2H2 −→ HO2 + C2H 1.21e+013 74520 0 279 O2 + C2H3 −→ O + CH2HCO 1.447e+015 3135 -0.78 280 O2 + C2H3 −→ HCO + CH2O 1.853e+023 3892 -3.29 281 O2 + C2H3 −→ H + CHOCHO 1.91e+018 2398 -2.25 282 HO2 + C2H2 −→ OH + CH2CO 6.09e+009 7948 0 283 OH + C2H2 −→ H2O + C2H 3.4e+007 14000 2 284 OH + C2H2 −→ CO + CH3 0.000484 -2000 4 285 OH + C2H2 −→ H + CH2CO 0.0002192 -1000 4.5 286 OH + C2H2 −→ H + HCCOH 505900 13500 2.3 287 OH + C2H2 −→ C2H2OH 1.925e+022 6596 -3.35 288 O + C2H2 −→ CO + CH2 19800 656 2.6 289 O + C2H2 −→ H + HCCO 46200 656 2.6 290 H + C3H2 −→ CH + C2H2 2.13799e+015 35929.7 0 291 H2 + C2H −→ H + C2H2 1.506e+013 3100 0 292 OH + C2H −→ H + HCCO 2e+013 0 0 293 OH + C2H −→ H2O + C2 4e+007 8000 2 294 O + C2H −→ CH + CO 1e+013 0 0 295 O2 + C2H −→ CO + HCO 2.41e+012 0 0 296 N2 + C2H3CO −→ N2 + CO + C2H3 8.6e+015 23000 0 297 O2 + C2O −→ O + 2CO 2e+013 0 0 298 OH + C2O −→ H + 2CO 2e+013 0 0 299 O + C2O −→ 2CO 5e+013 0 0 Continued on next page

290 Table A.1 – Continued from previous page No. Reaction A Ea n 300 H + C2O −→ CH + CO 5e+013 0 0 301 N2 + CH3 + C2H5 −→ N2 + C3H8 4.9e+014 0 -0.5 302 O2 + C3H2 −→ HCCO + HCO 3e+010 2870 0 303 OH + C3H2 −→ HCO + C2H2 5e+013 0 0 304 CH2 + C3H2 −→ H + NC4H3 3e+013 0 0 305 O2 + C3H3 −→ HCO + CH2CO 3.01e+010 2870 0 306 HO2 + C3H3 −→ H2O2 + C3H2 2e+012 0 0 307 OH + C3H3 −→ H2O + C3H2 5e+012 0 0 308 OH + C3H3 −→ HCCO + CH3 5e+012 0 0 309 O + C3H3 −→ HCO + C2H2 1.385e+014 0 0 310 O + C3H3 −→ CO + C2H3 4.615e+013 0 0 311 O + C3H3 −→ C2H + CH2O 4.615e+013 0 0 312 O + C3H3 −→ H + CO + C2H2 4.615e+013 0 0 313 H + C3H3 −→ H2 + C3H2 5e+013 3000 0 314 CH + C3H3 −→ H + NC4H3 7e+013 0 0 315 CH + C3H3 −→ H + NC4H3 7e+013 0 0 316 CH2 + C3H3 −→ H + C4H4 4e+013 0 0 317 C3H3 + CH3 −→ NC4H6 3.33e+012 0 0 318 C3H3O −→ C2H3CO 8.51e+014 14000 0 319 O2 + C3H3O −→ OH + HCCO + HCO 5.01e+012 19192 0 320 N2 + NC3H4 −→ H + N2 + C3H3 4.7e+018 80000 0 321 NC3H4 −→ C2H + CH3 4.2e+016 100000 0 322 O2 + NC3H4 −→ HO2 + C3H3 2.5e+012 51000 0 323 O2 + NC3H4 −→ OH + HCCO + CH2 1e+007 30100 1.5 324 HO2 + NC3H4 −→ OH + CO + C2H4 6.09e+009 7948 0 325 HO2 + NC3H4 −→ CH2O + CH3CO 3e+012 16000 0 326 HO2 + NC3H4 −→ HCO + CH3HCO 4.5e+012 16000 0 327 HO2 + NC3H4 −→ C3H3 + H2O2 5e+011 19000 0 328 OH + NC3H4 −→ H2O + C3H3 650 200 3 329 OH + NC3H4 −→ CH2CO + CH3 0.0002 -1000 4.5 330 OH + NC3H4 −→ HCO + C2H4 0.0001 -1000 4.5 331 OH + NC3H4 −→ CH2O + C2H3 0.0001 -1000 4.5 332 O + NC3H4 −→ CH2 + CH2CO 6.4e+012 2100 0 333 O + NC3H4 −→ HCO + C2H3 3.2e+012 2100 0 334 O + NC3H4 −→ HCCO + CH3 9.18e+012 2100 0 335 O + NC3H4 −→ CH2O + C2H2 3.2e+011 2100 0 336 O + NC3H4 −→ H + HCCO + CH2 3.2e+011 2010 0 337 H + NC3H4 −→ NC3H5 1.3e+013 1999 0 338 H + NC3H4 −→ H2 + C3H3 2e+014 15000 0 339 CH3 + NC3H4 −→ C3H3 + CH4 4e+011 7700 0 340 C2H3 + NC3H4 −→ C3H3 + C2H4 1e+011 7700 0 Continued on next page

291 Table A.1 – Continued from previous page No. Reaction A Ea n 341 C2H + NC3H4 −→ C2H2 + C3H3 4.2e+016 100000 0 342 O2 + C3H4O −→ HO2 + C3H3O 4e+013 61500 0 343 HO2 + C3H4O −→ OH + CH2O + CH2CO 1e+012 14340 0 344 HO2 + C3H4O −→ H2O2 + C3H3O 4e+012 17000 0 345 OH + C3H4O −→ H2O + C3H3O 4.79e+013 5955 0 346 C3H6 + NC4H5 −→ NC4H6 + NC3H5 1e+011 9800 0 347 H + C3H8 −→ H2 + NC3H7 10140 2931 2.88 348 CH3 + C3H8 −→ CH4 + NC3H7 8.07e+011 10110 0 349 C2H5 + C3H8 −→ C2H6 + NC3H7 5.01e+010 10400 0 350 C2H3 + C3H8 −→ C2H4 + NC3H7 1000 8829 3.1 351 C3H8 + NC3H5 −→ C3H6 + NC3H7 7.94e+011 16200 0 352 O2 + C3H8 −→ HO2 + NC3H7 4e+013 50872 0 353 O2 + C3H8 −→ HO2 + NC3H7 4e+013 47500 0 354 HO2 + C3H8 −→ H2O2 + NC3H7 47600 16494 2.55 355 HO2 + C3H8 −→ H2O2 + NC3H7 9640 13910 2.6 356 OH + C3H8 −→ H2O + NC3H7 3.16e+007 934 1.8 357 OH + C3H8 −→ H2O + NC3H7 7.08e+006 -159 1.9 358 O + C3H8 −→ OH + NC3H7 3.715e+006 5505 2.4 359 O + C3H8 −→ OH + NC3H7 549500 3140 2.5 360 CH3 + C3H8 −→ CH4 + NC3H7 3e+012 11710 0 361 C2H5 + C3H8 −→ C2H6 + NC3H7 3.16e+011 12300 0 362 C2H5 + C3H8 −→ C2H6 + NC3H7 5.01e+010 10400 0 363 C2H3 + C3H8 −→ C2H4 + NC3H7 600 10502 3.3 364 C2H3 + C3H8 −→ C2H4 + NC3H7 1000 8829 3.1 365 C3H8 + CH3O −→ NC3H7 + CH3OH 3.18e+011 7050 0 366 C3H8 + CH3O −→ NC3H7 + CH3OH 7.2e+010 4470 0 367 NC3H7 −→ CH3 + C2H4 1.14e+011 31110 0 368 NC3H7 −→ H + C3H6 3e+029 639758 -5.24 369 C2H2 + NC3H7 −→ CH3 + NC4H6 2.77e+010 6504 0 370 O2 + NC3H7 −→ HO2 + C3H6 2.754e+010 -2151 0 371 O2 + NC3H7 −→ HO2 + C3H6 1e+012 5020 0 372 C3H6 −→ H + NC3H5 1.45e+015 98060 0 373 C3H6 −→ C2H3 + CH3 1.1e+021 97720 -1.2 374 C3H6 −→ C2H2 + CH4 3.5e+012 70000 0 375 H + C3H6 −→ NC3H7 5.704e+009 874 1.16 376 H + C3H6 −→ H2 + NC3H5 170000 2492 2.5 377 CH3 + C3H6 −→ CH4 + NC3H5 1.6e+011 8800 0 378 C2H5 + C3H6 −→ C2H6 + NC3H5 1e+011 9800 0 379 C2H3 + C3H6 −→ CH3 + NC4H6 7.2e+011 5008 0 380 O2 + C3H6 −→ HO2 + NC3H5 1.95e+012 39000 0 381 HO2 + C3H6 −→ OH + C3H6O 1.02e+012 14964 0 Continued on next page

292 Table A.1 – Continued from previous page No. Reaction A Ea n 382 HO2 + C3H6 −→ H2O2 + NC3H5 3e+009 9930 0 383 OH + C3H6 −→ H2O + NC3H5 769800 622 2.214 384 O + C3H6 −→ HCO + C2H5 4.689e+007 -628 1.57 385 O + C3H6 −→ CH3 + CH2HCO 3.9e+007 -628 1.57 386 O + C3H6 −→ CH2O + C2H4 7.02e+007 -628 1.57 387 O + C3H6 −→ OH + NC3H5 1.75e+011 5884 0.7 388 NC3H5 −→ H + NC3H4 5.62e+012 43500 0 389 H + NC3H5 −→ H2 + NC3H4 1e+013 0 0 390 CH2 + NC3H5 −→ H + NC4H6 3e+013 0 0 391 CH3 + NC3H5 −→ CH4 + NC3H4 1e+011 0 0 392 HO2 + NC3H5 −→ OH + C3H5O 7.5e+011 0 0 393 C2H5 + NC3H5 −→ C2H6 + NC3H4 4e+011 0 0 394 C2H3 + NC3H5 −→ C2H4 + NC3H4 1e+012 0 0 395 C2H3 + NC3H5 −→ NC5H8 1e+013 0 0 396 O2 + NC3H5 −→ HCO + CH3HCO 4.335e+012 0 0 397 HO2 + NC3H5 −→ OH + CH2CO + CH3 4.5e+012 0 0 398 O + NC3H5 −→ CH2CO + CH3 1.807e+014 0 0 399 O2 + NC3H5 −→ CH2O + CH3CO 5.01e+012 19000 0 400 O2 + NC3H4 −→ HO2 + C3H3 4e+013 61500 0 401 HO2 + NC3H4 −→ C3H3 + H2O2 1.8e+013 19000 0 402 HO2 + NC3H4 −→ OH + CH2 + CH2CO 5e+011 19000 0 403 HO2 + NC3H4 −→ OH + CH2O + C2H2 5e+011 19000 0 404 OH + NC3H4 −→ H2O + C3H3 2.15e+012 -200 0 405 OH + NC3H4 −→ CH2CO + CH3 1e+012 -393 0 406 OH + NC3H4 −→ HCO + C2H4 1.25e+011 -393 0 407 OH + NC3H4 −→ CO + C2H5 1.875e+011 -393 0 408 O + NC3H4 −→ CO + C2H4 0.01125 -4243 4.613 409 O + NC3H4 −→ HCO + C2H3 0.0005 -4243 4.613 410 O + NC3H4 −→ CH2 + CH2CO 0.001 -4243 4.613 411 O + NC3H4 −→ CH2O + C2H2 0.0025 -4243 4.613 412 H + NC3H4 −→ NC3H5 1.2e+012 2700 0 413 H + NC3H4 −→ H2 + C3H3 1e+014 15009 0 414 N2 + NC3H4 −→ H + N2 + C3H3 2e+018 80000 0 415 CH3 + NC3H4 −→ C3H3 + CH4 1.333e+012 7700 0 416 NC3H4 + NC3H5 −→ C3H3 + C3H6 9e+011 7700 0 417 C3H3 + CH3 −→ NC4H6 3.33e+012 0 0 418 H + C3H3 −→ H2 + C3H2 5e+013 3000 0 419 C3H3 + NC3H4 −→ H + C6H6 1.4e+012 12800 0 420 2C3H3 −→ C6H6 3e+011 0 0 421 CH2 + C3H2 −→ H + NC4H3 3e+013 0 0 422 NC4H10 −→ 2C2H5 2e+016 81300 0 Continued on next page

