Prediction of NOx Emissions for a Hydrogen Fueled Industrial ...

Prediction of NOx Emissions for a Hydrogen Fueled

Industrial Gas Turbine Combustor

with Water Injection

Vorhersage der NOx Emissionen einer wasserstoffbetriebenen

Industriegasturbinenbrennkammer mit Wassereinspritzung

Von der Fakultat fur Maschinenwesen der

Rheinisch-Westfalischen Technischen Hochschule Aachen

zur Erlangung des akademischen Grades eines

Doktors der Ingenieurwissenschaften

genehmigte Dissertation

vorgelegt von

Daniel Kroniger

Berichter: Univ.-Prof. Dr.-Ing. habil. Manfred Christian Wirsum

Prof. Dr.-Ing. Thomas Sattelmayer

Tag der mundlichen Prufung: 23. Januar 2019

Diese Dissertation ist auf den Internetseiten der Universitatsbibliothek online verfugbar.

Acknowledgments

This doctoral thesis is based on studies I conducted as a researcher at the Institute of Power

Plant Technology, Steam and Gas Turbines (IKDG) at RWTH Aachen University. I would

like to take the opportunity to express my sincere thanks to all those who have contributed

to its completion.

First and foremost, I am thankful to my supervisor, Professor Manfred Wirsum, for his

supervision, the freedom of research in the field of hydrogen combustion and providing the

experimental infrastructure. I would also like to thank Professor Thomas Sattelmayer of the

Chair of Thermodynamics of Technical University of Munich, not only for being co-supervisor

of my thesis but also for being an experienced partner for scientific discussion.

My sincere thank goes to the Corporate Technology Division and the Gas Turbine Section

of the Machinery Division of Kawasaki Heavy Industries, Ltd. for the trustful collaboration

over the years. Only their financial and personnel support made this research possible. I

warmly thank the research team around Atsushi Horikawa, Kunio Okada, and Masahide

Kazari as well as Toshiaki Sakurazawa and Takeo Oda for their technical expertise and

overall support. I keep the long test days in my memory with pleasure.

High pressure combustion tests of this extent require a sophisticated test team, without

which the test campaigns could not have been run successfully. I would like to express my

gratitude to Germar Heibges for his strong commitment and his superior help with electrical

components, sensors and control systems all around the clock. I am also sincerely thankful

for the devoted support of Rene Schmitz for his assistance and Askin-Musaffer Karayazı for

construction. I grately thank Hans-Peter Nießen and his entire workshop team for their

support and the preparation of several test campaigns. I furthermore thank my intermittent

fellow Jonas Stutenkemper and the unconditional support of the students Kai Risthaus,

Robert Hausmann, Dominik Nieborg, Johannes Hirsch, Bastian Weiß, Nils Petersen and

Stephan Sagert. I also would like to thank all my colleagues for the warm atmosphere. Here,

special thanks are dedicated to Philipp Vinnemeier, Thomas Bexten and Moritz Lipperheide

for several inspiring discussions.

Last but not least I would like to thank all of my close friends Gregor, Robert, Piet, Kiki,

Yiqi, and Katja from Aachen for their great spiritual support. Lastly, I would like to hearty

thank my parents as well as my two sisters for their unconditional trust and everything they

have done for me, and Tomoka for her sympathetic support and love.

Aachen, March 2019 Daniel Kroniger

Abstract

Hydrogen is a promising alternative to fossil fuels for future gas turbines since it can be

produced using renewable energy sources and uses CO2-free combustion. However, due to

its higher reactivity (when compared to natural gas) it cannot be used with the state-of-the-

art premixed combustors developed for natural gas. Besides the risks of flame flashback and

autoignition, high hydrogen fuels tend to produce significantly higher NOx emissions due to

a the higher flame temperature. This study observes the effects of the use of a non-premixed

combustor and water injection, which can reduce emissions.

An industrial, non-premixed gas turbine combustor is measured and characterized at real

conditions with a high-pressure combustion test rig. The focus of this study is on the NOx

emissions as a function of the pressure, the hydrogen content in the fuel, and the amount

of directly-injected water. A significant increase in NOx emissions for increasing pressure

and hydrogen share is shown using the experimental combustor geometry. Furthermore, the

emissions increase at elevated air inlet temperature and reduced air inlet velocity. Water

directly-injected into the combustor significantly reduces the NOx emissions.

A numerical investigation of the combustion process within the combustor is performed with

a semi-empirical chemical reactor network model that derives the flame temperature and

residence time. The results show a good agreement with the experimentally determined

NOx emissions over the entire range of operating conditions. Besides the effect of the flame

temperature, a detailed analysis of the chemical effect of hydrogen and directly-injected

water on the combustion process shows a significant impact of the relevant radicals and NOx

formation mechanisms.

Furthermore, a simplified NOx model is derived based on the findings. With the help of

correlations for flame temperature and fluid residence time in the flame, a correlation is

developed that predicts the NOx emissions taking the substantial chemical effects of hydrogen

and water into account. This opens the possibility to estimate the NOx emissions at ‘real’

pressure conditions on the basis of low pressure combustion tests.

Kurzfassung

Wasserstoff als Gasturbinenbrennstoff der Zukunft ist ein vielversprechender Ersatz fur fossile

Energietrager, da er mit Erneuerbaren Energiequellen hergestellt werden kann und seine Ver-

brennung kein CO2 emittiert. Durch seine im Vergleich zu Ergas hohere Reaktivitat kann er

jedoch nicht in den aktuell fur Erdgas verwendeten Vormischbrennkammern eingesetzt wer-

den. Neben der zunehmenden Gefahr von Flammenruckschlag und Selbstzundung neigen

wasserstoffhaltige Brennstoffe zudem aufgrund ihrer hoheren Flammentemperaturen zu sig-

nifikant hoheren NOx-Emissionen. Diesen kann bei einer nicht-vorgemischen Brennkammer

mit Wassereinspritzung begegnet werden, die in der vorliegenden Arbeit Gegenstand der

Untersuchung ist.

Mittels eines Hochdruckbrennkammerprufstands wird eine industrielle nicht-vorgemische

Gasturbinenbrennkammer unter realen Bedingungen charakterisiert. Im Fokus steht hier-

bei das NOx-Emissionsverhalten in Abhangigkeit vom Druck, vom Wasserstoffanteil im

Brennstoff und von der Menge direkt eingespritzen Wassers. Fur die vorliegende Brennkam-

mergeometrie konnte ein signifikanter Anstieg der NOx-Emissionen bei Erhohung des Drucks

und des Wasserstoffanteils nachgewiesen werden. Des Weiteren erhohen sich die Emissionen

auch bei erhohter Lufteintrittstemperaturen und Verringerung der Lufteintrittsgeschwindig-

keit. Direkt in die Brennkammer eingespritztes Wasser senkt die NOx-Emissionen signifikant.

In einer numerischen Untersuchung wird der Verbrennungsprozess innerhalb der Brenn-

kammer mittels eines semi-empirischen chemischen Reaktornetzwerks modelliert und hiermit

Flammentemperatur und Aufenthaltszeit abgeleitet. Die Ergebnisse zeigen im gesamten Be-

triebsbereich eine gute Ubereinstimmung mit den experimentell ermittelten NOx-Emissionen.

Eine detaillierte Untersuchung der chemischen Wirkung von Wasserstoff und direkt einge-

spritzten Wassers auf den Verbrennungsprozess zeigt neben dem Einfluss auf die Flammen-

temperatur wesentliche Einflusse auf relevante Radikale und NOx-Bildungsmechanismen.

Aus den gegebenen Erkenntnissen wird ferner ein vereinfachtes NOx-Modell abgeleitet. Auf

der Basis von Korrelationen fur die Flammentemperatur und die Aufenthaltszeit des Fluids in

der Flamme wird eine Korrelation ermittelt, die die NOx-Emissionen unter der Berucksichtung

der wesentlichen chemischen Einflusse von Wasserstoff und Wasser hinreichend vorhersagt.

Damit wird die Moglichkeit geschaffen, NOx-Emissionen unter ,realen‘ Druckbedingungen

auf Basis von Verbrennungsversuchen bei niedrigerem Druck abzuschatzen.

Contents

Nomenclature xiii

1 Introduction 1

2 Wet hydrogen combustion in gas turbines 4

2.1 Gas turbine emissions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4

2.2 Hydrogen as gas turbine fuel . . . . . . . . . . . . . . . . . . . . . . . . . . . 7

2.2.1 Hydrogen combustion properties . . . . . . . . . . . . . . . . . . . . . 8

2.2.2 Fuel effects on gas turbine cycle . . . . . . . . . . . . . . . . . . . . . 9

2.2.3 State of the art high hydrogen gas turbines . . . . . . . . . . . . . . . 11

2.2.4 High Hydrogen combustor development . . . . . . . . . . . . . . . . . 11

2.3 Humidification as NOx abatement measure . . . . . . . . . . . . . . . . . . . 14

2.3.1 Historical review . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 14

2.3.2 Recent wet high-hydrogen applications research . . . . . . . . . . . . 15

2.3.3 Effects of water on combustion . . . . . . . . . . . . . . . . . . . . . 17

2.4 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 18

3 Objective and methods 19

3.1 Objective . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 19

3.2 Methods . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 20

4 High pressure combustion tests 23

4.1 Worldwide test rig overview . . . . . . . . . . . . . . . . . . . . . . . . . . . 23

4.2 Test rig set up . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 25

4.2.1 Air supply system . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 25

4.2.2 Fuel supply system . . . . . . . . . . . . . . . . . . . . . . . . . . . . 26

4.2.3 Water injection system . . . . . . . . . . . . . . . . . . . . . . . . . . 27

4.2.4 Combustion system . . . . . . . . . . . . . . . . . . . . . . . . . . . . 27

4.2.5 Measurement section . . . . . . . . . . . . . . . . . . . . . . . . . . . 28

4.2.6 Exhaust gas path . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 29

4.2.7 Control principles . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 30

4.3 Measurement techniques . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 30

4.3.1 Pressure Measurement . . . . . . . . . . . . . . . . . . . . . . . . . . 31

4.3.2 Temperature measurement . . . . . . . . . . . . . . . . . . . . . . . . 32

4.3.3 Mass flow measurement . . . . . . . . . . . . . . . . . . . . . . . . . 34

x Contents

4.3.4 Emission analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 35

4.4 Test combustor . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 36

4.5 Operation parameter and variation . . . . . . . . . . . . . . . . . . . . . . . 38

4.6 Experimental results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 40

4.6.1 Air fuel ratio verification . . . . . . . . . . . . . . . . . . . . . . . . . 40

4.6.2 Sample gas temperature . . . . . . . . . . . . . . . . . . . . . . . . . 42

4.6.3 Pressure loss of combustor . . . . . . . . . . . . . . . . . . . . . . . . 45

4.6.4 NOx emissions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 46

4.6.5 CO emissions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 53

4.7 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 57

5 Chemical reactor network model 59

5.1 Reactor model review . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 59

5.2 Methods . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 62

5.2.1 Ideal reactors and reactor networks . . . . . . . . . . . . . . . . . . . 63

5.2.2 Reaction kinetic and mechanisms . . . . . . . . . . . . . . . . . . . . 64

5.2.3 NOx formation mechanisms . . . . . . . . . . . . . . . . . . . . . . . 67

5.2.4 Validation and analysis methods . . . . . . . . . . . . . . . . . . . . . 68

5.2.4.1 Comparison criterion . . . . . . . . . . . . . . . . . . . . . . 69

5.2.4.2 Sensitivity analysis . . . . . . . . . . . . . . . . . . . . . . . 69

5.2.4.3 Pathway analysis . . . . . . . . . . . . . . . . . . . . . . . . 70

5.3 Chemical reactor network model . . . . . . . . . . . . . . . . . . . . . . . . . 70

5.3.1 Flame zone . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 72

5.3.1.1 General residence time formulation . . . . . . . . . . . . . . 73

5.3.1.2 Residence time for varying combustor outlet temperature . . 74

5.3.1.3 Residence time for varying pressure . . . . . . . . . . . . . . 75

5.3.2 Mixing zone . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 76

5.3.3 Burnout zone . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 76

5.4 Model validation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 77

5.4.1 Reaction scheme validation . . . . . . . . . . . . . . . . . . . . . . . . 77

5.4.2 NOx characteristic validation . . . . . . . . . . . . . . . . . . . . . . 81

5.4.2.1 Pressure and combustor outlet temperature . . . . . . . . . 81

5.4.2.2 Air inlet temperature . . . . . . . . . . . . . . . . . . . . . . 83

5.4.2.3 Air inlet velocity . . . . . . . . . . . . . . . . . . . . . . . . 84

5.4.2.4 Water injection . . . . . . . . . . . . . . . . . . . . . . . . . 86

5.4.3 Validation conclusion . . . . . . . . . . . . . . . . . . . . . . . . . . . 88

5.5 Model response analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 89

Contents xi

5.6 Chemical influence of hydrogen and water . . . . . . . . . . . . . . . . . . . 92

5.6.1 Review . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 92

5.6.2 Radical mix analysis . . . . . . . . . . . . . . . . . . . . . . . . . . . 93

5.6.2.1 NO sensitivity analysis . . . . . . . . . . . . . . . . . . . . . 94

5.6.2.2 H◦ radical analysis . . . . . . . . . . . . . . . . . . . . . . . 96

5.6.2.3 O◦ radical analysis . . . . . . . . . . . . . . . . . . . . . . . 99

5.6.2.4 OH◦ radical analysis . . . . . . . . . . . . . . . . . . . . . . 101

5.6.2.5 HO2◦ radical analysis . . . . . . . . . . . . . . . . . . . . . 104

5.6.2.6 CH◦ radical analysis . . . . . . . . . . . . . . . . . . . . . . 106

5.6.2.7 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . 108

5.6.3 NO formation analysis . . . . . . . . . . . . . . . . . . . . . . . . . . 109

5.6.3.1 Effects of radicals on NO pathways . . . . . . . . . . . . . . 110

5.6.3.2 NO pathway analysis . . . . . . . . . . . . . . . . . . . . . . 111

5.6.3.3 Conclusion . . . . . . . . . . . . . . . . . . . . . . . . . . . 115

5.7 Distribution of the effects of H2O on NOx . . . . . . . . . . . . . . . . . . . 116

5.8 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 119

6 Simplified NOx emissions model 121

6.1 NOx correlations review . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 121

6.2 Correlations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 123

6.2.1 Residence time correlation . . . . . . . . . . . . . . . . . . . . . . . . 123

6.2.2 Stoichiometric flame temperature correlation . . . . . . . . . . . . . . 125

6.2.3 NOx correlation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 126

6.3 Pressure estimation evaluation . . . . . . . . . . . . . . . . . . . . . . . . . . 129

6.4 Generalization evaluation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 132

6.5 Summary . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 133

7 Conclusion 135

A Appendix 138

B Previous Publications 139

List of Figures 140

List of Tables 143

Bibliography 145

Nomenclature

Latin Symbols

Name Unit Description

A [varies] Arrhenius pre-exponent factor

AFR [-] Air fuel ratio

B [-] Exponent factor

c [mol/mol] Concentration

cp [J/kgK] Heat capacity

d [m] Combustor outlet diameter

E [J] Arrhenius activation energy

f [-] Sensitivity factor

h [J/kg] Specific enthalpy

Hl [kWh/m3] Lower heating value

Hu [kWh/m3] Upper heating value

H [J/s] Enthalpy stream

i [-] Pressure exponent

k [varies] Reaction rate coefficient

kτ [-] Residence time factor

Kc [-] Equilibrium constant

L [m] Flame length

m [kg] Mass

M [g/mol] Molar mass

m [kg/s] Mass flow rate

n [-] Quantity

N [-] Number of reactions

NOx [ppm] NOx concentration

p [bar] Pressure

PLR [-] Pressure loss ratio

q [J/s] Heat sink

Q [mol/m3s] Progress rate

ri [-] Residence time correlation parameter

rRMSE [%] Relative root mean square error

R [J/kgK] Ideal gas constant

Re [-] Reynolds Number

t [s] Time

xiv Nomenclature


ti [-] Stoichiometric temperature correlation parameter

T [◦C] Combustor outlet temperature

Tfuel [◦C] Fuel temperature

Tair [◦C] Air inlet temperature

Tst [K] Stoichiometric flame temperature

U [J] Internal Energy

v [m/s] Air inlet velocity

V [m3] Volume

V [m3/s] Volume flow rate

w [-] Correlation parameter for water injection

x [vol.%] Mole fraction of hydrogen in fuel

X [mol/mol] Molar fraction

y [var.] Variable representative

Y [kg/kg] Mass fraction

Greek Symbols


αT [K] Temperature accuracy

ατ [ms] Residence time accuracy

β [-] Arrhenius temperature exponent

∆ [-] Difference

ε [-] Chaperon efficiency factor

λ [-] Air fuel equivalence ratio

ν [mol/mol] Molar stoichiometric coefficient

ν ′ [mol/mol] Molar stoichiometric coefficient in forward reaction

ν ′′ [mol/mol] Molar stoichiometric coefficient in reverse reaction

ω [mol/m3s] Net production rate

Φ [-] Equivalence ratio

ψ [kg/kg] Water-to-fuel ratio

ψn [kg/kg] Water-to-fuel ratio referenced to natural gas

% [kg/m3] Density

σ [-] First-order sensitivity coefficient

τ [ms] Residence time of PSR

ϑi [K] Stoichiometric temperature correlation parameter

Nomenclature xv

Subscripts

Name Description

a Air

comb Combustor

corr Correlation

CRN Chemical reactor network

exh Exhaust

evp Evaporation

exp Experiment

f Fuel

i Species index

in Inlet

init Initial

k Reaction index

NOx Nitrogen oxides

num Numeric

out Outlet

PFR Plug flow reactor

pr Primary

PSR Perfectly stirred reactor

ref Reference

rel Relative

sat Saturation

st Stoichiometric

t Turbulent

th Thermal

w Water

Abbreviations

Name Description

BR Batch reactor

CFD Computational fluid mechanics

CLD Chemiluminescence detection

CRN Chemical reactor network

EBC Environmental barrier coating

FID Flame ionization detection

FS Full scale

xvi Nomenclature

Name Description

IGCC Integrated gasification combined cycle

NDIR Nondispersive infrared sensor

NOx Nitric oxides

PFR Plug flow reactor

PLR Pressure loss ratio

PMD Paramagnetic detection

PSR Perfectly stirred reactor

rRMSE relative root mean square error

TBC Thermal barrier coatings

TIT Turbine inlet temperature

TOT Turbine outlet temperature

UHC Unburnt hydrocarbons

1 Introduction

World energy consumption is projected to grow by about 30% from 2018 to 2040. A growing

electrification of energy markets is predicted for worldwide end-uses of energy, with elec-

tricity generation making up 40% of the rise in consumption by 2040. Because of this, the

International Energy Agency (2017) expects rapid deployment and falling costs of renewable

energy technologies. By 2040, the combination of all renewable energies is predicted to make

up 40% of total global power generation, thereby limiting greenhouse gas emissions.

Renewableenergies

Many renewable energy sources, in particular wind turbines and photovoltaic systems, have

inherently fluctuating energy production. As these volatile supplies of renewable energy take

up an increasing share of the marketplace, the need for flexible backup facilities becomes

crucial for electrical grid stability. Facing the global challenges of a carbon-free electricity

production, power-to-gas systems are a promising grid stabilization technique (Gatzen 2008).

HydrogenHydrogen has the strong potential to replace fossil fuels and to reduce greenhouse gas emis-

sions in the electricity generation markets. Pollutants produced during natural gas combus-

tion, like carbon dioxide (CO2), carbon monoxide (CO), unburned hydrocarbons (UHC),

and soot, simply cannot form due to the lack of carbon atoms in the fuel. Within a dis-

tributed energy supply system with high penetration of renewable energy sources, hydrogen

production through electrolysis and re-electrification on demand can be used as a negative

power reserve and as a method of power arbitrage (Emonts & Stolten 2016).

Gasturbines

The fluctuation in the production of renewable power has to be smoothed, and periods

of instantaneous nonavailability of wind and solar power have to be compensated for by

conventional power generation plants. Stationary gas turbines have the potential to provide

this service, due to their inherently short start-up times, high load gradients, and wide part-

load characteristics. Their application in combined cycle power plants (IGCC) offers the

most efficient way to provide electric power with the highest thermal efficiency currently

available.

H2 as gasturbine fuel

Next generation gas turbine combustors aim to increase fuel flexibility while achieving stable

combustion within a wide range of operating conditions. Burning hydrogen or hydrogen-

containing fuels in gas turbines significantly affects the thermodynamic process and compo-

nent performance. The higher reactivity, higher burning velocity, and shorter ignition delay

time lead to a higher risk of flashback and auto-ignition. Thus, state-of-the-art dry low NOx

(DLN) combustors developed for natural gas are not directly applicable for burning hydro-

gen. The higher reactivity also leads to higher flame temperatures which increases the rate

of nitrogen oxide (NOx) generation. Thus abatement strategies for gas turbine applications

2 1 Introduction

are being developed throughout the world for use with hydrogen combustion technologies

that as well affect the thermodynamic process of the engine.

NOx

reductionConcerns about acid rain, smog, and stratospheric ozone depletion have made the reduction

of NOx emissions an important challenge within the design of gas turbine power plants and

propulsion systems (Correa 1993). In general, key drivers for NOx generated during gas

turbine combustion are a low mixing performance of fuel and oxidizer, long residence times,

and high flame temperatures. Among others, one NOx reduction measure for high hydrogen

fuel aims to reduce the initial reactivity of hydrogen by dilution with nitrogen or water and

water vapor. While for non-premixed combustors, the dilution aims at the reduction of the

stoichiometric flame temperature, for premixed combustors, the dilution aims to increase

the combustor stability by lowering the risk of flashback and autoignition.

Wetcombustion

Although non-premixed flames naturally tend to have higher NOx emissions than lean pre-

mixed flames due to the stoichiometric conditions of the flame and the consequently higher

flame temperature, they are considered a promising candidate for pure hydrogen fueling

due to their inherently higher operational stability. In the past, water injection has been

used in various industrial, non-premixed gas turbines for the purpose of NOx reduction and

power increase. However, wet combustion reduces the overall plant efficiency by lowering

the process temperatures and increasing the energy demand for auxiliary facilities (Gazzani

et al. 2014). Currently, the dry lean premixed combustor concepts are the state of the art

for natural gas combustion and achieve current NOx thresholds without dilution.

Wetcombustionresearch

The technique of combustion under high moisture has experienced a renaissance due to de-

mands for increased fuel flexibility. Examples of recent wet gas turbine cycles are the HAT

(Humidified Air Turbine) and the STIG (STeam Injected Gas turbine). The influence of wa-

ter on the reaction properties of premixed gas turbine combustion and fuel flexibility is being

investigated in ongoing research projects (e.g. by Paschereit, see Goke (2012) and Gockeler

(2015), and by Sattelmayer, see Lellek (2017), Lellek & Sattelmayer (2015, 2017), Stadl-

mair et al. (2017), both under atmospheric experimental conditions). The authors indicate

a chemical influence of water on the NOx production, beyond the influence of temperature

reduction. The chemical effect is attributed to the change in the radical mix of the reacting

regime that influences the reaction rates of elementary reactions and increases or restrains

certain pathways of NOx formation.

Recent high-pressure investigations with an industrial combustor were done by a collabora-

tion of ENEL and GE, who redesigned a natural-gas-based combustor for hydrogen applica-

tions (Cocchi et al. 2008, Cocchi & Sigali 2010). For the purpose of design improvements,

numerical CFD studies of a reacting flow have been done by (Masi et al. 2010) and (Marini

et al. 2010)). Besides the improvement of the cooling configuration, the potential for NOx

reduction by steam and water injection as well as nitrogen dilution was investigated.

1 Introduction 3

ObjectiveThis study aims to contribute to the technical development of hydrogen fueling of stationary

gas turbines. The objective is to provide a NOx prediction approach, based on correlations

developed using both experimental results and a generic numerical study of the underlying

chemistry. This study first examines the effect of transition to pure hydrogen fueling for a

gas turbine combustor on NOx emissions. Water injection is hereby used as NOx abatement

measure. The test specimen is a non-premixed unscaled industrial gas turbine combustor

with a thermal power of 10 MWth. This study furthermore intends to promote the un-

derstanding of the chemical interaction between the fuel and the oxidizer at gas turbine

conditions during the replacement of methane fuel with hydrogen, along with the additional

effects of water. Based on these findings, a NOx prediction tool is developed, aiming to open

the possibility of high pressure NOx prediction solely on the basis of low pressure combustion

test data. This model focuses on the transition from natural gas to hydrogen fuel and water

injection for the suppression of NOx.

OutlineThe outline of the thesis is as follows: Chapter 2 introduces the fundamentals of gas turbine

combustion and emissions with focus on the properties of hydrogen as a gas turbine fuel.

Thereafter, state-of-the-art high hydrogen gas turbine applications and recent combustor

development projects are presented. Finally, a review of wet combustion is given with focus

on the effects of water on the combustion process. Chapter 3 presents the objectives and

proceedings of this study in more detail. The methods of the experimental, numerical, and

statistical approaches are discussed. Chapter 4 introduces the experimental setup of the

study. Thereafter the experimental facility and the test combustor are introduced. Finally,

the emissions results are discussed for various operating parameter variations including a

feasibility evaluation of the experimental range of operation. Chapter 5 presents a numeric

reactor model based on a chemical reactor network (CRN) that allows for a detailed repre-

sentation of the chemical process and thermodynamic conditions inside the combustor. After

discussing the predicted NOx results, a chemical analysis shows the radical distribution for

selected parameter variations. Finally, the effects of hydrogen and water are described by

a NOx pathway study, and a quantification of the physical, thermo-physical, and chemical

effects of water is provided. Based on the experimental and numerical findings, Chapter 6

presents a tool consisting of relevant correlations that enables the NOx prediction of the non-

premixed combustor at gas turbine conditions for hydrogen applications and water injection.

Chapter 7 summarizes the findings and draws conclusions.

2 Wet hydrogen combustion in gas

turbines

This chapter presents the state of the art of hydrogen combustion in stationary gas turbines

and gives an overview of relevant research conducted over the past decades. In the first

section, fundamental issues of gas turbine combustion and emissions that are essential for

understanding the impact of fuel flexibility on the gas turbine cycle are presented. In the sec-

ond section, the application of hydrogen as a gas turbine fuel is discussed with a focus on its

combustion properties and its effects on the gas turbine cycle. A technical review introduces

state-of-the-art high hydrogen gas turbines and current research on advanced high-hydrogen

combustor technology. Humidification as a NOx abatement measure is presented in the last

section. The development of water injection is reviewed and recent research on humidified

high-hydrogen applications is introduced. Finally, the effects of water on the combustion

process are presented in more detail.

2.1 Gas turbine emissions

Gas turbinecombustors

Stationary gas turbines are an inherent part of today’s technical infrastructure for electricity

generation and heat supply. Due to high exhaust temperatures, typical applications are

combined cycle heat and power facilities in electricity generation. The requirements of

modern gas turbines, especially in recent renewable energy applications, are a high thermal

efficiency, low pollution emissions, fuel flexibility (e.g. high hydrogen fuels), low minimum

load, fast transient response, reliability, and cost. Therefore, the combustion process within

the engine has to fulfill several engineering criteria (e.g. combustion efficiency, stability

(flashback resistance especially with premixed flames and highly reactive fuels, autoignition

resistance), flexibility (fuel and operating limits), combustor integration, material, chemistry,

low emissions, safety, and reliability) which are all challenging due to the high power density,

high temperature, and high pressure environment, cf. Lieuwen & Yang (2013).

Technologically, conventional non-premixed combustors and lean premixed combustors are

distinguished from one another. The former type features separate injection of fuel and air

into the combustor, which results in stable combustion with damped oscillation character-

istics. Due to local stoichiometric conditions, high flame temperatures occur. This results

in higher NOx emissions and - in the case of carbon-based fuels - higher soot emissions.

The latter type premixes fuel and air before the injection into the flame zone. Significantly


lower local flame temperatures arise due to lean conditions, significantly suppressing NOx

generation. However, premixed flames face the risk of flame propagation upstream into the

mixing area, which may cause serious damage (flashback). A premixed flame also tends

to autoignite prior to the flame position and to experience combustion instabilities due to

acoustic pulsation. For all mentioned failure mechanisms, the fuel composition and operating

conditions are relevant and the mechanisms are promoted by higher hydrogen shares. More

details about these technical issues can be found in Lieuwen & Yang (2013) and Lefebvre &

Ballal (2010). Today, however, this technical distinctive feature does not apply to modern

high hydrogen combustor developments, as presented in the next section.

Stabiliza-tion

Flame stabilization is essential for safe, flexible, low-emission operation. Swirl-stabilized

flames have been established in both premixed and non-premixed combustion regimes due

to high power density. Vortex breakdown at sufficiently high swirls leads to intense inter-

nal recirculation. This efficient transport of combustion products towards the burner tip

provides a permanent, self-regenerating ignition of new inflowing reactants. This principle

allows higher downward velocities and thus a high power density of the combustor. More

information on flame stabilization strategies can be found in Lefebvre & Ballal (2010) and

Lieuwen & Yang (2013).

Combustordesigndemands

The combustor has only a small influence upon cycle efficiency or specific power of a station-

ary gas turbine. The combustion efficiency is an important consideration during combustor

development, but not the most difficult challenge. However, the need for more efficient en-

ergy conversion has driven the combustion temperatures higher in future cycle designs. The

combustor does have an important impact on feasibility of certain cycles e.g. steam addi-

tion, engine operating limits, emission performance, cleaner burning, and wider operation

(Lieuwen & Yang 2013).

EmissionsAir pollution and greenhouse gas emissions from stationary gas turbines arise from the

chemical conversion process of oxidizer and fuel. The substances emitted differ in their

effects on the environment and living things:

Carbonoxides

Carbon monoxide (CO) is a colorless and odorless gas that results from incomplete com-

bustion or flame quenching of carbon-containing fuels. CO is poisonous for humans; it acts

by blocking the oxygen receptors on hemoglobin and reducing the amount of oxygen carried

around the body. It arises mainly from the conversion of decomposed fuel within the flame

zone. Whenever the reaction conditions are unfavorable (e.g. in cold temperature regions,

when improper mixing occurs and the mixture strength is too weak to support combustion,

in over-rich regions, or when there is insufficient residence time) the CO does not have the

chance to further oxidize to CO2. The high levels of CO at low equivalence ratios are due

to the slow rates of oxidation associated with low combustion temperatures. However, at

temperatures higher than ∼1800 K, the production of CO by chemical dissociation of CO2

6 2 Wet hydrogen combustion in gas turbines

starts to become significant. Furthermore, an increase of the pressure reduces CO emission,

mainly by suppressing the chemical dissociation of CO2. The CO emissions of gas turbine

combustors significantly decrease by increasing air temperature. The wall cooling and dilu-

tion hole configuration of the combustor geometry is another important factor that influences

CO emissions. CO formed in primary combustion can travel toward the wall and penetrate

the cool injection streams, which causes the CO2 formation reaction to be quenched due to

the rapid local temperature decrease (Lefebvre & Ballal 2010).

Carbon dioxide (CO2) is a colorless gas with sharp acidic odor and is formed in the burnout

zone of the flame by further oxidation of CO to CO2. It thus completes the carbon oxidation

process. CO2 has an strong environmental impact via the CO2 cycle and the atmospheric

greenhouse effect, promoting anthropogenic global warming.

Water Water vapor (H2O) is generated by oxidation of hydrogen-containing fuels and is a green-

house gas; however, it has less of an effect than CO2, since its residence time in the at-

mosphere is much shorter and the quantity of human-made water vapor is negligible with

respect to the amount occurring naturally.

Nitrogenoxides

Nitrogen oxides (NOx) are the aggregation of NO and NO2 that have different impacts on

living organisms and the environment. NO (Nitric oxide, nitrogen (mono-)oxide) is a colorless

gas that is formed at high temperatures in combustors within the flame zone. NO is formed

by decomposition of molecular nitrogen from the air via multiple pathways, including the

thermal NO pathway at high temperature, the prompt NO pathway including reactive carbon

fuel fragments, the nitrous oxide (N2O) pathway, and the NNH pathway. Note that in the

case of natural gas combustion, NO generation via fuel-bond nitrogen is non-existent. More

details about the pathways are given in section 5.2.3. In the post-flame zone, further thermal

NO is generated as long as the temperature is high enough. The combustion design faces a

trade-off between CO and NO generation, particularly for non-premixed flames, since they

display opposite behavior when changing the mean flame temperature by excess air ratio

variation.

NO2 (nitrogen dioxide) is a brown toxic gas and is formed via a shift reaction from NO.

However, the temperature must not exceed 900◦C because NO2 dissociates above that tem-

perature limit. In non-premixed flames, NO2 is predominantly formed via direct emission,

when the mixing of fuel and air is poor and the mean temperature is low. In the atmosphere,

however, the time scale for the oxidation of NO to NO2 is on the order of hours and thus,

the dilution is stronger and the effects of NO2 are generally less significant than those of

directly emitted NO2.

The hazardous effects of NOx result from the formation of nitrous and nitric acid (HNO2,

HNO3) on contact with mucous membranes causing irritation and burns. Furthermore, NO2

is involved in the formation of all major secondary pollutants such as ground-level ozone,


acid rain, and smog (London Smog, Los Angeles Smog). The NOx emissions are therefore

legally limited to 15 ppm or 25 ppm, respectively, for stationary gas turbines.

Unburnt hy-drocarbons

Unburnt hydrocarbons (UHCs) contain many species ranging from unburned fuel like methane

(CH4) to large molecules and include several species classified as hazardous air pollutants.

The generation of UHCs is similar to CO formation. CH4 is a relevant greenhouse gas, while

non-methane hydrocarbons have not only negative heath impacts, but promote ground-level

ozone production and smog, where they are converted to organic radicals before forming

hazardous air pollutants in the presence of NO2 and sunlight.

ParticulateMatter

Particulate Matter (PM) includes nonvolatile solid particles and condensed volatile hydro-

carbons and sulfur oxides. PM mainly consists of soot and non-combustible fuel components.

It arises mostly from incomplete combustion of hydrocarbons in rich fuel regimes. Depending

on their size, they are retained in the nasal mucous membranes (> 15 µm) or, if they are fully

respirable (< 2.5 µm), penetrate into the lung tissue and can even enter the bloodstream.

Sulfuroxides

Sulfur oxides, consisting of gaseous SO2 and solid SO3 at standard ambient conditions, are

non-flammable, colorless, and originate from sulfur-containing fuels. However, since sulfur

compounds are usually removed from natural gas and the remaining concentration is in the

range around 10 ppm, they are mentioned here simply to complete the list of emissions.

2.2 Hydrogen as gas turbine fuel

HydrogenHydrogen is the first and thus the lightest chemical element on the periodic table. It con-

tains a single proton and a single electron. Usually, hydrogen gas occurs as the molecular

compound H2 with a single covalent bond. Pure hydrogen is naturally rare on Earth. Due to

its low weight, it cannot be held by gravity and passes through the atmosphere to leave the

planet. Thus hydrogen is not a primary energy source but has to be produced and considered

as an energy carrier.

Hydrogenproduction

Currently, industrial quantities of hydrogen are most economically derived from fossil sources

using steam reforming of natural gas, partial oxidation of methane, or coal gasification.

The product is a syngas containing varying amounts of CO and H2. However, within the

framework of greenhouse gas emissions reduction, several alternative production processes

are have recently become the focus of research. Methods of thermo-chemical hydrogen

production approach include solar-based thermal water decomposition, solar-based water

reforming, and biomass-based supercritical water gasification. Furthermore, hydrogen can

be produced using electricity by decomposing water into its compounds H2 and O2 via

electrolysis. Hydrogen from solar radiation is produced by direct photo-electrochemical water


Table 2.1: Fuel properties of hydrogen, natural gas, and methane.

Fuel/combustion property Unit Hydrogen Natural gas Methane

Density at 273 K, 1.1013 bar kg/m3 0.09 0.7−0.9 0.72

Flammability limits vol.% in air 4−75 4.5−13.5 5−15

Flammability limits (Φ) - 0.1−7.1 0.4−1.6

Specific lower heating value (LHV) MJ/kg 120 38.9−47.1 50

Molar lower heating value (LHV) MJ/m3 10.8 31−41 35.9

Maximum laminar flame speed m/s 3.25 0.45

Adiabatic flame temp. (Φ = 1) K 2370 2226

Lower Wobbe Index MJ/m3 40.7 46.5−48 47.9

Sources: Glassman et al. (2014), Dunn-Rankin (2008), Lackner et al. (2013), Abbot et al. (2009),

Hydrogen data (2018).

decomposition and photosynthetic hydrogen production via organisms like algae. Further

information on hydrogen production technologies can be found in Emonts & Stolten (2016).

2.2.1 Hydrogen combustion properties

Chemicalpropertiesof hydrogen

Pure hydrogen is a single-component fuel with particular chemical properties. With regard

to combustion, hydrogen has a significantly higher chemical reactivity compared to natural

gas. Thus, characteristic combustion properties like flame speed and flame temperature

are significantly higher (see Tab. 2.1). In the following section, the relevant combustion

properties are described in more detail.

Flammabilityrange

Hydrogen features a significantly wider volumetric flammability range compared to natural

gas. The lean extinction of hydrogen occurs at lower equivalence ratios due to the higher

chemical reactivity and diffusivity. Furthermore, hydrogen features a significantly higher

upper flammability limit in fuel-rich regimes.

Heatingvalue

The specific-mass-based lower heating value (LHV) of hydrogen is up to three times higher

than the specific LHV of natural gas. However, due to the low density of hydrogen, the molar

LHV is significantly lower. This has significant effects on the total engine, as described later.

Flamespeed andtemperature

The higher reactivity of hydrogen also leads to a higher laminar flame speed. The maximum

laminar flame speeds are reached at different equivalence ratios (Φ) for different gases/gas

mixtures. While the maximum laminar flame speed for hydrogen occurs at Φ = 1.80, it

occurs at Φ = 1.08 in the case of methane. Moreover, the laminar flame speed depends

more strongly on the equivalence ratio in the case of hydrogen and hydrogen admixtures

(Donohoe et al. 2013). Additionally, the adiabatic stoichiometric flame temperature of a

hydrogen flame is about 150 K higher than that of a natural gas flame.


Wobbeindex

The Wobbe index is an indicator of the interchangeability of fuel gases. It is used to compare

the combustion energy output of fuel gases with different compositions in an appliance with

a given fuel nozzle geometry. As the Wobbe index increases during fuel variation, a higher

degree of flexibility is required in the fuel control system to achieve the design heat input.

Since the Wobbe index of hydrogen is 15% lower than that of methane, the fuel control

system requires higher volumetric flows for the same thermal power.

Ignitiondelay

The ignition delay indicates the time available for fuel air mixing prior to the initiation of

ignition. This time depends on the fuel composition, the ambient temperature, and the

pressure. A longer ignition delay time shifts the flame downwards and allows for more

intense premixing. For all temperature conditions, the ignition delay times are significantly

shorter with increasing hydrogen content. Increasing the pressure results in shorter ignition

delay times in the high temperature (T > 1350 K) and the low temperature regimes (T <

1000 K). The opposite dependence has been observed in the moderate temperature regime

(see Keromnes et al. (2013)).

Autoigni-tion andflashback

Lean premixed combustion faces the problems of autoignition and flashback. Autoignition

is the spontaneous ignition of the fuel-air mixture without an external ignition source, such

as the hot flame region or a spark. Flashback is the upstream propagation of the flame that

takes place whenever the flame speed exceeds the flow velocity. Free-stream flashback and

boundary-layer flashback are considered separate phenomena. Flashback in the free stream

occurs, when the turbulent flame speed is higher than the local flow velocity. Boundary layer

flashback occurs through the retarded flow in a boundary layer (Lefebvre & Ballal 2010).

The tendency towards autoignition and flashback increases with the reactivity of the fuel

and thus with increased hydrogen content.

Consequen-ces foremissions

As a consequence, lean premixed combustion, being today’s most common combustion tech-

nology in stationary gas turbines, cannot be easily adapted to burn hydrogen in a safe and

environmentally-friendly way. The differences between the physical properties of hydrogen

and natural gas affects the entire gas turbine cycle, which is described in Sec. 2.2.2. In

Sec. 2.2.3, existing gas turbines which can already handle a certain hydrogen fraction are

presented. Major burner or combustor redesigns or completely new combustion technologies

are being developed to avoid pre-ignition and flashback and to meet the emission restrictions.

The current research is presented in Sec. 2.2.4.

2.2.2 Fuel effects on gas turbine cycle

Fuel conversion from natural gas to hydrogen has a strong effect on the gas turbine cycle.

Additionally, depending on the NOx control strategy (nitrogen, steam, or water dilution or

non-dilution strategy by removal from exhaust gases) the effects on the gas turbine cycle


differ significantly. Chiesa et al. (2005) analyzed the fundamental thermodynamic changes

in the gas turbine when switching from natural gas to pure hydrogen fuel. This discussion

focuses on the NOx reduction strategy of dilution and describes the major effects on the

gas turbine cycle. In particular, the following criteria are highlighted: compressor/turbine

matching, hot path material/blade cooling and the outlet casing.

Compressorturbinematching

Variations in the interaction of compressor and turbine arise from the higher volumetric fuel

flow rate of hydrogen, from the varied energy input, and from the exhaust gas composition.

A different operating point will arise for each of these variations, where mass flow rate and

pressure ratio will restore the fluid-dynamic equilibrium between the two turbo machines.

A higher compressor outlet pressure is a result of the higher volumetric flow rate and the

constant flow capacity of the turbine. Consequently, the increased compressor outlet pressure

moves the operating point of the compressor to instability or surge.

A global balance shows the mismatch of compressor and turbine when switching to hydrogen-

based fuel. Using the model of a choked turbine, the volume flow of the turbine inlet

(= the combustor outlet flow) remains constant when switching the fuel from methane

to hydrogen. A comparison at stoichiometric conditions shows an 8.4% reduction of the

thermal power and a 25% reduction of the air flow. Under realistic lean boundary conditions,

hydrogen combustion increases the enthalpy drop in the turbine by about 5% (Chiesa et al.

2005) compared to natural gas, due to the change in exhaust composition without any NOx

reduction measures.

Steam dilution requires a steam/H2 mass ratio of more than 8 in order to keep the stoichio-

metric flame temperature at the level of natural gas flames. The volume flow rate of the

turbine inlet increases up to 23%, under a constant turbine inlet temperature and combus-

tion air quantities compared to natural gas fueling. Steam dilution furthermore increases

the specific heat capacity of the turbine inlet flow, since H2O is a triatomic molecule. At

a given steam/H2 mass ratio, a higher water vapor content in the exhaust gas leads to an

increase in the isentropic enthalpy drop of the turbine by up to 13%. However, due to the

simultaneous decrease of the isentropic exponent, the temperature drop within the turbine

is decreased, leading to an increase in the turbine outlet temperature (TOT) (see Chiesa

et al. (2005)).

In contrast, the effect of nitrogen dilution on the turbine enthalpy drop is virtually negligible

since a large amount of nitrogen is already contained in the mixture. However, a N2/H2 mass

ratio of more than 16 is required in order to keep the stoichiometric flame temperature at the

level of a natural gas flame. Thus the volume flow rate of the turbine inlet increases up to

28% (Chiesa et al. 2005), again, under constant turbine inlet temperatures and combustion

air quantities compared to natural gas fueling.


Materialand cooling

The higher water vapor content in the exhaust gas in the case of hydrogen combustion with

steam dilution causes an increase in convective heat transfer from to fluid stream to the

outer side of the turbine blades. This changes the temperature distributions on the blades

and thus, a more efficient cooling system with larger cooling flow rates is required in order to

operate the gas turbine at the same turbine inlet temperatures (TIT) as are used for natural

gas systems. Higher vapor content in the exhaust gas also has a higher risk of hot corrosion

of the hot path material. Additionally, faster degradation of environmental barrier coatings

(EBC) and thermal barrier coatings (TBC) occurs (Gazzani et al. 2014).

Outletcasing

Despite the appreciable increase of the expansion enthalpy drop, the volumetric flow ratio

of exhaust gases and air increases by about 3.5% in the case of pure hydrogen fueling and

increases up to 20% and 16% for nitrogen and steam injection, respectively, in order to keep

the adiabatic flame temperature the same as with natural gas fuel. This variation implies

a major resizing of either the compressor or the turbine cross-sectional area (Gazzani et al.

2014). When the pressure ratio and TIT are held constant due to technical restrictions, the

increase of hydrogen in the fuel causes an increase of the turbine outlet temperature.

2.2.3 State of the art high hydrogen gas turbines

Combustors that are capable of burning syngas with hydrogen fractions up to 45 vol.%

have been developed by heavy duty gas turbine manufacturers for Integrated Gasification

Combined Cycle (IGCC) processes. They are derived from machines that were originally

designed for natural gas fueling, since the demand for hydrogen-fueled gas turbines is small.

Tab. 2.2 gives an overview of gas turbines, introduced in the years 2002-2008, that are capable

of burning hydrogen admixtures. Non-premixed flames are the prevailing technology, where

NOx generation is suppressed by flame zone cooling and fuel dilution. The necessary fluids,

such as water, steam or nitrogen, are available from the steam cycle or an air separation

unit of the IGCC process. More recently, Kawasaki demonstrated a pure hydrogen fueled

combined heat and power (CHP) gas turbine in a demonstration plant funded by NEDO

(2018). In 2018, a 1.5 MWel gas turbine plant was successfully operated using water injection

as the NOx abatement measure. The NOx emissions could be limited to 50 ppm. However,

no commercial gas turbine with pure hydrogen fueling is currently in operation.

2.2.4 High Hydrogen combustor development

NOx

reductionstrategy

Worldwide research and development is carried out by academia and gas turbine manu-

facturers aiming for dry and undiluted hydrogen combustion. With respect to emissions

reduction, the focus is on NOx prevention since hydrogen combustion does naturally not

emit CO and CO2. According to the general correlation provided by Lefebvre (1984), NOx


Table 2.2: State-of-the-art gas turbine technology for hydrogen

Name YearPel H2

NOx reduction via dilution[MW] [vol.%]

GE (former Alstom)1 2002 200 45 N2 dilution (55 vol.%)

GE2 2008 10-280 45 N2 and steam dilution

Siemens3 2006 10-250 41 Steam (22 vol.%), N2 (30 vol.%)

Ansaldo Energia4 2007 170 45 Steam dilution (50 vol.%)

Mitsubishi/Hitachi5 2002 250 20 N2 dilution

Kawasaki6 2018 1.7 100 Water injection

Sources: 1Reiss et al. (2002). 2Payrhuber et al. (2008). 3Gadde et al. (2006). 4Bonzani & Gobbo

(2007). 5Tajina (2002). 6NEDO (2018).

emissions are suppressed with better mixing, shorter residence times, and lower reactivity.

This relation is the main design criteria for NOx reduction in high hydrogen combustors

that are described in the following section. They can be classified into premixed, lean

direct injection, flame miniaturization, flameless oxidation, catalytic, and staged combustion

approaches.

Leanpremixedcombustion

Lean premixed combustion is the state-of-the-art for natural gas applications. It features a

comparatively low flame temperature and fewer hot spots in the reaction zone. GE developed

a premixed multi-tube mixer (York et al. 2013). Their concept comprise small-scale jet-in-

crossflow mixing in multiple tubes in order to face the high risk of flashback and auto-ignition.

Another lean premixed combustor has been developed by Siemens (Bradley & Marra 2012)

that verified stable combustion at hydrogen content up to 70 vol.% H2. It uses a modified

fuel injection system based on a premixed natural gas burner design. Siemens also developed

a triple fuel syngas burner capable of H2 with a premixed configuration (Wu et al. 2007)

within the Advanced Hydrogen Turbine Development Project with the U.S. Department

of Energy (Bancalari et al. 2006). The generation of NOx is suppressed by dilution (H2O

and N2) in this system. Flashback resistance also has been demonstrated. A low swirl

injector has been developed by Lawrence Berkley National Laboratory (Therkelsen et al.

2012). Here, the flame stabilization method for lean premixed combustion is low-swirled

annular flow centered on a non-swirled inner flow. This approach can be used with hydrogen

content of up to 90 vol.% H2. Wind et al. (2014) of GE (former Alstom) successfully operated

their two-stage combustor for the GT26 successfully at up to 30 vol.% H2 without change

of hardware. In the EV burner 45 vol.% H2 was possible until full load conditions (30 bar),

and at partial load conditions (16 bar) the share could be increased up to 60 vol.% H2. The

SEV burner featured maximal 45 vol.% H2. Furthermore, a conceptual design for a dry

premixed hydrogen combustor has been developed by Cerutti et al. (2014) under a cluster


arrangement for the analysis of flashback and flame holding resistance. Experimental tests

demonstrated the maturity for a full scale arrangement design.

Lean directinjection

Lean direct injection aims to combine the low NOx emissions of lean premixed flames and

the stability and flashback resistance of non-premixed flames. Fuel and air are separately

injected into the combustor in lean proportions and mixed rapidly enough in order to attain

the low NOx emissions typical for premixed combustion. A lean direct injection burner has

been developed by a cooperation of NASA, Parker Hannifin, and NETL (Marek et al. 2005).

This burner is characterized by multiple injection points and quick jet-in-crossflow mixing

inside air guiding tubes. This approach is similar to the flame miniaturization technique

because of the use of multiple fuel injection. With this method, however, no individual

flames develop.

Flameminiaturiza-tion

Flame miniaturization aims to reduce the residence time inside the hot flame region by

decreasing the flame size and increasing the number of flames. These flames are typically

diffusion-type flames for increased stability and flashback resistance. Kawasaki Heavy Indus-

tries is developing a micromix combustor (Horikawa et al. 2015), Mitsubishi Hitachi Power

Systems is developing a cluster burner (Asai et al. 2015), and Parker Hannifin and UCI are

developing a combustor of micro-mixing cups (Hollon et al. 2011) based on this principle.

Flamelessoxidation

The concept of flameless oxidation was originally developed for industrial ambient pressure

applications and its principle of operation lies in strong recirculation. The dilution of the

reaction zone with flue gas significantly reduces the temperature so that no visual flame ap-

pears. The flameless oxidation approach has been applied to elevated gas turbine conditions

and hydrogen admixtures by Lammel et al. (2010), Roediger et al. (2013), and fuel flexibility

up to pure hydrogen had been demonstrated under laboratory conditions.

Catalyticcombustion

The advantage of catalytic combustion is ability to carry out a stable reaction at low tem-

peratures. The necessary amount of diluent can be significantly reduced compared to non-

premixed approaches. Precision Combustion, Inc. developed a rich catalytic-based hydrogen

reactor for low NOx emissions in cooperation with Solar Turbines. The concept comprises

a rich catalytic stage, where hydrogen is injected, and a secondary lean burn stage (see

Alavandi et al. (2012)).

Stagedcombustion

Staged combustion offers the combination of different flame characteristics in order to achieve

stability and emission targets. A rich/lean staged combustion approach is developed by

FHNW with the Paul Scherrer Institute (Bolanos et al. 2013) featuring a rich first stage

with laminar flow and a secondary non-premixed flame lean burnout zone.


2.3 Humidification as NOx abatement measure

Humidification of the combustion regime is a method to reduce NOx emissions of gas turbine

combustors by lowering the flame temperature and reducing the reactivity of the mixture.

Humidification can also be used to increase the power of a gas turbine, since the volume

flow through the turbine is increased while the pressure ratio of the compressor is nearly

constant, requiring no additional energy for air compression. Furthermore, the presence of

water increases the heat capacity of the fuel-air mixture. As a result, the temperature of

the combustion products is significantly lower compared to combustion without water with

an equivalent fuel quantity. The equivalence ratio in the combustion chamber can therefore

be increased in order to reach similar combustor outlet temperatures as are seen without

water injection. The focus of this thesis is humidification as a NOx abatement measure. This

section presents the historical background and recent research activities of water injection

techniques and the effect of water on the combustion process.

2.3.1 Historical review

Earlyresearch

First attempts to reduce NOx emissions of conventional non-premixed and natural gas fueled

combustion systems by steam and water injection were reported by Dibelius et al. (1971)

and Hilt & Johnson (1972), who conducted investigations with entire gas turbine engines at

constant speed and variable loads. They verified the potential of NOx emission reduction

for steam and water injection. Later, Wilkes & Russell (1978) analyzed the impact of fuel-

bound nitrogen on NOx emissions and the effect of water injection on both suppression of

NOx emissions via fuel-bound nitrogen and the thermal NOx pathway. In a sub-scale test

facility, Mulik et al. (1981) discovered that the effectiveness of water injection decreased

as the fuel-bound nitrogen content increased. Farrell & Thomas (1981) analyzed the NOx

emissions for a variety of inlet conditions and water injection methods using both engine tests

and downscaled tests at reduced pressure and air flow. The test specimen comprised of a dual

fuel nozzle that was suitable for use with both natural gas and liquid fuel. Combustor exit

temperature variations simulated operating conditions in part loads. Bahr & Lyon (1984)

applied water injection on an annular combustor of an aircraft derivate. NOx emissions

could be reduced with increasing water fuel weight ratio for both, steam and water injection

and both, liquid and natural gas combustion. Furthermore, they verified the effect of water

injection on another engine. Antos & Emmerling (1984) did water injection engine tests but

could not find a dynamic response of the combustor to water injection.

Analyticalmethods

Over the years, evaluation methods have been improved and combined. Theoretical methods

arose from analytical based proceedings resulting in formulations of quantitative correlations.

Shaw (1975) performed a theoretical analysis of the effect of water by quantitatively esti-


mating the effectiveness of water injection for NOx reduction based on measurements. He

presented a correcting equation for different combustion air humidities in aviation engines.

Hung & Agan (1985) tested two 7 MW industrial gas turbines to verify the influence of

ambient air temperature, humidity, and water injection on NOx emissions. They set up an

a-priori NOx prediction model including the above mentioned parameters. They also pre-

sented a correction model in order to compare NOx emissions appearing at varying ambient

conditions. In this model, they applied both the constant fuel flow and the constant turbine

inlet temperature (TIT) operating strategies.

Displace-ment bypremixers

With the development of lean premixed gas turbine combustors, the research and develop-

ment of wet combustion has stalled. Dry lean premixed flames suppress NOx generation by

limiting the reactivity and the flame temperature by fuel air premixing. Currently, premixed

combustors are the state-of-the-art for natural gas fueling.

Advancedmethods

With progress in the analysis methods, detailed chemical analysis became possible. Wet

combustion for methane diffusion flames has been studied by Zhao et al. (2002), who per-

formed a numerical analysis on the influence of the OH◦ radical on the NOx formation with

steam addition. This study performed fundamental research on the progress of chemical

reactions in the advanced reaction scheme (GRI-Mech, compare Sec. 5.2.2) and counter-

flow diffusion flame approach. They found that the production rate of OH◦ decreased with

increasing addition of steam due to a significant reduction of the flame temperature. For

the same flame temperature, however, the OH◦ concentration increases with the presence of

water.

2.3.2 Recent wet high-hydrogen applications research

RenaissanceRecently, the technique of wet combustion has undergone a renaissance due to the increased

demand for high reactivity fuel combustion (syngas and hydrogen) in gas turbines. Investi-

gations of ignition delay times, flame speed, flame temperature, and higher NOx emissions

are carried out with improved methods, e.g. chemical reactor network (CRN) modeling and

computational fluid dynamics (CFD) analysis. In the following section, relevant studies are

presented.

Combustorredesign

Redesigning an existing, mostly non-premixed natural gas turbine for hydrogen application is

a common approach to developing high-hydrogen technology. High combustion temperatures

are prevented by fuel dilution or direct water/steam injection into the combustion zone.

An 11 MWth GE 10 gas turbine with a silo-type combustion chamber and non-premixed

combustion type is operated with 100% hydrogen fuel in a dry and wet configuration by GE

and ENEL (see Cocchi et al. (2008), Benovsky et al. (2008), and Cocchi & Sigali (2010)).

Steam is injected directly into the cold side combustion air stream in order to concentrate


the dilution effect in the primary zone. Steam injection could reduce the NOx emission by

up to 94% compared to the dry configuration. Further numerical studies were performed

to optimize the combustion chamber design (Marini et al. 2010, Masi et al. 2010). GE also

performed combustion tests with a hydrogen content between 8.6 and 61.9 vol.% H2 using

N2, H2O, and CO2 injection, as well as fuel moisturization (Brdar & Jones 2000). Direct

injection into a non-premixed combustor configuration assured controllability, efficiency, and

limited system costs.

Humid gasturbinecycle

The humid gas turbine cycle is a novel gas turbine combustion technology that is based on

the injection of large amounts of steam into the combustor. This steam is generated by the

hot exhaust gas at the turbine outlet. Thus, in humid gas turbines the steam is injected

back into the gas turbine instead of running a steam cycle like in conventional combined-

cycle power plants (Jonsson & Yan 2005). According to Goke (2012), this procedure offers

efficiencies similar to those of state-of-the-art combined cycle power plants, but featuring

lower installation and electricity production costs. Furthermore, this ultra-wet gas turbine

requires very short start-up times, and thus, it is suited to support the fluctuating power

generation from renewable energies. Goke analyzed different combustor concepts (premixed,

non-premixed, and Rich-Quench-Lean configurations) with both methane and hydrogen fu-

els numerically and verified the premixed results with a laboratory atmospheric premixed

combustor. He showed a strong impact of steam on the combustion chemistry and thus on

the formation of NOx.

Wetpremixedflames

Premixed flames are also the focus of current research on water injection in gas turbine

combustors in order to allow increased flexibility in the power output and start-up speed

with simultaneous control of NOx emissions. Lellek (2017) transferred these requirements

to a premixed combustor at lab scale and analyzed the feasibility in experiments. He found

a strong influence of water droplet size on NOx and CO formation for constant operating

conditions.

The specific issues of acoustic challenges of water injection on the exhaust system were

addressed by Schmid et al. (2011). They found that the sonic velocity does not change

in their current test rig because the droplets are too big to follow the acoustic fluctuations.

Thus the influence of water injection on the acoustic properties, and therefore on the stability

behavior, is very sensitive to the injection conditions, especially the droplet diameter. Finally,

Stadlmair et al. (2017) analyzed the impact of water injection on premixed combustion.

In single-swirled burner experiments with an atmospheric test stand at elevated air inlet

temperatures, they analyzed the stability and acoustic dampening characteristic of different

water injection levels.


The studies discussed above are advanced studies for actual combustor development and

focus on application. However, theoretical studies analyze the effect of water on the reaction

process. These studies are discussed in the next section.

2.3.3 Effects of water on combustion

Water injection alters the reactivity of the underlying combustible mixture. The effect of

additional water on combustion can be distinguished into three fundamental phenomena:

the physical, the thermo-physical, and the chemical effects. These effects, and secondary

effects that stem from the fundamental phenomena, are described in this section.

Physicaleffect

By injecting additional H2O, the concentration of all other species is reduced, in particu-

lar the NOx concentration, while the concentration of H2O, a combination of combustion-

produced H2O and injected H2O, increases. This phenomenon can also be called ‘dilution’.

Note that dilution inherently reduces concentration of active species, e.g. chemical radi-

cals in the combustion zone. This directly causes a reduction of the underlying reactivity.

Furthermore, a non-premixed flame is shifted from a diffusion-type to a premixed-type flame.

Thermo-physicaleffect

The addition of any diluent affects the global energy balance of the reaction zone. Primarily

depending on the heat capacity of the diluent, the adiabatic flame temperature is reduced. In

the case of H2O injection, additional heating is required to bring up the overall temperature

of the combustible mixture, and additional energy for evaporation is needed in the case of

liquid water. This resulting temperature drop affects the reactivity of the reacting fluid and

thus the final composition of all species.

Chemicaleffect

The chemical effect results from the active participation of H2O in the reactions. This

influence is comprised of two effects:

First, H2O itself is a reactive species. A change of its concentration leads to a shift of the

chemical conversions within the reaction pathways. It has been reported that the effect of

H2O on the radicals are non-negligible (Gockeler 2015). Their presence is essential for the

initialization of reaction pathways and their propagation.

Second, H2O is an effective inert collision partner in three-body reactions. In the water-

enriched regime, the collision efficiency (also called the ‘chaperon efficiency’) of the general

collision species M is affected by a higher H2O concentration. This results in a change in the

reaction rates and the subsequent concentration distributions, which impacts the reactivity

of the combustible mixture.

Secondaryeffects

These three effects describe the basic variation of fundamental combustion properties in the

presence of H2O. In general, these variations in reactivity affect the flame speed (Gockeler

2015, Mazas et al. 2011, Xie et al. 2014) and the ignition delay time (Le Cong & Dagaut


2008). Water-induced quenching leads to an increased concentration of unburnt fuel, as well

as CO and UHC (Lefebvre 1995). In the case of pure hydrogen combustion, however, this

aspect is not relevant due to the absence of carbon. The chemical effects of water/vapor on

the flame speed and the ignition delay times of the combustion have been investigated as well

by Park et al. (2007) for a counterflow diffusion flame model and Le Cong & Dagaut (2008)

for a flue gas recirculation approach. Furthermore, the physical properties of the fluid, e.g.

the thermal diffusivity, density, and heat capacity, are affected by dilution.

2.4 Summary

Substitution of natural gas by hydrogen in stationary gas turbines is a challenging issue

due to the increase in fuel reactivity and, as a consequence, elevated burning velocity, flame

temperature, and NOx emissions. Research and combustor development is being performed

worldwide to apply hydrogen to gas turbines with conventional premixed and non-premixed

designs as well as new developments, such as flame miniaturization, lean direct injection,

flameless oxidation, and catalytic combustion.

Water injection is a method for power upgrading and NOx abatement for premixed and

non-premixed stationary gas turbines. In particular, within the framework of increasing the

fuel flexibility of gas turbines, this method is facing a renaissance with new developments

for high-hydrogen applications. Recently, there is research ongoing for both fundamental

combustion analytics and advanced full-scale combustor developments. The effect of water on

the combustion process can be categorized into a physical effect (dilution), thermo-physical

effect (cooling), and chemical effect (H2O as reaction partner and collision partner).

In summary, the relevance of hydrogen combustion in stationary gas turbines has been

demonstrated and related fundamental issues have been discussed. On this basis, the fol-

lowing chapter introduces the objective of this thesis and justifies the underlying methods.

3 Objective and methods

This chapter identifies the current demand for research relating to water injection for hydrogen-

fueled gas turbines. On the basis of the fundamentals given in the last chapter, the first

section describes the objectives of this study in more detail. The second section outlines the

structure of this thesis and discusses the selection of the applied methods.

3.1 Objective

Concluding from Chapter 1 and Chapter 2, a relevant number of academic studies have

analyzed the effects of hydrogen and water on combustion processes. Studies approved

the general feasibility of high hydrogen fueling of advanced gas turbine combustors at real

operation conditions. However, there is still a gap between the fundamental research on the

influence of hydrogen and water injection and the technical application. This is addressed

here by the key objectives:

• Effects of hydrogen and water on the combustion process at real gas turbine conditions

• Combining of these findings to create a NOx correlation

that are explained in the following chapter. Current open questions in the underlying field of

research are discussed, the relevance is pointed out, and application possibilities are given.

Effects ofhydrogenand water

The effect of hydrogen and humidification on the combustion process has recently been

analyzed by Goke (2012) and Gockeler (2015) for premixed flames at ambient pressure con-

ditions. The authors indicated the chemical influence of water on NOx production. Lellek

(2017) referred to a disagreement found in the scientific community about the importance

of the chemical effect of water on the combustion process. Gockeler (2015) stated that the

chemical effect is strongly influenced by the choice of operating conditions. A detailed chem-

ical analysis with liquid water injection and pure hydrogen fueling, that is transfered to high

pressure regimes occurring in gas turbine applications, does not exist.

This study aims to show the effects of hydrogen and water on a non-premixed flame by a

radical analysis in the style of Goke (2012). However, this study is performed at gas turbine

conditions, using both a congruent NOx pathway study and a quantification of the effects of

water (thermal, chemical, dilution) on the NOx reduction. Within the framework of elevated

operational conditions, this study adds further parameters to the analysis, selected based on

a practical combustor control strategies. These parameter variations are comprised of the

combustor pressure, the combustor outlet temperature, the fuel composition (natural gas,

20 3 Objective and methods

hydrogen, and mixtures of both), the air inlet temperature, the air inlet velocity, and the

water mass flow. These parameters are increased up to real gas turbine conditions and also

contain part load conditions.

The practical relevance of this study is high due to the industrial design of the experiment

and the use of an actual gas turbine combustor. The limited accessibility of the combustor

and the provision of the operational conditions create unique challenges. The choice of

experimental boundary conditions and parameter variations, inspired by full and part load

operation, was made in a close agreement to real rather than academic operational conditions.

Combiningof findings

Gas turbine manufacturers make significant use of atmospheric pretests of a combustor in

order to keep the development costs at reasonable levels. Recent high hydrogen gas turbine

combustor concept developments, that may also include water injection, require an advanced

NOx correlation for estimating the corresponding high pressure NOx emissions. A correlation

for dry conditions and low hydrogen containing fuels was developed by Visser & Bahlmann

(1994). However, a correlation as a link between low and high pressure regimes that is vali-

dated for the chemical influence of hydrogen and water does not currently exist. This study

provides a correlation model for NOx prediction that is verified for high hydrogen content

fuel applications and water injection. The formulation is based on prevailing correlations in

literature (e.g. Rizk & Mongia (1994)) including correlations for the flame temperature and

the flame residence time.