293 Table A.1 – Continued from previous page No. Reaction A Ea n 423 NC4H10 −→ CH3 + NC3H7 1e+017 85400 0 424 CH3 + NC4H10 −→ CH4 + NC4H9 2.189e+011 11400 0 425 NC4H10 −→ H + NC4H9 1e+016 95000 0 426 O2 + NC4H10 −→ HO2 + NC4H9 3.981e+013 47600 0 427 H + NC4H10 −→ H2 + NC4H9 4.777e+006 7369 2.5 428 O + NC4H10 −→ OH + NC4H9 428000 2583 2.6 429 OH + NC4H10 −→ H2O + NC4H9 4.14e+007 753 1.73 430 HO2 + NC4H10 −→ H2O2 + NC4H9 4e+012 18150 0 431 CH3 + NC4H10 −→ CH4 + NC4H9 2.189e+011 11400 0 432 C2H3 + NC4H10 −→ C2H4 + NC4H9 7.943e+011 16800 0 433 C2H5 + NC4H10 −→ C2H6 + NC4H9 1e+011 10400 0 434 NC3H5 + NC4H10 −→ C3H6 + NC4H9 3.162e+011 16400 0 435 CH3O + NC4H10 −→ CH3OH + NC4H9 6e+011 7000 0 436 NC4H9 −→ H + C4H8 1.26e+013 38600 0 437 NC4H9 −→ CH3 + C3H6 1.26e+012 27100 0 438 CH3 + NC4H9 −→ CH4 + C4H8 2e+012 0 0 439 O2 + NC4H9 −→ HO2 + C4H8 3.8e+010 -2000 0 440 O + NC4H9 −→ OH + C4H8 4.16e+014 0 0 441 OH + NC4H9 −→ H2O + C4H8 2.4e+013 0 0 442 H + NC4H9 −→ H2 + C4H8 1.25e+013 0 0 443 HO2 + NC4H9 −→ H2O2 + C4H8 2.41e+013 0 0 444 C2H5 + NC4H9 −→ C2H6 + C4H8 1.6e+012 0 0 445 NC3H5 + NC4H9 −→ C3H6 + C4H8 1e+012 0 0 446 C4H8 −→ CH3 + NC3H5 1.5e+019 73400 -1 447 CH3 + C4H8 −→ CH4 + C4H7 1e+011 7300 0 448 HO2 + C4H8 −→ OH + CH2O + C3H6 2.5e+012 14340 0 449 O + C4H8 −→ C2H4 + CH3HCO 1.3e+013 850 0 450 O + C4H8 −→ C2H5 + CH3CO 1.625e+013 850 0 451 H + C4H8 −→ H2 + C4H7 1.72e+014 8000 0 452 H + C4H8 −→ CH3 + C3H6 1.72e+013 3600 0 453 HCO + C4H8 −→ CH2O + C4H7 3.3e+011 6210 0 454 C2H3 + C4H8 −→ C2H4 + C4H7 1e+013 13000 0 455 NC3H5 + C4H8 −→ C3H6 + C4H7 7.94e+011 20500 0 456 C4H7 −→ C2H3 + C2H4 5e+013 38000 0 457 CH3 + C4H7 −→ CH4 + NC4H6 7.94e+012 0 0 458 C4H7 −→ CH3 + NC3H4 2e+013 51460 0 459 HO2 + C4H7 −→ OH + C4H7O 4.5e+012 0 0 460 NC4H6 −→ C2H2 + C2H4 1e+014 75000 0 461 NC4H6 −→ H + NC4H5 1.58e+016 110000 0 462 H + NC4H6 −→ CH3 + NC3H4 1e+013 15000 0 463 CH3 + NC4H6 −→ CH4 + NC4H5 2e+014 24800 0 Continued on next page

294 Table A.1 – Continued from previous page No. Reaction A Ea n 464 C2H3 + NC4H6 −→ C2H4 + NC4H5 632 18010 3.13 465 O2 + NC4H6 −→ HO2 + NC4H5 1.4e+012 50600 0 466 HO2 + NC4H6 −→ H2O2 + NC4H5 1e+011 9920 0 467 OH + NC4H6 −→ H2O + NC4H5 8e+006 3744 2 468 O + NC4H6 −→ OH + NC4H5 4.53e+015 57028 -0.47 469 H + NC4H6 −→ H2 + NC4H5 5e+015 22800 0 470 CH3 + NC4H6 −→ CH4 + NC4H5 4e+014 22800 0 471 C2H3 + NC4H6 −→ C2H4 + NC4H5 632 18010 3.13 472 OH + NC4H6 −→ HOC4H6 3.5e+012 -994 0 473 NC4H5 −→ C2H3 + C2H2 5e+013 44000 0 474 H + NC4H5 −→ H2 + C4H4 1e+014 0 0 475 C2H2 + NC4H5 −→ C2H + NC4H6 1.8e+011 17300 0 476 O2 + NC4H5 −→ CH2O + C2H3CO 2e+011 14000 0 477 O + NC4H5 −→ C2H3 + CH2CO 1e+013 2000 0 478 HO2 + NC4H5 −→ OH + C2H3 + CH2CO 1e+012 2000 0 479 HO2 + NC4H5 −→ OH + C4H5O 1e+013 2000 0 480 O + NC4H5 −→ CO + NC3H5 1e+013 2000 0 481 OH + NC4H5 −→ H2O + C4H4 2e+007 1000 2 482 HO2 + NC4H5 −→ OH + CO + NC3H5 1e+013 0 0 483 H + NC4H5 −→ H2 + C4H4 1e+014 0 0 484 C2H2 + NC4H5 −→ C2H + NC4H6 1.8e+011 17300 0 485 CH2O + NC4H5 −→ HCO + NC4H6 4e+012 8840 0 486 C2H2 + NC4H5 −→ H + C6H6 2800 1400 2.9 487 CH3 + NC4H5 −→ NC5H8 2e+013 0 0 488 C4H4 −→ 2C2H2 3.2e+013 77100 0 489 N2 + C4H4 −→ H + N2 + NC4H3 1e+020 99300 0 490 C4H4 −→ H + NC4H3 6.3e+013 87100 0 491 C4H4 −→ H2 + C4H2 1.26e+015 94700 0 492 H + C4H4 −→ H2 + NC4H3 333000 9240 2.53 493 CH3 + C4H4 −→ CH4 + NC4H3 16.6 9499 3.7 494 O2 + C4H4 −→ HO2 + NC4H3 1e+013 44640 0 495 HO2 + C4H4 −→ H2O2 + NC4H3 1e+011 9920 0 496 HO2 + C4H4 −→ OH + HCO + C3H3 4e+011 8000 0 497 OH + C4H4 −→ H2O + NC4H3 2e+007 2000 2 498 O + C4H4 −→ CO + NC3H4 3e+013 1810 0 499 H + C4H4 −→ H2 + NC4H3 333000 9240 2.53 500 C2H + C4H4 −→ C2H2 + NC4H3 3.98e+013 0 0 501 CH3 + C4H4 −→ CH4 + NC4H3 16.6 9499 3.7 502 C4H4 + NC4H5 −→ NC4H6 + NC4H3 4e+011 5000 0 503 C2H3 + C4H4 −→ H + C6H6 1.9e+012 3510 0 504 C6H6 −→ C2H2 + C4H4 9e+015 107430 0 Continued on next page

295 Table A.1 – Continued from previous page No. Reaction A Ea n 505 2C3H3 −→ H + C6H5 1e+013 0 0 506 C2H2 + NC4H3 −→ C6H5 2800 1400 2.9 507 H + NC4H3 −→ H2 + C4H2 5e+013 0 0 508 O2 + NC4H3 −→ HCCO + CH2CO 1e+012 0 0 509 OH + NC4H3 −→ H2O + C4H2 3e+013 0 0 510 O + NC4H3 −→ C2H + CH2CO 2e+013 0 0 511 O + NC4H3 −→ H + H2C4O 2e+013 0 0 512 H + NC4H3 −→ H2 + C4H2 5e+013 0 0 513 CH2 + NC4H3 −→ C2H + NC3H4 2e+013 0 0 514 OH + NC4H3 −→ H2O + C4H2 3e+013 0 0 515 C2H3 + NC4H3 −→ C6H6 2.87e+014 817 0 516 N2 + C4H2 −→ H + N2 + C4H 3.5e+017 80065 0 517 OH + C4H2 −→ H + H2C4O 6.66e+012 -410 0 518 O + C4H2 −→ CO + C3H2 1.2e+012 0 0 519 O2 + C4H −→ 2CO + C2H 1e+014 0 0 520 NC16H34 −→ CH3 + NC15H31 1e+017 84360 0 521 NC16H34 −→ C2H5 + NC14H29 1e+017 82540 0 522 NC16H34 −→ NC3H7 + NC13H27 1e+017 83480 0 523 NC16H34 −→ NC4H9 + NC12H25 1e+017 82340 0 524 NC16H34 −→ NC11H23 + NC5H11 1e+017 83410 0 525 NC16H34 −→ NC10H21 + NC6H13 1e+017 82290 0 526 NC16H34 −→ NC9H19 + NC7H15 1e+017 83190 0 527 NC16H34 −→ 2NC8H17 5e+016 78650 0 528 NC16H34 −→ H + NC16H33 1e+015 100000 0 529 NC16H34 + H −→ H2 + NC16H33 5.6e+007 7700 2 530 NC16H34 + O −→ OH + NC16H33 1128 1653 3.27 531 NC16H34 + O2 −→ HO2 + NC16H33 4e+013 47600 0 532 NC16H34 + OH −→ H2O + NC16H33 1.13e+006 -1391 2 533 NC16H34 + CH3 −→ CH4 + NC16H33 1.3e+012 11600 0 534 NC16H34 + C2H5 −→ C2H6 + NC16H33 1e+012 10400 0 535 NC16H34 + HO2 −→ H2O2 + NC16H33 4.88e+012 18500 0 536 NC16H34 + C2H3 −→ C2H4 + NC16H33 8e+011 16800 0 537 NC16H34 + NC3H5 −→ C3H6 + NC16H33 8e+011 16800 0 538 NC16H34 + NC3H7 −→ C3H8 + NC16H33 1e+012 10400 0 539 NC16H34 + NC4H7 −→ C4H8 + NC16H33 8e+011 16800 0 540 NC16H34 + C6H5 −→ C6H6 + NC16H33 8e+011 16800 0 541 NC16H33 −→ C4H8 + NC12H25 2e+013 28700 0 542 NC16H33 −→ CH3 + NC15H30 2e+013 31000 0 543 NC16H33 −→ C2H5 + NC14H28 2e+013 29700 0 544 NC16H33 −→ NC10H21 + NC6H12 5e+013 28700 0 545 NC15H31 −→ CH3 + NC14H28 2e+013 31000 0 Continued on next page