The outcome of this thesis could reduce the development time and cost for future gas turbines

by improving the prediction quality of high pressure NOx emissions solely on the basis of

low pressure test data. Consequently, combustion test results under real conditions can be

acquired more specifically, and thus more cost-effectively. Furthermore, many gas turbine

simulation codes have been developed to estimate power plant performance in both design

and off-design conditions in order to establish an adequate control criteria or to determine

possible cycle improvements. The estimation of pollutant emissions is also important for

optimal performance and satisfying legal emission restrictions. During commercial operation,

a comparison between measured and predicted NOx emissions from a reliable correlation

would further serve as an important tool for detecting deterioration in the combustion process

at an early stage.

3.2 Methods

The objectives of this thesis require an experimental study (Chapter 4), a numerical study

(Chapter 5), and finally a correlation approach, which combines both findings (Chapter 6).

In the following section, the reasons for the selection of these methods are discussed and

their potential is highlighted.

3 Objective and methods 21

Highpressurecombustiontests

The experimental study aims to evaluate the use of hydrogen and natural gas combustion

with water injection as a NOx abatement measure under real conditions. The test speci-

men is a non-premixed can-type gas turbine combustor with a maximum thermal power of

10 MWth. It features a real industrial geometry and size to achieve the highest relevance

of the experimental results. The emissions characteristics of this combustor is measured on

the basis of a representative operation map. Besides full load condition, the characterization

map includes individual variations of global parameters according to part load operation

strategies of real gas turbines and parameters affecting relevant internal combustion proper-

ties, such as residence time and flame temperature. The parameter variations are comprised

of high and low pressure regimes, an adjustment of the air inlet temperature, the combustor

outlet temperature, and the air inlet velocity. Focus, however, is on the effects of hydrogen in

the fuel (0-100%) and of water injection, which are investigated for selected pressure stages

and fuel compositions.

Chemicalreactornetworkmodel

The real gas turbine combustor used for these experiments has a limited access for instrumen-

tation and thus experimental data are reduced to the global input and output parameters.

An internal investigation, for example, of the flame length and shape, the local flame tem-

perature, or the composition distribution, is not practical. Since these quantities are crucial

for understanding the chemical process, this lack of information is addressed in Chapter 5 by

a numerical reactor model. This generic semi-empirical model, set up in Cantera (Goodwin

et al. 2016) for MATLAB, distinguishes the flame zone and the dilution zone by a simple

chemical reactor network (CRN). The main benefit of the CRN approach is the ability to

model intermediate products with a non-equilibrium reaction process, which is a precondi-

tion for NOx emission computations, while the computational effort is relatively low. The

model requires the NOx results from Chapter 4 at a reference point in order to allocate the

residence time. The model is validated over the entire set of experimental data and allows for

the understanding of internal physical and chemical impacts of fuel composition and water

injection on NOx emissions. Focus of the numerical analysis is on the radical distribution,

the NOx pathways, and the quantification of the effects of water injection. The results also

form the basis of the NOx prediction tool developed in the next chapter.

SimplifiedNOx

emissionsmodel

For a general and simplified description of the underlying physics, the findings of the numeri-

cal model are presented in a simplified structure capturing the relevant influences. Chapter 6

presents this simplification of the NOx prediction process. The outcome is a set of correla-

tions which can estimate the residence time, the flame temperature, and the NOx emissions

of the combustor. The focus of the multi-dimensional NOx prediction correlation is on fuel

flexibility in terms of high hydrogen fuels and water injection at various pressure condi-

tions. Thus, it builds the bridge between fundamental combustion analytics and full-scale

combustor testing and enables high pressure NOx emission prediction solely on the basis

22 3 Objective and methods

of low pressure test data. The correlations are the outcome of the numerical study done

in Chapter 5 in terms of the selection of containing quantities for each target value and

the parametrization process itself. Finally, the application possibility of the correlations to

similar non-premixed gas turbine combustors is verified with a different combustor.

Outline In the following three chapters, the above mentioned methods are applied. Each of these

chapters starts with a review of the underlying methods based on a specific literature study.

Thereafter, the methods are described in detail. To the end of each chapter, the results are

presented, validated, and discussed in terms of the objectives of this thesis. A summary

concludes the general findings of each method.

4 High pressure combustion tests

The Institute of Power Plant Technology, Steam and Gas Turbines (IKDG) operates a high

pressure combustion test rig for 12 MWth reverse-flow silo gas turbine combustors from in-

dustrial gas turbines. Since 2009, characterization of state-of-the-art gas turbine combustors

and development of future combustion concepts have been done in close collaboration with

Kawasaki Heavy Industries Ltd. The test results are an abstraction of a real gas turbine

conditions due to the elevated pressure conditions like in an actual engine. Thus the test rig

can be considered as the last stage of development before the testing of the equipment in

actual engines. The test rig allows fuel variation of natural gas and hydrogen mixtures and

direct water injection.

This chapter describes the specification of this test rig with reference to the worldwide set of

similar high pressure combustion test rigs. Thereafter, the set up of the test rig is explained in

more detail focusing on the facilities, control principles, and measurement instrumentation.

Next, the test combustor and its operating range is presented before the experimental results

are given and discussed.

4.1 Worldwide test rig overview

In general, high pressure gas turbine combustion test rigs are limited to countries where

gas turbine manufacturers are located. The air pressure is an essential influencing factor

that is often scaled down in order to simplify combustion tests. Reproducing high pressure

conditions is a significant cost for combustion tests. The entire test rig has to be designed

for high pressure application and corresponding air mass flows. Additionally, the supply

facilities are comparatively greater and the heat transfer issues in the exhaust duct are

significantly more challenging to handle. Many of the test facilities are located in public or

national laboratories like DLR, JAXA, and NASA and are frequently used by gas turbine

manufacturers for research and development. Gas turbine manufacturers also usually run in-

house atmospheric combustor facilities. Tab. 4.1 shows a selection of high pressure combustor

test facilities including their rated capabilities. Criteria of this list is a pressure exceeding

5 bar and air mass flow exceeding 5 kg/s which are considered to be the lower thresholds

for industrial gas turbines. The underlying classifications for gas turbines are: micro gas

turbines at 1-500 kW, industrial gas turbines at 1-70 MW, and heavy duty gas turbines

70-500 MW power output.

24 4 High pressure combustion tests

Table 4.1: Overview of high pressure combustion test rigs for gas turbine combustors.

NameAir inlet conditions Outlet

Fuels / diluentsp [bar] mair [kg/s] Tair [◦C] T [◦C]

Europe

ENEL1 25 42 1230 NG, H2, CH4, CO2, C3H8,

C4H10, CO, N2, Liquid fuels

HBK 22 40 30 700 2130 Kerosine, oil, naphta, NG,

syngas

HBK 33 40 7 700 Kerosine, oil, NG

HBK 44 40 45 700 Kerosine, oil, naphta, NG,

Syngas, H2, CO, CH4, CO2,

N2

HBK 55 35 70 700 Kerosine, gas

HPAF6 23 10.5 753 NG, CO2, N2, H2, CO,

C3H8

HPCR7 16 5 630 H2, CH4, CO2

IKDG 24 12 550 1350 NG, H2, H2O

North America

DGTC8 10 0.75 330 NG, Liquid fuel

GTTL9 31 50 510 1440

GE10 34.5 4.5 540

GEAE11 24* 49* 500* 1430*

HPCRF12 40 15.4 593 1430* NG, Liquid fuel

NASA CE-5B13 19 5.5 730 1760

NASA ASCR14 62 22.7 650 1870

NRC-TC115 21.5 21.7 650 NG, Kerosine, oil

NRC-TC215 24 21.7 650 NG, Kerosine, oil

NRC-TC315 6.4 19.3 650 NG, Kerosine, oil

Solar16 7.5 16 316

PSM17 24 27 650 1930 NG, H2, CH4, C2H6, C3H8,

C4H10

Asia

JAXA18 50 4 730 1730 H2, CH4, Kerosene

MHI19 16 50 1500 NG, Oil, Methanol

*estimated. Sources: 1Cocchi et al. (2007, 2008), Cocchi & Sigali (2010), Riccio et al. (2007),

Tomczak et al. (2002). 2German Aerospace Center (2017a). 3German Aerospace Center (2017b).4Gadde et al. (2006), German Aerospace Center (2017b), Wu et al. (2007). 5Hassa (2014). 6Lam& Parsania (2016), Liu et al. (2012, 2013). 7Bagdanavicius et al. (2010), Gas Turbine ResearchCenter (2007), Shelil et al. (2010). 8Sidwell & Straub (2017). 9Feigl et al. (2006), Karim et al.(2017), Myers et al. (2003). 10Brdar & Jones (2000), Goldmeer et al. (2017), Maughan et al.(1994). 11Mahajan et al. (2000). 12Wright-Patterson (2016). 13Severt (2018a,b). 14DeLaat (2009),Severt (2018a,c). 15NRC Canada (2015). 16Szweda et al. (2005). 17Oumejjoud & Stuttaford(2007), Stuttaford et al. (2010, 2016). 18Shimodaria et al. (n.d.). 19Hashimoto et al. (2009),Tsukuda et al. (2001).


4.2 Test rig set up

The following section describes the set up of the combustion test facility. The infrastructure

can be distinguished into the air supply unit, the fuel supply unit, the water supply unit

for direct water injection and exhaust gas cooling, the combustion system, the exhaust gas

analysis section, and the exhaust gas path. Fig. 4.1 shows a simplified component and flow

diagram of the test rig. In the following section, each subsystem is described in more detail.

M

Exhaust gas handling

Air supply system

Fuel system

Exhaust gas analysisCombustion

system

H2NG

Water injection system

Q,T

H2O

10

H2

8

65

4

7

3

21 9

11

12

14

1516 17 18

1920

21

22

23

26

25

24

13

(1) Inlet guide vanes, (2) Intercooled compressor (3) Compressor bypass valve, (4) Controlvalve I, (5) Blower, (6) Gas combustor, (7) Air preheater, (8) Measuring orifice, (9)

Natural gas tanks, (10) Hydrogen trailers, (11) Pressure reduction valves, (12) Mixingsystem, (13) Mixture pressure reduction valve, (14) Fuel control valves, (15) Water tank,(16) Pump, (17) Water control valve, (18) Flow measurement, (19) Air plenum, (20) Test

combustor, (21) Temperature sensor, Gas analyzer, Camera, (22) Pump, (23) Quenchcooler, (24) Control valve II, (25) Silencer, (26) Stack.

Figure 4.1: Flow diagram of the test rig.

4.2.1 Air supply system

Com-pression

The air supply system, shown in Fig. 4.1 (1)-(8), provides pressurized air to the combustion

system. The six-stage radial gear compressor (2) is driven by an electric engine at constant

rotational speed. It features intermediate cooling on the first four stages. The compressor


mass flow is controlled by variable inlet guide vanes (1) for the first three stages. The air

mass flow rate to the combustion system can additionally be adjusted by venting a part of

the compressed air through a bypass valve (3) behind the compressor.

Control Control valve I (4), which is located between the compressor and the air heater, is used to

control both the mass flow and the air pressure in order to fulfill the combustor requirements.

This expands the range of air conditions to lower pressures and mass flows since the air

compressor must be operated within its choke and surge limits, and therefore the compressor

outlet pressure may be significantly higher than the combustor operating pressure.

Preheatingandhumidity

The air is discharged after the sixth stage of the compressor with a temperature of around

160 ◦C. In order to reach typical combustor inlet temperatures, the air temperature is further

increased in a 10 MWth natural gas fired air preheater (5)-(7) via convective and radiative

heat transfer. The humidity of the air at preheated conditions is negligible. Assuming a

saturated air condition downstream of the last intercooler behind the fourth compressor

stage (10 bar, 35 ◦C), the air is further pressurized and heated so that the relative humidity

is less than 0.4% at full combustor conditions.

Piping andmeasure-ment

The piping system is insulated both to reduce the thermal heat loss and to reduce the

preheating time. This also allows for efficient air temperature control at the combustion

system. The combustor inlet air mass flow is measured via an orifice (8) just before entering

the combustion system.

4.2.2 Fuel supply system

Specifica-tion

The fuel supply unit, shown in Fig. 4.1 (9)-(14), provides hydrogen, natural gas or mixtures

of both to the combustion system and enables control of the fuel flow. The fuel mixing

system has been designed for enough fuel of each component to produce up to 12 MWth.

The maximum hydrogen flow rate is 4000 Nm3/h (360 kg/h) and the maximum natural gas

volume flow rate is 1200 Nm3/h (970 kg/h).

Storage Compressed natural gas is taken from the local grid and stored in a 16 m3 storage tank (9)

at a pressure of 200 bar. Hydrogen is delivered by external suppliers. The hydrogen docking

station is capable of connecting to two mobile trailers (10) at 200 bar. Simultaneous trailer

operation and trailer switching is possible and thus an unlimited hydrogen supply is feasible.

Pressure reduction valves (11) allow the natural gas injection pressure and mixture pressure

to be adjusted according to the requirement of the fuel control valves at the combustor unit.

Mixing The mixing system is capable of supplying fuel mixtures of natural gas and hydrogen from

0 up to 100 vol.% H2. The mixture is controlled by a cascaded volume flow measurement,

each featuring a control valve. The system controls the mixing facility outlet pressure by a

mixture pressure control valve (13).


ControlThe mixture pressure control valve (13) and the natural gas pressure control valve (11) can

deliver fuel pressures up to 50 bar and 40 bar, respectively. Fuel control valves (14) allow

for the continuous adjustment of the fuel inlet mass flows. Four individual fuel lines for

each fuel path (natural gas and mixture) access up to four individually-controllable injection

positions of a combustor. All eight fuels lines feature mass flow measurement and control.

4.2.3 Water injection system

The water injection system, shown in Fig. 4.1 (15)-(18), supplies water for direct combustor

injection. The system is capable of providing a water mass flow rate of up to 1000 kg/h

at a maximum pressure of 75 bar. Demineralized water is used in order to avoid mineral

depositions on the sensitive combustor structure. It is produced with an in-house water

demineralization plant that is capable of producing 2 m3/h of demineralized water and

consists of a water-softening stage, a reverse osmosis system, and a mixed media filtration

system. The demineralized water is stored in a 100 m3 water tank (15). In operation, the

water is pumped towards the combustor (16) while the water mass flow is adjusted by a

water control valve (17) and measured with an impeller measuring device (18).

4.2.4 Combustion system

Air plenumThe combustion system, shown in Fig. 4.1 (19)-(20), consists of an air plenum, the test

combustor, and the transition duct. The schematic set up of the combustion system is shown

in Fig. 4.2. The air plenum (19) collects the combustion air and guides it to the annular

passage of the combustor inlet geometry. The plenum features three major functions: First,

it decelerates the air to establish a uniform annular combustor inlet flow. Second, it is set

up as heat exchanger to cool the transition duct which separates the flows of the combustion

air inside the air plenum from the exhaust gas. Third, it provides the connection to the

combustor (20) and the exhaust measuring system.

Transitionduct

The combustor itself is a can-type combustor with reverse-flow configuration. It will be

described in detail in Sec. 4.4. The hot exhaust gases of the combustor pass through the

transition duct before entering the measuring section. The transition duct is a nickel-based

alloy tube with an inner thermal barrier coating. It is followed by the measurement section

at the downstream side.


T

Qmair

Temperature probe

Exhaust sampling

probePlenum

CombustorMeasuring

sectionTransition duct

Exhaust gas

Fuel

Air

p

Tair

Gas analyzer

Water

Δpcomb

Figure 4.2: Overview of the combustion system and the exhaust measuring section, including

the gas analysis system.

4.2.5 Measurement section

The exhaust gas analysis duct, as seen in Fig. 4.1 (21), is connected to the air plenum.

The hot combustion gases enter this section from the transition duct. Fig. 4.2 gives a

schematic overview of the arrangement. The measuring section consists of a cone-shaped

duct and is equipped with an exhaust gas temperature probe and access for the exhaust gas

sampling probe. It additionally features optical access to the combustor. With d denoting

the combustor exit diameter, the position of the temperature probe is about 6.0d behind

the combustor exit. The distance between the temperature probe and the exhaust sampling

probe is about 2.5d.

Combustoroutlettemperature

The combustor outlet temperature is measured via a water-cooled probe made by Yamari

Industries Ltd. It comprises five R-type thermocouples placed along the diameter of the

exhaust duct, as shown in Fig. 4.2. The combustor outlet temperature, T , is the arithmetic

mean temperature of all five temperature measuring points of this probe. For the purpose

of determining the combustor outlet temperature, the influence of the heat exchanger (19)

is considered by estimating the temperature differences based on heat transfer calculations.

As a results, the temperature drop in the transition duct is below 10 K. The overall accuracy


of the hot gas temperature measurement is about ±10 K. Details regarding to the accuracy

are given in Sec. 4.3.

Exhaust gasanalysis

A sample gas stream is extracted by a water-cooled exhaust gas sampling probe in order to

perform exhaust gas composition analysis. The exhaust gas sample is collected through five

radially distributed ports along the flow stream. The gas sampling line is a heated flexible

tube that is kept at a temperature of ∼200◦C to prevent condensation and loss of NO2 by

absorption in water. The composition of the exhaust gas is measured with the Mexa One

gas analysis system from Horiba. This system is capable of measuring the concentration of

CO, CO2, O2, unburnt hydrocarbons (UHC), and NO/NO2. Details regarding the accuracy

of these measurements are given in Sec. 4.3.

Opticalaccess

The flame inside the combustor can be observed and investigated by cameras through an

optical access point downstream of the combustor (see Fig. 4.1 (21)).

CoolingThe measuring section and the pressurized vessels of the exhaust gas path have extremely

high thermal loads at elevated power operation. Thus, these parts feature active wall cooling.

Cooling media are both water and shop air.

4.2.6 Exhaust gas path

The main functions of the exhaust gas path, shown in Fig. 4.1 (23)-(26), are to simulate of

the pressure drop in the turbine and to discharge the exhaust gas into the atmosphere. The

temperature of the exhaust gas is reduced upstream of the pressure reduction valve in order

to meet the temperature specifications of the downstream devices.

Quenchcooler

A quench cooler (23), located downstream of the exhaust gas analysis duct, is used to

decrease the exhaust gas temperature to the permissible temperature of the control valve

II (24) and the exhaust gas path. Demineralized water from the tank (15) is pumped (22)

towards the quench cooler and injected into the hot exhaust stream via multiple nozzles.

Controlvalve II

The control valve II (24) is a multi-stage steam valve which is located downstream of the

quench cooler. It features a maximum temperature rating of 550◦C. It reduces the pressure

of the exhaust gas flow and thereby represents the pressure drop of the turbine. This config-

uration allows for control of pressure level in the combustor independent of the combustor

mass flow rate.

Gasdischarge

The exhaust gas path furthermore ensures a safe discharge of the expanded exhaust gas into

the atmosphere through the stack (26). A perforated cone silencer (25) is positioned inside

the exhaust duct for acoustic damping to lower noise emissions.


4.2.7 Control principles

In summary, the control of the inlet air consist of the variable guide vane control of the air

compressor (1), the bypass at the air compressor exit (3), control valve I in the main line

between the air compressor and the air preheater (4), and control valve II (24) downstream of

the quench cooler, shown in Fig. 4.1. The air temperature is controlled via the air preheater.

The interaction of all control devices allow for the efficient operation of the test rig.

Air pressure The air mass flow rate and pressure are controlled by the combined adjustment of the

compressor inlet guide vane control, control valve I (4) in the air main line, and control

valve II (24) in the exhaust tract. Although the compressor bypass valve (3) is primarily

used to start the compressor, it can also be used as a control device for part load operation.

The advantage of using a back-pressure control valve is the ability to adjust the combustor

pressure independently of the exhaust gas mass flow and exhaust gas temperature. Note that

the volume flow upstream of control valve II (24) affects the pressure and mass flow, which

are also significantly influenced by the combustor outlet temperature, T , and the mass flow

of the injection water into the quench cooler. The flexibility of the control system, however,

allows for operation with constant parameters at all times.

Airtemperature

The air preheating setup allows for the combustor air inlet temperature to be adjusted

independently from the air pressure. This allows for simulating gas turbine compressors

with varying polytropic efficiencies as well as gas turbine processes featuring an intercooler

or recuperation.

Variablefueloperation

The setup of the fuel supply system allows for combustor operation with natural gas and

a fuel mixture (0-100 vol.% H2) at the same time. For example, while the pilot flame is

operated with natural gas, the main flame or a staged flame can be operated simultaneously

with hydrogen elevated mixtures or even pure hydrogen.

4.3 Measurement techniques

The measurement techniques are essential for both monitoring conditions for operation con-

trol and acquiring test data for evaluation. Occasionally, these applications may have dif-

ferent requirements for measurement accuracy. In the following section, the relevant mea-

surement systems are introduced and their accuracy is discussed. Note that only stationary

operating conditions have been recorded. In order to have a precise image of the combustion

state, the operating conditions must show stable values before the measurement is initial-

ized. Tab. 4.2 contains a summary of measuring devices used for this study, including both

auxiliary intermediate and final measurement sensors. For the latter, the actual absolute

accuracy is given in the last column. The discussion of the measurement techniques includes


Table 4.2: Summary of underlying measurement accuracies

Variable Measurement principle Range relative (FS)/

absolute accuracy

pambient Absolute pressure sensor 0 - 1300 mbar 0.1%

p Combustor pressure sensor 0 - 25 bar 0.1% / 2.5 mbar

pair Air pressure sensor 0 - 69 bar 0.075%

pfuel Fuel pressure sensor 0 - 63 bar 0.1%

∆pair Differential pressure sensor 16 - 1600 mbar 0.1%

∆pfuel Differential pressure sensor 2.5 - 250 mbar 0.1%

∆pNG,∆pH2Differential pressure sensor 2.5 - 250 mbar 0.1%

Tair Thermocouple type K -270 - 1300◦C ±8 K

Tfuel Thermocouple type K -270 - 1300◦C ±8 K

T Thermocouple probe type S -50 - 1768◦C ±10 K

mair Orifice ±2.0% / ±0.34 kg/s

mfuel Orifice NG: ±2.1% / 16.8 kg/h

H2: ±2.1% / 6.3 kg/h

mwater Impeller 2.0 - 20 l/min ±1% / max 0.2 l/min

x Orifices 0 - 100 vol.% H2 ±0.5 vol.% H2

cNOxCLD 0 - 100 ppm ±2%/± 2 ppm

0 - 1000 ppm ±2%/± 20 ppm

cO2PMD 0 - 20.5 vol.% ±2%/±0.41 vol.%

cCO2NDIR 0 - 15 vol.% ±1.75%/± 0.26 vol.%

cCO NDIR 0 - 1000 ppm ±2.0%/± 20 ppm

0 - 50 ppm ±2.0%/± 1 ppm

the absolute, relative, and differential pressure measurement in Sec. 4.3.1, the temperature

measurement for air and exhaust gas in Sec. 4.3.2, the mass flow measurement via the volume

flow for air, fuel, and water in Sec. 4.3.3 and gas composition measurement for fuel mixture

and exhaust gas in Sec. 4.3.4.

4.3.1 Pressure Measurement

The pressure measurement is necessary for both independently determining the fluid pressure

and for determining derived variables like the mass flow rate via an orifice. In general,

absolute, relative, and differential pressure measurements are performed separately. Note

that in this thesis, combustor pressure, p, is referred to as the absolute pressure. An absolute

pressure sensor is used to determine the atmospheric pressure. The other measurement


devices for fluid pressure measurement are relative pressure sensors. Relative pressure sensors

are used to determine the pressures of the air and fuel. The air supply pressure is determined

in the plenum while the relative pressure for mass flow determination is measured within the

pipe, as defined by standards.

Differential pressure sensors are used to measure the static-to-static pressure difference in-

duced by a flow through an orifice in order to determine the mass flow rate. Orifice measure-

ment is used for the rate measurement of the air and the fuel mass flow. The fuel composition

is also determined using the mass flow measurement of its components via orifices.

4.3.2 Temperature measurement

The temperature is a measure of the thermal motion of the fundamental particles in a fluid.

It is indirectly measured via a second system that is in thermodynamically equilibrium with

the fluid. A static and total temperature have to be distinguished for moving fluids. In

this thesis, the temperature is additionally an indirect measurement of the density during

the mass flow rate measurement and of the fluid entering and exiting the combustor. Due

to the principles of fluid dynamics, the recovery factor has to be considered in temperature

measurements.

Recoveryeffect andradiativeheattransfer

Within the framework of this thesis, contact thermocouples are used. In a flow, however,

the impact of the kinetic energy of the flow (recovery effect) and radiative heat transfer

have to be considered. The recovery effect is correction due to the adiabatic damming up

of the flow. If a streaming flow is slowed down (e.g. by friction in the shear layer of a

temperature sensor) a portion of the kinetic energy is converted into heat. By assuming

that particles are complete decelerated to 0 m/s, a complete conversion of kinetic energy

into heat with adiabatic conditions results in increase of the temperature from the (static)

temperature to the total temperature. The radiative heat transfer between the thermocouple

and its environment (e.g. the surrounding walls or another radiation source) may also lead

to a significant measurement error. More details about these fundamentals can be found

in Bernhard (2014).

Airtemperature

The air inlet temperature measurements inside the plenum via thermocouples are not sig-

nificantly influenced by the aforementioned phenomena. The influence of the heat radiated

by the transition duct on the temperature measurement is negligible because the thermo-

couples are protected by the combustor inflow geometry, as seen in Fig. 4.2. The transition

duct partially exchanges the heat via radiation with the plenum wall. However, the differ-

ence between the highest and lowest wall temperature measurements of the material within

line-of-sight was not more than 3 K.


Gas sampling probe

Temperature probe Thermocouple Water jacket

Ceramic coatingHeat shield

Figure 4.3: Temperature probe. Left: Temperature probe and gas sampling probe mounted

inside the measuring section. Right: Detail of a single thermocouple integration.

The radiative heat transfer is neglected for air and fuel temperature measurements taken via

an orifice in the air supply line. In case of air temperature measurements, due to the pipe

insulation, it is assumed that the fluid temperature and the inner wall temperature of the

pipe do not differ significantly. The recovery factor is neglected, as well. In case of the fuel

temperature measurement, the fuel temperature is approximately the ambient temperature,

and thus significant heat transfer is not expected. The recovery factor also is neglected due

to comparably low flow velocities.

Probedesign

The combustor outlet temperature, T , is the arithmetic mean temperature of all five tem-

perature measurement points on the exhaust temperature probe, as shown in Sec. 4.2. In

the following, the measurement accuracy is discussed. Fig. 4.3 shows the temperature probe

in more detail. On the left side, the integration of temperature probes into the measuring

section is shown. This picture has been taken from an upstream position, with a view in

the direction of the exhaust gas flow. The arrangement of a single R-type thermocouple,

embedded in a ceramic-coated heat shield, is displayed on the right. In the following, the

radiative heat transfer and the recovery effect are discussed.

RadiationAs shown in Fig. 4.2, the thermocouples are directly facing the radiated heat from the flame.

Due to the non-premixed combustor configuration, the radiation source temperature is the

stoichiometric adiabatic flame temperature. A 1D heat transfer estimate shows that the

influence of the radiated heat is negligible since the surface of the thermocouple is small


compared to the size of the flame and the distance. Heat transfer via convection has also

been neglected because the temperature difference between the sensor and the flow velocity

is low, as the thermocouple itself is not placed in the core flow but protected inside in a

slot in the sensor. Each thermocouple tip resides in a decelerated flow area of the sensor.

The right side of Fig. 4.3 shows the tip of a thermocouple protected inside in a slot of the

probe. Since the heat shield features a thermally-insulating ceramic coating, the temperature

difference between the thermocouple and the heat shield is negligible and thus heat transfer

between the two can be neglected. The heat release of the thermocouple via radiation to

the walls of the measuring section and the heat transfer of the thermocouple casing pipe are

also negligible since the temperature difference (the driving force of the heat transfer) has

been determined to be very small.

Recoveryeffect

A flow field analysis shows that the velocity of the exhaust gas is comparably small. The

recovery effect is thus negligible because of the small Mach numbers of the exhaust gas flow

(Mach≤ 0.1). The thermal influence of the kinetic energy of the flow is thus lower than

1.3 K.

Transitionduct

The thermocouple probe measures a slightly reduced exhaust gas temperature caused by

the heat transfer from the hot exhaust gas to the inlet air within the pressure vessel, (see

Fig. 4.1 (19)). For the purpose of determining the exhaust gas temperature at the exhaust

gas temperature probe, the influence of the heat exchanger is determined by temperature

differences estimated from heat transfer calculations. A 1D calculation has been set up,

providing a perpendicular flow in the downstream portion of the transition duct and an

axial flow in the upstream part. Two sections of the transition duct have been modeled. In

the first section, the air streams axially along the outer wall of the transition duct, while

in the second section, the air has been implemented as a radial cross flow (see Fig. 4.2).

The resulting temperature difference between the exhaust gas in the transition duct and the

combustor outlet to the temperature probe is less than 10 K.

Probeaccuracy

In summary, the accuracy of the R-type thermocouple is given by the manufacture as ±2 K.

Furthermore, the electronic data transducer features an uncertainty of ±1 K (due to volt-

age drift), with a cold junction compensation by a semiconductor linear temperature sensor

(±5 K uncertainty estimated) and interfering frequency suppression. Thus, the total accu-

racy of the hot gas temperature measurement is about ±10 K, which is less than 1% FS.

4.3.3 Mass flow measurement

The mass flow measurement of combustion air and fuel is performed via individually-designed

orifices. The design and accuracy evaluation is based on the standard DIN EN ISO 5167

(2004). Almost all inner pipe diameters of the fuel orifices do not meet these standards due


to their small size, but experience has shown that smaller orifices are sufficiently accurate.

For mass flow measurements, the fluid density is computed from the pressure and temper-

ature measurement. The density is computed using the reference fluid thermodynamic and

transport properties database Refprop by NIST (2013). An accuracy evaluation based on

these standards shows that the measurement uncertainty of the orifices is below ±2.0% of full

scale (FS) for combustion air and below ±2.1% FS for fuel. The accuracies of the pressure

and temperature measurements are thereby included. As a result, the maximum uncertainty

of the combustor air mass flow rate is ±0.34 kg/s. The uncertainties of the fuel mass flows

are max. 16.8 kg/h for the natural gas fuel system and 6.3 kg/h (H2) and 16.9 kg/h (NG)

for the blend fuel system. Note that the relative humidity of the combustor air is negligible

within the entire combustor pressure range, as already been outlined in Sec. 4.2.1. The

accuracy of the fuel composition is based on the accuracy of the individual mass flow mea-

surements of the two fuel components natural gas and hydrogen. The accuracy of the fuel

mixture composition is kept within a range of ±0.5 vol.% H2 for the entire mixing range.

The accuracy of the water mass flow measurement by the impeller is ±1% of the measured

value.

4.3.4 Emission analysis

Exhaustsampling

In order to analyze the chemical composition of the exhaust gas, a sample gas stream is

extracted by a water-cooled exhaust gas sampling probe, shown in Fig. 4.2. The exhaust gas

sample is collected through five ports which are radially distributed along the flow stream,

see Fig. 4.3 (left). The temperature of the extracted exhaust gas is influenced by water

cooling. In order to correct the extraction temperature introduced by the probe, exhaust

gas is bypassed via a control valve in order to adjust the mass flow rate of the sample gas.

Water condensation also needs to be prevented at any time, since condensed water affects

NOx measurements due to NO2 absorption in water. The condensation threshold is also

dependent on the operating pressure, air inlet temperature, the fuel composition, and the

amount of water injection.

Exhausttransport

The extracted gas is filtered before a temperature-approved spring regulator valve reduces

the sample pressure to ∼2 bar. As the pressure decreases, as does the lower temperature

threshold for condensation prevention. The sampling gas is then fed into a heated, flexible

tube that maintains at a temperature of nearly 200 ◦C to prevent condensation and loss of

NO2. The flexible tube finally feeds into the gas analyzing system.

Gasanalyzer

The composition of the dry exhaust gas is measured with the Mexa One gas analyzer manu-

factured by Horiba. This system is capable of measuring the concentration of CO, CO2, O2,

Unburnt Hydrocarbons (UHCs) and NOx. The gas analyzer consists of a hot wet measuring


line for UHC and NO/NO2 and a cold dry measurement line for CO2, CO, and O2. For

the dry line, H2O is removed by a cooler with a condensate trap. The wet measuring line

is heated and feeds into the line for UHCs and NO/NO2 measurement. NO/NO2 (nitro-

gen monoxide and nitrogen dioxide) are measured by chemiluminescence detection (CLD).

The CLA-01HV model analyzer used for this study has a measurement uncertainty of max.

±2 ppm for NOx < 100 ppm and max. ±20 ppm for NOx < 1000 ppm, respectively, depend-

ing on the actual operating point. In this study, the NO and NO2 emissions are summed

and thus one value represents the NOx concentration. The O2 concentration is measured by

paramagnetic detection (PMD). The MPA-01 model used in this study has a measurement

uncertainty of max. ±0.41 vol.% with 20.5 vol.% full scale. The CO2 and CO concentra-

tions are measured by non-dispersive infrared spectroscopy (NDIR). In the case of CO2, the

AIA-32 model analyzer used in this study has a measurement uncertainty of ±0.26 vol.% FS

(noise ±0.75% FS and linearity ±1% FS) where full scale is 15 vol.% for CO2. The CO

concentration is measured by a AIA-11 model analyzer that has a measurement uncertainty

of max. ±10 ppm for CO < 1000 ppm and max. ±1 ppm for CO < 50 ppm, respectively,

depending on the actual operating point (noise ±1% FS and linearity ±1% FS).

More details about combustion engine emission measurement principles and the general

operation of the above mentioned analyzers can be found in Adachi & Nakamura (2014).

Drift checks have been performed after each test. The drift check showed less than 1%

variation for all channels.

NOx

referenceIn gas turbine research and commissioning, NOx values are usually given in reference to the

residual oxygen concentration in the exhaust in order to make comparisons between different

engines. Usually, a value of 15 vol.% oxygen concentration within the dry exhaust gas is used

as reference. In this study, however, the NOx emissions have not been referenced in order

to eliminate distorting influences during the data comparison with the numerical results in

Chapter 5 and during the set up of a physical correlation for NOx in Chapter 6.

4.4 Test combustor

The test combustor, shown in Fig. 4.4, features a non-premixed design and is capable of

producing thermal energy of about 10 MWth. This can-type combustor has a reverse flow

configuration and consists of a pressure casing, a flame tube with coated liner and dilution

holes, an air guiding plate, and the air inlet swirler.

Typical of reverse-flow combustors, air enters the combustor from the downstream side

through an annular channel between the combustor casing and the liner. The wall of the

liner is louvered for cooling and contains arrays of dilution holes. A portion of the air in the

combustor penetrates the liner through a set of dilution holes and annular gaps for cooling


Liner

Swirler

Dilution hole

Fuel nozzle Liner cooling CasingWater nozzle

End plate

Figure 4.4: Combustor drawing courtesy of Kawasaki Heavy Industries, Ltd.

and for adjusting the temperature of the burned gas to the desired turbine inlet temperature.

The remaining combustion air is bent both towards the center and back at the end plate

and enters the combustion volume through a swirler.

The fuel is injected into the shear layer between the incoming air and the recirculated hot

products via the main fuel nozzle. In Fig. 4.4, an inner pilot fuel nozzle and an outer main

fuel nozzle are shown. The use of pure hydrogen fuel is possible due to an additional air slit

to cool the tip of the fuel nozzle.

The water injection nozzle is located in the center of the fuel nozzle. This nozzle sprays the

water axially into the combustion volume at a specific cone geometry.

The hot exhaust gases are mixed with the cooling air before leaving the combustor. Note

that all geometric dimensions are normalized on the outlet diameter d of the combustor in

this thesis.


4.5 Operation parameter and variation

The NOx emissions of the combustor are characterized at various load conditions, pres-

sure conditions, fuel concentrations, and water injection levels (for NOx abatement). The

characterization is done by varying the most influential variables (listed in Tab. 4.3) in the

given range. The conditions during the variation away from the reference point are briefly

described now for all input parameters.

Table 4.3: Operational parameter ranges and reference point.

Name Unit Description Parameter range Steps* Reference

p [bar] Pressure 3 - 24 4 16

T [◦C] Combustor outlet temp. 900 - 1300 5 1300

x [vol.% H2] Hydrogen fraction of fuel 0 - 100 6 0

vrel [-] Relative air inlet velocity 0.6 - 1.2 4 1

Tair [◦C] Air inlet temperature 400 - 500 3 500

ψ [-] Water fuel ratio 0 - 4 6 0

*max. number of steps covering the parameter range

Pressure The pressure is varied from ignition conditions (around 3 bar) to the maximum pressure of

the combustor (24 bar). The combustor air inlet velocity is kept constant by adjusting the air

mass flow rate. The fuel mass flow rate is adjusted to set the combustor outlet temperature

to a specific value. The reference pressure of 16 bar is used as the best representation of

test rig operation ability, as 3 bar tests are more difficult due to significantly longer control

response times. The upper limit of 24 bar is used as it is near the upper limit of the test rig.


The combustor outlet temperature is mainly controlled by the fuel flow rate, while all other

operating parameters remain at their target values. While varying the temperature, the air

inlet velocity and thus the air mass flow rate are kept constant. The combustor pressure is

also kept constant and consequently, the combustor outlet volume flow is adapting due to

the changing density.

Fuelcomposition

The reference fuel for the tests is high calorific natural gas. Its composition is a monthly

average given in the Appendix A. The hydrogen fuel has a given purity level of ≥ 99.9 vol.%.

The comparison of different fuels is made by keeping the power output of the combustor

constant. Consequently, the combustor outlet temperature is kept constant. Due to the

lower specific LHV of hydrogen, the fuel mass flow rate is significantly lower during hydrogen

operation. Note that the energetic 50% mixture of natural gas and hydrogen is at about 80

vol.% H2.


Air inletvelocity

The air inlet velocity is defined as the air inlet volume flow rate relative to the outlet surface

of the combustor given by the diameter d. The air inlet velocity is related to a reference value

vref that is characterized by an overall pressure loss ratio of 4.81% with all other parameters

being at reference conditions. Variation of the air inlet velocity significantly affects the air

inlet mass flow rate. Consequently, this variation mainly affects the pressure loss ratio of

the combustor.

Airtemperature

The reference air inlet temperature (500◦C) represents the compressor outlet temperature

at full pressure conditions of 24 bar. The lower limit (400◦C) is the compressor outlet

temperature at a reference compressor outlet pressure of 16 bar under typical gas turbine

compressor efficiency. In order to keep the air inlet velocity constant, the air inlet mass

flow rate is adjusted according to change the density. The reduced air inlet temperature

is compensated for by increasing the fuel mass flow in order to keep a constant combustor

outlet temperature.

Waterinjection

The amount of injected water is described by the water-to-fuel mass flow ratio, ψ. There

are basically two operating strategies for water injection. First, the combustor outlet tem-

perature is kept constant during an increase in the water mass flow rate. Second, the water

mass flow can be increased without adjusting the fuel mass flow. Consequently, the com-

bustor outlet temperature decreases. In this study, the first strategy is applied in order to

keep the measured combustor outlet temperature constant, similarly to all other parameter

variations. In general, ψ is increased until the maximum supply mass flow rate of the water

injection system is reached or indications of combustion instabilities are observed.

Dataacquisitioncriteria

Data acquisition was performed at stationary operation conditions. Main indicators of a

stable operating point is a constant rate of NOx emissions and a constant exhaust gas

temperature. All data acquired are mean values over a 10 s interval. The data sampling

rate is 1 data point per second. Due to the large consumption of fuel and electricity, the

parameter set was strategically reduced. While the reference pressure is characterized in

detail, parameter variations were reduced at other pressure stages without major effects on

the quality of the data. However, all threshold values were always acquired with the following

exceptions. At 24 bar, the reference combustor outlet temperature was not acquired. The

air inlet temperature and air inlet velocity variations were only acquired at the reference

pressure and at 3 bar. The water injection is only applied at reference pressure and at 3 bar,

and at reference velocity. In total, 187 data points were acquired.


4.6 Experimental results

In this section the experimental results are presented. In the first part, measurements are

verified for consistency. The air fuel ratio is determined by two independent methods to com-

pare the mass flow and emissions measurements. Furthermore, the sample gas temperature

is shown to meet and to exceed the threshold value defined by the saturation temperature

in order to prevent condensation in the gas measuring line. Finally, the pressure loss ratio

of the combustor is shown. The second part contains the emissions results of the combustor

for the defined operating parameter range. First, the NOx emissions at every operating

parameter in this study is shown, namely pressure, combustor outlet temperature, hydrogen

content, air velocity, air inlet temperature, and amount of water injection. Finally, the CO

emissions at select operating conditions are compared to the NOx emissions.

4.6.1 Air fuel ratio verification

The air fuel ratio can be determined by two independent measurement techniques. Thus,

this quantity allows for verification of the consistency of the combustor inlet mass flow

measurements and the emissions measurement. The air fuel ratio can be directly determined

via air and fuel mass flow rate measurements (AFR). Additionally, the air fuel ratio can

be calculated from the emissions measurements, namely the CO2 concentration and the

residual oxygen concentration in the exhaust (AFRexh). This procedure is possible for all fuel

mixtures since the H2O is eliminated from the sample gas and the CO2 and O2 concentrations

refer to the dry conditions. Both procedures are briefly described here.

Definition The air fuel ratio can be calculated according to its definition

AFR =ma

mf, (4.1)

where m is the mass flow rate of air (a) and fuel (f). With the assumption of complete

combustion and inert nitrogen chemistry, the air fuel ratio can also be determined solely on

the basis of the residual O2 concentration and the CO2 concentration in the exhaust gas.

Assuming the fuel components to be H2 and CH4 only, the stoichiometric oxygen demand

per mole of fuel is (nO2,st

nf

)= x νst,H2

+ (1− x) νst,CH4, (4.2)

where νst is the required stoichiometric demand of O2 with νst,H2= 0.5 and νst,CH4

= 2. The

amount of oxygen that is consumed during the combustion is determined via oxygen balance

of the combustor by

∆nO2= naXO2,a − nexh,dryXO2,exh,dry, (4.3)


where XO2,exh,dry is the molar fraction of oxygen within the dry exhaust gas and XO2,a = 0.21

is the molar oxygen content in the combustion air. Note that the exhaust gas is dried for

the purpose of measurement, as described earlier. The molar exhaust flow is determined via

the nitrogen balance nN2,a = nN2,exh of the combustor

nexh,dry =nN2,exh

XN2,exh,dry=

nN2,a(1−XO2,exh,dry −XCO2,exh,dry

) . (4.4)

The resulting molar fuel flow per mole of air is

nfna

=∆nO2

nO2,st. (4.5)

Finally, the AFRexh via the exhaust gas composition measurement is

AFRexh =naMa

nfMf. (4.6)

ComparisonThe comparison of the two calculated air fuel ratios is given in Fig. 4.5. The abscissa

shows the measured mass flow fraction (Eq. 4.1) and the ordinate shows the air fuel ratio

determined via the exhaust composition (Eq. 4.6). The values show consistent agreement.

Note that the AFR of the reference point is about 40 and increases up to 100 during hydrogen

fueling due to the higher specific heat of hydrogen. During maximum water injection with

iso-combustor outlet temperature, the AFR lowers to 34 with natural gas fuel (ψ = 1.25)

and to 81 with hydrogen fuel (ψ = 4). With leaner conditions the scattering of the values

indicate a stronger a deviation of both values. However, this condition manifests solely for

lowest combustor outlet temperatures during hydrogen fueling.

ConclusionIn conclusion, the comparison of the AFRs determined via both the combustor inlet mass

flow measurements and the combustor outlet composition indicates consistent measurements.

This confirms the measurements of the inlet mass flow rates of fuel and water and excludes

systematic errors in the determination of the exhaust gas composition.


Figure 4.5: Comparison of AFR and AFRexh for all operation points.

4.6.2 Sample gas temperature

In order to ensure an accurate NOx measurement, it is essential to keep the sample tem-

perature higher than the saturation temperature of water in the exhaust gas. The following

section evaluates this criteria. Therefore, the dew point is determined for the prevailing

partial pressure of H2O by Refprop for MATLAB (NIST 2013).

Saturationtemperaturecharacteris-tic

Fig. 4.6 shows saturation temperature (the dew point) for various operating points without

water injection. The main influencing parameters are the combustor outlet temperature and

the hydrogen fraction of the fuel. Both increase the fraction of water in the exhaust gas

which increases the saturation temperature, as seen in (a). The sample gas pressure also has

a significant influence. With increasing pressure, the saturation temperature also increases.

A decrease of the air inlet temperature also causes a small increase of the saturation tem-

perature, as seen in (b). Due to the higher fuel demand at lower air temperature, the water

fraction in the exhaust gas increases. Changing the air velocity does not have an impact on

the saturation temperature since the exhaust gas composition remains constant, as shown

in (c).


(a) (b)

(c)

Figure 4.6: Saturation temperature Tsat for (a) different combustor outlet temperatures and

sets of pressures and fuels, (b) different air inlet temperatures and sets of pres-

sures and fuels, (c) different relative air inlet velocities and a set of fuels.


(a) (b)

Figure 4.7: Saturation temperature Tsat over the water fuel ratio for (a) sets of fuels and

air inlet temperatures, (b) sets of fuels, pressures and combustor outlet temper-

atures.

The water injection greatly influences the saturation temperature, since it significantly in-

creases the fraction of water in the exhaust gas, as seen in Fig. 4.7. The saturation temper-

ature increases by nearly 20◦C when the maximum water fuel ratio is reached. Additionally,

the highest saturation temperature in this study, 136◦C, appears at highest water injec-

tion rate for hydrogen fuel at an air inlet temperature of 400◦C, highest combustor outlet

temperature, and a pressure of 16 bar.

In summary, the pressure and the water injection rate have the highest influence on the

saturation temperature. At reference conditions the saturation temperature is 104◦C. The

global minimum temperature for this data set is at lowest pressure, occurs with natural gas

fuel and dry conditions, and is about 46◦C.

Evaluation The actual sample gas temperature must not fall below the above mentioned threshold val-

ues for each operating point of the combustor. Fig. 4.8 shows a comparison of the saturation

temperature and the actual sample gas temperatures of all data points. Each marker indi-

cates an operating point. It can be seen that the gas sampling temperature was above the

dew point for all operating points.


Figure 4.8: Verification of sample gas temperature exceeding the saturation temperature for

all operation points.

4.6.3 Pressure loss of combustor

The pressure loss ratio (PLR) is a key design issue that affects the efficiency of the gas

turbine. The lower threshold of the PLR is set by the ability to effectively mix the fuel and

air and the size limitations of the combustor. Typical values range from 4 to 8% (Lefebvre

& Ballal 2010).

DefinitionThe pressure loss ratio is given by

PLR =∆pcomb

p. (4.7)

Here, ∆pcomb is the static pressure difference between the input and output pressure of the

combustor and p is the static combustor inlet pressure.

ResultsThe PLR is primarily dependent on the combustor inlet air velocity and, to a lesser extent,

on the combustor outlet temperature due to its strong impact of the volume flow of the

combustor. The PLR of the combustor used in these experiments is displayed in Fig. 4.9 as

a function of the relative air inlet velocity. The combustor outlet temperature is indicated

by the color. The trend of the PLR curve shows an increase as the volume flow increases

similar to a standard nozzle. The shift to higher PLRs is also caused by the reduction of the

air inlet temperature to 400◦C, and to a lesser extent, a lower pressure level, both of which

affect the inlet volume flow and viscosity. The PLR at the reference point of this study,

defined in Tab. 4.3, is 4.81%. Thus the pressure loss ratio of the test combustor is within

the design criteria given in the literature.


Figure 4.9: Pressure loss ratio against the relative combustor velocity for a set of combustor

outlet temperatures.

4.6.4 NOx emissions

The characterization of the NOx emissions is the key aspect of this thesis and fundamental

for the verification of the ensuing numerical analysis. Based on the reference point defined

earlier in Tab. 4.3, the NOx emissions of the combustor are shown at various conditions.

In the following, results from varying the pressure and the combustor outlet temperature,

followed by the results from changing the fuel from natural gas to hydrogen are presented

and discussed. Next, the results from changing the air inlet temperature and the air velocity

are shown. Finally, the results from using water injection are presented for various pressures,

fuels, air inlet temperatures, and combustor outlet temperatures. Note that experimental

results may appear in more than one figure in order to show a complete comparison of all

varied parameters.

Pressure The change in the NOx emissions due to changing combustor pressure is given in Fig. 4.10 for

a set of combustor outlet temperatures. The diagrams (a) to (c) show this trend for natural

gas, 80 vol.% H2, and pure hydrogen fueling, respectively. Note that data at 24 bar and the

reference combustor outlet temperature has not been acquired, as well as at 3 bar for a 80%

mixture. The NOx emissions with all fuels and sets of combustor outlet temperatures show

a consistent trend with changing pressure. The test results indicate an increase in the NOx

emissions with increasing pressure whereas the gradient decreases with increasing pressure.

The outcome is a general root function that is typical for non-premixed flames and can be

verified theoretically according to Joos (2006). Furthermore, the influence of the pressure on

the NOx is higher with greater hydrogen concentration in the fuel. Thus, a pressure increase


(a) (b)

(c)

Figure 4.10: NOx emissions as a function of combustor pressure for a set of combustor outlet

temperatures and (a) natural gas fuel, (b) 80 vol.% H2 and (c) 100 vol.% H2.

with high hydrogen content fuels has a stronger impact on the NOx emissions compared to

a low hydrogen content fuel.



The change in NOx emissions with changing combustor outlet temperature is shown in

Fig. 4.11 for a set of pressures. The diagrams show the results for (a) natural gas, (b)

80 vol.% H2, (c) and pure hydrogen fuel. Thus, the same dataset as in Fig. 4.10 is displayed.

For all given conditions, the NOx emissions increase with the combustor outlet temperature.

Additionally, higher pressure and more hydrogen in the fuel result in a steeper gradient.

(a) (b)

(c)

Figure 4.11: NOx emissions as a function of air inlet temperatures for a set of fuel composi-

tions and pressures.


(a) (b)

(c) (d)

Figure 4.12: NOx emissions as a function of hydrogen content in the fuel for a set of com-

bustor outlet temperatures.

FuelThe change of the NOx emissions due to the fuel hydrogen fraction is shown in Fig. 4.12

for a set of combustor outlet temperatures and at (a) 3 bar, (b) 8 bar, (c) 16 bar, and

(d) 24 bar. In general, the NOx emissions increase with the hydrogen fraction of the fuel.

The NOx emissions increase less significantly in the region with low caloric hydrogen mixtures

< 80 vol.% H2 than in the region with high hydrogen admixtures. Consequently, the gradient

of the NOx emissions as a function of hydrogen admixtures increases with increasing hydrogen

fractions. This trend is seen for all pressures except for 3 bar due to a low number of data


(a) (b)

(c)

Figure 4.13: NOx emissions as a function of air inlet temperatures for a set of fuel composi-

tions and pressures.

points. This trend is furthermore valid for all combustor outlet temperatures. The increase

in the gradient of the NOx as a function of increasing hydrogen fraction is in general higher

at higher combustor outlet temperatures.

Airtemperature

The change in NOx emissions due to variation of air inlet temperature is displayed in

Fig. 4.13, with sets of different fuel compositions and pressures. The combustor outlet

temperatures are (a) 1300 ◦C, (b) 1100 ◦C, and (c) 900 ◦C, respectively. At the reference

pressure, the NOx emissions reduce nearly linearly with the air inlet temperature for all


(a) (b)

(c) (d)

Figure 4.14: NOx emissions as a function of relative air inlet velocity for a set of fuel com-

positions.

fuel mixtures. Because of this, at high NOx emission levels the potential for absolute NOx

reduction is higher. Despite the lower NOx levels, the pressure does not have an impact on

this trend. A decrease of the air inlet temperature typically leads to a reduction of the NOx

emissions since the stoichiometric flame temperature decreases simultaneously.

Air velocityFig. 4.14 shows the NOx emissions as a function of the air inlet velocity for various fuel

compositions. While diagram (a) is at the reference conditions, (b) represents a reduced air

inlet temperature, (c) a reduced combustor outlet temperature, and (d) a reduced pressure


(a) (b)

Figure 4.15: NOx emissions as a function of water injection for a set of fuel compositions,

air inlet temperatures and pressures.

and air inlet temperature. Variation of the air inlet velocity causes a change in the volume

flow of the combustor and can be regarded as a flame residence time variation. As velocity

increases, the residence time of the reactants within the hot flame temperature decreases

and the mixing process improves due to the significantly higher pressure loss ratio, compare

Fig. 4.9. Consequently, the NOx emissions decrease. An increasing rate of NOx formation

with decreasing velocity can be observed for all fuel compositions. Furthermore, the relative

increase in the NOx concentration is higher for elevated hydrogen fractions.

Waterinjection

Fig. 4.15 shows the effect of water injection as a NOx reduction measure for multiple fuel

compositions at different pressure levels, combustor outlet temperatures, and air inlet tem-

peratures. Note that the ordinate scale is logarithmic. Herein, (a) shows results at 16 bar

and a combustor outlet temperature of 1300◦C, while (b) shows results at an air inlet tem-

perature of 500◦C. Both diagrams refer to the reference air inlet velocity. Note that the

water fuel ratio ψ reaches higher values at increasing hydrogen fraction in the fuel since hy-

drogen features a higher specific LHV. The actual maximum injected water mass flow rate,

however, is the same for all curves of the same pressure. As a result, the NOx emissions

can be limited to significantly below 50 ppm by water injection for all parameter studies.

The relative potential of water injection is significantly higher for hydrogen than for natural

gas. Starting from a higher NOx level, the emissions for hydrogen fuel are reduced to similar

absolute values to those seen with natural gas. Reducing the air inlet temperature reduces

the initial dry NOx emission and thereby increases the (negative) gradient. Furthermore,

the potential of NOx reduction by water injection is lower at 3 bar compared to the refer-


ence pressure due to a strong increase of the CO emissions, as shown in the next section.

Consequently, lower absolute NOx emissions can be achieved with water injection at higher

pressures.

Closingremark

In general, the level of the NOx emissions exceeds today’s emissions regulations, see Sec. 2.1.

However, these levels are typical for non-premixed combustors. The data at dry conditions

were acquired without any NOx reduction strategies. The key driving phenomena, however,

can be investigated even with this elevated level of NOx emissions.

4.6.5 CO emissions

This section presents the trends for CO emissions from the experimental combustor, along

with a comparison to the NOx emissions. The CO analysis allows for an additional validation

of the possible operating range of the combustor since elevated CO emissions are indicative

of incomplete combustion. In the following section, the experimentally-determined CO emis-

sions are given for changes in the combustor outlet temperature, the pressure, the air inlet

temperature, the hydrogen content of the fuel, and amount of water injection.


(a) (b)

(c) (d)

Figure 4.16: CO and NOx emissions as a function of the combustor outlet temperature.

Combustoroutlet tem-perature,pressure

The CO emissions trend for changing combustor outlet temperatures is shown in Fig. 4.16

for a set of pressures from (a) 3 to (d) 24 bar. The CO emissions show a minimum at a

combustion temperature of about 1100 ◦C and slightly increase beyond this temperature

due to dissociation of CO2. At higher pressure, however, the suppression of dissociation

is stronger and thus the CO emissions increase by a lower amount. In general, the CO

emissions are lower at higher pressure. The compromise between NOx and CO emissions

can be observed clearly. At reduced combustor outlet temperature, the lower NOx emissions

come with a dramatic increase in the CO emissions.


(a) (b)

(c) (d)

Figure 4.17: CO and NOx emissions as a function of the air inlet temperature.

Air inlettemperature

The CO emissions as a function of the air inlet temperature are shown in Fig. 4.17. In general,

the CO emissions decrease with increasing air inlet temperature at lower combustor outlet

temperatures, while they remain nearly constant at higher combustor outlet temperatures.

Besides shifting the absolute level of emissions, the pressure does not change the overall trend.

Varying the inlet air temperature mainly influences the stoichiometric flame temperature,

which supports the conversion of CO to CO2 before dissociation initiates.


(a) (b)

(c) (d)

Figure 4.18: CO and NOx emissions as a function of the hydrogen fraction x within the fuel.

Hydrogencontent

The higher the hydrogen fraction of the fuel, the lower the carbon content and thus, the CO

emissions naturally decrease to zero with pure hydrogen fuel, as shown in Fig. 4.18. Here,

the CO emissions are given for a set of pressures from (a) 3 bar to (d) 24 bar. However,

increasing the hydrogen content in the fuel causes the flame temperature to increase and

favors residual CO generated by dissociation. The most detailed data for emissions as a

function of the fuel composition were acquired at the reference pressure of 16 bar in (c).

With increasing hydrogen content in the fuel, the CO emissions first decrease but reach a

minimum at 80 vol.% H2. Remarkably, the CO emissions increase between 80 and 90 vol.%


(a) (b)

Figure 4.19: CO and NOx emissions as a function of the water injection ψ at 16 bar.

H2 before the lack of carbon becomes the determining effect that limits CO. The other

pressure levels do not show this effect due to the lower resolution of the data.

Waterinjection

Fig. 4.19 shows the effect of water injection on the CO emissions for natural gas fuel. The

CO emissions dramatically increase at lower pressures due the significant reduction of the

stoichiometric flame temperature. Due to the low combustor outlet temperature, the CO

burnout is suppressed until the combustor exit. At the reference combustor outlet temper-

ature, however, the CO emissions slightly decline, indicating a sufficient CO burnout before

the combustor exit. Because of the decrease of NOx with the increasing water fuel ratio, an

operating point that minimizes both NOx and CO emissions can be achieved in the case of

high combustor outlet temperatures.

Closingremark

Note that for the non-premixed flames a different trend is measured than usually seen for

premixed flames. In the case of premixed flames, the flame temperature is generally signifi-

cantly lower and highly dependent on the equivalence ratio. For the non-premixed combustor

used in these experiments, the flame is always stoichiometric and thus the flame temperature

remains constant for varying combustor outlet temperatures.

4.7 Summary

A high pressure combustion test rig was used to characterize an industrial non-premixed

can-type gas turbine combustor regarding the feasibility of high hydrogen fueling and water

injection as a NOx abatement measure. In order to evaluate full and part load conditions,


other studied variables are the pressure, the combustor outlet temperature, the air inlet

temperature, and the air inlet velocity.

The comparison of the directly measured air fuel ratio and the same value determined solely

on the basis of emissions measurements shows a good agreement between the two and indi-

cates a congruent measurement system. The sample gas temperature is furthermore proven

to be appropriate for NOx emission measurements. The pressure loss ratio of the experi-

mental conditions meets the common value of today’s gas turbine designs.

The NOx trends shows congruence with the known phenomena of non-premixed flames.

With increasing pressure, the emissions increase by a general root function. This effect is

based on the higher reactivity caused by the collision frequency of fuel and oxidizer. The

emissions furthermore increase with increasing combustor outlet temperatures. Increased

hydrogen content significantly increases the flame temperature due to its higher reactivity

and consequently leads to a significant increase in NOx emissions. The air temperature has a

direct effect on the flame temperature, that alone affects the NOx emissions. The air velocity

mainly affects the residence time of the reactants in the hot flame region and thus leads to

a reduction of NOx production. Finally, water injection reduces the flame temperature and

thereby significantly suppresses the NOx generation. For hydrogen flames, the potential of

NOx emissions reduction is significantly higher than for conventional natural gas flames.

The physical and chemical phenomena underlying this result are investigated in detail in the

next chapter.

The CO effects limit the operating range of the combustor. Natural gas operation at low

combustor outlet temperatures between 900 and 1000◦C leads to a significant increase in

CO emissions. Particularly, a further reduction of the air inlet temperature and water

injection leads to an increase in the CO emissions. The effect of increasing CO production

from hydrogen mixtures due to the increase of the flame temperature does not exceed the

initial CO emissions from natural gas fueling. Dissociation effects cause an increase of the

CO emissions at highest flame temperatures. In conclusion, since the CO emissions are

generally in an acceptable range, the combustor is technically verified to be sufficient for

general natural gas combustion and for the purposes of this thesis.

This data acquisition is the basis for the following theoretical investigations. In the upcoming

chapter, the phenomena inside the combustor are analyzed by a numerical investigation.

With the help of these results, more precise explanations of the underlying phenomena are

given. Chapter 6 then bridges the gap between the two methods and summarizes all results

in a simplified model approach.

5 Chemical reactor network model

In this chapter, a chemical reactor network is described, which was used to model the non-

premixed flame of the experimental gas turbine combustor. Based on the combustor design

and experimental results of the last chapter, a chemical reactor network (CRN) was designed

and validated. The model allows for an inside view into the combustor by computing the

stoichiometric flame temperature and reproducing the residence time in the flame, that are

mayor influencing parameters for the NOx formation. Subsequently, the chemical influence

of hydrogen and water injection on the combustion process is studied. The detailed chem-

ical analysis furthermore investigates the influence of global operating parameters, such as

the effects of pressure, combustor outlet temperature, air inlet velocity, and the air inlet

temperature on the internal NOx formation process.

OutlineThis chapter is structured as follows: First, the existing models for non-premixed combustors

from relevant literature are reviewed. After introducing the fundamental components of a

reactor network including the implementation of the chemistry, the analysis and validation

methods of this study are described. The third section presents the underlying network

model of the combustor used in this study. Then, the response of internal model parameters

to external operating parameter variations chosen based on experiments are discussed. The

fifth section contains the detailed validation of the model’s results with the experimental

data on the basis of the NOx emissions including the determination of the most suitable

reaction scheme. Next, the influence of hydrogen and water injection on the chemical reaction

progress in general and in particular on the NOx formation pathways is investigated. Finally,

the distribution of the thermal, chemical, and dilution effect of water injection is identified.

5.1 Reactor model review

Methodsreview

Combustion modeling reproduces the real reaction processes by a mathematical model so

that its numerical or analytical solution represents the physical processes in the combustor.

Within the wide field of combustion simulation, the simplest approach is a one-step reaction,

which transforms fuel and oxygen directly and completely into products. However, this

approach cannot deal with intermediate reaction species. The next step of sophistication is

the chemical equilibrium model that is capable of handling intermediate species, radicals and

dissociation effects. However, it does not take time-dependent effects into account. Since

NOx formation is temporally driven, it can be addressed by chemical reactor networks, that

provide a 0D or 1D representation of the real combustor. The chemical reaction progress

60 5 Chemical reactor network model

and the residence time result from the reactor volume, mass flow and density. However, in

a CRN the focus is on chemistry. Flow field and boundary conditions are approximated by

general definitions. The final level of sophistication are numerical combustion models that

discretize the real geometry of the combustor and combine 2D or 3D flows with the reaction

scheme as an additional degree of freedom. The efforts for such studies are high since the

transport equation for each species has to be solved and this method is currently limited to

a few reactions and species.

CRN forgas turbinecombustion

Chemical reactor networks (CRN) are commonly used for modeling of combustion processes

since the middle of the last century. Over the years, they have been developed into a strong

tool for detailed combustion analysis. Thereby, the most important reactors are the perfectly

stirred reactor (PSR) with zero dimensional homogeneous conditions and the plug flow reactor

(PFR). The latter is one dimensional. Its composition changes in the axial direction while

it is homogeneous in the radial direction. These ideal reactors are connected to networks

that represent characteristic areas of a complex combustor. The application of PSR and

PFR for real combustion environments, especially for gas turbine combustion, has to be

addressed carefully since perfectly homogenized conditions are not realistic. While a degree

of inhomogeneity is usually defined for premixed flames, the assumption of a homogeneous

distribution of fuel and air appears contradictory for non-premixed flames.

Non-premixedapplication

In general, non-premixed flame combustors are characterized by the assumption of a sig-

nificantly faster chemical reaction than flow mixing (i.e. mixed = burned). The turbulent

mixing time is the rate determining phenomenon and the chemical reaction is regarded to be

infinitely fast compared to the turbulent mixing time scale. Although the ideal reactor fea-

tures a homogeneous distribution in the reactor and thus no mixing phenomena are resolved,

it is still applicable for non-premixed flames but requires special attention. In non-premixed

flames, NOx generation mainly takes place in the flame zone (see Sec. 2.1). At the highest

flame temperature, a homogeneous distribution can be assumed. With these assumptions,

the reactor volume represents the flame volume and the definition of this volume is directly

connected to the residence time within the flame.

Residencetime

In general, for non-equilibrium chemical problems, the residence time of the reactive fluid

within the combustor is one of the most important parameters. The residence time can be

investigated by different approaches. First, it can be determined experimentally. Here, the

combustor can be discretized in space and time on the basis of a flow field analysis inside the

real combustor geometry. Thereby, water flow tests (Guethe et al. 2009) and measurements

of the OH radical distribution can be used. Second, the residence time can be expressed

by an analytical approach, for example, the flame length approximation of Feitelberg et al.

(2000). The advantage of this technique is a formulation which is potentially independent

from the geometry allowing for an efficient simulation. This procedure is limited to simple


chemical reactor models that are capable of evaluating the emission trends for gas turbine

combustor applications including the effects of parameters of interest using detailed chemical

kinetic mechanisms. Third, the residence time can be determined by a reactive flow CFD

simulation with strongly reduced reaction mechanism, where the focus is on the flow field

and the heat release. The resulting flow pattern may be sectioned into appropriate iso-

temperature regions, e.g. the flame zone, an inner/outer recirculation zone, and a postflame

zone, that each can be represented by a reactor model. The residence time is determined by

the fluid velocity and the volume of the regions. Drawbacks of this method, however, are

its complexity and need to specify the combustion geometry and considered operating point

(Hackney et al. 2016). In the following, published models that pay special attention to the

CRN set up and the residence time formulation and that are used with non-premixed flames

and NOx emissions prediction are reviewed.