296 Table A.1 – Continued from previous page No. Reaction A Ea n 546 NC15H31 −→ C2H5 + NC13H26 2e+013 29700 0 547 NC15H31 −→ NC7H15 + NC8H16 2e+013 28700 0 548 NC15H31 −→ H + NC15H30 5e+012 40000 0 549 NC15H30 −→ NC3H5 + NC12H25 1e+016 71000 0 550 H + NC15H30 −→ H2 + NC15H29 1e+012 3900 0 551 CH3 + NC15H30 −→ CH4 + NC15H29 2e+011 7300 0 552 O2 + NC15H30 −→ HO2 + NC15H29 1.4e+013 31900 0 553 OH + NC15H30 −→ H2O + NC15H29 1e+012 1230 0 554 O + NC15H30 −→ OH + NC15H29 1e+012 4000 0 555 HO2 + NC15H30 −→ H2O2 + NC15H29 1e+011 17000 0 556 C2H3 + NC15H30 −→ C2H4 + NC15H29 2e+011 7300 0 557 NC15H29 −→ NC4H6 + NC11H23 2.5e+013 35000 0 558 H + NC15H29 −→ NC15H30 1e+013 0 0 559 NC14H29 −→ CH3 + NC13H26 2e+013 31000 0 560 NC14H29 −→ C2H5 + NC12H24 2e+013 29700 0 561 NC14H29 −→ NC7H15 + NC7H14 2e+013 28700 0 562 NC14H28 −→ NC3H5 + NC11H23 1e+016 71000 0 563 H + NC14H28 −→ H2 + NC14H27 1e+012 3900 0 564 CH3 + NC14H28 −→ CH4 + NC14H27 2e+011 7300 0 565 O2 + NC14H28 −→ HO2 + NC14H27 1.4e+013 31900 0 566 HO2 + NC14H28 −→ H2O2 + NC14H27 1e+011 17000 0 567 OH + NC14H28 −→ H2O + NC14H27 1e+012 1230 0 568 O + NC14H28 −→ OH + NC14H27 1e+012 4000 0 569 C2H3 + NC14H28 −→ C2H4 + NC14H27 2e+011 7300 0 570 NC14H27 −→ NC4H6 + NC10H21 2.5e+013 35000 0 571 H + NC14H27 −→ NC14H28 1e+013 0 0 572 NC13H27 −→ CH3 + NC12H24 2e+013 31000 0 573 NC13H27 −→ C2H5 + NC11H22 2e+013 29700 0 574 NC13H27 −→ NC6H13 + NC7H14 2e+013 28700 0 575 NC13H27 −→ H + NC13H26 5e+012 40000 0 576 NC13H26 −→ NC3H5 + NC10H21 1e+016 71000 0 577 H + NC13H26 −→ H2 + NC13H25 1e+012 3900 0 578 CH3 + NC13H26 −→ CH4 + NC13H25 2e+011 7300 0 579 O2 + NC13H26 −→ HO2 + NC13H25 1.4e+013 31900 0 580 HO2 + NC13H26 −→ H2O2 + NC13H25 1e+011 17000 0 581 OH + NC13H26 −→ H2O + NC13H25 1e+012 1230 0 582 O + NC13H26 −→ OH + NC13H25 1e+012 4000 0 583 C2H3 + NC13H26 −→ C2H4 + NC13H25 2e+011 7300 0 584 NC13H25 −→ NC4H6 + NC9H19 2.5e+013 35000 0 585 H + NC13H25 −→ NC13H26 1e+013 0 0 586 NC12H25 −→ CH3 + NC11H22 2e+013 31000 0 Continued on next page

297 Table A.1 – Continued from previous page No. Reaction A Ea n 587 NC12H25 −→ C2H5 + NC10H20 2e+013 29700 0 588 NC12H25 −→ NC6H13 + NC6H12 2e+013 28700 0 589 NC12H25 −→ H + NC12H24 5e+012 40000 0 590 NC12H24 −→ NC3H5 + NC9H19 1e+016 71000 0 591 H + NC12H24 −→ H2 + NC12H23 1e+012 3900 0 592 CH3 + NC12H24 −→ CH4 + NC12H23 2e+011 7300 0 593 O2 + NC12H24 −→ HO2 + NC12H23 1.4e+013 31900 0 594 HO2 + NC12H24 −→ H2O2 + NC12H23 1e+011 17000 0 595 OH + NC12H24 −→ H2O + NC12H23 1e+012 1230 0 596 O + NC12H24 −→ OH + NC12H23 1e+012 4000 0 597 C2H3 + NC12H24 −→ C2H4 + NC12H23 2e+011 7300 0 598 H + NC12H23 −→ NC12H24 1e+013 0 0 599 NC12H23 −→ NC4H6 + NC8H17 2.5e+013 35000 0 600 NC11H23 −→ NC5H11 + NC6H12 2e+013 28700 0 601 NC11H23 −→ NC6H13 + NC5H10 2e+013 28700 0 602 NC11H22 −→ NC3H5 + NC8H17 1e+016 71000 0 603 H + NC11H22 −→ H2 + NC11H21 1e+012 3900 0 604 CH3 + NC11H22 −→ CH4 + NC11H21 2e+011 7300 0 605 O2 + NC11H22 −→ HO2 + NC11H21 1.4e+013 31900 0 606 HO2 + NC11H22 −→ H2O2 + NC11H21 1e+011 17000 0 607 OH + NC11H22 −→ H2O + NC11H21 1e+012 1230 0 608 O + NC11H22 −→ OH + NC11H21 1e+012 4000 0 609 C2H3 + NC11H22 −→ C2H4 + NC11H21 2e+011 7300 0 610 NC11H21 −→ NC4H6 + NC7H15 2.5e+013 35000 0 611 H + NC11H21 −→ NC11H22 1e+013 0 0 612 NC10H21 −→ H + NC10H20 5e+012 40000 0 613 NC10H21 −→ CH3 + NC9H18 2e+013 31000 0 614 NC10H21 −→ C2H5 + NC8H16 2e+013 29700 0 615 NC10H21 −→ NC5H11 + NC5H10 2e+013 28700 0 616 NC10H20 −→ NC3H5 + NC7H15 1e+016 71000 0 617 H + NC10H20 −→ H2 + NC10H19 1e+012 3900 0 618 CH3 + NC10H20 −→ CH4 + NC10H19 2e+011 7300 0 619 O2 + NC10H20 −→ HO2 + NC10H19 1.4e+013 31900 0 620 HO2 + NC10H20 −→ H2O2 + NC10H19 1e+011 17000 0 621 OH + NC10H20 −→ H2O + NC10H19 1e+012 1230 0 622 O + NC10H20 −→ OH + NC10H19 1e+012 4000 0 623 C2H3 + NC10H20 −→ C2H4 + NC10H19 2e+011 7300 0 624 NC10H19 −→ NC4H6 + NC6H13 2.5e+013 35000 0 625 H + NC10H19 −→ NC10H20 1e+013 0 0 626 NC9H19 −→ CH3 + NC8H16 2e+013 31000 0 627 NC9H19 −→ C2H5 + NC7H14 2e+013 29700 0 Continued on next page

298 Table A.1 – Continued from previous page No. Reaction A Ea n 628 NC9H19 −→ H + NC9H18 5e+012 40000 0 629 NC9H18 −→ NC3H5 + NC6H13 1e+016 71000 0 630 H + NC9H18 −→ H2 + NC9H17 1e+012 3900 0 631 CH3 + NC9H18 −→ CH4 + NC9H17 2e+011 7300 0 632 O2 + NC9H18 −→ HO2 + NC9H17 1.4e+013 31900 0 633 HO2 + NC9H18 −→ H2O2 + NC9H17 1e+011 17000 0 634 OH + NC9H18 −→ H2O + NC9H17 1e+012 1230 0 635 O + NC9H18 −→ OH + NC9H17 1e+012 4000 0 636 C2H3 + NC9H18 −→ C2H4 + NC9H17 2e+011 7300 0 637 NC9H17 −→ NC4H6 + NC5H11 2.5e+013 35000 0 638 H + NC9H17 −→ NC9H18 1e+013 0 0 639 NC8H17 −→ CH3 + NC7H14 2e+013 31000 0 640 NC8H17 −→ C2H5 + NC6H12 2e+013 29700 0 641 NC8H17 −→ NC3H7 + NC5H10 2e+013 28700 0 642 NC8H17 −→ H + NC8H16 5e+012 40000 0 643 NC8H16 −→ NC3H5 + NC5H11 1e+016 71000 0 644 H + NC8H16 −→ H2 + NC8H15 6e+006 5000 2 645 CH3 + NC8H16 −→ CH4 + NC8H15 4e+011 9500 0 646 O2 + NC8H16 −→ HO2 + NC8H15 1.4e+013 31900 0 647 HO2 + NC8H16 −→ H2O2 + NC8H15 1e+011 17000 0 648 OH + NC8H16 −→ H2O + NC8H15 1e+012 1230 0 649 O + NC8H16 −→ OH + NC8H15 1e+012 4000 0 650 H + NC8H16 −→ H2 + NC8H15 1e+012 3900 0 651 CH3 + NC8H16 −→ CH4 + NC8H15 2e+011 7300 0 652 C2H3 + NC8H16 −→ C2H4 + NC8H15 2e+011 7300 0 653 H + NC8H15 −→ NC8H16 1e+013 0 0 654 NC8H15 −→ C2H4 + NC6H11 1e+013 30000 0 655 NC7H15 −→ H + NC7H14 1e+013 37900 0 656 NC7H15 −→ CH3 + NC6H12 2e+013 31000 0 657 O2 + NC7H15 −→ HO2 + NC7H14 2.759e+010 -3532 0 658 HO2 + NC7H15 −→ H2O2 + NC7H14 1e+013 0 0 659 OH + NC7H15 −→ H2O + NC7H14 2.43e+013 0 0 660 NC7H14 −→ C2H5 + NC5H9 1.55e+019 73400 -1 661 H + NC7H14 −→ H2 + NC7H13 1.6e+014 3400 0 662 O2 + NC7H14 −→ HO2 + NC7H13 1.4e+013 31900 0 663 HO2 + NC7H14 −→ H2O2 + NC7H13 1e+011 17060 0 664 O + NC7H14 −→ OH + NC7H13 8e+013 4000 0 665 CH3 + NC7H14 −→ CH4 + NC7H13 4e+011 6800 0 666 C2H3 + NC7H14 −→ C2H4 + NC7H13 4e+011 6800 0 667 NC7H13 −→ NC3H7 + NC4H6 1e+011 37000 0 668 H + NC7H13 −→ NC7H14 1e+013 0 0 Continued on next page

299 Table A.1 – Continued from previous page No. Reaction A Ea n 669 NC6H13 −→ H + NC6H12 2e+013 40400 0 670 NC6H13 −→ CH3 + NC5H10 2e+013 31000 0 671 NC6H13 −→ C2H5 + C4H8 2e+013 29700 0 672 H + NC6H13 −→ H2 + NC6H12 1.81e+012 0 0 673 CH3 + NC6H13 −→ CH4 + NC6H12 1e+012 0 0 674 C2H5 + NC6H13 −→ C2H6 + NC6H12 1e+012 0 0 675 O2 + NC6H13 −→ HO2 + NC6H12 2e+012 2000 0 676 O2 + NC6H13 −→ OH + CH3HCO + C4H8 2.1e+011 6858 0 677 OH + NC6H13 −→ H2O + NC6H12 2.43e+013 0 0 678 O + NC6H13 −→ CH2O + NC5H11 1.61e+013 0 0 679 HO2 + NC6H13 −→ OH + CH3HCO + NC4H9 1e+013 0 0 680 O + NC6H13 −→ CH3HCO + NC4H9 1.61e+013 0 0 681 NC6H12 −→ NC3H7 + NC3H5 8.162e+015 70830 0 682 NC6H12 −→ 2C3H6 3.98e+012 57700 0 683 H + NC6H12 −→ H2 + NC6H11 6.55e+012 4445 0 684 CH3 + NC6H12 −→ CH4 + NC6H11 2e+011 6800 0 685 C2H3 + NC6H12 −→ C2H4 + NC6H11 2e+011 6800 0 686 O2 + NC6H12 −→ HO2 + NC6H11 4e+012 40000 0 687 HO2 + NC6H12 −→ H2O2 + NC6H11 1e+011 17060 0 688 OH + NC6H12 −→ H2O + NC6H11 2e+013 2600 0 689 O + NC6H12 −→ OH + NC6H11 4e+013 4000 0 690 H + NC6H11 −→ NC6H12 1e+013 0 0 691 O2 + C6H6 −→ HO2 + C6H5 6.3e+013 60000 0 692 HO2 + C6H6 −→ H2O2 + C6H5 1.52e+011 17000 0 693 OH + C6H6 −→ H2O + C6H5 2.11e+013 4570 0 694 H + C6H6 −→ H2 + C6H5 2.5e+014 16000 0 695 O + C6H6 −→ H + C6H5O 35.6 940 3.8 696 CH3 + C6H6 −→ CH4 + C6H5 2e+012 15000 0 697 C2H5 + C6H6 −→ C2H6 + C6H5 2e+012 15000 0 698 C2H3 + C6H6 −→ C2H4 + C6H5 2e+012 15000 0 699 NC3H5 + C6H6 −→ C3H6 + C6H5 2e+012 15000 0 700 NC4H5 + C6H6 −→ NC4H6 + C6H5 2e+012 15000 0 701 H + C6H5 −→ C6H6 2.2e+014 0 0 702 NC5H11 −→ C2H4 + NC3H7 4.5e+013 28700 0 703 NC5H11 −→ H + NC5H10 4.26e+013 38600 0 704 NC5H11 −→ C2H5 + C3H6 4.5e+013 29700 0 705 H + NC5H11 −→ H2 + NC5H10 5e+013 0 0 706 O2 + NC5H11 −→ HO2 + NC5H10 2.759e+010 -2000 0 707 OH + NC5H11 −→ H2O + NC5H10 2.43e+013 0 0 708 O + NC5H11 −→ CH2O + NC4H9 5e+014 0 0 709 HO2 + NC5H11 −→ OH + NC3H7 + CH3HCO 2.43e+013 0 0 Continued on next page

300 Table A.1 – Continued from previous page No. Reaction A Ea n 710 OH + NC5H11 −→ H2O + NC5H10 4.8e+012 2000 0 711 NC5H10 −→ C2H5 + NC3H5 1e+016 70746 0 712 NC5H10 −→ C2H3 + NC3H7 1.376e+016 93720 0 713 H + NC5H10 −→ H2 + NC5H9 1.95e+013 4445 0 714 CH3 + NC5H10 −→ CH4 + NC5H9 1e+011 7300 0 715 C2H5 + NC5H10 −→ C2H6 + NC5H9 1e+011 7300 0 716 O2 + NC5H10 −→ HO2 + NC5H9 2.7e+013 37000 0 717 OH + NC5H10 −→ H2O + NC5H9 2.25e+013 2217 0 718 HO2 + NC5H10 −→ H2O2 + NC5H9 1e+011 17060 0 719 O + NC5H10 −→ OH + NC5H9 9.615e+012 1968 0 720 H + NC5H10 −→ H2 + NC5H9 1.95e+013 4445 0 721 H + NC5H9 −→ NC5H10 1e+013 0 0 722 NC5H9 −→ CH3 + NC4H6 2.5e+013 45000 0 723 NC5H9 −→ H + NC5H8 8e+013 47000 0 724 H + NC5H9 −→ H2 + NC5H8 3.16e+013 0 0 725 CH3 + NC5H9 −→ CH4 + NC5H8 1e+013 0 0 726 C2H5 + NC5H9 −→ C2H6 + NC5H8 4e+012 0 0 727 O2 + NC5H9 −→ HO2 + NC5H8 2.1e+009 0 0 728 C2H3 + NC5H9 −→ C2H4 + NC5H8 4e+012 0 0 729 C2H5 + NC5H9 −→ C2H6 + NC5H8 4e+012 0 0 730 NC3H5 + NC5H9 −→ C3H6 + NC5H8 4e+012 0 0 731 NC5H8 −→ C2H2 + C3H6 2.5e+013 62970 0 732 H + NC5H8 −→ H2 + NC5H7 7e+006 0 2 733 H + NC5H8 −→ C2H4 + NC3H5 3.35e+008 2000 1.5 734 CH3 + NC5H8 −→ CH4 + NC5H7 2e+014 22800 0 735 HO2 + NC5H8 −→ OH + CH2O + NC4H6 5e+011 19000 0 736 HO2 + NC5H8 −→ H2O2 + NC5H7 2e+011 12600 0 737 OH + NC5H8 −→ CH3HCO + NC3H5 6e+012 -393 0 738 OH + NC5H8 −→ H2O + NC5H7 7e+006 0 2 739 O + NC5H8 −→ CH2O + NC4H6 4.5e+008 -858 1.45 740 H + NC5H7 −→ NC5H8 1e+014 0 0 741 NC5H7 −→ H + CPD 2e+011 18000 0 742 H + CPD −→ H2 + NC5H5 2.19e+008 3000 1.77 743 H + CPD −→ C3H3 + C2H4 3e+012 0 0 744 CH3 + CPD −→ CH4 + NC5H5 3.11e+011 5500 0 745 C2H5 + CPD −→ C2H6 + NC5H5 3.11e+011 5500 0 746 O2 + CPD −→ HO2 + NC5H5 2e+013 25000 0 747 HO2 + CPD −→ H2O2 + NC5H5 2e+012 11660 0 748 OH + CPD −→ H2O + NC5H5 3.43e+009 -447 1.18 749 O + CPD −→ OH + NC5H5 1.81e+013 3080 0 750 C2H3 + CPD −→ C2H4 + NC5H5 3.11e+011 5500 0 Continued on next page

301 Table A.1 – Continued from previous page No. Reaction A Ea n 751 NC4H5 + CPD −→ NC4H6 + NC5H5 3.11e+011 5500 0 752 C6H5 + CPD −→ C6H6 + NC5H5 3.11e+011 5500 0 753 O2 + CPD −→ CH2CO + C3H4O 3e+008 3000 0 754 H + NC5H5 −→ CPD 2e+014 0 0 755 OH + NC5H5 −→ H + C5H4OH 3e+014 0 0 756 OH + NC5H5 −→ CO + NC4H6 1.5e+014 0 0 757 O + NC5H5 −→ CO + NC4H5 2e+014 0 0 758 C5H4OH −→ H + C5H4O 2.1e+013 48000 0 759 O2 + CO −→ O + CO2 2.49979e+012 47801.1 0 760 O + CO2 −→ O2 + CO 1.69983e+013 52581.3 0 761 HO2 + CO −→ OH + CO2 1.5e+014 23650 0 762 OH + CO −→ H + CO2 4.4e+006 -740 1.5 763 H + CO2 −→ OH + CO 1.6e+014 26291.5 0 764 N2 + O + CO −→ N2 + CO2 2.83e+013 -4540 0

302 APPENDIX B

SPECIATION OF LIQUID FUELS

This appendix comprises the speciation of the diesel and bio-diesel fuels used in this dissertation. The main purpose of the speciation is to have an idea of the initial compounds and their complexity. Figures B.1 and B.2 present the spectrogram in terms

of relative abundance of diesel and bio-diesel respectively. Abundance is understood

as a quantiﬁcation in terms of the hydrocarbon with the largest composition. For ex-

ample, benzene,1-butenyl (1-Phenyl-1-butene, C10H12, CAS # 824-90-8) presents the

largest concentration for diesel fuel, therefore it represents 100%. Diesel fuel presents

a wide variety of hydrocarbons as expected. Hexadecane, used in this thesis, presents

about a 60% abundance (retention time = 21.61 min). Tetradecane (2,6,10,trimethyl,

C17H36, CAS# 14905-56-7) presents the largest second composition.