TouchtonTouchton (1984) presented a reactor network model for non-premixed flames consisting of

three consecutive PSRs representing the flame zone, the post flame zone, and the dilution

zone for methane fuel. The first reactor solves the global oxidation equations for methane,

assuming a residence time that guarantees the equilibrium of the oxygen dissociation at

the exit. Stoichiometric flame conditions are not defined but the flame equivalence ratio is

kept as a variable parameter. Furthermore, steam injection is applied as a NOx abatement

measure. The hot exhaust is fed into the second reactor, where NOx is generated on the basis

of the Zeldovich mechanism. By fitting with experimental data, Touchton found suitable

residence times on the order of a few milliseconds that are directly proportional to the NOx

emissions. Finally, in the third reactor, the reaction progress is frozen by excess air injection

through the sudden drop in temperature. Doing so, he achieved a closed NOx prediction

equation including gas turbine part load conditions and steam injection as NOx reduction

measure, where mixing effects are empirically accounted for.

AndreiniandFacchini

Andreini & Facchini (2004) presented a reactor network model for a diffusion flame. The

primary zone is modeled by a flame reactor model that was developed by Broadwell & Lutz

(1998) for free turbulent jet flames containing two PSRs to simulate the central region (core)

and the flame sheet zone where stoichiometric conditions are predominant. The secondary

zone is simulated by a center PSR and an outer PSR and represent the admixture of a portion

of excess air through first cooling holes. At that point, the flame looses its diffusive character

entirely and the mixture faces a homogenization because of air cooling. The dilution zone is

modeled by a PFR and a wall PSR for analogous reasons. The authors do not comment on

detailed definitions of reactor volume or the resulting residence times; however, the model,

embedded in a gas turbine power plant simulation code, could reproduce measured engine

NOx emissions satisfactorily.


Feitelberg Feitelberg et al. (2000, 2001) developed a semi-empirical model to estimate the residence

time on the basis of measured NOx emissions. The model assumes a confined turbulent

diffusion flame length that is proportional to the ratio of the fuel and stoichiometric air flow.

Here, the residence time is assumed to be proportional to the flame length. A PSR reactor

at stoichiometric conditions is used to determine the flame residence time corresponding

to measured NOx concentration. Applied to a standard and an improved combustor, the

differences in NOx emissions as well as the effect of a combustor outlet temperature variation

on the NOx emissions are compared using turbulent flame length arguments.

Iyler Iyer et al. (2005) presented an approach for predicting the NOx emissions of a non-premixed

combustor for syngas applications. The combustion process is represented by a network of

three consecutive PSR reactors with individual analytical residence time formulations. The

first reactor is operated stoichiometrically and its residence time represents the mixing time

scales of fuel, air, and diluent. Its residence time is obtained based on the mixing length

model for turbulent antisymmetric jets. In the second reactor, the excess air is added and its

residence time is calculated by estimating the volume of the jet between the stoichiometric

contour and the contour of full entrainment, divided by the total volume flow rate. The

third reactor represents the dilution zone and its residence time is chosen to match the

overall combustor residence time. Iyer et al. analyzed the influence of the combustor exit

temperature, the syngas fuel composition and diluents (N2 and steam) on the NOx emissions

by a model response analysis without a comparison to measured data.

Summary Summarizing, chemical reactor network modeling has been proven to be a valuable tool for

the investigation of NOx formation in the non-premixed combustion processes even though

the assumptions of the ideal reactor formulations more accurately represent the conditions

of premixed flames. Here, the residence time definition in the flame zone is crucial for an

accurate simulation of NOx emissions. Different approaches for predicting the residence

time have been investigated, including analytical expressions, flame length correlation, and

determination based on experimental data.

5.2 Methods

This section introduces the methods used for combustion modeling and results analysis.

First, the components of a chemical reactor network are explained. Then, the implementation

of the chemistry is presented by the introduction of reaction schemes and the mechanisms of

NOx generation. Finally, methods of validating the numerical results are presented, including

the chemical investigation via the reaction kinetics analysis.


5.2.1 Ideal reactors and reactor networks

Ideal reactors provide a simple frame to apply detailed chemical kinetics for numerical com-

putations. In general, perfectly stirred reactors and plug flow reactors are the two main

types of ideal reactors used in this thesis (see Fig. 5.1). Networks are formed by combining

these basic reactors.

PSRThe perfectly stirred reactor is the simplest reactor that takes reaction kinetics into account.

Within PSRs, there are two main types: perfectly stirred batch reactors (BR) (Fig. 5.1 (a)),

and perfectly stirred flow reactors (PSR) (Fig. 5.1 (b)). Batch reactors represent a homoge-

neous non-flow reaction in a closed volume and with a temporally varying fluid composition.

Such reactors are thus not practical for steady-state combustor simulation. In contrast, per-

fectly stirred (flow) reactors assume a uniform mixing of the components upon entering the

ideal reactor. The exiting species composition is the same as of the initial fluid. It is calcu-

lated using the reaction mechanism and the resulting residence time inside the reactor. This

zero-dimensional reactor is completely described by the following two differential equations.

The mass conversion equation of a species i for a reactor with the volume V is

dmi

dt= MiV ωi +

∑(mYi)in − (mYi)out, (5.1)

where Mi is the molar mass and Yi is the mass fraction of a species i. While ωi denotes

the net production rate of a species i, m is an inflowing or outflowing mass flow rate. The

energy balance for an ideal gas reactor is

dU

dt= q +

∑(mh)in − (mh)out, (5.2)

where h is the specific enthalpy of an inflowing or outflowing mass flow rate m and q is an

energetic heat source. Note that the specific work is neglected. The residence time of the

reacting fluid within the reactor is determined from the reactor volume V . Since the reactor

(a) Perfectly stirred batch reactor (BR)

Feed Feed

Product Product

(b) Perfectly stirred flow reactor (PSR)

(c) Plug flow reactor (PFR)

Uniformly mixed

Figure 5.1: Types of ideal reactors.


contains the mass distribution of the outlet flow, the residence time of the fluid inside the

reactor is given by

τ =V · %m

, (5.3)

where m is the outflowing mass flow rate and % is the exit density. Note that the PSR is

based on stationary inlet and outlet flows.

PFR The plug flow reactor, see Fig. 5.1 (c), is a one-dimensional reactor that is used to describe

chemical reactions in continuous stationary flows. Composition and properties of the fluid

change within the flow stream. Perpendicular to the flow direction, a homogeneous mixture

is assumed. The residence time of the fluid is not only dependent on the reactor volume and

the mass flow rate, but also on the change in density % within the flow, which depends on the

reaction rate in addition to changing temperature. Note that the PFR can be approximated

by a consecutive series of PSRs.

Networks Ideal reactors can be combined into a chemical reactor network (CRN) in order to approxi-

mate the flow within a real combustor geometry. The more information that can be gathered

about the inner flow structure and temperature distribution in the real reactor, the higher

the level of sophistication can be for a specific application. In practice, flame regions, recir-

culation regions, cooling regions, dilution zones, and post flame zones can be distinguished

and represented by ideal reactors. For further information about ideal reactors and reactor

networks, the reader is referred to Levenspiel (1999). Each reactor of a CRN is character-

ized by thermodynamic and chemical properties of all containing and evolving species. This

information is provided by the reaction mechanism, that are introduced in the following

section.

5.2.2 Reaction kinetic and mechanisms

This section introduces the reaction kinetics that give the net production rates ωi required in

Eq. 5.1. The explanation is based on the Cantera Manual (Goodwin et al. 2016). Thereafter,

the underlying reaction mechanisms are introduced.

Reactionkinetic

The chemically-induced species conversion in chemical reactions is quantified by the reaction

rate ωi that describes rate of formation of species i per unit space and time. It is the sum

of all individual rates ωik = νikQk of each reaction k in the reaction scheme, according to

ωi =∑k

νikQk. (5.4)

Thus, νik = ν ′ik − ν′′ik are the net stoichiometric coefficients of a species i and a reversible

reaction k, where ν ′ is the coefficient of the forward reaction and ν ′′ is the coefficient of the


reverse reaction. The progress rate Qk of the reaction k is described by

Qk =

(kfk

∏i

[Xk]ν′ik − krk

∏i

[Xk]ν′′ik

)(∑i

εik [Mk]

), (5.5)

where [Xk] is the concentration of the reacting species, and kfk and krk are the forward

and reverse reaction rate coefficient, respectively. The inert collision partner Mk in third-

body reactions is an unspecified species at the concentration [Mk] that carries away excess

energy for stabilization or supplies activation energy to break a chemical bond. This effect

is quantified by the collision efficiency εik which is also called the chaperon efficiency factor.

The standard value is εik = 1, representing the efficiency of nitrogen. Deviations from this

values are defined for each species.

Tempera-turedependence

The quantitative dependence of the reaction rate coefficient kfk on the temperature can be

described by the Arrhenius equation

kfk = Ak · T βk · exp

(−EkRT

), (5.6)

where Ak is a pre-exponential factor that represents the overall collision frequency, βk is the

temperature exponent, and Ek is the activation energy for the reaction k. These parameters

are defined by the reaction mechanism. The reverse reaction rate coefficient krk can be

derived with

krk =kfk

Kck, (5.7)

where Kck is the equilibrium constant that can be determined from thermodynamic data.

Pressuredependence

The reaction rate coefficient kf and kr of some third-body reactions are highly pressure

dependent. For these reactions, two sets of reaction rate coefficients are provided, one for

the low pressure limit and for high pressure limit, k∞. The region between these pressures are

characterized by functions defined by empirical parameters, such as the Lindemann approach

or the Troe form (see e.g. Warnatz et al. (2001)).

Reactionmechanisms

Chemical kinetics are provided by reaction mechanisms that contain required thermody-

namic information about the relevant element and/or species i and a set of elementary

reaction equations in order to calculate the net production rate ωi of species i (cf. Eq. 5.1).

Reaction schemes describe the complex chemical processes of species conversion and exist

for various applications where chemical reactions are involved. For combustion modeling,

several reaction schemes are available. For this study, the reaction scheme must support

methane and hydrogen combustion and nitrogen chemistry including NOx formation. Four

reaction mechanisms that fulfill the requirement are listed in Tab. 5.1.


Table 5.1: Reaction mechanisms for numerical study.

NameCondition Ranges

Fuel speciesp [bar] T [◦K] Φ [−]

1GRI Mech 3.0 0.013 - 10.13∗ 1000 - 2500∗ 0.1 - 5∗ H2, CH4, C2H6, CO

2Aramco Mech† 1.0 - 60.0∗∗ 600 - 1400∗∗ 0.05 - 5.0∗∗ CH4, C2H6, H2, C3H8,

C4H10, CO

3UC San Diego CH4, C2H6, H2, C3H8,

C4H10, CO

4Konnov CH4, C2H6, H2, C3H8, CO

†Aramco Mech 1.3 updated by Bohon et al. (2017) incl. detailed NOx chemistry of

GDF-Kin 3.0 NCN by El-Bakali et al. (2006).

Sources: 1Smith et al. (2000), 2Metcalfe et al. (2013), 3San Diego (2016), 4Konnov (2000).

Validation: ∗Henry J. Curran (2004), ∗∗Data validated in Flow Reactor by Metcalfe et al. (2013)

GRI Mech3.0

The GRI Mech 3.0 (Smith et al. 2000) mechanism has been developed by University of

California at Berkeley, by Stanford University, by University of Texas at Austin, and by

SRI International. It has been especially designed to model methane and ethane flames at

high temperatures. Additionally, hydrogen chemistry and NOx formation is included. It

includes information on the five elements O, H, C, N, and Ar and includes 53 species and

325 reactions. The nitrogen chemistry alone contains 124 reactions. It is popular in gas

turbine combustion research due to the balance between computational time and accuracy.

It is not recommended for use in excess of 10 bar or and 2500 K (Henry J. Curran 2004).

AramcoMech

The Aramco Mech (Metcalfe et al. 2013) mechanism has been developed by the Combustion

Chemistry Centre in NUI Galway, Ireland. It uses a detailed chemical kinetic mechanism

that characterizes the kinetic and thermochemical properties of a large number of C1 to C4

based hydrocarbon and oxygenated fuels. It includes the six elements C, H, N, O, Ar, and He

and features 346 species and 1542 reactions. The original version of the Aramco Mech table

does not include nitrogen chemistry. However, Bohon et al. (2017) recently updated this

mechanism and El-Bakali et al. (2006) included the detailed NOx chemistry of the GDF-Kin

3.0 NCN mechanism. The Aramco Mech has been validated up to 60 bar and 1400 K in a

flow reactor and up to 260 bar and 2500 K in a shock tube (Metcalfe et al. 2013).

UC SanDiego Mech

The UC San Diego mechanism has been developed by San Diego (2016), and includes the

six elements N, Ar, He, H, O, and C, 70 species, and 321 equations. Combustion of carbon-

based fuels from methane up to butane and hydrogen combustion can be considered. The


university also provides a nitrogen chemistry sub-mechanism that is added to the complete

mechanism. The application rage of this mechanism is not provided.

KonnovMech

The Konnov mechanism has been developed by Konnov (2000) for the combustion of small

hydrocarbons. It contains the five elements H, C, O, N, and Ar, and features 129 species,

and 1231 equations. A nitrogen chemistry mechanism is included. However, the application

range is not given.

In the remainder of this study, these reaction schemes are referred to as GRI 3.0, Aramco,

San Diego, and Konnov, respectively.

5.2.3 NOx formation mechanisms

The nitrogen chemistry within reaction mechanisms is highly complex and results in a net-

work of mutually dependent pathways for NO formation. However, NOx pathways always

initiate by cracking the triple bond of molecular nitrogen. In gas turbine flames, NOx for-

mation takes place mainly via four different pathways, which are introduced in this section.

Here, the reaction numbering refers to the GRI 3.0 mechanism.

ThermalNO

Thermal NO or Zeldovich NO, first described by Zeldovich (1946), is dominant at high

temperatures above about 1000◦C. This is typically the strongest pathway for NO formation

in undiluted stoichiometric flames. Thermal NO generation is not an equilibrium process

and is highly dependent on the residence time in the hot temperature region. The original

thermal pathway has been expanded and now comprises the following three reactions:

N2 + O◦ NO + N◦ (GRI -178)

N◦ + O2 NO + O◦ (GRI 179)

N◦ + OH◦ NO + H◦ (GRI 180)

The first reaction is the dissolution of atmospheric nitrogen into the final product NO and

a nitrogen radical. This reaction is dependent upon the oxygen radical concentration. The

atomic nitrogen is then further oxidized by molecular oxygen O2 or the hydroxyl radical

OH◦ in the second and third reactions.

Prompt NOPrompt NO or Fenimore NO was firstly postulated by Fenimore (1979). He showed that N2

can be broken up by the CH◦ radical formed during the combustion carbon containing fuels.

The rate limiting chemical reaction is

CH◦ + N2 HCN + N◦ (GRI 240)

where HCN is formed that can then oxidate to NO. The activation energy is lower than

the reaction GRI -178 of the thermal NO pathway and thus it is most relevant in the low

temperature regimes in either rich or lean flames. Obviously, the presence of carbon in the


fuel is precondition for this reaction scheme. Furthermore, prompt NO is sensitive to the

equivalence ratio since the CH◦ radical is the reactant.

NNHmechanism

The NNH mechanism was postulated by Konnov et al. (2000), by which nitrogen is broken

up by an atomic hydrogen radical forming a NNH◦ radical.

H◦ + N2 (+M) NNH◦ (+M) (GRI 204/-205)

NNH◦ + O◦ NH◦ + NO (GRI 208)

Due to the required H◦ radical, this mechanism is promoted in high-hydrogen flames.

N2Omechanism

The N2O mechanism also breaks up the nitrogen with an O◦ radical similar to the thermal

mechanism, but forms a single molecule N2O in a three-body reaction (Malte & Pratt 1974).

This further reacts with another O◦ radical to form NO.

N2 + O◦ + M N2O + M (GRI 185)

N2O + O◦ 2 NO (GRI 182)

This pathway does not require a high activation energy and thus is particularly significant

at low temperatures. Since free O◦ radicals are required, it is an important formation path

in lean conditions. The initial three-body reaction is promoted at high pressures.

Otherpathways

There are more possible pathways to generate NOx emissions, e.g. via the NH◦ radical or

the NCO◦ radical. However, these alternative pathways do not play a significant role in this

study, as will be shown later, and are thus not discussed in detail here. Finally, only fuels

without nitrogen are considered in this study, thus a pathway accounting for NO produced

by oxidation of fuel-bound nitrogen is not required.

5.2.4 Validation and analysis methods

This section presents the methods for model validation and analysis of reaction kinetics. The

numerical results are validated with the experimental outcome. The deviation of the model

results from the experimental results are therefore quantified by an appropriate compari-

son criterion. Furthermore, the underlying reaction mechanisms are complex and consist of

multiple species and several hundred reaction equations. In order to identify the relevant re-

actions and species for NOx analysis, a sensitivity analysis of the NO formation is performed.

Finally, the NOx pathway analysis is introduced as an approach to outline the complex NO

formation process.


5.2.4.1 Comparison criterion

The comparison of the numerical results with the experimental results is done using a root

mean square error definition. The validation criteria is the relative root mean square error

(rRMSE). The relative root mean square error for n data points is defined as

rRMSE =

√√√√ n∑i=1

(∆yi

ymax − ymin

)2

/n (5.8)

where yi is a placeholder for the variable compared at data point i, and ymax and ymin are

the maximum and minimum values of the variable in the data set, respectively.

5.2.4.2 Sensitivity analysis

Sensitivity analysis is a method that is used to quantify how much the output of a model is

affected by changes of different input parameters. In chemical kinetics, the input parameters

are the chemical properties of the reaction, comprised by the progress rate Qk for a reaction

k. The output is the concentration ci of the species i. Using this approach, the most

relevant rate-limiting reactions are identified by a high sensitivity coefficient. The first-order

sensitivity coefficients are defined as

σi,k =∂ci

∂Qk(5.9)

where ci is the concentration of the species i and Qk is progress rate of the reaction k. The

relative form of this definition can be simplified according to

σreli,k =

Qkci

∂ci

∂Qk≈ Qk

ci

∆ci

∆Qk. (5.10)

The computation of one specific operating point has to be done N + 1 times, where N is the

number of reactions in the reaction scheme. The factor f is used to successively manipulate

the Qk of reaction k, while keeping all other reactions unchanged. A factor variation study

showed that f = 1.05 is an appropriate value that itself does not influenced the result of

the sensitivity study. Finally, reactions featuring a sensitivity coefficient σreli,k > 3% are

considered as important for the NO formation. More general information about this method

can be found in Warnatz et al. (2001).


5.2.4.3 Pathway analysis

Chemical reactions connect reactants with intermediate species (e.g. radicals) and products

and thereby form a complex network of chains. The pathway analysis is applied in this

study to identify the formation and consumption pathways of NO. The major NO formation

pathways are the thermal, prompt, NNH and N2O mechanism (as described in Sec. 5.2.3).

They can be distinguished by the rate of their initial reaction. Since these initial reactions

feature the breaking up of the triple-bonded N2, they require the highest activation energies

and are thus the rate-limiting reactions.

Review In literature, the contribution of individual NOx formation pathways have been investigated

by consecutively disabling the initial reactions, cf. Bhargava et al. (2000), Guethe et al.

(2009), Fackler et al. (2011), Monaghan et al. (2012), Goke (2012) and Gockeler (2015). The

difference in the resulting NOx concentrations from the original mechanism was attributed

to the corresponding pathway. However, the result of this popular method is biased since

the manipulation of the mechanism influences the balance of the thermal energy that is then

free for other reactions.

Implemen-tation

Despite the experience in literature, the influence of deactivating these reactions was not

found to be negligible for the relevant network and operating conditions. Instead, in the cur-

rent study, all N2-breaking reactions are simultaneously computed without reaction scheme

manipulation. The allocation of broken-up nitrogen molecules to one of the four forma-

tion mechanisms is done on the basis of the net consumption rate of N2 of the N2-breaking

reactions. The allocation to the thermal NO is done via the nitrogen break-up reaction

(GRI -178). The contribution to the NNH mechanism and the N2O mechanism is done if

NNH◦ and N2O are products of the respective break-up reactions. For the contribution of

prompt NO, a carbon atom has to be involved in the break-up of the nitrogen and neither

NNH◦ nor N2O are produced. Finally, all other nitrogen N2-breaking reactions are summa-

rized and attributed to the expression ‘else’, compare e.g. GRI 196 and GRI 198, where NO

is produced via NH◦. A systematic error due to the manipulation of the reaction scheme is

excluded by this implementation.

5.3 Chemical reactor network model

A chemical reactor network of ideal reactors was used to model the combustion process and

NOx emissions of an industrial non-premixed flame combustor with a focus on the transition

to pure hydrogen fueling and water injection for NOx reduction. The influence of other

parameters like the pressure, combustor outlet temperature, air inlet temperature, and air

inlet velocity are also investigated. The parameter variation of this model was done according


Table 5.2: Input parameters and model results

Input Output

Variable Name Variable Name

p Pressure NOx NOx emissions

T Temperature cO2, cCO2

Exhaust composition

x Fuel composition Tst Stoichiometric flame temperature

vrel Relative air speed mair,st Stoichiometric air mass flow

Tair Air temperature mfuel Fuel mass flow

ψ Water fuel ratio τ Flame residence time

VPSR,init Reference flame volume τtotal Total residence time

to the experimental procedure. The underlying model focuses on simplicity for the purpose

of trend analysis.

Modeldesign

The model is designed to simulate the NOx emissions of an industrial non-premixed com-

bustor based on typical conditions in a real gas turbine. The model results are obtained

for various combustor operating conditions that are adjusted by changing the model input

parameters given in Tab. 5.2 on the left. These parameters are selected based on the ex-

perimental test conditions and thus represent the actual combustion conditions in real gas

turbine operation. The most important simulation results of the model are given in Tab. 5.2

on the right. By design, the model simulates aspects of the combustor that are challenging to

measure within the constraints of industrial combustor design and real operating conditions.

Model setup

The model abstracts the complex flow and reaction processes with a simple arrangement of

ideal reactors. The layout of the network is shown in Fig. 5.2, with the generic regions labeled

as the flame zone, the excess air admixture zone, and the burnout zone of the combustor. The

network includes two PSRs representing the flame zone and the mixing zone, respectively,

and a PFR representing the burnout zone. The generic model uses a semi-empirical approach

that is tuned at a single reference point to the residence time in the flame zone. This model

has been set up in the open-source chemistry software Cantera (Goodwin et al. 2016).

In the following sections, these zones are introduced with focus on the abstraction principles

and the specific implementation of each zone. Special attention is given to the residence

time calculation in the reacting zones.


Fuel

AirΦ=1

PFR

PSR 1(Flame zone)

Water (g)

PSR 2

Heat Loss

(Mixing zone) (Burnout zone)

Figure 5.2: Generic chemical reactor network for the non-premixed combustor.

5.3.1 Flame zone

A single PSR represents the flame zone of the combustor (PSR 1). It features a flexible

volume VPSR1 and has three inlet mass flows for air, fuel, and water and a single outlet mass

flow. Furthermore, a heat release enables energy balance during water injection.

Massbalance

Fuel of a composition x and a temperature Tfuel is fed into the reactor. The fuel mass flow

rate, mfuel, is determined by the combustor outlet temperature, T , the air inlet speed, vrel,

and the air density, %air. The assumption that the reaction takes place after mixing and at

ideal stoichiometric conditions (Φ = 1) leads to a stoichiometric air mass flow mair that is

fed into the flame zone (PSR 1). A water inlet mass flow, mw, allows for water injection.

The fluid exits the reactor with the stoichiometric flame temperature Tst.

Heat release The PSR 1 includes a heat loss mechanism to satisfy energy conservation during water

injection. In the experiments, liquid water is injected in to the combustor. However, the

simulated PSR 1 is a homogeneous single phase reactor and the water needs to be fed in

as gas. A physically analogous model uses the injection of water vapor at the evaporation

temperature Tevp and the actual combustor pressure p. In order to fulfill energy conservation,

a heat stream Hw is discharged from the reactor. It represents the enthalpy h∆T required for

heating the water from the water inlet temperature Tw to the evaporation temperature Tevp

and the evaporation enthalpy ∆hevp that is necessary for the phase change at the evaporation

temperature. This discharged heat stream can be described by

Hw = mw (h∆T + ∆hevp) . (5.11)


The enthalpy ∆hevp is given by

∆hevp = h′′w(Tevp)− h′w(Tevp), (5.12)

where h′′ and h′ are the enthalpy of the gas and liquid phases, respectively. The enthalpy

h∆T is required for heating the liquid phase of the injected water and is given by

h∆T =

∫ Tevp

Tw

cp(T )dT, (5.13)

where cp(T ) is the heat capacity of the water. The integral is simplified by using the arith-

metic mean heat capacity of both temperatures. The further heating of the water from Tevp

to the stoichiometric temperature Tst is inherently captured by the conditions of the ideal

reactor. Required thermodynamic data are taken from IAPWS (2007).

5.3.1.1 General residence time formulation

Referenceresidencetime

As already mentioned in Sec. 5.1, the PSR 1 represents the stoichiometric heat release zone

of the non-premixed flame. At the reference point (see Tab. 4.3), the PSR 1 volume VPSR1 is

tuned to match the experimental NOx emissions at the combustor exit. The results of this

parametrization (VPSR1,init) are shown in Tab. 5.3 for all reaction mechanisms. The flame

residence time τ of the PSR is then given by

τ =VPSR1 · %

mair,st + mf + mw(5.14)

according to the general description in Eq. 5.3. Note that the residence times determined

by the particular reaction schemes differ from each other significantly. This is due to the

above mentioned determination of VPSR1, that underlies the experimental NOx emissions

as a target value. The total residence time of the fluid in the reactor network and the

resulting stoichiometric flame temperature are also given in Tab. 5.3 for the reference point.

A significantly higher temperature was found compared to the data from literature shown

in Tab. 2.1 due to the elevated air inlet temperature.

Non-referenceresidencetime

In operating conditions other than the reference point, the residence time is modeled via a

flame length approach. The residence time is assumed here to be reciprocal to the flame

length. A function based on the length of a swirled non-premixed flame is thus used to

describe the variation of the flame length as the input parameters change. The different

formulations for varying combustor outlet temperature and pressure are described in the

following. For all other parameter variations, the residence time is modeled using Eq. 5.14.


Table 5.3: Flame volume, flame and total residence time, and stoichiometric temperature at

reference conditions*

VPSR,init τ τtotal Tst

10−4 [m3] [ms] [ms] [K]

GRI 3.0 4.31 0.381 8.66 2464

Aramco 6.03 0.540 8.71 2462

Konnov 4.00 0.273 8.62 2443

San Diego 3.03 0.357 8.66 2468

*compare Tab. 4.3

5.3.1.2 Residence time for varying combustor outlet temperature

Flamelengthapproach

The combustor outlet temperature variation is essentially a variation of the equivalence

ratio. At combustor outlet temperatures other than the reference point, the flame length is

adjusted by an empirical approach described by Chen & Driscoll (1988), who showed that

the length of a swirler-stabilized flame with a swirl number higher than 0.4 is proportional

to the fraction mfuel/mair,pr. The differentiation of primary mair,pr and stoichiometric mair,st

air inlet mass flow is necessary to guarantee stoichiometric conditions in the PSR 1 for all

operating conditions. In this model, the primary air mass flow mair,pr is defined as the

stoichiometric air mass flow at the reference combustor outlet temperature. Following the

approach of Feitelberg et al. (2000, 2001), the turbulent diffusion flame length L is related

to the fuel flow rate mfuel and the primary air flow rate mair,pr according to

L ∝ mfuel

mair,pr∝ τ. (5.15)

Here, the primary air mass flow mair,pr is the air mass flow that is required to completely

consume the total fuel mass flow at the reference combustor outlet temperature. The initially

determined reactor volume VPSR,init is used not only to determine the residence time at

reference conditions but also to determine the residence times of all other conditions with

T = 1300 ◦C. Thus, the flame volume is constant for all operating conditions with reference

combustor outlet temperature. The residence time is determined by Eq. 5.14 and varies

with the mass flow rates and the density. This assumption is justified because turbulent

mixing is the rate-determining effect of the swirler-based non-premixed flame and thus the

flow properties primarily determine the shape of the flame.

Non-ref.temperature

The residence time of the PSR in the case of T < 1300 ◦C is connected to the mass flow

ratio according to

τ = kτmfuel

mair,pr(5.16)


with the proportional factor kτ . This factor is determined at reference combustor outlet

temperature T = 1300 ◦C, when the primary and stoichiometric air mass flow rates are equal

by definition. The factor kτ is kept constant for all combustor outlet temperatures < 1300◦C

in order to determine the residence time via Eq. 5.16. The stoichiometric conditions of

the PSR 1 are not affected by this residence time adjustment for lower combustor outlet

temperatures. Note that the accuracy criterion ∆τ < aτ with aτ = 0.1 · 10−3 ms is used.

5.3.1.3 Residence time for varying pressure

The pressure has a macroscopic influence on the combustion process that affects the flame

length and thus the residence time. A change in pressure of a non-premixed swirler-stabilized

flame can be addressed by the application of the flame surface density model that gives the

burning rate of turbulent flames. Via the flame surface density model, a connection between

the turbulent mixing and the burning rate can be established. Initially, Lachaux et al. (2005)

experimentally determined a relation for a premixed Bunsen flame using a quantification of

the burning rate via the flame surface density model. It can be applied in this study since

the reactions primarily take place after the fuel and air are mixed and are thus present in

perfectly stirred conditions. In this case, the theory and assumptions of premixed flames

are applicable on the microscopic scales of air and fuel interaction, as in Janus (2005). The

burning rate is inversely proportional to the flame volume and thus can be applied to quantify

the change in residence time due to the change in the turbulent Reynolds number Ret caused

by a change in pressure.

Non-ref.pressure

The quantitative relationship between the turbulent Reynolds number and the burning rate

has been experimentally investigated by Lachaux et al. (2005). They found the burning rate

to be proportional to the square root of the turbulent Reynolds number as in

Re0.5t ∝ δtΣ, (5.17)

wherein the product of δt (turbulent flame brush thickness) and Σ (flame surface density) is

proportional to the burning rate. With the turbulent Reynolds number textcolorredas the

dependent variable and directly proportional to the pressure, the initial volume of the flame

VPSR1,init is corrected via the factor f(p) according to

VPSR1,p = VPSR1,init · f−1(p). (5.18)


in this model. The relationship between Ret ∝ p/pref and the burning rate δtΣ ∝ f(p), as

determined experimentally in Lachaux et al. (2005), is described by

f(p) = a+ b

(p

pref

)0.5

. (5.19)

Here, the constant parameters a and b transfer the data given in Lachaux et al. (2005) to

the relevant pressure range. And, since a + b = 1, they transfer the experimental data to a

relative expression for the pressure dependency of the flame volume. From this, it follows

that f(pref) = 1 and thus the initial volume is not affected at the reference pressure. Their

values a = 0.3046 and b = 0.6954 have been determined by fitting the experimental data of

Lachaux et al. (2005).

Closingremark

The change in the flame reactor volume due to the variation in pressure is in line with physical

expectations. Accordingly, an increase in pressure results in a decrease of the flame reactor

volume and a simultaneous decrease of the flame residence time. This decrease in volume is

due to the increasing reactivity of the combustible mixture with increasing pressure. Note

that the reaction kinetics are also influenced by the pressure, as discussed in Sec. 5.2.2.

5.3.2 Mixing zone

The mixing zone connects the flame zone and the burnout zone and is represented by a second

PSR 2, compare Fig. 5.2. It mainly allows for the adiabatic dilution of the flame reactor

outlet flow with excess air. The outlet stream of the mixing zone is an uniformly mixture of

both inlet streams. For that purpose, the chemical reaction progress within this reactor is

suppressed. Therefore the volume of the reactor and its residence time are insignificant.

5.3.3 Burnout zone

The burnout zone of the combustor is modeled by an adiabatic PFR that represents the

movement of the diluted exhaust gas from the mixing zone to the combustor exit. The

inlet stream of this PFR is the outlet stream of the mixing zone. The outlet stream of this

PFR represents the combustor outlet flow and contains the main results of the model: the

exhaust composition and the total residence time. The combustor outlet temperature, T , is

iteratively approximated by a variation of the fuel mass flow rate, using an accuracy criterion

of ∆T < aT with aT = 2 K.

Referenceresidencetime

The PFR volume is chosen in a way that the sum of the flame zone reactor (PSR 1) volume

and the burnout zone reactor PFR volume match the total geometric volume of the exper-

imental test combustor. The total residence time, τtotal, at reference conditions is slightly


lower than the typical residence time of gas turbine combustors (usually ∼10 ms in literature,

see also Tab. 5.3).

5.4 Model validation

This section validates the model results for the four underlying reaction mechanisms. By a

residual investigation, the quality of the results of the four reaction scheme are compared

with the experimental results. Thereafter, the numerical NOx results of the four reaction

schemes are shown in comparison to the experimental data and their trends with different

pressures, combustor outlet temperatures, air inlet temperatures, air inlet velocities, and

amounts of water injection are analyzed. For all validations, special attention is given to

the fuel composition variation from natural gas to hydrogen. Finally, the most suitable

mechanism is identified for the subsequent chemical analysis.