For bio-diesel (B-100), 9,12-Octadecadienoic acid, C19H34O2 (Linoleic acid - methyl

ester, CAS # 112-63-0), presents the largest composition. Its second largest compound

is 9-Octadecenoic acid, C19H36O2 (methyl9-octadecenoate, CAS # 2462-84-2), while

the third one is Hexadecanoic acid, C17H34O2 (Palmitic acid - methyl ester, CAS #

112-39-0).

303 304

Figure B.1. Liquid Speciation: GC-MS Diesel Fuel BP ECD-1 305

Figure B.2. Liquid Speciation: GC-MS Bio-Diesel Fuel B-100 APPENDIX C

SPECIATION OF REFORMATE

This appendix comprises the speciation of the reformate via GC-TID from the three diﬀerent fuels tested in the experimental section Chapter 6: hexadecane, diesel and bio-diesel. The speciation was carried out at the Transportation Research Center.

Explanation of these compounds and their concentrations is given in Chapter 7.

C.1 Hexadecane Cases

Class Compound CAS # H1 H2 p Methane 00074-82-8 31.13 20.23 o Ethylene 00074-85-1 1.78 5.68 o Acetylene (Ethyne) 00074-86-2 53.59 62.31 p Ethane 00074-84-0 0.00 0.00 o Propene 00115-07-1 0.00 0.05 p Propane 00074-98-6 0.13 0.00 o Allene (Propadiene) 00463-49-0 0.24 0.63 o Propyne 00074-99-7 0.00 0.00 I 2-Methylpropane 00075-28-5 0.00 0.00 o 2-Methylpropene & 1-Butene 00115-11-7 0.04 0.00 +00106-98-9 o 1,3-Butadiene 00106-99-0 0.06 0.06 p n-Butane 00106-97-8 0.00 0.00 Continued on next page

Table C.1. Speciation for Hexadecane

306 Table C.1 – Continued from previous page Class Compound CAS # H1 H2 i 2,2-Dimethylpropane 00463-82-1 0.00 0.93 o t-2-Butene 00624-64-6 0.00 0.00 o 1-Butyne 00107-00-6 0.00 0.00 o c-2-Butene 00590-18-1 0.00 0.00 o 3-Methyl-1-butene 00563-45-1 0.00 0.00 p Methane 00074-82-8 31.13 20.23 o Ethylene 00074-85-1 1.78 5.68 o Acetylene (Ethyne) 00074-86-2 53.59 62.31 p Ethane 00074-84-0 0.00 0.00 o Propene 00115-07-1 0.00 0.05 p Propane 00074-98-6 0.13 0.00 o Allene (Propadiene) 00463-49-0 0.24 0.63 o Propyne 00074-99-7 0.00 0.00 I 2-Methylpropane 00075-28-5 0.00 0.00 o 2-Methylpropene & 1-Butene 00115-11-7 0.04 0.00 +00106-98-9 o 1,3-Butadiene 00106-99-0 0.06 0.06 p n-Butane 00106-97-8 0.00 0.00 i 2,2-Dimethylpropane 00463-82-1 0.00 0.93 o t-2-Butene 00624-64-6 0.00 0.00 o 1-Butyne 00107-00-6 0.00 0.00 o c-2-Butene 00590-18-1 0.00 0.00 o 3-Methyl-1-butene 00563-45-1 0.00 0.00 i 2-Methylbutane (Isopentane) 00078-78-4 0.00 0.02 o 1-Pentene & 2-Butyne 00109-67-1 0.00 0.01 +00503-17-3 o 2-Methyl-1-butene 00563-46-2 0.00 0.00 p n-Pentane 00109-66-0 0.00 0.02 o 2-Methyl-1,3-butadiene 00078-79-5 0.00 0.00 o t-2-Pentene 00646-04-8 0.00 0.00 o 3,3-Dimethyl-1-butene 00558-37-2 0.00 0.00 o c-2-Pentene 00627-20-3 0.00 0.00 o 2-Methyl-2-butene 00513-35-9 0.00 0.00 o Cyclopentadiene 00542-92-7 0.04 0.22 i 2,2-Dimethylbutane 00075-83-2 0.00 0.00 o Cyclopentene 00142-29-0 0.00 0.00 o 3 & 4-Methyl-1-Pentenes 00691-37-23 0.01 0.00 +00760-20- n Cyclopentane 00287-92-3 0.00 0.00 e MTBE 01634-04-4 0.00 0.00 Continued on next page

307 Table C.1 – Continued from previous page Class Compound CAS # H1 H2 i 2,3-Dimethylbutane 00079-29-8 0.00 0.00 i 2-MePentane & 4-Me-c-2-Pentene 00691-38-3 0.00 0.00 +00107-83-5 o 4-Methyl-t-2-pentene 00674-76-0 0.00 0.00 i 3-Methylpentane 00096-14-0 0.00 0.00 o 2-Methyl-1-pentene & 1-Hexene 00592-41-6 0.00 0.00 +00763-29-1 p n-Hexane 00110-54-3 0.00 0.00 o t-3-Hexene & c-3-Hexene 13269-52-8 0.00 0.00 +07642-09-3 o t-2-Hexene 04050-45-7 0.00 0.02 o 3-Methyl-t-2-pentene 00616-12-6 0.00 0.00 o 2-Methyl-2-pentene 00625-27-4 0.00 0.00 o c-2-Hexene & 3-MeCyclopentene 07688-21-3 0.00 0.00 +01120-62-3 e ETBE 00637-92-3 0.00 0.00 o 3-Methyl-c-2-pentene 00922-62-3 0.00 0.00 i 2,2-Dimethylpentane 00590-35-2 0.00 0.00 n Methylcyclopentane 00096-37-7 0.00 0.01 i 2,4-Dimethylpentane 00108-08-7 0.00 0.00 i 2,2,3-Trimethylbutane 00464-06-2 0.00 0.00 o 1-Methylcyclopentene 00693-89-0 0.00 0.00 a Benzene 00071-43-2 0.62 5.98 i 3,3-Dimethylpentane 00562-49-2 0.00 0.00 o 3-Me-1-Hexene 03404-61-3 0.00 0.00 n Cyclohexane 00110-82-7 0.00 0.00 i 2-Methylhexane 00591-76-4 0.00 0.00 i 2,3-Dimethylpentane 00565-59-3 0.00 0.00 i Cyclohexene & 3-Methylhexane 00110-83-8 0.00 0.00 +00589-34-4 x Unknown #3 0.00 0.00 n c-1,3-Dimethylcyclopentane 02532-58-3 0.00 0.00 n t-1,2-Dimethylcyclopentane 00822-50-4 0.00 0.21 i 2,2,4-TriMePentane (IsoOctane) 00540-84-1 0.00 0.00 o 1-Heptene 00592-76-7 0.00 0.00 o t-3-Heptene 14686-14-7 0.00 0.00 p n-Heptane 00142-82-5 0.00 0.00 o 2-Methyl-2-Hexene & c-3-Heptene 02738-19-4 0.00 0.00 +07642-10-6 o 3-Me-t-3-Hexene & t-2-Heptene 03899-36-3 0.00 0.00 +14686-13-6 o 3-Ethyl-c-2-Pentene 00816-79-5 0.00 0.00 Continued on next page

308 Table C.1 – Continued from previous page Class Compound CAS # H1 H2 o 2,4,4-TMe-1- & 2,3-DMe-2-Pentene 00107-39-1 0.00 0.00 +10574-37-5 o c-2-Heptene 06443-92-1 0.00 0.00 x Unknown #4 0.00 0.00 i 2,2-DiMeHexane 00590-73-8 0.00 0.00 n Methylcyclohexane 00108-87-2 0.00 0.00 o 2,4,4-Trimethyl-2-Pentene 00107-40-4 0.00 0.00 i 2,5-DiMeHexane & EtCyPentane 00592-13-2 0.00 0.00 +01640-89-7 i 2,4-Dimethylhexane 00589-43-5 0.00 0.00 i 3,3-Dimethylhexane 00563-16-6 0.00 0.00 i 2,3,4-Trimethylpentane 00565-75-3 0.00 0.00 i 2,3,3-Trimethylpentane 00560-21-4 0.00 0.00 a Toluene 00108-88-3 0.07 0.13 i 2,3-DiMeHexane & 2,3-MeEtPentane 00584-94-1+ 0.00 0.00 i 2-Methylheptane 00592-27-8 0.00 0.00 i 1-MeCyHexene & 4-MeHeptane 00591-49-1 0.00 0.00 +00589-53-7 x Unknown #5 0.00 0.00 i 3-Methylheptane 00589-81-1 0.00 0.00 n 1c-2t-3-TriMeCyPentane 15890-40-1 0.00 0.00 n c-1,3-Dimethylcyclohexane 00638-04-0 0.00 0.00 n t-1,4-Dimethylcyclohexane 02207-04-7 0.00 0.00 i 2,2,5-Trimethylhexane 03522-94-9 0.00 0.00 o 1-Octene 00111-66-0 0.01 0.00 n 1,1-Dimethylcyclohexane 00590-66-9 0.00 0.00 x Unknown #6 . 0.00 0.00 o t-4-Octene 14850-23-8 0.01 0.00 x Unknown #7 0.00 0.00 p n-Octane 00111-65-9 0.00 0.00 n t-2-Octene & t-1,2-DiMeCyHexane 13389-42-9 0.00 0.00 +06876-23-9 n t-1,3 & c-1,4-DiMeCyHexane 02207-03-6 0.00 0.00 +00624-29-3 o c-2-Octene 07642-04-8 0.00 0.00 i 2,3,5-Trimethylhexane 01069-53-0 0.00 0.00 i 2,4-Dimethylheptane 02213-23-2 0.00 0.00 x Unknown #8 0.75 0.00 n c-1,2-Dimethylcyclohexane 02207-01-4 0.00 0.01 n Ethylcyclohexane 01678-91-7 0.00 0.00 i 3,5-Dimethylheptane 00926-82-9 0.00 0.00 x Unknown #9 0.00 0.00 Continued on next page