5.4.1 Reaction scheme validation

The four reaction schemes are compared to the experimental results by the rRMSE criterion

shown in Eq. 5.8. With this criterion, the most important variable is the level of NOx

emissions. Furthermore, the residual oxygen in the exhaust gas and the CO2 emissions are

to verify the chemical conversion. Finally, the fuel mass flow rate is used to confirm that

energy balance has been maintained. The database used for this comparison is the complete

set of experimental operating points.

rRMSEvalidation

The relative difference between the experimental and numerical results are summarized in

Tab. 5.4 for NOx, residual O2, CO2, and fuel mass flow rates. Here, the total number of

experimental operating points are referred to as ‘total’. The conglomeration of operating

points excluding and including water injection are referred to as ‘dry’ and ‘wet’, respectively.

GRI 3.0 is the most accurate reaction mechanism according to the results for NOx emissions

for all operating conditions. At wet conditions, however, the rRMSE is higher than at dry

conditions. This increase is observable for all reaction mechanisms. The rRMSE characteris-

tic of the O2 emissions, CO2 emissions, and fuel mass flow, however, give very similar results

without a clear indication of the most favorable reaction mechanism. Note that in case of

the fuel mass flow rate, the rRMSE in wet conditions is lower than the value for the dry con-

ditions. In summary, GRI 3.0 is the most appropriate reaction scheme reaction mechanisms

among the others with regard to the NOx emissions. However, the rRMSE criterion does

not give information about systematic errors and divergent trends of the solutions. Thus the

deviations are displayed and analyzed in more detail.


Table 5.4: Comparison of the underlying reaction schemes for NOx and CO2 emissions, resid-

ual O2, and fuel mass flow rate.

NOx emissions

rRSME [%] total dry wet

GRI 3.0 3.2 3.5 5.2

Aramco 3.4 3.7 5.7

Konnov 3.7 4.0 5.8

San Diego 3.9 4.1 7.3

Residual O2


GRI 3.0 4.9 4.4 6.7

Aramco 4.7 4.5 6.8

Konnov 4.8 4.5 6.8

San Diego 4.8 4.5 6.8

CO2 emissions


GRI 3.0 5.7 5.8 8.1

Aramco 5.8 5.9 8.3

Konnov 5.6 5.6 8.1

San Diego 5.8 5.8 8.2

Fuel mass flow mfuel


GRI 3.0 4.3 4.6 4.5

Aramco 4.4 4.6 4.5

Konnov 4.4 4.7 4.5

San Diego 4.4 4.7 4.6

Comparisonvalidation

Fig. 5.3 (a) to (d) is a more detailed view of the deviation from experiment of the four sim-

ulation mechanisms: GRI 3.0, Aramco, Konnov, and San Diego, respectively. The figures

show the experimental NOx results (abscissas) and the corresponding model results (ordi-

nate). The marker color differentiates test conditions according to the fuel composition and

amount of water injection. The variations of pressure, combustor outlet temperature, air

inlet velocity, and air temperature are implicitly contained within this data but not differ-

entiated in this representation. The logarithmic axis allows for an detailed evaluation of the

various operational regimes, in particular, the low NOx emission region.

While the dry numerical results of GRI 3.0 show good agreement with the combustor opera-

tion data, a systematic deviation can be observed for the results with water injection, as seen

in Fig. 5.3 (a). In wet conditions, the numerical results overestimate the NOx emissions with

natural gas fueling and underestimate the emissions with hydrogen fueling. Such a deviation

is not visible at dry conditions with the same level of NOx emissions (about 70 ppm). Thus,

the deviation can clearly be attributed to water injection. This pattern is also observed

with the Konnov mechanism and the San Diego mechanism with the same trend. Aramco,

however, shows a different trend. Here, the deviation from experimental results during water

injection is observable but significantly smaller and not influenced by the fuel type. The most

accurate set of points from the Aramco mechanism is the dataset with water injection at an

air inlet temperature of 400◦C, all other conditions being at reference conditions. However,

the lower the NOx emissions are, the more the experimental data is underestimated for all


(a) (b)

(c) (d)

Figure 5.3: Comparison of numerical and experimental results of NOx emissions for

(a) GRI 3.0, (b) Aramco, (c) Konnov and (d) San Diego mechanism. The exper-

imental conditions are given in Tab. 4.3.


(a) (b)

(c)

Figure 5.4: Comparison of numerical and experimental results for the mfuel, residual O2, and

CO2 emissions for GRI 3.0. The experimental conditions are given in Tab. 4.3.

fuels. This systematic deviation also occurs at dry conditions for natural gas fueling, so this

phenomenon cannot be attributed to the water injection.

Interimconclusion

As interim conclusion, the GRI 3.0 is the most promising reaction mechanism of this study,

despite the fact that the application range of the pressure conditions is exceeded, com-

pare Tab. 5.1 since this mechanism does not show significant deviation at higher pressures.

Aramco’s higher pressure operation range does not show a significant better outcome re-

garding to the NOx emission in this study. However, GRI 3.0 has a systematic limitation


in matching the experimental NOx emission at increasing water fuel ratio. In the following,

this deviation characteristic is addressed in more detail.

Comparisonfor GRI 3.0

The deviation of the remaining parameters from their experimental values, namely the fuel

mass flow rate, the residual O2, and the CO2 emissions, are shown in Fig. 5.4. The experi-

mental and numerical fuel mass flow rates show excellent agreement. Even at wet conditions,

the fuel mass flow rate is computed accurately. Thus the energy balance is fulfilled for both

dry and wet conditions and all fuel compositions. From these results, it can be concluded

that the heat reduction of the water injection is adequately simulated since the correct fuel

mass flow is used to compensate for the temperature drop. The residual O2, displayed in

Fig. 5.4 (b), shows an overestimation of the model especially at increased water injection

levels. As a consequence, the CO2 emissions are underestimated (see Fig. 5.4 (c)) because

less fuel is converted into the global reaction products. Note that experimental residual O2

and CO2 emissions are measured in dry conditions. Thus the numerical results are also given

as dry conditions. In summary, those trends are minor and independent from the fuel.

5.4.2 NOx characteristic validation

This section presents the numerical results by displaying and analyzing the NOx emissions

trends with changing pressure p, combustor outlet temperature T , hydrogen fraction in the

fuel x, air inlet velocity v, air inlet temperature Tair, and water fuel ratio ψ. These repre-

sentations allow for a more detailed comparison of the experimental data and the numerical

results and allow for the observation of any deviations. Although the diagrams partly show

data that has previously been displayed, they are given in order to observe the trends of the

influencing variables on NOx emissions.

5.4.2.1 Pressure and combustor outlet temperature

The effect of pressure changes on dry NOx emissions is shown in Fig. 5.5 (a) for 0 vol.%

H2 and (b) for 100 vol.% H2 and a set of combustor outlet temperatures. The remaining

parameters are at reference point conditions. The experimental data are displayed by the

symbols and the numeric results are displayed by a set of different lines types depending

on the underlying reaction mechanisms. The colors distinguish sets of combustor outlet

temperatures.

Fig. 5.5 (a) shows that the change in NOx emissions due to pressure variation is best char-

acterized by a mathematical root function. This experimental trend can be matched best

by GRI 3.0 and San Diego, while Aramco and Konnov significantly underestimate the ex-

perimental NOx emissions at low pressure. With decreasing combustor outlet temperature

and all other conditions being at reference levels, the experimental NOx emissions are sys-


(a) (b)

Figure 5.5: NOx emissions as a function of the pressure, p, for a set of combustor outlet

temperatures, T , at (a) 0 vol.% H2 and (b) 100 vol.% H2.

tematically underestimated. Aramco underestimates the NOx emissions most significantly,

followed by Konnov. The deviation from experiment increases with decreasing pressure. As

the hydrogen fuel content increases, shown in Fig. 5.5 (b), both the numerical NOx levels

and trends conform to the experimental results. With pure hydrogen fueling and all re-

maining parameters being at reference conditions, the Aramco mechanism overestimates the

experimental NOx emissions while the other three mechanisms underestimate the emissions.

(a) (b)

Figure 5.6: NOx emissions as a function of the hydrogen content, x, for (a) a set of combustor

outlet temperatures, T , at 16 bar and (b) a set of pressures, p, at T = 1300◦C.


The NOx characteristic of hydrogen admixing is given in Fig. 5.6 (a) for a set of combustor

outlet temperatures at reference pressure and (b) for a set of combustor pressures at reference

combustor outlet temperature. In between the pure fuel conditions, the NOx emissions are

overestimated for all remaining parameter at reference conditions. Thereby, Aramco shows

the highest deviation. Significant deviations can also be observed for pure hydrogen fueling

at low combustor outlet temperature and elevated pressures.

5.4.2.2 Air inlet temperature

Fig. 5.7 shows trend in NOx emissions as the air inlet temperature is changed for a set of

combustor outlet temperatures for (a) natural gas and (b) pure hydrogen fueling. When

decreasing the air inlet temperature from reference conditions while all other influencing

variables remain constant, Aramco shows the best match to experimental data. The other

reaction mechanisms overestimate the NOx emissions with decreasing Tair. In the case of

hydrogen fueling, this trend is enhanced the effects of hydrogen leading to a greater under-

estimation especially at low combustor outlet temperatures.

In the case of pressure reduction with all other variables at reference conditions, see Fig. 5.7

(c) and (d), GRI 3.0 and San Diego overestimate the NOx emissions, while Aramco and

Konnov underestimate significantly. In the case of low pressure and hydrogen fueling, all

reaction mechanisms are close to the experimental data due to superposing effects.


(a) (b)

(c) (d)

Figure 5.7: NOx emissions as a function of the air inlet temperature, Tair, at p = 16 bar for

a set of combustor outlet temperatures, T , at (a) 0% and (b) 100 vol.% H2 and

at T = 1300◦C for a set of pressures, p, at (c) 0 vol.% H2 and (d) 100 vol.% H2.

5.4.2.3 Air inlet velocity

The NOx emissions trends due to the variation of the relative air inlet velocity for a set

of different combustor outlet temperatures are shown in Fig. 5.8 for (a) natural gas fuel

and (b) pure hydrogen fueling. When the velocity is less than the reference point, Aramco

shows the smallest deviation between the experimental data and model results. The other

mechanisms underestimate the NOx emissions to a similar degree. Reduced combustor outlet

temperatures continue this trend. The hydrogen fueling in (b) shows a similar characteristic


(a) (b)

Figure 5.8: NOx emissions as a function of the relative air inlet velocity, v/vref , at p = 16 bar

for a set of combustor outlet temperatures, T , at (a) 0 vol.% H2 and (b) 100 vol.%

H2.

for the reduction of the inlet velocity, however, at lower T the underestimation is very large

for all mechanisms.

(a) (b)

Figure 5.9: NOx emissions as a function of the relative air inlet velocity, v/vref , at p = 16 bar

and T = 1300 ◦C for a set of combustor inlet temperatures, Tair, at (a) 0 vol.%

H2 and (b) 100 vol.% H2.


Fig. 5.9 shows the NOx emissions trend for the air inlet velocity at reduced air inlet tem-

perature. At Tair = 400 ◦C, the NOx emissions’ results from different numerical mechanisms

merge under the experimental emissions value for natural gas fueling. In the case of hydro-

gen, the model predictions underestimate the experimental results, with Aramco showing

the smallest deviation.

5.4.2.4 Water injection

In order to study water injection as active NOx reduction measure, Fig. 5.10 shows the

NOx emissions as a function of the water injection ratio for a set of fuels, (a) at reference

conditions and (b) at a decreased combustor outlet temperature of T = 900◦C. Note the

logarithmic scale of the ordinate. At reference conditions, GRI 3.0, San Diego and Konnov

overestimate the experimental NOx emissions, while Aramco underestimates the emissions.

During pure hydrogen combustion, however, all mechanisms underestimate the NOx emis-

sions at higher water injection ratios. This phenomena has already been shown in Fig. 5.3.

At the lower combustor outlet temperature, shown in (b), and natural gas fueling, Aramco

underestimates the experimental data, while all others match the NOx emissions. During

hydrogen combustion, however, all reaction mechanisms significantly underestimate the ex-

perimental results due to the superposed underestimation of the influence of the combustor

outlet temperature, as seen in Fig. 5.7 (b).

(a) (b)

Figure 5.10: NOx emissions as a function of the water fuel ratio, ψ, for a set of hydrogen

fractions x at p = 16 bar and (a) T = 1300 ◦C and (b) T = 900 ◦C.


(a) (b)

Figure 5.11: NOx emissions as a function of the water fuel ratio, ψ, for a set of pressures, p,

at T = 1300 ◦C and (a) 0 vol.% H2 and (b) 100 vol.% H2.

Fig. 5.11 shows the NOx emissions as a function of the water injection ratio for different

pressures down to 3 bar for (a) natural gas and (b) hydrogen fueling, with all other param-

eters remaining at reference conditions. While the underestimation of Aramco and Konnov

increases at low pressure, the underestimation of GRI 3.0 and San Diego decreases. Only

GRI 3.0 captures the greater NOx emissions reduction at the reference pressure compared to

lower pressures, though the transition point is shifted to a higher water fuel ratio (ψ = 1.25

in the numerical solution compared to about 0.8 in the experiment). While the emissions

in dry conditions are overestimated at 3 bar for hydrogen fuel (compare Fig. 5.11 (b) and

Fig. 5.5 (b)) all wet NOx emissions are underestimated.

The effects of the air inlet temperature variation and water injection on NOx emissions

are shown in Fig. 5.12 for (a) natural gas fuel and (b) hydrogen fuel. At reduced air inlet

temperature, Aramco shows the best results with natural gas fueling, while the other reaction

mechanisms overestimate the NOx emissions. During hydrogen combustion, the GRI 3.0,

Aramco, and Konnov mechanisms underestimate to a similar degree while the San Diego

mechanism shows significantly greater underestimation.


(a) (b)

Figure 5.12: NOx emissions as a function of the water fuel ratio, ψ, for a set of combustor

inlet temperatures, Tair, at p = 16 bar and T = 1300 ◦C and (a) 0 vol.% H2

and (b) 100 vol.% H2.

5.4.3 Validation conclusion

rRMSE andNOx

deviation

Using the rRMSE approach to compare the deviations of all mechanisms from the exper-

iment indicates that GRI 3.0 has the best overall matching for NOx emissions while the

performance for residual O2, CO2 emissions and fuel mass flow rate is satisfactory for all

mechanisms. At dry conditions, the comparison of deviations from experiment for NOx

emissions showed that all reaction mechanisms performed satisfactorily. GRI 3.0 showed the

best matching to experimental results. At wet conditions, GRI 3.0 showed the closest match

in terms of the rRMSE criterion but showed a systematic overestimation for natural gas

and an underestimation for hydrogen. Aramco showed a systematic underestimation with

increasing water fuel ratio that is apparent at dry low NOx emissions conditions, as well.

Thus, this deviation is not limited to the water injection. Konnov and San Diego showed

higher deviation in both dry and wet conditions.

NOx char-acteristic

The emissions characteristic investigation showed that the numerical results match the ex-

pected general NOx trends of the experimental non-premixed combustion system for all pa-

rameter studies remarkably well. In case of pressure variation, the emissions have a square

root functional dependence on pressure, and GRI 3.0 and San Diego performed best. How-

ever, comparably high deviations from experimental results can be seen for all reaction mech-

anisms at dry combustion conditions and reduced combustor outlet temperatures, especially

with high hydrogen fuel content. The linear relationship between the air inlet tempera-

ture, and the slightly convexly-curved relationship between the air inlet velocity and NOx


emissions have also been successfully described by the numerical data. Therefore, the per-

formances of all reaction mechanisms were acceptable. With reference to water injection,

the general trends of the NOx emissions are sufficiently captured by all mechanisms. The

large relative deviations for high water fuel ratios can be attributed to the large distance

from the reference point.

ConclusionIn conclusion, GRI 3.0 is the best choice for the further detailed chemical investigations.

Since this model is capable of capturing the trends rather than the exact conditions of the

this individual combustor, the systematic deviation is acceptable, which is the lowest among

the four reaction mechanisms. Thus, GRI 3.0 is selected as the highest performing mechanism

for both methane and hydrogen at dry and wet conditions. The identified model limitation

at low combustor outlet temperatures is acceptable as the combustor outlet temperature

during real gas turbine operation is kept as high as possible for efficiency.

5.5 Model response analysis

Fundamental studies on combustion in the literature analyze the effects of the individual re-

action scheme on radical distribution, pathways, and emissions at constant adiabatic reactor

temperatures. In those studies, the effects of parameter variations, like equivalence ratio,

fuel composition, and humidity, on the reactor temperature are eliminated by adjusting the

air inlet temperature to keep the reaction temperature constant. However, the focus of this

study is on real operating conditions and thus the flame temperature and residence time are

results of the model and vary significantly with the operating conditions of the combustor.

This section investigates the response of the flame temperature and flame residence time to

the operating parameters. The focus is on the flame reactor PSR 1, since the NOx emissions

are mostly formed in the flame of PSR 1. The burnout of PFR has a minor contribution

to NOx formation (lower than 4 ppm) since the chemical reactions for NOx formation are

quenched by the excess air injection.

Flametemperature

The dependence of the stoichiometric flame temperature, Tst, on the operating point varia-

tions is given in Fig. 5.13 (a). Here, the response of Tst to variations in individual parameters

relative to their reference value is displayed. Note that the parameter variations are inde-

pendent of each other: for a specific variation all other input parameters remain constant at

their reference values (= ceteris paribus). The pressure increase from 3 to 24 bar results in a

significant increase of Tst mainly due to an increase in reactivity. The impact of the pressure

has the same magnitude as that of the fuel composition change. Compared to natural gas,

the flame temperature increases by nearly 8% or 185 K ceteris paribus for pure hydrogen fu-

eling. Changes in the combustor outlet temperature and the air inlet velocity have a minor,


(a) (b)

Tst/Tst,ref Single parameter

varied*

x-axis

p/pref lower

x lower

T/Tref upper

v/vref upper

Tair/Tair,ref upper

*All other parameters remain at reference conditions,

compare Tab. 4.3

Figure 5.13: Model response to varying model input parameters p, T , x, vrel, Tair for (a) the

relative stoichiometric air temperature, Tst/Tst,ref , and (b) the PSR 1 residence

time, τ/τref . The diagrams are comprised of five parameter variations, where

only single parameters are varied to show their individual influence on Tst and

all other operating conditions explicitly remain constant. The table assigns the

line styles to the parameters and indicates which parameters correspond to each

x-axis. The reference conditions are given in Tab. 4.3.

respectively, increasing and decreasing influence on Tst. Furthermore, an increase of air inlet

temperature causes a significant, 2.6% (or 65 K) increase in Tst, showing a linear trend.

The response of the flame temperature to water injection is shown in Fig. 5.14. Note that

the water injection is quantified here by the normalized water fuel ratio, ψn. It is normalized

to the thermal heat input of natural gas and thus allows for a simple comparison between


Figure 5.14: Response of the model NOx emissions, flame temperature, Tst, and residence

time, τ , to changes in the water fuel ratio, ψn, and hydrogen fuel content, x.

different fuels. This is done to avoid the high deviations of ψ for different fuels caused by

different fuel mass flow rates at similar heat inputs. By definition, at a constant ψn, the

quotients of the water mass flow rate and the heat input of the fuel are equal for any fuel.

The flame temperature reduces significantly during water injection due to the heat loss to

the cooling and evaporation of water. The underlying trend is linear and the temperature

reduction is 15.5% (more than 380 K), as shown in Fig. 5.14. In the case of hydrogen

combustion, the water cooling is slightly more efficient (17.2% or 425 K).

Residencetime

The effects of different parameters on the residence time in the flame are given in Fig. 5.13 (b).

A decrease in pressure increases the residence time by 86%, primarily caused by the model

definition of the reactivity decrease. The residence time decreases linearly to about 50%

of the reference value when the combustor outlet temperature is reduced to 900◦C, which

is also caused by the model definition. The fuel gas composition has a negligible effect

on the residence time. Here, the elevated stoichiometric flame temperature, Tst, seen with

hydrogen fueling causes a lower density, which is nearly compensated for by the decreasing

stoichiometric air mass flow rate, mair,st, of Eq. 5.14. The air inlet velocity is a measure of

the volume flow into the combustor and thus has a significant effect on the residence time.

The flame residence time drops with increasing velocity. Since the variation of the air inlet

temperature affects the adiabatic flame temperature, the residence time is simultaneously


affected due to the change in the density. When the air inlet temperature is decreased to

400◦C, the residence time decreases by 22%.

The residence time for water injection is given in Fig. 5.14. Water injection adds additional

mass flow to the reactor and, in order to keep the combustor outlet temperature, T , constant,

the fuel mass flow rate is increased. This leads to a linear decrease of the residence time by

20% during natural gas fueling. In the case of hydrogen fueling, the decease of the residence

time is 21.5%. Thus, the increase in the total mass flow rate and the resultant reduction

of residence time is the predominating effect of the density increase due to the reduction

of the flame temperature, Tst. Note that this model does not agree with Syed (2013), who

found a constant residence time for water injection. In this study, the initial flame volume

is regarded as constant leading to a flexible residence time during parameter variations.

Waterinjection

Fig. 5.14 furthermore indicates a higher potential for the use of water injection as a NOx

reduction mechanism when using hydrogen fuel. Quantitatively, the reduction factor of

hydrogen is about four times higher than the reduction factor of natural gas. This significant

factor is partially due to the higher potential for flame temperature reduction. However, the

underlying chemistry plays another significant role in this effect and is investigated in the

following section.

5.6 Chemical influence of hydrogen and water

The chemical effects of hydrogen and water cause a significant change in the reaction progress,

particularly in the NOx formation pathways. In the following, methods and results of the

literature are reviewed, and the important radicals of this study are formally identified.

Thereafter, the trend effects of hydrogen and water injection on the radical mix are in-

vestigated. Finally, a pathways study investigates the effects of radicals and model input

parameters on NOx formation.

5.6.1 Review

Although the chemical analysis of hydrogen and water has been described in literature, no

consistent trend regarding the radical distribution can be identified. This is mainly because

the various applications fundamentally differ in their boundary conditions. In the following,

the influence of hydrogen and water addition is reviewed.

Influence ofhydrogen

The literature agrees on an increase in radical concentration when increasing the amount of

hydrogen in the fuel. For a non-premixed flame, Park et al. (2007) identified an increase in

OH◦ and H◦, and a slight increase in O◦ radical concentrations. Goke (2012) and Gockeler

(2015) reported an increase of the H◦ radical concentration and to a lesser extent an increase


in the O◦ and OH◦ radical concentrations for a premixed configuration and constant flame

temperature.

Influence ofwater

In literature, the focus of the effect of water is on O◦, H◦, and OH◦ radicals. It is reported

that water injection influences the radical concentrations differently in premixed and non-

premixed flames. In non-premixed flames, Zhao et al. (2002) investigated both decreasing

and constant flame temperature conditions for a counterflow diffusion flame. They reported

a significant decrease in OH◦ concentration in the case of a decreasing flame temperature

and an increase in OH◦ concentration in the case of a constant flame temperature. Park

et al. (2007) confirmed the trend of decreasing radical concentrations of OH◦, H◦ and O◦

for a decreasing flame temperature. In contrast, for a flame with a constant molar oxygen

share (20% O2), Suh & Atreya (1995) and Atreya et al. (1999) reported an increase of OH◦

when increasing the water content, while also reporting a slight increase in the flame tem-

perature. In premixed flames during decreasing flame temperature applications, Guo et al.

(2008) reported a decrease of H◦, O◦, and OH◦ for counterflow premixed flames at Φ = 1

and an increasing water mass flow rate. This is confirmed by Le Cong & Dagaut (2008)

who stated that the chemical effect of water results in a reduction of the concentrations of

the main radicals H◦, O◦, and OH◦. In contrast, Mazas et al. (2011) showed that, for a

constant flame temperature, during a comparison of water injection with inert dilution of

the same properties, water reduces the H◦ and O◦ concentration, while the OH◦ concentra-

tion increases. For a premixed flame at constant flame temperature, Goke (2012) found a

slightly increasing OH◦ concentration at low steam content that can also decrease again for

higher dilution levels depending on the equivalence ratio. Here, the O◦ radical concentration

is significantly reduced at wet conditions and the CH◦ concentration slightly increases in

natural gas flames.

ConclusionIn conclusion, while it is clear that the chemical influence of hydrogen causes a universal

increase in the radical concentrations, the effect of water on the radical distribution is highly

dependent on the operating conditions. In particular, the OH◦ concentration is highly de-

pendent on different boundary and operating conditions. Thus, an individual analysis of the

relevant application is necessary, which is presented in the following section.

5.6.2 Radical mix analysis

Here, the effects of increasing the hydrogen content in the fuel and the water injection mass

flow rate on the most important radicals are evaluated. A sensitivity analysis identifies the

most significant reactions and radicals relevant to the NO formation. Thereafter, the concen-

tration dependences on the most relevant radicals are investigated for various hydrogen fuel


concentrations and water injection ratios. Finally the results of this analysis are evaluated

for their effects on NOx emissions.

The chemical effect of water is investigated by introducing an artificial species, referred

to as iH2O (= inert H2O). It has the same thermodynamic properties as H2O, but does

not participate in any chemical reactions. This approach is very common in literature, see

e.g. Goke (2012), Guo et al. (2008), Le Cong & Dagaut (2008), Mazas et al. (2011) and Park

et al. (2007). Water is also a significant inert collision species M in three-body reactions.

It enhances chemical reaction rates due to its high chaperon efficiency. In this study, the

influence of H2O as third body is investigated by neutralizing its efficiency. Therefore, the

chaperon efficiency is set to unity, which is the efficiency of nitrogen.

5.6.2.1 NO sensitivity analysis

The sensitivity analysis identifies the most important reactions of the NOx generation pro-

cess. The equations with the highest sensitivity are the rate-determining reactions and thus,

their kinetic data must be known accurately. At the same time, the sensitivity analysis

identifies the most important radicals. The higher the absolute value of the sensitivity co-

efficient σrelNO,k, the larger effect the reaction k will have on the formation of NO. Thus, a

positive coefficient indicates enhanced NO formation, while a negative coefficient indicates

suppressed NO formation.

Figure 5.15: Sensitivity coefficients σ for NO formation in different hydrogen admixtures.


Figure 5.16: Sensitivity coefficients σ for NO formation with natural gas fuel and different

water injection levels.

Hydrogensensitivity

Fig. 5.15 shows the NO sensitivity coefficients with σrelNO,k > 4% of the maximum value.

The set of different hydrogen mixtures are represented by different shades of gray. All other

input parameters remain at reference conditions. With natural gas fueling, half of reactions

contain a carbon atom (namely GRI 10, 93, 97, 99, 125, 126, 127, and 132), while three

reactions belong to the H2/O2 mechanism (GRI 38, 43, and 287). The remaining reactions

(GRI 178, 180, 199, 208, and 240) are attributed to the nitrogen mechanism. The most

important reactions regarding to NO formation during natural gas fueling are GRI 38 (a

chain branching reaction transforming H◦ into O◦ and OH◦), GRI 99 (a chain propagating

reaction for CO oxidation via OH◦ forming an H◦ radical), and GRI 178 (the nitrogen

break-up reaction via the thermal NO pathway). During the fuel switch, the importance of

reactions containing carbon naturally reduces to zero and consequently, GRI 178 and most

of the minor reactions gain importance. The most important radicals for the fuel switch

analysis are H◦, O◦, OH◦, and CH◦ radicals.

Watersensitivity

The sensitivity coefficients for a set of water injection levels are shown in Fig. 5.16 with

natural gas fueling. Compared to hydrogen fueling, the set of important reactions is extended

to include GRI 35, and 95, while reactions GRI 43, 199, 208, and 132 are excluded. Except

for GRI 126, 178, and 180, all reactions show an increase in significance at higher water

injection levels. While the thermal NO pathway decreases in significance (GRI 178), the

significance of the prompt nitrogen break-up reaction (GRI 240) increases. Due to the


increasing significance of GRI 287 and 35, the HO2◦ radical is added to the list of important

radicals.

Conclusion In conclusion, the radicals H◦, O◦, and OH◦ are important in chain-branching reactions and

thus have a strong influence on the combustion process and NO formation. CH◦ is important

for the methane oxidation pathway and HO2◦ gains importance in high H2O regimes. Thus,

the focus during this radical distribution study is on this set of radials. Note that the rate

coefficients of the high sensitivity reactions are crucial for the performance of the scheme

and have to be well-known. The remaining reactions are comparably fast and thus have a

minor influence on the simulation results.

5.6.2.2 H◦ radical analysis

The production and consumption trends of the atomic hydrogen radical H◦ are significantly

influenced by the hydrogen fraction of the fuel and injected water mass flow rate. Fig. 5.17

shows the production rate of the H◦ radical for the most important reactions (those with

an absolute value higher than 3% of the maximum value). A negative production value

represents the consumption of the radical. While (a) shows the change for an increasing

hydrogen fraction, (b) shows an increasing water injection mass flow rate during natural gas

fueling.

In natural gas flames, the reaction

CO + OH◦ H◦ + CO2 (GRI 99)

is the most important reaction of the H◦ pathway that contains carbon. Additionally im-

portant H◦ production reactions

H2O + HCO CO + H◦ + H2O (GRI 166)

and

HCO + M CO + H◦ + M (GRI 167)

contain carbon, as well. As the hydrogen content of the fuel increases, these reactions are

replaced by the reactions

H2 + O◦ H◦ + OH◦ (GRI 3)

and

H2 + OH◦ H◦ + H2O (GRI 84)

that do not contain carbon and mainly produce H◦ radicals by breaking up molecular hy-

drogen using O◦ and OH◦ radicals. The main consumption reactions are

H◦ + O2 O◦ + OH◦ (GRI 38)


(a)

(b)

Figure 5.17: The H◦ net production rate for significant reactions at the reference parameters

for changing (a) hydrogen fraction of the fuel and (b) water injection in the case

of natural gas fuel. All other operating parameters p, T , vref , and Tair are at

reference conditions according to Tab. 4.3.


(a) (b)

Figure 5.18: The H◦ concentration at reference point conditions and selected variations of p,

T , vrel, and Tair: (a) shows results for varying hydrogen fractions in the fuel and

(b) shows results for water injection at reference point conditions (Tab. 4.3).

and

H◦ + H2O + O2 H2O + HO2◦ (GRI 35)

breaking up O2 and producing O◦, OH◦, and HO2◦ radicals. Thus, the H◦ radical has a

chain branching effect on the reaction process that increases at first but then decreases at

higher hydrogen levels. Also,

H◦ + OH◦ + M H2O + M (GRI 43)

is important as chain termination reaction that increases its H◦ consumption at higher

hydrogen levels.

The overall H◦ radical concentration in the flame is shown in Fig. 5.18, where (a) presents

the results of increasing hydrogen levels at reference conditions. Additionally, fuel variations

at points off the reference conditions are shown. The indicated variable has been varied while

all other input variables remain at reference conditions. As the hydrogen fraction of the fuel

increases, the H◦ radical increases exponentially. The H◦ radical concentration decreases

with pressure, and the pressure influence has the same magnitude as the hydrogen effect.

The H◦ radical concentration also decreases with combustor outlet temperature and shows

a minor increase with air inlet speed and air inlet temperature.

The relative change of the H◦ radical concentration under the influence of water injection

is shown in Fig. 5.18 (b). The reference values of natural gas and hydrogen fuel can be


found in (a). For the natural gas flame, the H◦ radical concentration decreases to about half

of its initial concentration when water injection is increased. For the hydrogen flame, the

concentration drops to less than 40% of the original value. Because of this, in both natural

gas and hydrogen flames, the chemical effect of the water plays a significant role. The

chemical activity of H2O leads to a decrease in the H◦ radical concentration. The specific

collision efficiency of water accounts for about one fifth of the chemical effect (iH2O effect)

of water.

5.6.2.3 O◦ radical analysis

The most important reactions for production and consumption of the atomic oxygen radical

O◦ are displayed in Fig. 5.19 (a) for fuel variation and (b) for water injection increase. The

major production path of O◦ is GRI 38 for both methane and hydrogen flames. This reaction

is not significantly affected during fuel variation from natural gas to hydrogen. The major

consumption pathway is along reaction GRI 3 and

2 OH◦ H2O + O◦. (GRI 86)

Further consumption pathways exist via the of decomposition of CH3 in

CH3 + O◦ CH2O + H◦ (GRI 10)

and

CH3 + O◦ CO + H◦ + H2 (GRI 284)

and the decomposition of methane via

CH4 + O◦ CH3 + OH◦. (GRI 11)

While the consumption rate of reactions GRI 10, 11, and 284 vary with the amount of

hydrogen share in the fuel due to the presence of carbon, the consumption via GRI 86 is not

significantly affected by the hydrogen fuel variation.