309 Table C.1 – Continued from previous page Class Compound CAS # H1 H2 x Unknown #10 0.00 0.00 x Unknown #11 0.00 0.00 a Ethylbenzene 00100-41-4 0.01 0.00 i 2-MeOctane & 2,3-DiMeHeptane 03074-71-3 0.00 0.00 +03221-61-2 a meta- & para-Xylenes 00108-38-3 0.17 0.59 +00106-42-3 i 4-Methyloctane 02216-34-4 0.00 0.00 i 3-Methyloctane 02216-33-3 0.00 0.00 x Unknown #12 0.00 0.00 a Styrene 00100-42-5 0.19 0.07 x Unknown #13 0.00 0.00 a ortho-Xylene 00095-47-6 0.01 0.00 o 1-Nonene 00124-11-8 0.04 0.00 o c- & t-4-Nonene 02198-23-4 0.00 0.00 p n-Nonane 00111-84-2 0.00 0.00 o t-2-Nonene 06434-78-2 0.00 0.00 a Isopropylbenzene (Cumene) 00098-82-8 0.00 0.00 i 2,2-Dimethyloctane 15869-87-1 0.00 0.00 x Unknown #14 0.00 0.00 i 2,4-DiMeOctane+AlBenz+PrCyHexane 04032-94-4 0.02 0.00 +00300-57-2 +01678-92-8 x Unknown #15 0.00 0.00 a n-Propylbenzene 00103-65-1 0.00 0.00 a 1-Methyl-3-Ethylbenzene 00620-14-4 0.00 0.00 a 1-Methyl-4-Ethylbenzene 00622-96-8 0.00 0.00 a 1,3,5-Trimethylbenzene 00108-67-8 0.00 0.00 x Unknown #16 0.01 0.00 x Unknown #17 0.00 0.00 a 1-Ethyl-2-Methylbenzene 00611-14-3 0.00 0.00 i 3-Methylnonane 0.00 0.00 a 1,2,4-TriMeBenz & t-Butylbenzene 00095-63-6 0.13 0.00 +00098-06-6 p n-Decane 00124-18-5 0.05 0.00 a Isobutylbenzene 00538-93-2 0.00 0.00 a sec-Butylbenzene 00135-98-8 0.00 0.00 a 1-Methyl-4-Isobutylbenzene 05161-04-6 0.02 0.00 a 1,2,3-Trimethylbenzene 00526-73-8 0.00 0.00 a 4-Isopropyltoluene (p-Cymene) 00099-87-6 0.00 0.00 a Indan 00496-11-7 0.01 0.00 a 1,3-Diethylbenzene 00141-93-5 0.00 0.00 Continued on next page

310 Table C.1 – Continued from previous page Class Compound CAS # H1 H2 a 1-Methyl-3-Propylbenzene 01074-43-7 0.00 0.00 a 1,4-Diethylbenzene 00105-05-5 0.01 0.00 a 1,2-Diethylbenzene 00135-01-3 0.00 0.00 a n-Butylbenzene 00104-51-8 0.00 0.00 a 1-Methyl-2-Propylbenzene 01074-17-5 0.00 0.00 a 1,4-Dimethyl-2-Ethylbenzene 01758-88-9 0.03 0.00 a 1,3-Dimethyl-4-Ethylbenzene 00874-41-9 0.03 0.00 a 1,2-Dimethyl-4-Ethylbenzene 00934-80-5 0.02 0.00 a 1,3-Dimethyl-2-Ethylbenzene 02870-04-4 0.03 0.00 o 1-Undecene 00821-95-4 0.30 0.00 p n-Undecane 01120-21-4 0.16 0.00 x Unknown #18 0.00 0.00 x Unknown #19 0.00 0.00 a 1,2,4,5-Tetramethylbenzene 00095-93-2 0.00 0.00 a 1,2,3,5-Tetramethylbenzene 00527-53-7 0.11 0.00 x Unknown #20 0.00 0.00 x Unknown #21 0.00 0.00 a Methylindan 27133-93-3 0.08 0.00 a 1,3-Diisopropylbenzene 00099-62-7 0.32 0.00 a 1,2,3,4-TetMeBenzene & Amylbenz 00488-23-3 0.26 0.01 +00538-68-1 x Unknown #22 0.12 0.00 x Unknown #23 0.20 0.00 a 1,4-Diisopropylbenzene 00100-18-5 0.08 0.00 x Unknown #24 0.09 0.00 a Naphthalene 00091-20-3 5.12 0.18 o 1-Dodecene 00112-41-4 0.04 0.00 x Unknown #25 0.10 0.00 x Unknown #26 0.09 0.00 p n-Dodecane 00112-40-3 0.60 0.00 x Others 2.78 2.39 e Methanol 0.29 0.19 e Ethanol 0.00 0.00

311 C.2 Diesel Cases

Class Compound CAS # D1 D2 D3 p Methane 00074-82-8 19.68 27.86 19.66 o Ethylene 00074-85-1 26.05 1.40 11.39 o Acetylene (Ethyne) 00074-86-2 45.31 64.90 0.00 p Ethane 00074-84-0 0.00 0.00 49.18 o Propene 00115-07-1 0.04 0.00 0.00 p Propane 00074-98-6 0.00 0.00 0.16 o Allene (Propadiene) 00463-49-0 0.27 0.36 0.30 o Propyne 00074-99-7 0.59 0.00 0.67 I 2-Methylpropane 00075-28-5 0.00 0.00 0.00 o 2-Methylpropene & 1-Butene 00115-11-7 0.03 0.02 0.01 +00106-98-9 o 1,3-Butadiene 00106-99-0 0.00 0.00 0.16 p n-Butane 00106-97-8 0.00 0.00 0.00 i 2,2-Dimethylpropane 00463-82-1 0.35 0.16 0.92 o t-2-Butene 00624-64-6 0.00 0.00 0.00 o 1-Butyne 00107-00-6 0.00 0.00 0.00 o c-2-Butene 00590-18-1 0.00 0.00 0.00 o 3-Methyl-1-butene 00563-45-1 0.00 0.00 0.00 i 2-Methylbutane (Isopentane) 00078-78-4 0.02 0.00 0.00 o 1-Pentene & 2-Butyne 00109-67-1 0.01 0.01 0.00 +00503-17-3 o 2-Methyl-1-butene 00563-46-2 0.00 0.00 0.00 p n-Pentane 00109-66-0 0.01 0.00 0.01 o 2-Methyl-1,3-butadiene 00078-79-5 0.00 0.00 0.00 o t-2-Pentene 00646-04-8 0.00 0.00 0.00 o 3,3-Dimethyl-1-butene 00558-37-2 0.00 0.00 0.00 o c-2-Pentene 00627-20-3 0.00 0.00 0.00 o 2-Methyl-2-butene 00513-35-9 0.00 0.00 0.00 o Cyclopentadiene 00542-92-7 0.16 0.04 0.32 i 2,2-Dimethylbutane 00075-83-2 0.00 0.00 0.00 o Cyclopentene 00142-29-0 0.00 0.00 0.00 o 3 & 4-Methyl-1-Pentenes 00691-37-2 0.00 0.00 0.00 +00760-20-3 n Cyclopentane 00287-92-3 0.00 0.00 0.00 e MTBE 01634-04-4 0.00 0.00 0.00 i 2,3-Dimethylbutane 00079-29-8 0.00 0.00 0.00 Continued on next page

Table C.2. Speciation for Diesel ECD-1

312 Table C.2 – Continued from previous page Class Compound CAS # D1 D2 D3 i 2-MePentane & 4-Me-c-2-Pentene 00691-38-3 0.00 0.00 0.00 +00107-83-5 o 4-Methyl-t-2-pentene 00674-76-0 0.00 0.00 0.00 i 3-Methylpentane 00096-14-0 0.00 0.00 0.00 o 2-Methyl-1-pentene & 1-Hexene 00592-41-6 0.00 0.00 0.00 +00763-29-1 p n-Hexane 00110-54-3 0.00 0.01 0.00 o t-3-Hexene & c-3-Hexene 13269-52-8 0.00 0.01 0.01 +07642-09-3 o t-2-Hexene 04050-45-7 0.02 0.00 0.00 o 3-Methyl-t-2-pentene 00616-12-6 0.00 0.00 0.00 o 2-Methyl-2-pentene 00625-27-4 0.00 0.01 0.00 o c-2-Hexene & 3-MeCyclopentene 07688-21-3 0.00 0.00 0.00 +01120-62-3 e ETBE 00637-92-3 0.00 0.00 0.00 o 3-Methyl-c-2-pentene 00922-62-3 0.00 0.00 0.00 i 2,2-Dimethylpentane 00590-35-2 0.00 0.00 0.00 n Methylcyclopentane 00096-37-7 0.00 0.00 0.00 i 2,4-Dimethylpentane 00108-08-7 0.00 0.00 0.01 i 2,2,3-Trimethylbutane 00464-06-2 0.00 0.00 0.00 o 1-Methylcyclopentene 00693-89-0 0.00 0.00 0.00 a Benzene 00071-43-2 4.55 2.29 10.35 i 3,3-Dimethylpentane 00562-49-2 0.00 0.00 0.00 o 3-Me-1-Hexene 03404-61-3 0.00 0.00 0.00 n Cyclohexane 00110-82-7 0.00 0.00 0.00 i 2-Methylhexane 00591-76-4 0.00 0.00 0.00 i 2,3-Dimethylpentane 00565-59-3 0.00 0.00 0.00 i Cyclohexene & 3-Methylhexane 00110-83-8 0.00 0.00 0.00 +00589-34-4 x Unknown #3 0.00 0.00 0.00 n c-1,3-Dimethylcyclopentane 02532-58-3 0.00 0.00 0.01 n t-1,2-Dimethylcyclopentane 00822-50-4 0.02 0.02 0.00 i 2,2,4-TriMePentane (IsoOctane) 00540-84-1 0.00 0.00 0.00 o 1-Heptene 00592-76-7 0.00 0.00 0.00 o t-3-Heptene 14686-14-7 0.00 0.00 0.00 p n-Heptane 00142-82-5 0.00 0.00 0.00 o 2-Methyl-2-Hexene & c-3-Heptene 02738-19-4 0.00 0.00 0.00 +07642-10-6 o 3-Me-t-3-Hexene & t-2-Heptene 03899-36-3 0.00 0.00 0.00 +14686-13-6 Continued on next page