The concentration of the oxygen radical O◦ increases with increasing hydrogen fraction of

the fuel, as seen in Fig. 5.20 (a). However, the strongest effect on the O◦ concentration

is from the pressure, which causes an increase of the radical concentration by a factor of

more than three for the experimental pressure reduction. Decreasing the combustor outlet

temperature leads to a slight increase, while a reduction of the air inlet temperature and the

reduction of the air inlet velocity each cause a decrease of the O◦ radical concentration.

In the case of water injection (see Fig. 5.20 (b)), the O◦ radical concentration decreases with

an increase in injected water. This shows that the reduction characteristic is similar for both

fuels when comparing similar water injection mass flow rates. In both cases, the radical

concentration decreases to about one third of their dry reference values. The inert H2O


(a)

(b)

Figure 5.19: The O◦ net production rate for significant reactions at the reference parameters

for changing (a) hydrogen fraction of the fuel and (b) water injection in the case

of natural gas fuel. All other operating parameters p, T , vref , and Tair are at

reference conditions according to Tab. 4.3.


(a) (b)

Figure 5.20: The O◦ concentration at reference point conditions and selected variations of p,

T , vrel, and Tair: (a) shows results for varying hydrogen fractions in the fuel and

(b) shows results for water injection at reference point conditions (Tab. 4.3).

simulations show significantly higher O◦ radical concentrations and thus indicate a strong

chemical effect. Thus, the chemical effect with natural gas fuel is stronger than with hydrogen

fuel. H2O as a collision partner, however, slightly reduces the radical concentration. Overall,

the chemical influence of water causes a significant reduction of the O◦ radical concentration.

5.6.2.4 OH◦ radical analysis

The production and consumption rates of the hydroxyl radical OH◦ are linked to the O◦

and H◦ radical concentrations. Fig. 5.21 (a) shows the net production rate of OH◦ when the

hydrogen concentration is increasing while (b) shows the net production rate during water

injection. The OH◦ radical is strongly integrated in the combustion process as indicated by

a large number of important production and consumption reactions. The main production

reactions are GRI 3, 38, and 86. While the net production rate of the first reaction increases

with increasing hydrogen share, the net production rates of the latter remain almost constant.

The OH◦ radical is mainly consumed by the non-carbon containing reactions GRI 43 and

84, whose consumption increases significantly with the hydrogen fraction. The reactions

2 OH◦ (+M) H2O2 (+M), (GRI 85)

H2O2 + OH◦ H2O + HO2◦, (GRI 89)

and


(a)

(b)

Figure 5.21: The OH◦ net production rate for significant reactions at the reference param-

eters for changing (a) hydrogen fraction of the fuel and (b) water injection in

the case of natural gas fuel. All other operating parameters p, T , vref , and Tair

are at reference conditions according to Tab. 4.3.


(a) (b)

Figure 5.22: The OH◦ concentration at reference point conditions and selected variations of

p, T , vrel, and Tair: (a) shows results for varying hydrogen fractions in the fuel

and (b) shows results for water injection at reference point conditions (Tab. 4.3).

HO2◦ + OH◦ H2O + O2 (GRI 287)

remain nearly constant when increasing the hydrogen concentration. Furthermore, OH◦

consumption takes place in the carbon oxidation reactions

CH3 + OH◦ CH2(S) + H2O, (GRI 97)

CH4 + OH◦ CH3 + H2O, (GRI 98)

GRI 99, and

CH2O + OH◦ H2O + HCO, (GRI 101)

where the importance decreases down to zero due to the absence of carbon during pure

hydrogen fueling.

The OH◦ radical concentration significantly increases with the hydrogen concentration to

nearly twice of the value at natural gas fueling (see Fig. 5.22 (a)). The concentration of the

OH◦ radicals is the highest among all the radicals. With decreasing pressure, the radical

concentration again increases significantly. With increasing combustor outlet temperature

and decreasing air inlet temperature and velocity, the OH◦ concentration decreases.

With increasing water injection, the OH◦ concentration deceases to about half of its original

value in both fuel scenarios, as shown in Fig. 5.22 (b). In contrast to the behavior of O◦

and H◦ radical, the chemical effect of water results in an increase of the OH◦ concentration.


Neglecting the effect of H2O as collision partner, however, eliminates the chemical effect of

H2O.

5.6.2.5 HO2◦ radical analysis

As pointed out earlier, the HO2◦ radical gains significance in the case of water injection. In

general, this radical has a chain terminating effect on the combustion process. It is mainly

produced by GRI 35, 36, and 89 and thereby, OH◦ and H◦ are consumed, see Fig. 5.23. An

increase in the hydrogen level does not have a significant effect on the net production rate.

HO2◦ is significantly consumed by GRI 287,

H◦ + HO2◦ 2 OH◦, (GRI 46)

and GRI 87. Note that equation GRI 87 and 287 include the same reactants but have

different Arrhenius coefficients in the reaction mechanism. While hydrogen has a minor

effect on the net production rates, water injection has increases the rate of the production

reaction GRI 35 and decreases the rate of GRI 89. The consumption of the HO2◦ radical

increases with water injection.

The HO2◦ concentration is at the ppm level, shown in Fig. 5.24 (a). When increasing the

hydrogen concentration, the radical concentration reaches a maximum at around 80% H2

and then decreases again to the initial concentration with pure hydrogen fuel. In contrast

to the H◦, O◦, and OH◦ radicals, the pressure increases the HO2◦ concentration. While

increasing the combustor outlet temperature decreases the HO2◦ concentration, the radical

concentration decreases with increasing air inlet velocity. Furthermore, the lower the hydro-

gen fraction, the higher the stronger the effect of a decrease in the HO2◦ concentration with

increasing air inlet temperature.

With increasing water injection mass flow rate, the radical concentration of HO2◦ initially

decreases but reaches a the minimum at ψn ≈ 0.25 for natural gas and ψn ≈ 0.60 for hydrogen

fuel. Thereafter, the concentration increases, reaching global peak values for natural gas.

The chemical effect further increases the radical concentration. The collision efficiency of

H2O causes a decrease in the radical concentration for both fuels.


(a)

(b)

Figure 5.23: The HO2◦ net production rate for significant reactions at the reference param-





(a) (b)

Figure 5.24: The HO2◦ concentration at reference point conditions and selected variations of



5.6.2.6 CH◦ radical analysis

The CH◦ radical contains atomic carbon and hence can only appear with natural-gas-

containing fuels. With increasing hydrogen concentration, the significance of all relevant

reactions decreases to zero, shown in Fig. 5.25 (a). CH◦ is mainly produced by

CH2 + OH◦ CH◦ + H2O (GRI 93)

and

CH◦ + H2 CH2 + H◦. (GRI 126)

While GRI 93 produces less CH◦ via the OH◦ pathway with an increasing water injection

mass flow rate, the pathway via the H◦ radical in GRI 126 becomes more productive, shown

in Fig. 5.25 (b). The main consumption pathways of CH◦ are

CH◦ + O2 HCO + O◦, (GRI 125)

CH◦ + H2O CH2O + H◦, (GRI 127)

and

CH◦ + CO2 CO + HCO. (GRI 132)

The consumption of CH◦ in GRI 125 and 132 decreases with increasing water mass flow rate,

while the consumption in GRI 127 increases.


(a)

(b)

Figure 5.25: The CH◦ net production rate for significant reactions at the reference param-





(a) (b)

Figure 5.26: The CH◦ concentration at reference point conditions and selected variations of



The overall level of the CH◦ concentration is the lowest among all considered radicals, com-

pare Fig. 5.26 (a). The concentration of the CH◦ radical decreases with increasing extent

down to zero for hydrogen combustion due to the absence of carbon. An increase of the

combustor pressure, the combustor outlet temperature, and the air inlet temperature has a

reducing effect on the CH◦ radical concentration. In contrast, an increase in the air inlet

velocity has an increasing effect on the radical concentration.

With increasing water injection, the CH◦ radical concentration decreases to significantly

below 60% of its original value, shown in Fig. 5.26 (b). Here, the chemical effect of the

water is the greatest of all the considered radicals. In the case of inert water injection, the

relative CH◦ concentration only decreases to about 85% of its original value. The impact of

the collision efficiency further supports the radical concentration decrease.

5.6.2.7 Summary

For the boundary conditions, given in Sec. 5.5, similar trends for the formation of O◦, H◦,

and OH◦ radical species have been found. Briefly, an increased hydrogen concentration in the

fuel increases the concentration of these radicals, while the concentration of CH◦ decreases

and the concentration of HO2◦ is nearly constant. An increased water injection mass flow

rate reduces all radical concentrations except for that of the HO2◦ radical. The chemical

effect of water further reduces the concentrations of H◦, O◦, and CH◦ but increases the


concentrations of OH◦, and HO2◦. The collision efficiency of H2O marginally decreases all

considered radical concentrations.

The decrease in concentration with increasing pressure is caused by Le Chatelier’s principle,

described e.g. in Turns (2000). The pressure significantly affects the competition between

the chain branching reaction, chain propagating reaction, and chain termination reactions.

At lower pressure, the chain branching reactions dominate leading to the increase in the H◦,

O◦, OH◦, and CH◦ radical concentrations. At higher pressures, three-body reactions take

over that recombine the radicals via an inert third body collision partner, e.g. forming the

HO2◦ radical. Thus, H2O dissociation is inherently affected by the pressure, as well. Note

that a variation in the flame temperature and residence time only marginally influences the

radical concentrations of O◦, H◦, and OH◦.

These results are in agreement with the literature with respect to hydrogen behavior. For

a non-premixed flame, Park et al. (2007) identified an increase in the OH◦ and H◦ radical

concentrations, and a slight increase in the O◦ radical concentration for an increase in hydro-

gen fueling. When flame temperature is decreased during water injection, Park et al. (2007)

found a decrease in the H◦, O◦, and OH◦ concentrations. This is confirmed for OH◦ by Zhao

et al. (2002). Qualitatively, the results from the model created for this study identify the

general radical formation trends. However, a quantitative comparison is difficult due to the

elevated gas turbine conditions necessary for this study.

The distribution of the radicals is essential knowledge for a deeper understanding of the

NOx pathway analysis. The next section presents the significance of the radicals for NOx

generation and evaluates the response of the pathways with varying operational parameters.

5.6.3 NO formation analysis

Besides the thermo-physical conditions within the flame, the presence and distribution of

the radicals play a significant role in NO formation. This section connects the outcome of

the radical concentration investigation with the NOx generation trends found in this study.

The remainder of this section is as follows: First, the effect of the radical distribution on the

NO pathways is discussed, based on the most significant equations (based on the sensitivity

analysis) and their behavior during changes in fuel and an increase in the water injection

rate. Thereafter, the contributions of the NO pathways are evaluated for a variation of all

global parameters.


5.6.3.1 Effects of radicals on NO pathways

This section evaluates the effects of the radical concentrations on the NO formation process

with a focus on the pathways. Since this study simulates the flame characteristics based

on operational combustor parameter variations, a detailed investigation in the variation

of each single radical is not intended. Thus overall phenomena are described with the

known responses of internal parameters, like flame temperature and residence time. The NO

formation chains of the thermal, prompt, NNH, and N2O pathways are addressed in this

study with a focus on the effects of hydrogen and water injection.

Thermalpathway

The thermal NO pathway implies the presence of the O◦ radical in the rate-determining

reaction GRI 178. Furthermore, either molecular oxygen O2 (GRI 179) or OH◦ radicals

(GRI 180) are required for the oxidation of the N◦ radical, and these reactions then produce

an O◦ radical or H◦ radical. However, the first reaction is rate-limiting and thus the O◦

radical chemically controls the formation of thermal NOx. However, the flame temperature,

Tst, significantly determines the behavior of the thermal NO pathway as well, since a higher

flame temperature provides more activation energy for breaking up the N2. The significant

decrease of the O◦ radical concentration with a decrease in pressure is superposed by the

significant increase of the flame temperature, and thus compensates for the O◦ radical reduc-

tion. During hydrogen combustion, an increase in both the O◦ radical concentration and Tst

is observed. For water injection, both the O◦ concentration and flame temperature decrease.

The residence time has only a minor effect on the O◦ radical concentration, as shown in

Sec. 5.6.2.3.

Promptpathway

The prompt NO pathway implies the presence of the CH◦ radical in order to break up N2

and form HCN via GRI 240. The products HCN and N◦ are then oxidized to form NO.

In this pathway, a high flame temperature and a long residence time are not required. The

sensitivity analysis shows that CH◦ radicals have an impact on NO production, because of

the significant number of CH◦-containing reactions that participate in the formation (see

Fig. 5.15). CH◦ is present in six relevant reactions. This indicates that the formation

pathways of NO significantly change during fuel transition from natural gas to hydrogen.

The increased combustor pressure leads to a significant decrease in the concentration of

CH◦ radicals. The radical concentration decreases by 50% during the combustor outlet

temperature increase. The prompt mechanism strongly depends on the presence of CH◦

radicals that are delivered only by the fuel. For pure hydrogen fuel, CH◦ is non-existent

and the prompt pathway is thus totally blocked. For carbon-containing fuels, the CH◦

concentration significantly increases with increasing air inlet velocity. An air temperature

increase results in a slight decrease in the CH◦ concentration. Water injection significantly

reduces the CH◦ radical concentration by more than 40%, as shown in Sec. 5.6.2.6.


NNHpathway

Within the NNH pathway, an H◦ radical is involved in the formation of NNH◦ from molecular

nitrogen, N2. The sole origin of the H◦ radical is the fuel, as both natural gas and hydrogen

contain bound hydrogen atoms. H◦ radicals can be considered as unburnt fuel and therefore

the NNH pathway competes with oxidizing reactions mainly via GRI 38, which is the most

important H◦-consuming reaction.

N2Opathway

Within the N2O pathway, the N2O molecule is formed by a reaction of molecular nitrogen N2

with an O◦ radical from the same reactants as are in the initial thermal pathway reaction. In

a next step, another O◦ radical reacts with N2O forming two NO molecules. This pathway

is thus significantly dependent on the O◦ radical concentration whose effects on the model

parameters are given in Sec. 5.6.2.3.

5.6.3.2 NO pathway analysis

The following section presents the results of the NO pathway analysis. The changes in

the influence of the four NO pathways during the variation of all the relevant operating

parameters are characterized. As such, the effects on the stoichiometric flame temperature,

Tst, and the residence time, τ , within the combustor are addressed. Note that the results of

this pathway analysis are based on instantaneous nitrogen conversion at the flame reactor

exit. Thus the results do not represent an integral temporal development of the reaction

scheme from the ignition to end of the flame.

(a) Natural gas (x=0% H2) (b) Natural gas (x=0% H2)

Figure 5.27: The NOx formation pathways for (a) combustor pressure p and (b) combustor

outlet temperature T . All other operating parameters p, T , x, v, Tair, and ψ

are at reference conditions as listed in Tab. 4.3.


(a) Various fuels (b) Natural gas (x=0% H2)

Figure 5.28: The NOx formation pathways for (a) hydrogen fraction x and (b) relative air

inlet velocity vrel. All other operating parameters p, T , x, v, Tair, and ψ are at

reference conditions as listed in Tab. 4.3.

Referenceconditions

At reference conditions, the thermal NOx pathway is the most important mechanism for

Nx formation, producing nearly half (47%) of the NO, followed by the prompt mechanism

(42%). While the NNH pathway is responsible for 4% of NO production, the N2O pathway

produces about 7%, see Fig. 5.27 (a). The relative contributions to the total emissions are

indicated by bars, while the absolute emissions are given in units of parts per million [ppm].

Pressure When the pressure is increased while holding all other parameters at reference conditions, the

relative contribution of the prompt NO pathway is partially replaced by the thermal and N2O

pathway, while the contribution of the NNH pathway remains constant, see Fig. 5.27 (a). Al-

though all radical concentrations dramatically decrease, as seen earlier, there is an absolute

increase in NOx emissions. In addition to the decrease of all relevant radical concentrations,

the reactivity and thus the stoichiometric flame temperature, Tst, increase with the pressure

promoting NO formation despite the trends of decreasing radical concentration and decreas-

ing the residence time, τ . Overall, the pressure causes an increase of production via the

thermal NOx pathway by 77%, while production via the prompt NOx pathway increases by

33%. While absolute NO formation via the NNH pathway remains constant, NO via the

N2O pathway increases by 120%. In summary, the effect of increasing the flame temperature

is dominant during the pressure increase.


Increasing the combustor outlet temperature mainly causes an increase in the flame residence

time (PSR 1) and a minor increase in the flame temperature. Fig. 5.27 (b) shows the pathway

distribution for the combustor outlet temperature. While the relative contribution of the


thermal pathway increases with increasing combustor outlet temperature, the contributions

of NNH and the N2O pathway remain constant and the contribution of the prompt pathway

decreases. The significant relative increase in the contribution of prompt pathway is in

conjunction with the large increase in the CH◦ radical concentration at reduced combustor

outlet temperatures. However, the total NO formation for all NOx pathways decreases

significantly.

Hydrogenfraction

Increasing the hydrogen fraction in the fuel results in a significant increase in the flame

temperature and a significant increase in the radical concentration of H◦, O◦, and OH◦, while

the residence time remains nearly constant. Consequently, the thermal pathway contribution

increases up to 70% of the total NO production during pure hydrogen fueling, as shown

in Fig. 5.28 (a). In contrast, the shortage of CH◦ during the fuel switch from natural

gas to hydrogen reduces the contribution from the prompt mechanism until its complete

disappearance with pure hydrogen fuel. The prompt mechanism is replaced by the NNH

pathway and, to a lesser extent, by the N2O pathway. The total NO formation of the NNH

and N2O pathways increases by a factor of about 3.5, which is significantly more than the

increase of thermal pathway contribution, which increases by a factor of 2.8.

Air inletvelocity

Varying the air inlet velocity mainly affects the residence time, while the remaining param-

eters, especially the radical concentration and the stoichiometric flame temperature, remain

nearly constant. The NO pathway distribution is shown in Fig. 5.28 (b). With increasing

air inlet velocity, the contribution from the thermal NO pathway is replaced by the prompt

NO pathway, while the contribution from the NNH and N2O pathways remain constant.

The absolute formation of NO, the thermal NO formation, and, to a lesser extent, the N2O

formation are all affected by the residence time decrease, while the prompt and NNH path-

ways remain nearly constant. Further, both the NNH and N2O pathways remain mostly

unaffected by residence time variations since the radical distribution and the stoichiometric

flame temperature are nearly constant in the reactor.


Figure 5.29: The NOx formation pathways for air inlet temperature, Tair. All other operating

parameters p, T , x, vref , and ψ are at reference conditions as listed in Tab. 4.3.

Air inlettemperature

A similar trend can be observed when increasing air inlet temperature, shown in Fig. 5.29.

Although the variation of the air inlet temperature mainly affects the flame temperature,

the residence time also increases with increasing Tair. Increasing the air inlet temperature

leads to an increase in NO formation that can almost completely be attributed to the thermal

pathway. While the absolute NO formation along the prompt, NNH, and N2O pathways pro-

duce nearly the same amounts of NO, NO formation via the thermal NO pathway increases

by about 80% over the given range. Thus, the small influence of the radical concentration

on the thermal pathway that was stated above, can be also seen during air inlet temperature

variation.

Waterinjection

Finally, water injection plays a significant role in the NO pathway distribution. Fig. 5.30 (a)

shows the pathway contributions for natural gas fuel for an increasing water fuel ratio. All

pathways form less NO with increasing water mass flow rates. As such, the thermal NO

pathway dramatically decreases in importance from nearly 50% of NO production down to

less than 10%. This change is compensated for mainly by an increase in the contribution of

the prompt pathway, while the share of NNH and N2O nearly remain constant.

The effects of water injection for pure hydrogen fuel operation is shown in Fig. 5.30 (b).

Prompt NO formation is not included here due to the absence of CH◦ radicals. For hydrogen

fueling, thermal NO formation dramatically reduces with an increasing water mass flow rate,

and this contribution is compensated for by the NNH and the N2O pathways. Thus, mainly

due to the lack of the prompt mechanism, the potential of NO reduction with water injection

is higher at hydrogen fueling than in natural gas fueling.


(a) Natural gas (x=0% H2) (b) Hydrogen (x=100% H2)

Figure 5.30: The NOx formation pathways for different water fuel ratios for (a) natural gas

and (b) hydrogen fuel. All other operating parameters p, T , x, v, and Tair are

at reference conditions as listed in Tab. 4.3.

5.6.3.3 Conclusion

Thermalpathway

The thermal NO pathway domincates NO formation in the experimental non-premixed flame

structure. In terms of absolute NO formation, the other pathways follow the trend of the

thermal pathway with exception of the prompt mechanism, which decreases for increasing

hydrogen levels. The stoichiometric flame temperature has a strong effect on the ability of the

initial thermal pathway reaction GRI 178 to overcome its activation energy. The existence of

O◦ radicals is thus essential, which are theoretically completely consumed at stoichiometric

conditions and full fuel conversion. However, due to dissociation and recombination processes

at high temperatures, the O◦ radical concentration is always sufficient for NO production.

Second, with increasing residence time, the solution is approaching equilibrium, where NO

concentration is dramatically higher. Hydrogen fueling and water injection significantly

affect thermal and temporal properties of the flame and thus dominate the NO formation

trends via the thermal pathway.

Promptpathway

Prompt NO is the second important NO formation pathway, as shown in this study. As

is commonly known and indicated by these results, the relevance of the prompt forma-

tion pathway decreases with increasing flame temperature, Tst. The prompt NO pathway

gains importance at low flame temperatures, in particular at low pressure and low air inlet

temperature. In contrast to the thermal pathway, the prompt pathway shows a significant

dependence on the CH◦ radical concentration for initiating the N2 break-up. As fuel com-


position approaches pure hydrogen, the radical concentration dominates by suppressing and

finally eliminating NO formation via the prompt mechanism. In the case of water injection,

the contribution of the prompt pathway increases up to 77% of the total NO production

during natural gas fueling. During hydrogen fueling, however, the absence of CH◦ eliminates

this dominant pathway. Consequently, the potential to reduce NO emissions via water in-

jection is significantly greater for hydrogen fueling, even though the flame temperature does

not decrease significantly more in the case of hydrogen fueling. The temporal effect on the

prompt NO pathway is minor.

NNHpathway

The NNH mechanism showed the lowest contribution in this study. A significant thermal

or temporal dependence was not observable, indicating the sole dependence on the radical

concentrations distribution. The H◦ radical distribution is most significant and shows a

significant increase for decreasing pressure, as well as an increase due to increased hydrogen

fuel content, and a significant decrease for water injection. However, the NNH pathway is

not affected by the pressure variation, despite the fact that the H◦ radical concentration is

significantly affected. The absolute NO formation via NNH is most relevant in the case of

hydrogen fueling (47 ppm), where this pathway is significantly promoted by the increasing

H◦ radical concentration. The highest relative contribution (22%) is observed for hydrogen

fuel and water injection due to the suppression of the thermal and prompt pathways.

N2Opathway

NO formation via the N2O pathway requires sufficient concentrations of the O◦ radical, as

discussed earlier. Here, formation is supported by a high stoichiometric flame temperature,

as indicated by the following: the absolute formation of NO increases with the pressure, the

combustor outlet temperature, and the hydrogen share, while it nearly remains constant or

decreases for an increase in the air inlet velocity and the inlet temperature. With increasing

water injection, the relative contribution increases while the total NO formation decreases.

Hence, the characteristic of the N2O pathway is similar to the thermal pathway production,

however to a lesser extent.

5.7 Distribution of the effects of H2O on NOx

The chemical influence of water has been identified to be relevant in some operating states of

the combustor. However, the effect has not yet been quantitatively described. This section

quantifies the physical, thermo-physical, and chemical effects of water, shown in Sec. 2.3.3,

on the reduction of the NOx emissions for varying operating parameters.

Method The dilution of the reactor fluid with water is studied by taking advantage of the fact that

the water concentration does not significantly vary when assuming the injected water is

chemically active (H2O) or is inert (iH2O). The maximum deviation is 0.7%. The effect of


water dilution on NOx emissions reduction can be described by using the concentration of

the iH2O according to

NOx

∣∣no dil.

= NOx1

1− ciH2O, (5.20)

since iH2O distinguishes the injected water from the combustion product H2O. The chemical

effect is determined via

NOx

∣∣no dil.,no chem.

= NOx

∣∣no chem.

1

1− ciH2O, (5.21)

where NOx

∣∣no chem.

is the NOx emission value for the iH2O combustion, which includes

dilution. Finally, the thermal effect is indirectly measured by elimination the other two

effects from the global difference. Therefore, the global difference

∆NOx = NOxψ=0 − NOxψ 6=0 , (5.22)

is defined as the difference between emissions in dry and wet conditions and thus represents

the total NO-reducing effect of water injection.

Finally, the distribution of the three influences is given by the following set of equations:

∆NOx

∣∣dilution

= NOx

∣∣no dil.

− NOxψ 6=0 (5.23)

∆NOx

∣∣chemical

= NOx

∣∣no dil.,no chem.

− NOx

∣∣no dil.

(5.24)

∆NOx

∣∣thermal

= ∆NOx −∆NOx

∣∣dilution

−∆NOx

∣∣chemical

. (5.25)

ResultsFig. 5.31 (a) shows the effects of water injection for the reference case. At reference condi-

tions, the thermal effect dominates with about 86% of the total reduction, while the chemical

effect makes up 10%, and the dilution effect makes up 5% of the total reduction in the case

of minor water injection (ψ = 0.25). At a high water injection rate (ψ = 1.5), the thermal

effect gains importance and replaces the dilution effect. In the case of hydrogen fuel, the

thermal effect contribution is initially higher and the contribution from the dilution effect

exceeds that of the chemical effect. At a high water injection rate, the trend is similar to

natural gas fuel. Since the prompt mechanism does not exist for natural gas fuel, the thermal

effect gains significance.


(a) (b)

(c) (d)

(e)

Figure 5.31: The distribution of the thermal, chemical and dilution effects of water on the

NO reduction for (a) reference conditions, (b) reduced pressure (p = 3 bar),

(c) reduced combustor outlet temperature (T=900 ◦C), (d) reduced air inlet

velocity (vrel = 0.6), and (e) reduced air inlet temperature (Tair = 400 ◦C).


At lower pressures (b) and lower combustor outlet temperatures (c), the thermal effect is

slightly reduced, and the fuel switch to hydrogen has basically the same trend as seen at

reference conditions. A lower air inlet velocity (d), however, increases the thermal effect for

both natural gas fueling and hydrogen fueling. A lower air inlet temperature (e) gives results

comparable to those at reference conditions, with a slightly reduced thermal influence.

The quantification of the three effects of water identifies the thermal effect as dominant.

The chemical effect has a contribution on the order of 10% of the total reduction, which

decreases significantly for hydrogen fuel. The contribution of the dilution effect decreases

with an increasing water injection mass flow rate.

5.8 Summary

A simple CRN model for a non-premixed gas turbine combustor flame characterizes the

influence of hydrogen and water injection on the combustion progress. The dependences of

the NOx emissions have been explained on the basis of the internal response of the combustor

state, e.g. the flame temperature, the flame residence time, a radical distribution within the

flame, and a pathway analysis. Special attention is given on the chemical effects of hydrogen

fuel and water injection and their influence on the NOx emissions.

The internal state of the combustor reacts to the variation of the model input data, namely

combustor pressure, combustor outlet temperature, hydrogen fraction of the fuel, combustor

inlet velocity and temperature, and the rate of water injection. The complexity of the applied

combustor does not allow for an individual analysis of parameters such as the effect of the

flame temperature or the residence time, since radical distributions also change.

Among other input variables of the model, the hydrogen fraction and the water injection

significantly affect the radical distribution and thus change the oxidation and NO formation

reaction chains. While H◦, O◦, and OH◦ radicals increase significantly with increasing hydro-

gen levels, CH◦ decreases. The highest concentration increase is that of the H◦ radical (by

nearly a factor of 3), caused by the fact that the H atom is available at higher concentrations

during hydrogen fueling. HO2◦, however, remains nearly at the same levels. Water injection

results in a decrease of all listed radicals. The chemical effect of H2O significantly influences

the distribution of the radical concentrations. The chemical effect of water, verified via the

injection of inert water (iH2O), further decreases the radical concentrations of H◦, O◦, and

CH◦ and increases the radical concentrations of OH◦, and HO2◦. The high collision factor

of H2O in three-body reactions has a minor influence on the radical concentration.

It can been concluded that the O◦ radical distribution is not rate-determining for NO for-

mation via the thermal pathway. However, the CH◦ radical concentration is crucial as the


main driver for the prompt mechanism, which is the pathway that dominates the NO for-

mation at low flame temperatures. The NNH and N2O pathways are of minor importance

and basically follow the trend of thermal pathway. The significantly higher ability to reduce

of NO formation using water injection during hydrogen combustion has been attributed to

the absence of the CH◦ radical which suppresses the prompt pathway completely.

The thermo-physical effect of water injection on the NOx emissions dominates both the

chemical effect and the physical effect (dilution). However, the chemical effect and the

physical effect are not negligible over the entire parameter range. The relative fraction of

NO reduced via the chemical effect is higher when using natural gas than with hydrogen fuel

and decreases with more water injection. The physical effect is the weakest among the three.

Its relative contribution is also higher with natural gas and decreases with the addition of

water injection.

In conclusion, the results of this chapter give relevant insights into the experimental com-

bustor. In the next chapter, the findings of this section with regards to the NOx formation

mechanisms are used to develop a simplified NOx model based on analytical equations de-

scribing to the thermal, temporal and chemical influences.

6 Simplified NOx emissions model

In this chapter, a general NOx correlation is provided, particularly designed for applications

including hydrogen fuel and water injection as a NOx reduction measure. To accomplish

this, existing correlations are parametrized on the basis of the chemical investigation of the

last chapter.

The outline of the chapter is as follows: First, relevant NOx correlations from literature

are introduced and the state-of-the-art is identified. Then the NOx correlation is presented

and parametrized on the basis of the model data. Since NOx correlations contain internal

parameters of the model that are not measured in the experiment, correlations for the res-

idence time τ and the stoichiometric air temperature Tst are developed and parametrized

separately. The validity of the approach is shown by comparison with the experimental NOx

emissions results. Next, the applicability of high pressure NOx predictions solely on the basis

of low pressure experimental data by this approach is evaluated. To test the general applica-

bility of this model, the correlation is finally applied to the results of a second non-premixed

industrial combustor that has been characterized at the test facility.

6.1 NOx correlations review

NOx correlations are transfer functions for efficiently predicting the emission of combustion

applications based on a set of operation parameters. They have been used in the past to

identify the key driving parameters of NOx formation in gas turbine applications. This

review intends to introduce the basic ideas of the correlations and to outline the various

applications. For stationary gas turbines, NOx correlations were developed to verify and to

quantify major influencing parameters that have been identified by theoretical approaches.

In the aviation industry, NOx correlations have been developed to compare the emissions at

high altitudes to those at sea level. NOx correlations can be classified into two main groups.

The first group uses cycle variables and physical correlations throughout the engine. These

are called direct prediction correlations. The second group of correlations are reference-type

equations that use relative variables determined by experiments.

The basic set up of the NOx correlations is similar, however, their coefficients differ signif-

icantly in different applications. The main criterion is the flame type, which distinguishes

non-premixed flames and premixed flames. The interested reader can find relevant non-

premixed correlations at Lefebvre (1984), Rizk & Mongia (1994), Visser & Bahlmann (1994),

and Dopelheuer & Lecht (1998). Correlations for premixed flames were developed by Nicol

122 6 Simplified NOx emissions model

et al. (1993), Malte et al. (1994), Becker & Perkavec (1994), Steele et al. (1998), Biagioli &

Guthe (2007), Lacarelle et al. (2010), and Han et al. (2014).

In the following, a summary of the most important influencing variables including the ar-

rangement and magnitude of their coefficients is given. The main variables for a non-

premixed combustion regime are pressure, flame temperature, and residence time. Addi-

tionally, hydrogen fueling and water injection are relevant for this study.

Pressure The pressure has a significant effect on the reaction rates. Higher pressure typically leads

to higher NOx emissions. In literature, the influence of pressure on the NOx emissions is

described by a power function NOx ∝ pi with a constant exponent i that has been identified

to be within −0.45 < i < 1.5, depending on the underlying correlation and test conditions.

In general, the literature agrees to i = 0.5 (Correa 1993, Shaw 1975) for non-premixed flames,

which can be analytically determined (Joos 2006). Note that the negative pressure exponents

have been found by Biagioli & Guthe (2007), who considered prompt NOx generation only

and by Steele et al. (1998) for a premixed jet stirred flame.

Flametemperature

In hot flame regimes, the thermal NOx pathway is the dominating mechanism (Correa 1993).

Non-premixed flames in particular follow this trend due to the stoichiometric flame temper-

ature in the primary zone. In literature, the dependence of NOx emissions on the flame

temperature is described by a natural exponential function. Here, the temperature is con-

sidered as the activation energy criterion in accordance with the Arrhenius equation (Joos

2006), compare Eq. 5.6.