313 Table C.2 – Continued from previous page Class Compound CAS # D1 D2 D3 o 3-Ethyl-c-2-Pentene 00816-79-5 0.00 0.00 0.00 o 2,4,4-TMe-1- & 2,3-DMe-2-Pentene 00107-39-1 0.00 0.00 0.00 +10574-37-5 o c-2-Heptene 06443-92-1 0.00 0.00 0.00 x Unknown #4 0.00 0.00 0.00 i 2,2-DiMeHexane 00590-73-8 0.00 0.00 0.00 n Methylcyclohexane 00108-87-2 0.00 0.00 0.00 o 2,4,4-Trimethyl-2-Pentene 00107-40-4 0.00 0.00 0.00 i 2,5-DiMeHexane & EtCyPentane 00592-13-2 0.00 0.00 0.00 +01640-89-7 i 2,4-Dimethylhexane 00589-43-5 0.00 0.00 0.00 i 3,3-Dimethylhexane 00563-16-6 0.00 0.00 0.00 i 2,3,4-Trimethylpentane 00565-75-3 0.00 0.00 0.00 i 2,3,3-Trimethylpentane 00560-21-4 0.00 0.00 0.00 a Toluene 00108-88-3 0.11 0.03 0.47 i 2,3-DiMeHexane & 2,3-MeEtPentane 00584-94-1+ 0.00 0.00 0.00 i 2-Methylheptane 00592-27-8 0.00 0.00 0.00 i 1-MeCyHexene & 4-MeHeptane 00591-49-1 0.00 0.00 0.00 +00589-53-7 x Unknown #5 0.00 0.00 0.00 i 3-Methylheptane 00589-81-1 0.00 0.00 0.00 n 1c-2t-3-TriMeCyPentane 15890-40-1 0.00 0.00 0.00 n c-1,3-Dimethylcyclohexane 00638-04-0 0.00 0.00 0.00 n t-1,4-Dimethylcyclohexane 02207-04-7 0.00 0.00 0.00 i 2,2,5-Trimethylhexane 03522-94-9 0.00 0.00 0.00 o 1-Octene 00111-66-0 0.00 0.00 0.00 n 1,1-Dimethylcyclohexane 00590-66-9 0.00 0.00 0.00 x Unknown #6 0.00 0.00 0.00 o t-4-Octene 14850-23-8 0.00 0.00 0.00 x Unknown #7 0.00 0.00 0.00 p n-Octane 00111-65-9 0.00 0.00 0.00 n t-2-Octene & t-1,2-DiMeCyHexane 13389-42-9 0.00 0.00 0.00 +06876-23-9 n t-1,3 & c-1,4-DiMeCyHexane 02207-03-6 0.00 0.00 0.00 +00624-29-3 o c-2-Octene 07642-04-8 0.00 0.00 0.00 i 2,3,5-Trimethylhexane 01069-53-0 0.00 0.01 0.00 i 2,4-Dimethylheptane 02213-23-2 0.00 0.00 0.00 x Unknown #8 0.00 0.00 0.02 n c-1,2-Dimethylcyclohexane 02207-01-4 0.02 0.17 0.00 n Ethylcyclohexane 01678-91-7 0.00 0.00 0.00 i 3,5-Dimethylheptane 00926-82-9 0.00 0.00 0.00 Continued on next page

314 Table C.2 – Continued from previous page Class Compound CAS # D1 D2 D3 x Unknown #9 0.00 0.00 0.00 x Unknown #10 0.00 0.00 0.00 x Unknown #11 0.00 0.00 0.00 a Ethylbenzene 00100-41-4 0.00 0.00 0.00 i 2-MeOctane & 2,3-DiMeHeptane 03074-71-3 0.00 0.00 0.00 +03221-61-2 a meta- & para-Xylenes 00108-38-3 0.47 0.12 1.25 +00106-42-3 i 4-Methyloctane 02216-34-4 0.00 0.00 0.00 i 3-Methyloctane 02216-33-3 0.00 0.00 0.00 x Unknown #12 0.00 0.00 0.00 a Styrene 00100-42-5 0.05 0.01 0.55 x Unknown #13 0.00 0.00 0.00 a ortho-Xylene 00095-47-6 0.00 0.00 0.01 o 1-Nonene 00124-11-8 0.00 0.01 0.00 o c- & t-4-Nonene 02198-23-4 0.00 0.00 0.00 p n-Nonane 00111-84-2 0.00 0.00 0.00 o t-2-Nonene 06434-78-2 0.00 0.00 0.00 a Isopropylbenzene (Cumene) 00098-82-8 0.00 0.00 0.00 i 2,2-Dimethyloctane 15869-87-1 0.00 0.00 0.00 x Unknown #14 0.00 0.00 0.00 i 2,4-DiMeOctane+AlBenz+PrCyHexane 04032-94-4 0.00 0.00 0.01 +00300-57-2 +01678-92-8 x Unknown #15 0.00 0.00 0.00 a n-Propylbenzene 00103-65-1 0.00 0.00 0.00 a 1-Methyl-3-Ethylbenzene 00620-14-4 0.00 0.00 0.00 a 1-Methyl-4-Ethylbenzene 00622-96-8 0.00 0.00 0.00 a 1,3,5-Trimethylbenzene 00108-67-8 0.00 0.00 0.00 x Unknown #16 0.00 0.00 0.02 x Unknown #17 0.00 0.00 0.00 a 1-Ethyl-2-Methylbenzene 00611-14-3 0.00 0.00 0.01 i 3-Methylnonane 0.00 0.00 0.00 a 1,2,4-TriMeBenz & t-Butylbenzene 00095-63-6 0.00 0.01 0.02 +00098-06-6 p n-Decane 00124-18-5 0.00 0.01 0.00 a Isobutylbenzene 00538-93-2 0.00 0.00 0.00 a sec-Butylbenzene 00135-98-8 0.00 0.00 0.00 a 1-Methyl-4-Isobutylbenzene 05161-04-6 0.00 0.00 0.00 a 1,2,3-Trimethylbenzene 00526-73-8 0.00 0.00 0.00 a 4-Isopropyltoluene (p-Cymene) 00099-87-6 0.00 0.00 0.00 a Indan 00496-11-7 0.00 0.00 0.00 Continued on next page

315 Table C.2 – Continued from previous page Class Compound CAS # D1 D2 D3 a 1,3-Diethylbenzene 00141-93-5 0.00 0.00 0.00 a 1-Methyl-3-Propylbenzene 01074-43-7 0.00 0.00 0.01 a 1,4-Diethylbenzene 00105-05-5 0.00 0.01 0.01 a 1,2-Diethylbenzene 00135-01-3 0.00 0.00 0.01 a n-Butylbenzene 00104-51-8 0.00 0.00 0.01 a 1-Methyl-2-Propylbenzene 01074-17-5 0.00 0.00 0.00 a 1,4-Dimethyl-2-Ethylbenzene 01758-88-9 0.00 0.01 0.01 a 1,3-Dimethyl-4-Ethylbenzene 00874-41-9 0.00 0.00 0.00 a 1,2-Dimethyl-4-Ethylbenzene 00934-80-5 0.00 0.00 0.00 a 1,3-Dimethyl-2-Ethylbenzene 02870-04-4 0.00 0.00 0.01 o 1-Undecene 00821-95-4 0.00 0.00 0.01 p n-Undecane 01120-21-4 0.00 0.04 0.03 x Unknown #18 0.00 0.00 0.01 x Unknown #19 0.00 0.00 0.00 a 1,2,4,5-Tetramethylbenzene 00095-93-2 0.01 0.01 0.01 a 1,2,3,5-Tetramethylbenzene 00527-53-7 0.00 0.01 0.01 x Unknown #20 0.00 0.00 0.00 x Unknown #21 0.00 0.00 0.00 a Methylindan 27133-93-3 0.00 0.02 0.01 a 1,3-Diisopropylbenzene 00099-62-7 0.00 0.04 0.01 a 1,2,3,4-TetMeBenzene & Amylbenz 00488-23-3 0.00 0.04 0.04 +00538-68-1 x Unknown #22 0.01 0.05 0.01 x Unknown #23 0.00 0.03 0.03 a 1,4-Diisopropylbenzene 00100-18-5 0.00 0.02 0.01 x Unknown #24 0.00 0.02 0.01 a Naphthalene 00091-20-3 0.47 0.62 0.89 o 1-Dodecene 00112-41-4 0.00 0.01 0.00 x Unknown #25 0.00 0.00 0.01 x Unknown #26 0.00 0.02 0.01 p n-Dodecane 00112-40-3 0.01 0.12 0.05 x Others 1.54 1.26 3.13 e Methanol 0.16 0.18 0.06 e Ethanol 0.02 0.00 0.00

316 C.3 Bio-Diesel Cases

Class Compound CAS # B1 B2 B3 p Methane 00074-82-8 24.15 20.82 19.58 o Ethylene 00074-85-1 7.49 2.38 3.47 o Acetylene (Ethyne) 00074-86-2 53.10 66.21 64.00 p Ethane 00074-84-0 0.00 0.00 0.00 o Propene 00115-07-1 1.00 0.09 0.00 p Propane 00074-98-6 0.00 0.00 0.04 o Allene (Propadiene) 00463-49-0 0.92 0.53 0.02 o Propyne 00074-99-7 0.00 0.00 0.53 I 2-Methylpropane 00075-28-5 0.00 0.00 0.00 o 2-Methylpropene & 1-Butene 00115-11-7 0.10 0.02 0.00 +00106-98-9 o 1,3-Butadiene 00106-99-0 0.57 0.05 0.04 p n-Butane 00106-97-8 0.02 0.01 0.00 i 2,2-Dimethylpropane 00463-82-1 0.61 0.63 0.67 o t-2-Butene 00624-64-6 0.00 0.00 0.00 o 1-Butyne 00107-00-6 0.02 0.00 0.00 o c-2-Butene 00590-18-1 0.02 0.00 0.00 o 3-Methyl-1-butene 00563-45-1 0.00 0.00 0.00 i 2-Methylbutane (Isopentane) 00078-78-4 0.00 0.01 0.00 o 1-Pentene & 2-Butyne 00109-67-1 0.02 0.01 0.00 +00503-17-3 o 2-Methyl-1-butene 00563-46-2 0.00 0.00 0.00 p n-Pentane 00109-66-0 0.01 0.02 0.01 o 2-Methyl-1,3-butadiene 00078-79-5 0.01 0.00 0.00 o t-2-Pentene 00646-04-8 0.00 0.00 0.00 o 3,3-Dimethyl-1-butene 00558-37-2 0.00 0.00 0.00 o c-2-Pentene 00627-20-3 0.00 0.00 0.00 o 2-Methyl-2-butene 00513-35-9 0.01 0.00 0.00 o Cyclopentadiene 00542-92-7 0.66 0.13 0.16 i 2,2-Dimethylbutane 00075-83-2 0.00 0.00 0.00 o Cyclopentene 00142-29-0 0.01 0.00 0.00 o 3 & 4-Methyl-1-Pentenes 00691-37-2 0.00 0.00 0.00 +00760-20-3 n Cyclopentane 00287-92-3 0.00 0.00 0.00 e MTBE 01634-04-4 0.00 0.00 0.00 i 2,3-Dimethylbutane 00079-29-8 0.00 0.00 0.00 Continued on next page

Table C.3. Speciation for Bio-Diesel 100

317 Table C.3 – Continued from previous page Class Compound CAS # B1 B2 B3 i 2-MePentane & 4-Me-c-2-Pentene 00691-38-3 0.00 0.00 0.00 +00107-83-5 o 4-Methyl-t-2-pentene 00674-76-0 0.00 0.00 0.00 i 3-Methylpentane 00096-14-0 0.00 0.00 0.00 o 2-Methyl-1-pentene & 1-Hexene 00592-41-6 0.01 0.00 0.00 +00763-29-1 p n-Hexane 00110-54-3 0.05 0.01 0.00 o t-3-Hexene & c-3-Hexene 13269-52-8 0.00 0.00 0.02 +07642-09-3 o t-2-Hexene 04050-45-7 0.02 0.01 0.00 o 3-Methyl-t-2-pentene 00616-12-6 0.00 0.00 0.00 o 2-Methyl-2-pentene 00625-27-4 0.00 0.00 0.00 o c-2-Hexene & 3-MeCyclopentene 07688-21-3 0.00 0.00 0.00 +01120-62-3 e ETBE 00637-92-3 0.00 0.00 0.00 o 3-Methyl-c-2-pentene 00922-62-3 0.00 0.00 0.00 i 2,2-Dimethylpentane 00590-35-2 0.00 0.00 0.00 n Methylcyclopentane 00096-37-7 0.01 0.01 0.00 i 2,4-Dimethylpentane 00108-08-7 0.00 0.00 0.01 i 2,2,3-Trimethylbutane 00464-06-2 0.00 0.00 0.00 o 1-Methylcyclopentene 00693-89-0 0.00 0.00 0.00 a Benzene 00071-43-2 6.58 3.36 7.99 i 3,3-Dimethylpentane 00562-49-2 0.00 0.00 0.00 o 3-Me-1-Hexene 03404-61-3 0.00 0.00 0.00 n Cyclohexane 00110-82-7 0.00 0.00 0.00 i 2-Methylhexane 00591-76-4 0.00 0.00 0.00 i 2,3-Dimethylpentane 00565-59-3 0.00 0.00 0.00 i Cyclohexene & 3-Methylhexane 00110-83-8 0.00 0.00 0.00 +00589-34-4 x Unknown #3 0.00 0.00 0.00 n c-1,3-Dimethylcyclopentane 02532-58-3 0.01 0.00 0.00 n t-1,2-Dimethylcyclopentane 00822-50-4 0.01 0.00 0.02 i 2,2,4-TriMePentane (IsoOctane) 00540-84-1 0.00 0.00 0.00 o 1-Heptene 00592-76-7 0.00 0.00 0.00 o t-3-Heptene 14686-14-7 0.00 0.00 0.00 p n-Heptane 00142-82-5 0.00 0.00 0.00 o 2-Methyl-2-Hexene & c-3-Heptene 02738-19-4 0.00 0.00 0.00 +07642-10-6 o 3-Me-t-3-Hexene & t-2-Heptene 03899-36-3 0.00 0.00 0.00 +14686-13-6 Continued on next page

318 Table C.3 – Continued from previous page Class Compound CAS # B1 B2 B3 o 3-Ethyl-c-2-Pentene 00816-79-5 0.00 0.00 0.00 o 2,4,4-TMe-1- & 2,3-DMe-2-Pentene 00107-39-1 0.00 0.00 0.00 +10574-37-5 o c-2-Heptene 06443-92-1 0.00 0.00 0.00 x Unknown #4 0.00 0.00 0.00 i 2,2-DiMeHexane 00590-73-8 0.00 0.00 0.00 n Methylcyclohexane 00108-87-2 0.00 0.00 0.00 o 2,4,4-Trimethyl-2-Pentene 00107-40-4 0.00 0.00 0.00 i 2,5-DiMeHexane & EtCyPentane 00592-13-2 0.00 0.00 0.00 +01640-89-7 i 2,4-Dimethylhexane 00589-43-5 0.00 0.00 0.00 i 3,3-Dimethylhexane 00563-16-6 0.00 0.00 0.00 i 2,3,4-Trimethylpentane 00565-75-3 0.00 0.00 0.00 i 2,3,3-Trimethylpentane 00560-21-4 0.00 0.00 0.00 a Toluene 00108-88-3 0.30 0.03 0.07 i 2,3-DiMeHexane & 2,3-MeEtPentane 00584-94-1+ 0.00 0.00 0.00 i 2-Methylheptane 00592-27-8 0.00 0.00 0.00 i 1-MeCyHexene & 4-MeHeptane 00591-49-1 0.00 0.00 0.00 +00589-53-7 x Unknown #5 0.00 0.00 0.00 i 3-Methylheptane 00589-81-1 0.00 0.00 0.00 n 1c-2t-3-TriMeCyPentane 15890-40-1 0.06 0.03 0.01 n c-1,3-Dimethylcyclohexane 00638-04-0 0.00 0.00 0.00 n t-1,4-Dimethylcyclohexane 02207-04-7 0.00 0.00 0.00 i 2,2,5-Trimethylhexane 03522-94-9 0.00 0.00 0.00 o 1-Octene 00111-66-0 0.00 0.00 0.00 n 1,1-Dimethylcyclohexane 00590-66-9 0.00 0.00 0.00 x Unknown #6 0.00 0.00 0.00 o t-4-Octene 14850-23-8 0.00 0.00 0.00 x Unknown #7 0.00 0.00 0.00 p n-Octane 00111-65-9 0.01 0.00 0.00 n t-2-Octene & t-1,2-DiMeCyHexane 13389-42-9 0.00 0.00 0.00 +06876-23-9 n t-1,3 & c-1,4-DiMeCyHexane 02207-03-6 0.00 0.00 0.00 +00624-29-3 o c-2-Octene 07642-04-8 0.00 0.00 0.00 i 2,3,5-Trimethylhexane 01069-53-0 0.00 0.00 0.00 i 2,4-Dimethylheptane 02213-23-2 0.00 0.00 0.00 x Unknown #8 0.00 0.00 0.00 n c-1,2-Dimethylcyclohexane 02207-01-4 0.30 0.29 0.05 n Ethylcyclohexane 01678-91-7 0.00 0.00 0.00 i 3,5-Dimethylheptane 00926-82-9 0.00 0.00 0.00 Continued on next page

319 Table C.3 – Continued from previous page Class Compound CAS # B1 B2 B3 x Unknown #9 0.00 0.00 0.00 x Unknown #10 0.00 0.00 0.00 x Unknown #11 0.00 0.00 0.00 a Ethylbenzene 00100-41-4 0.01 0.00 0.00 i 2-MeOctane & 2,3-DiMeHeptane 03074-71-3 0.00 0.00 0.00 +03221-61-2 a meta- & para-Xylenes 00108-38-3 0.49 0.06 0.19 +00106-42-3 i 4-Methyloctane 02216-34-4 0.00 0.00 0.00 i 3-Methyloctane 02216-33-3 0.00 0.00 0.00 x Unknown #12 0.00 0.00 0.00 a Styrene 00100-42-5 0.13 0.03 0.03 x Unknown #13 0.00 0.00 0.00 a ortho-Xylene 00095-47-6 0.01 0.00 0.00 o 1-Nonene 00124-11-8 0.00 0.00 0.00 o c- & t-4-Nonene 02198-23-4 0.00 0.00 0.00 p n-Nonane 00111-84-2 0.00 0.00 0.00 o t-2-Nonene 06434-78-2 0.00 0.00 0.00 a Isopropylbenzene (Cumene) 00098-82-8 0.01 0.00 0.00 i 2,2-Dimethyloctane 15869-87-1 0.00 0.00 0.00 x Unknown #14 0.00 0.00 0.00 i 2,4-DiMeOctane+AlBenz+PrCyHexane 04032-94-4 0.01 0.00 0.00 +00300-57-2 +01678-92-8 x Unknown #15 0.00 0.00 0.00 a n-Propylbenzene 00103-65-1 0.00 0.00 0.00 a 1-Methyl-3-Ethylbenzene 00620-14-4 0.00 0.00 0.00 a 1-Methyl-4-Ethylbenzene 00622-96-8 0.00 0.00 0.00 a 1,3,5-Trimethylbenzene 00108-67-8 0.00 0.00 0.00 x Unknown #16 0.00 0.00 0.00 x Unknown #17 0.00 0.00 0.00 a 1-Ethyl-2-Methylbenzene 00611-14-3 0.00 0.00 0.00 i 3-Methylnonane 0.00 0.00 0.00 a 1,2,4-TriMeBenz & t-Butylbenzene 00095-63-6 0.02 0.00 0.00 +00098-06-6 p n-Decane 00124-18-5 0.00 0.00 0.00 a Isobutylbenzene 00538-93-2 0.00 0.00 0.00 a sec-Butylbenzene 00135-98-8 0.00 0.00 0.00 a 1-Methyl-4-Isobutylbenzene 05161-04-6 0.00 0.00 0.00 a 1,2,3-Trimethylbenzene 00526-73-8 0.00 0.00 0.00 a 4-Isopropyltoluene (p-Cymene) 00099-87-6 0.00 0.00 0.00 a Indan 00496-11-7 0.00 0.00 0.00 Continued on next page

320 Table C.3 – Continued from previous page Class Compound CAS # B1 B2 B3 a 1,3-Diethylbenzene 00141-93-5 0.00 0.00 0.00 a 1-Methyl-3-Propylbenzene 01074-43-7 0.00 0.00 0.00 a 1,4-Diethylbenzene 00105-05-5 0.01 0.02 0.00 a 1,2-Diethylbenzene 00135-01-3 0.00 0.00 0.00 a n-Butylbenzene 00104-51-8 0.00 0.00 0.00 a 1-Methyl-2-Propylbenzene 01074-17-5 0.00 0.00 0.00 a 1,4-Dimethyl-2-Ethylbenzene 01758-88-9 0.00 0.00 0.00 a 1,3-Dimethyl-4-Ethylbenzene 00874-41-9 0.00 0.00 0.00 a 1,2-Dimethyl-4-Ethylbenzene 00934-80-5 0.00 0.00 0.00 a 1,3-Dimethyl-2-Ethylbenzene 02870-04-4 0.01 0.04 0.00 o 1-Undecene 00821-95-4 0.01 0.00 0.00 p n-Undecane 01120-21-4 0.01 0.00 0.00 x Unknown #18 0.00 0.00 0.00 x Unknown #19 0.02 0.03 0.01 a 1,2,4,5-Tetramethylbenzene 00095-93-2 0.00 0.00 0.00 a 1,2,3,5-Tetramethylbenzene 00527-53-7 0.00 0.00 0.00 x Unknown #20 0.00 0.00 0.00 x Unknown #21 0.00 0.00 0.00 a Methylindan 27133-93-3 0.00 0.00 0.00 a 1,3-Diisopropylbenzene 00099-62-7 0.00 0.04 0.00 a 1,2,3,4-TetMeBenzene & Amylbenz 00488-23-3 0.02 0.05 0.01 +00538-68-1 x Unknown #22 0.03 0.02 0.00 x Unknown #23 0.00 0.00 0.00 a 1,4-Diisopropylbenzene 00100-18-5 0.00 0.00 0.00 x Unknown #24 0.00 0.00 0.00 a Naphthalene 00091-20-3 0.80 2.06 0.66 o 1-Dodecene 00112-41-4 0.00 0.00 0.00 x Unknown #25 0.05 0.11 0.02 x Unknown #26 0.00 0.00 0.00 p n-Dodecane 00112-40-3 0.03 0.04 0.01 x Others 2.12 2.60 2.23 e Methanol 0.14 0.26 0.12 e Ethanol 0.00 0.00 0.00

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