Residencetime

In non-premixed-dominated regimes, the residence time is crucial for the thermal NO gen-

eration pathway. Rizk & Mongia (1992) performed a residence time study on the basis of

a CRN network resulting a linear dependence of the NOx emissions on the residence time.

This linear trend was later confirmed by Joos (2006).

Fuelflexibility

Correlations that explicitly include fuel variations are limited in literature. Since a variation

of the fuel goes along with a variation in the flame temperature, influences are inherently

captured as long as other chemical influences are minor. For a non-premixed combustor,

Visser & Bahlmann (1994) considered the effect of various natural gas compositions on the

NOx emissions. They found a linear relationship between the lower heating value and the

NOx emissions. At increased hydrogen fractions in the fuel, however, their relation was

found to no longer be valid.

Waterinjection

The first investigation on the dependence of NOx emissions on injected water was done by

Shaw (1975), who considered different humidity levels and corrected the NOx emissions of

aviation gas turbines to standard conditions. Within the frame of water injection develop-

ment, the first NOx emissions correlation for gas turbine applications was reported by Wilkes

& Russell (1978), who conducted water injection experiments with a complete gas turbine


engine. They described the reduction of NOx emissions by injecting water with an exponen-

tial dependence on the combustor outlet temperature. Newburry & Mellor (1995) estimated

the reduction of the non-premixed flame temperature based on a modified reaction equation

including water. They then defined a semi-empirical correlation including a characteristic

kinetic time that decreases for higher flame temperatures and thereby increases the NOx

concentration.

For many years, much research has been done to identify the parameters that influence NOx

emissions. The most important parameters have been identified recently. The next section

presents the correlations relevant to this study.

6.2 Correlations

The dominant parameters for non-premixed flames have been described in the literature

review. These are the pressure, the residence time, and the flame temperature. Since the

residence time, τ , and the flame temperature, Tst, were not measured in the experiment

directly, the CRN model is used to determine their quantities. However, the computa-

tion framework of this study ultimately aims to apply the NOx correlation without prior

calculations of the network model. Therefore, correlations for the residence time and the

stoichiometric air temperature are determined as a preliminary step to allow a quantification

of their value solely based of the experimental data. In a second step, these correlations are

used to finally parametrize the NOx correlation. All correlations are fitted with the linear

optimization tool in MATLAB.

6.2.1 Residence time correlation

The residence time of the fluid in the flame zone is significantly dependent on the combustor

operating parameters with various characteristics. On the basis of these characteristics, the

empirical equation

τ = τref

(p

pref

)r1 (T

Tref

)r2Tair

Tair,ref

vref

v(1 + r3ψn) (6.1)

is defined according to Fig. 5.13 (b) and parametrized on the basis of the network model

results. The residence time, τ , is dependent on the combustor outlet temperature, T , as

defined in Eq. 5.15. Further, the pressure plays a key role as described in Eq. 5.19. The

air inlet temperature, Tair, the combustor velocity, v, and the normalized water fuel ratio,

ψn, also have a significant effect on the residence time due to their effect on the mass flow

of PSR 1. Their linear dependences are also derived from Fig. 5.13 (b). Only the fuel


Figure 6.1: The correlation results of the resi-

dence time, τ . The conditions are

given in Tab. 4.3.

Table 6.1: Parameters and rRMSE of the cor-

relation fitting for the residence

time, τ .

τ correlation

Variable Value

τref 0.38 ms

r1 -0.3354

r2 1.9029

r3 -0.1459

rRMSE 1.03%

composition has a negligible effect on the residence time. Taking the reference residence

time of τref = 0.38 ms that was determined in the CRN (compare Tab. 5.3), the results of

the optimization process are given in Tab. 6.1 and represented by Fig. 6.1. With an accuracy

of rRMSE = 1.03%, the formulation is a sufficient approximation for this application.

The solution parameters show a high degree of robustness against a strategic reduction

of the experimental data. During the strategic reduction, the influence of the fuel has

been successively been eliminated and the response of the other fitting parameter has been

validated. The data subsets used for the robustness study include: dry and wet data only,

dry and wet natural gas data, and reference-pressure-only data for dry and wet conditions.

As a result, the deviation of all fitting parameters is less than 9% for all sets. This low

variation indicates that the resulting parameters well describe the underlying effects without

significant suppression of other influences.


6.2.2 Stoichiometric flame temperature correlation

The semi-empirical correlation for the stoichiometric flame temperature is an absolute corre-

lation and is based on an approximation of the Burke-Schumann approach for non-premixed

flames at stoichiometric conditions

Tst = Tair +−∆hmYf,u

cpνfMf, (6.2)

which is based on a heat balance and basic simplifications, such as adiabatic conditions,

constant pressure, no technical work, neglecting kinetic and potential energy, complete fuel

conversion, and a constant specific enthalpy, cp, at unburnt and burnt conditions (Peters

2000). Here, Y is the initial mass fraction of the unburnt fuel and ∆hm the molar enthalpy

production of a postulated global single reaction with νf as the stoichiometric coefficient of

the fuel.

This physical approach is the basis of the correlation for the stoichiometric flame temper-

ature. Besides the two terms in Eq. 6.2 representing the heat input and chemical heat

conversion, a third term is necessary to capture the heat reduction effect of water injection:

Tst = Tair + ϑ1 exp

(t1

(Hfuel

Hfuel,ref− 1

))(p

pref

)t2− ϑ2ψn. (6.3)

While the thermal heat input is taken directly from Eq. 6.2 (Tair in K), the influence of the

chemical fuel conversion term has been modified in order to capture the underlying influ-

ences. The influence of the chemical fuel conversion on stoichiometric flame temperature

is mainly dependent on the hydrogen fraction, x, in the fuel and the combustor pressure,

p, which have contributions of an equal magnitude as shown in Fig. 5.13 (a). This term

describes both the chemical heat input via the heating value of the fuel and the significant

influence of the pressure on the reactivity as described via the referenced pressure. The first

parameter ϑ1 [K] describes the temperature increase due to the heat output of the chem-

ical conversion of natural gas fuel at reference conditions. The parameter t1 describes the

exponential increase of the temperature during the transition from natural gas to hydrogen.

The reactivity increase caused by the pressure and subsequent increase in the stoichiometric

temperature is described by parameter t2. Finally, the third term describes linear reduction

of the stoichiometric flame temperature via water injection, as seen in Fig. 5.14. The nor-

malized water fuel ratio ψn is used since it refers to the actual mass flow of injected water

independent from the fuel composition. The parameter ϑ2 describes the linear decrease of

Tst scaled to a unit of ψn due to water injection.


Figure 6.2: The correlation results for the st.

flame temperature, Tst. The con-

ditions are given in Tab. 4.3.

Table 6.2: Parameters and rRMSE of the cor-

relation fitting for the stoichiometic

flame temperature, Tst.

Tst correlation

Variable Value

ϑ1 1684.6 K

ϑ2 260.4 K

t1 0.0582

t2 0.0309

rRMSE 5.98%

The results of the optimization process are shown in Fig. 6.2 and given in Tab. 6.2. The

accuracy is rRMSE = 5.98%. Fig. 6.2 indicates that hydrogen admixtures have a minor

deviation from the pure fuels. However, overall, this correlation is sufficient for the develop-

ment of the NOx correlation. The correlation proved to be robust against limitations of the

data. The parameter sets of t and ϑ vary by a maximum of 10% when considering parts of

the entire data set, using the same procedure as the robustness investigation of the residence

time.

6.2.3 NOx correlation

The correlations for residence time and stoichiometric flame temperature allow the inner

combustion parameters to be transfered to a simplified NOx prediction method. The under-

lying correlation

NOx

NOxref

=

(p

pref

)i·(

τ

τref

)· exp

(BTst − Tst,ref

Tst,ref

)· (1 + wψ)−1 (6.4)

is a physics-based formulation adapted from the Arrhenius equation. It is based on the

kinetic approach of the Zeldovich mechanism for the thermal NOx pathway that is dominant

in non-premixed flames. The pressure exponent, i, characterizes the pressure dependence

of the NOx emissions. The PSR 1 residence time, τ , is included due to the non-premixed

nature of the flame. The linear influence is traced back to the definition in Eq. 5.16. The


factor, B, quantifies the exponential nature of the influence of the stoichiometric flame

temperature. The fourth factor is introduced especially for the effect of water injection on

the NOx emissions, which is expressed by a reciprocal function and quantified by the fitting

parameter, w. Thus, the product, wψ, describes the impact of dilution and the chemical

influence of water. It inherently contains the higher potential for NOx emissions reduction

during hydrogen fueling, as described in Sec. 5.5.

Fittingprocess

The parameters of Eq. 6.4 are simultaneously determined by a non-linear optimization in

MATLAB while the correlations of the residence time (Eq. 6.1) and stoichiometric flame

temperature (Eq. 6.3) provide the internal parameters. Thus, the fitting of the underlying

NOx correlation is based on experimental data only. The parameters are determined for the

entire application range of the experimental study.

Totalresults

The resulting parameters for the entire data set are given in the first column of Tab. 6.3. The

pressure exponent, i, is remarkably close to the prevailing theoretical value of literature (0.5)

for a non-premixed flame. The thermal description via Tst is capable of the prediction of the

NOx emissions results during the transition from natural gas to hydrogen by the constant

factor, B. This factor inherently contains the chemical influence of hydrogen. However, it is

not sufficient during the drastic NOx reduction during water injection due to the significant

chemical influence. Therefore, the factor of water injection influence, w, is of a significant

magnitude. The water injection effect has been shown to be best approximated by the

reciprocal function rather than by a linear or exponential approach. It leads to a NOx

reduction factor of 0.51 in the case of natural gas and 0.28 in the case of hydrogen fueling.

Three aspects of the simulation of water injection that are inherently captured by the cor-

relation: the dilution effect, the reduction of significant radicals due to water injection, and

the absence of CH◦ in the case of hydrogen fueling. First, the dilution effect contributes up

to 6% of the total effect of water injection. Also, the absolute dilution is higher in the case

of hydrogen combustion, since the most relevant stoichiometric air mass flow is here reduced

by 20% at reference conditions. Second, the radicals responsible for NOx generation are sig-

Table 6.3: Parameters and rRMSE of the NOx correlation for different data sets.

Parameter total dry, x = 0 dry, x = 1 dry, all x wet, all x

i 0.4313 0.4012 0.4426 0.4313 0.2979

B 5.5015 6.2095 4.9197 5.2841 7.7911

w 0.6471 - - - 0.5194

rRMSE (model) [%] 6.80 15.15 9.37 8.02 7.74

rRMSE (exp) [%] 5.28 6.06 6.92 5.73 2.25


(a) (b)

Figure 6.3: A comparison of numerical and experimental and correlation results for NOx

emissions. The experimental conditions are given in Tab. 4.3.

nificantly reduced by water injection. While the O◦ radical concentration decreases by about

70% for both fuels (NG and H2), the H◦ radical reduces by about 50% for natural gas and

60% for H2 fuel. Note that the chemical effect of water plays an additionally important role

for the significant reduction of the radicals. Third, the difference between natural gas and

hydrogen arising from the lack of CH◦ in hydrogen fuel is described by using ψ instead of ψn.

This is a significant contributor to the production of prompt NOx during the combustion of

natural gas that simply does not exist during hydrogen fueling.

The NOx correlation results are displayed against the numerical results and the experimental

results in Fig. 6.3 (a) and (b) respectively. The correlation describes the model results well,

as seen in (a), with a rRMSE of 6.80%. When comparing the correlation and experimental

results in (b), overprediction the NOx at wet natural gas conditions and underprediction

at wet hydrogen conditions can be observed, as seen in Fig. 5.3 (a). The relative root

mean square error for the fitting with respect to the experimental data is 5.28%. Thus,

the transition of natural gas fuel to hydrogen is captured sufficiently well, even though the

importance of the prompt mechanism decreases significantly down to zero. Since the trend

of the absolute NO formation via the NNH and N2O pathways is similar to the thermal NOx


pathway, both pathways are not independently addressed but contained inherently in the

given correlation.

Partial setresults

The fitting of the correlation with partial set of experimental data allows for an evaluation

of the stability of the fitting parameters. At the same time, the prediction quality can be

increased by applying the correlation to a reduced data set only. The fitting results for the

evaluated parts are given in Tab. 6.3 for dry natural gas fuel, dry hydrogen fuel, and total

dry and total wet conditions. The last two sets include the complete fuel range. At wet

condition, the pressure exponent, i, decreases significantly from the total value while the

exponential factor, B, increases. Thus, the effect of decreasing stoichiometric temperature

gains significance over the pressure effect.

6.3 Pressure estimation evaluation

Until this section, the reference pressure has been 16 bar. Tests at elevated pressure require

high resource consumption and complicated test infrastructure. This section verifies the

method of estimating high pressure NOx emissions on the basis of low pressure combustion

test data.

MethodIn order to evaluate the feasibility of 3 bar as a pressure reference condition, it is assumed

that experimental high pressure data are not available. Thus, the initial PSR 1 volume

Vinit is determined at 3 bar and all numerical results are achieved based on this volume.

The new initial volume is Vinit,3bar = 6.30 · 10−4 m3. It is 31% larger than the value of

the former reference point at 16 bar. The resulting residence time is τref,3bar = 0.59 ms.

Based on the new numerical results, that contain no information about the high pressure

test data, the correlations for the residence time (Eq. 6.1) and flame temperature (Eq. 6.3)

are parametrized and finally, the NOx correlation (Eq. 6.4) is fitted based on the new results,

which are solely based on experimental tests at 3 bar.


Figure 6.4: Comparison of numerical and experimental NOx results for a reference pressure

of 3 bar. The experimental conditions are given in Tab. 4.3.

Comparisonto experi-mentalresults

Fig. 6.4 shows the new CRN model results compared to the experimental data. The compar-

ison criterion between the model results and the experimental results is rRMSE = 8.22% and

thus twice the value of the criterion found from the results with 16 bar reference conditions.

The results show a deviation at highest NOx emissions whose pressure condition is 24 bar

and thus featuring the highest distance from the reference pressure causing this inaccuracy

for both fuels. However, the general effects and water injection are predicted sufficiently.

Results of τand Tst

The fitting results for the residence time and flame temperature correlation are given in

Tab. 6.4. For the residence time correlation, only the pressure exponent r1 is significantly

affected, which is caused by the reference pressure change. For the flame temperature cor-

relation, the temperature difference from unburnt to burnt conditions increases by 180 K,

while the exponential factor of the fuel effect increases and the pressure exponent decreases.

The influence of water injection remains constant. The different reference pressure levels

thus leads to a quantitative shift in the representation of the flame temperature in the corre-

lation. In summary, the accuracy of both correlations do not significantly deviate from the

16 bar reference pressure.


Table 6.4: Results of correlations that fit τ and Tst for 3 bar reference conditions.

τ correlation

Variable Value

τref,3bar 0.59 ms

r1 -0.4101

r2 1.9012

r3 -0.1399

rRMSE 0.92%

Tst correlation

Variable Value

ϑ1 1864.8 K

ϑ2 260.6 K

t1 0.0819

t2 0.0266

rRMSE 6.06%

NOx correlation

Variable Value

i 0.4017

B 5.7892

w 0.5431

rRMSE (model) 6.35%

rRMSE (exp) 11.16%

NOx resultsThe NOx correlation results indicate a similar trend to the 16 bar correlation, as shown in

Tab. 6.4. The fitting parameters are similar in value to the 16 bar reference pressure. The

highest deviation (20%) is for the the water injection parameter w. The comparison of nu-

merical results and the correlation is given in Fig. 6.5 (a) and the comparison of experimental

results and the correlation is given in (b), which both show a notable deviation in the high

pressure regime for hydrogen fuel.

(a) (b)

Figure 6.5: Comparison of numerical and experimental results with the correlation developed

for 3 bar reference conditions. The experimental conditions are given in Tab. 4.3.


Evaluationof feasibility

In general, the reduction of the reference pressure to a lower value does not preclude the

possibility of predicting NOx emission trends. The accuracy deceases at higher pressure

regimes for dry conditions and for the highest NOx emissions. However, the predicted

emissions of water injection as a NOx abatement measure are sufficient and not significantly

affected by the reference pressure decrease.

6.4 Generalization evaluation

The underlying NOx correlations are parametrized on the basis of a real applications. As a

consequence, the parameters are highly dependent on the design of the combustor and thus

the transferability is limited to similar designs. Tsalavoutas et al. (2007) stated that predic-

tions using different NOx correlations that exist in literature may differ from the expected

results by orders of magnitude, when applied with their original parameters. Moreover, on

the same engine but at different operating conditions than used to create the correlation, a

set of parameters may be drastically wrong. The data presented in this study so far are, no

doubt, representative only of the specific combustor design used for the tests. Thus, in the

following, these models (CRN and simplified model) are applied to the experimental results

of a second combustor of a different gas turbine in order for the evaluation of transferability.

Operationfundamen-tals

The second combustor is also a non-premixed industrial gas turbine combustor of a common

size featuring a reverse-flow configuration. The geometric setups of flow guidance, flame

stabilization, and cooling configuration are similar to the former combustor. Its maximum

thermal power is about 60% power of the base combustor, presented in Sec. 4.4. Due to the

geometric similarity, it is appropriate for the evaluation of transferability. This combustor has

been operated with a variety of hydrogen fuel fractions up to pure hydrogen, water injection

ratios, and combustor outlet temperature variations. The constant elevated pressure, the

constant inlet velocity, and the constant air inlet temperature operation, however, limits the

full scope of parameters.

Methods In order to analyze the transferability of the simplified model, the three correlations for resi-

dence time (Eq. 6.1), flame temperature (Eq. 6.3), and NOx correlation (Eq. 6.4), parametrized

in Sec. 6.3, are applied to the experimental data of the new combustor without any further

application of the CRN. The new reference point differs in all parameters from the original

reference point. The required reference residence time is derived using the residence time

correlation applied to the new reference point. Underlying this new reference residence time,

all other residence times are estimated. The flame temperature correlation can be applied

without any further considerations. The reference NOx emission value is taken from the

experimental results.


Figure 6.6: Comparison of correlation results with the experimental results of a second com-

bustor of similar design.

ResultsThe quality of the predictions of NOx emissions is shown via a comparison of the correlation

results with the experimental results in Fig. 6.6. Note that the emission values are related

to the reference values for both axis. While the dry NO emissions data are well matched, a

deviation is observed at wet conditions with the same trend as the former combustor, that

is, overestimation of the emissions from natural gas fuel. This deviation is, however, within

acceptable levels. The rRMSE is 5.29% for the correlation results using reference conditions

at 3 bar.

ConclusionBy applying the simplified model to a second industrial combustor featuring similar geo-

metrical conditions, the transferability of the model has been successfully shown. The NOx

prediction is robust against different application geometries.

6.5 Summary

A simplified NOx model has been established based on the numerical results. The model

design allows it to be applied to similar combustor types without solving the chemical reactor

network model, since all correlation parameters are known from experiments. Unavailable

internal combustor states, namely the residence time and the stoichiometric flame tempera-

ture, are previously determined by correlations and are based on CRN results.

The residence time correlation is bases on the model response analysis. The most influential

variables under dry conditions are the combustor pressure, the combustor outlet temperature,

the air inlet temperature (which lengthen the residence time), and the air inlet velocity


(which shortens the residence time). For wet conditions, the water fuel ratio, normalized

to natural gas fuel, was used as input parameter. The residence time prediction and the

numerical results from the last chapter are in excellent agreement. The flame temperature

correlation is based on the energy balance. It includes the air inlet temperature, the lower

heating values of the fuel, and the pressure. An increase in any of these parameters increases

the flame temperature in dry conditions. The amount of flame cooling is given by the

water fuel ratio normalized to natural gas conditions. Although the comparison between the

prediction and the numerical results shows a small disparity, the correlation is appropriate

for the prediction of NOx emissions.

The NOx correlation is based on the CRN model results presented in the last chapter. The

findings on the chemical effects of water injection are applied to the necessary extent. The

correlation includes the pressure and the results of the two previously mentioned correlations

for the residence time and the flame temperature. For wet conditions, the effect of injected

water is determined by the inclusion of the water fuel ratio. The different effects of natural

gas and hydrogen are captured by the difference in the water fuel ratios for both fuels at

constant water mass flow. The resulting NOx prediction shows a good agreement to the

experimental results. This correlation has been verified to capture the physical phenomena

underlying the NOx formation process.

Furthermore, it has been proven that the current version of the model is capable of predict-

ing NOx emissions based on only low pressure test data. This correlation has an acceptable

limitation at highest NOx emissions, which occur at high pressure and pure hydrogen com-

bustion. This model contributes to the development process of gas turbine combustors by

offering the ability to reduce the number and length of tests for non-premixed combustors.

The correlation was finally applied to another non-premixed gas turbine combustor. Because

this combustor uses a similar combustion concept but a different geometry, it could be shown

that the trend of the NOx emissions is captured by the correlation.

7 Conclusion

Water injection for NOx reduction in high hydrogen non-premixed gas turbine applications

has been verified as a promising NOx reduction measure. Water injection can thus be used

as a potential technology for the transition to carbon-free electricity generation. Due to the

higher reactivity of hydrogen compared to natural gas, the main challenges in using pure

hydrogen fuel in gas turbine combustors are safe operation and NOx emission limits. An

experimental study combined with numerical analysis was performed to investigate the fea-

sibility of hydrogen combustion with water injection and the chemical effects of hydrogen

and water. The numerical model gives insight into the combustor physics, while the exper-

imental combustor is a black box, as is usually the case for high pressure combustion tests

for combustor development. Finally, a simplified model was developed, and its prediction

accuracy and application scope verified.

Experimen-talresults

The feasibility of fuel variation from natural gas to pure hydrogen combustion was demon-

strated in a high pressure, full-scale experimental tests using real gas turbine conditions. The

characteristics of the non-premixed flame combustor were acquired with a focus on the vari-

ation of external combustor parameters that represent full and part load conditions during

gas turbine operation including constant turbine outlet and inlet temperatures. Hydrogen

combustion significantly increases NOx emissions when compared to natural gas operation.

During water injection, the potential of NOx reduction is significantly higher at hydrogen

conditions. Reasonable operating points that limit CO emissions are identified, which gen-

erally require high combustor outlet temperatures.

Numericalresults

In the numerical study, internal combustor states that influence the NOx generation signif-

icantly were analyzed. A simple semi-empirical chemical reactor network has been set up

and evaluated with the experimental data. The model successfully captured all trends of

the experimental parameter studies and proved its feasibility for NOx emissions modeling.

The results quantify the stoichiometric flame temperature and the residence time, which

are both significant in the NOx formation process. A further study of the chemical effects,

especially for variations of the pressure, the hydrogen share in the fuel, and the addition of

water, showed the following results:

High pressures cause a significant reduction of the radical concentrations. This, however,

does not lead to a decrease in NOx formation since the simultaneous increase in the reactivity

increases the flame temperature. At higher pressure, the thermal pathway, prompt pathway,

and N2O pathway increase their contributions to NOx formation. In general, an increasing

flame temperature results in a higher thermal NO production, while the prompt, NNH, and

136 7 Conclusion

N2O pathways do not play an important role. Shortening the flame residence time showed

a significant reduction of the effects of the thermal NO pathway, while the influence of the

other pathways is minor. In general, variation of the residence time has a minor effect on

the radical concentrations.

Hydrogen has a significant chemical effect on the formation of NOx, beyond the NOx forma-

tion increase attributed to the rise in the flame temperature. An increasing percentage of

hydrogen leads to an increase in the concentrations of the H◦, O◦, and OH◦ radicals. The

H◦ radical shows the highest increase since the H◦ atom is more frequent during hydrogen

fueling. In contrast, CH◦ is naturally decreasing, due to the absence of carbon in the fuel,

while HO2◦ remains at the same level. Consequently, during the transition to pure hydrogen

fueling, NOx formation via the thermal, NNH, and N2O pathways increases, while formation

via the prompt pathway ceases.

The thermo-physical effect of H2O can reduce the flame temperature which also abates NOx

formation by blocking the activation energy for the nitrogen break-up. Reducing the flame

temperature reduces the radical concentrations (except for that of HO2◦) which supports the

further suppression of NOx formation. Thus, the relative H◦ concentration is reduced more

during hydrogen combustion (as compared to natural gas), while the relative O◦ and OH◦

concentration reduction is similar during both natural gas and hydrogen combustion. Since

the significance of the thermal pathway decreases during water injection, the transition to

hydrogen leads to a higher potential for NOx reduction because of the absence of CH◦ and

the dominance of the prompt pathway in wet regimes for natural gas fueling. At higher rates

of hydrogen fueling and wet conditions, the N2O pathway also contributes to the total NOx

formation.

The chemical effects of H2O as reaction partner influences the radical concentrations during

water injection. Shown by a comparison to an inert water molecule (iH2O) with the same

thermodynamic properties as H2O, water injection promotes the reverse reaction rates of

important chain branching reactions, and thus H◦, O◦, HO2◦, and CH◦ radical concentrations

are lower compared to inert dilution. For all radicals, the collision efficiency of H2O decreases

the radical concentration. However, this has only a minor effect on the radical concentrations

and thus NOx formation.

The physical effect of water injection (dilution) is the weakest of the three influences. The

relative effect of this phenomenon reaches its maximum importance at low pressure, natural

gas fueling and moderate water injection mass flow rates.

Simplifiedmodelresults

A simplified NOx emissions model based on correlation equations has been formulated and

parametrized with reactor network model results. Two correlations for the residence time

and the flame temperature allow the prediction of these parameters in advance, solely based

on global experimental operating parameters. Based on these two internal combustor op-

7 Conclusion 137

eration variables, a simplified model for NOx emissions has been developed, comprised of

the combustor pressure, residence time, and flame temperature as well as the water-to-fuel

ratio. Besides an adequate prediction of the numerical results for the reference pressure of

16 bar, the ability to fit a correlation at low pressure conditions has been verified. This

opens the possibility of predicting high pressure NOx emissions solely on the basis of low

pressure tests. Furthermore, the simplified model proved to be applicable for combustors

with similar designs. To verify the model’s general applicability, it was successfully applied

to a different non-premixed combustor.

ClosureThe thesis shows the potential of water injection as a NOx reduction measure in hydrogen-

fueled industrial gas turbine combustors at real operating conditions. The presented tools

allow a cost-effective NOx prediction of low and high pressure testing.

A Appendix

Table A.1: Natural gas composition of November 2016 by Thyssengas (2016)

General properties

Hu [kWh/m3] 11.41

Hl [kWh/m3] 10.303

% [kg/m3] 0.7912

Wobbe-Index [kWh/m3] 14.58

Composition

CO2 [vol.%] 1.51

N2 [vol.%] 1.18

CH4 [vol.%] 90.88

C2H6 [vol.%] 5.51

C3H8 [vol.%] 0.65

C4H10 [vol.%] 0.213

C5H12 [vol.%] 0.042

C6+ [vol.%] 0.015

B Previous Publications

Parts of this thesis have already been published by the author in conference proceedings and

journal papers. The complete list of these publication is given here.

02/2014 Kroniger, D., Rudolf, C., Vinnemeier, P., Wirsum, M., Oda, T., Horikawa,

A., 2014, Implementation history of a full-scale high pressure gas turbine

combustor test rig, 15th International Symposium on Transport Phenomena

and Dynamics of Rotating Machinery, ISROMAC-15, TU202.

06/2014 Kroniger, D., Vinnemeier, P., Rudolf, C., Wirsum, M., 2014, High pressure

combustion test rig for 10 MW full scale gas turbine combustors, Proceedings

of ASME Turbo Expo 2014, GT2014-26736.

11/2015 Kroniger, D., Wirsum, M., Horikawa, A., Okada, K., Kazari, M., 2015,

Investigation of the pressure dependence of NOx emissions of an industrial

gas turbine combustor with high hydrogen content fuels, Proceedings of

International Gas Turbine Congress 2015 Tokyo, p. 122-130, Nov. 15-20,

2015.

06/2016 Kroniger, D., Wirsum, M., Horikawa, A., Okada, K., Kazari, M., 2016, NOx

correlation for an industrial 10 MW non-premixed gas turbine combustor

for high hydrogen fuels, Proceedings of ASME Turbo Expo 2016, GT2016-

56189.

06/2017 Kroniger D., Lipperheide, M., Wirsum, M., 2017, Effects of Hydrogen Fu-

eling on NOx Emissions: A Reactor Model Approach for an Industrial

Gas Turbine Combustor, Proceedings of ASME Turbo Expo 2017, GT2017-

64401.

07/2017 Kroniger, D., Stutenkemper, J., Wirsum, M., 2017, Test Facility for Re-

search and Development of Hydrogen Capable Gas Turbine Combustion

Technology, Journal of the Gas Turbine Society of Japan, Volume 45(4),

Pages 42-49.

List of Figures

4.1 Flow diagram of the test rig. . . . . . . . . . . . . . . . . . . . . . . . . . . . 25

4.2 Overview of the combustion system and the exhaust measuring section, in-

cluding the gas analysis system. . . . . . . . . . . . . . . . . . . . . . . . . . 28

4.3 Temperature probe. Left: Temperature probe and gas sampling probe mounted

inside the measuring section. Right: Detail of a single thermocouple integration. 33

4.4 Combustor drawing courtesy of Kawasaki Heavy Industries, Ltd. . . . . . . 37

4.5 Comparison of AFR and AFRexh for all operation points. . . . . . . . . . . . 42

4.6 Saturation temperature Tsat for (a) different combustor outlet temperatures

and sets of pressures and fuels, (b) different air inlet temperatures and sets of

pressures and fuels, (c) different relative air inlet velocities and a set of fuels. 43

4.7 Saturation temperature Tsat over the water fuel ratio for (a) sets of fuels

and air inlet temperatures, (b) sets of fuels, pressures and combustor outlet

temperatures. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 44

4.8 Verification of sample gas temperature exceeding the saturation temperature

for all operation points. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 45

4.9 Pressure loss ratio against the relative combustor velocity for a set of com-

bustor outlet temperatures. . . . . . . . . . . . . . . . . . . . . . . . . . . . 46

4.10 NOx emissions as a function of combustor pressure for a set of combustor

outlet temperatures and (a) natural gas fuel, (b) 80 vol.% H2 and (c) 100 vol.%

H2. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 47

4.11 NOx emissions as a function of air inlet temperatures for a set of fuel compo-

sitions and pressures. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 48

4.12 NOx emissions as a function of hydrogen content in the fuel for a set of com-

bustor outlet temperatures. . . . . . . . . . . . . . . . . . . . . . . . . . . . 49

4.13 NOx emissions as a function of air inlet temperatures for a set of fuel compo-

sitions and pressures. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 50

4.14 NOx emissions as a function of relative air inlet velocity for a set of fuel

compositions. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 51

4.15 NOx emissions as a function of water injection for a set of fuel compositions,

air inlet temperatures and pressures. . . . . . . . . . . . . . . . . . . . . . . 52

4.16 CO and NOx emissions as a function of the combustor outlet temperature. . 54

4.17 CO and NOx emissions as a function of the air inlet temperature. . . . . . . 55

4.18 CO and NOx emissions as a function of the hydrogen fraction x within the fuel. 56

4.19 CO and NOx emissions as a function of the water injection ψ at 16 bar. . . . 57

List of Figures 141

5.1 Types of ideal reactors. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 63

5.2 Generic chemical reactor network for the non-premixed combustor. . . . . . 72

5.3 Comparison of numerical and experimental results of NOx emissions for (a) GRI 3.0,

(b) Aramco, (c) Konnov and (d) San Diego mechanism. The experimental

conditions are given in Tab. 4.3. . . . . . . . . . . . . . . . . . . . . . . . . . 79

5.4 Comparison of numerical and experimental results for the mfuel, residual O2,

and CO2 emissions for GRI 3.0. The experimental conditions are given in

Tab. 4.3. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 80

5.5 NOx emissions as a function of the pressure, p, for a set of combustor outlet

temperatures, T , at (a) 0 vol.% H2 and (b) 100 vol.% H2. . . . . . . . . . . . 82

5.6 NOx emissions as a function of the hydrogen content, x, for (a) a set of

combustor outlet temperatures, T , at 16 bar and (b) a set of pressures, p, at

T = 1300◦C. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 82

5.7 NOx emissions as a function of the air inlet temperature, Tair, at p = 16 bar

for a set of combustor outlet temperatures, T , at (a) 0% and (b) 100 vol.%

H2 and at T = 1300◦C for a set of pressures, p, at (c) 0 vol.% H2 and (d)

100 vol.% H2. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 84

5.8 NOx emissions as a function of the relative air inlet velocity, v/vref , at p =

16 bar for a set of combustor outlet temperatures, T , at (a) 0 vol.% H2 and

(b) 100 vol.% H2. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 85

5.9 NOx emissions as a function of the relative air inlet velocity, v/vref , at p =

16 bar and T = 1300 ◦C for a set of combustor inlet temperatures, Tair, at

(a) 0 vol.% H2 and (b) 100 vol.% H2. . . . . . . . . . . . . . . . . . . . . . . 85

5.10 NOx emissions as a function of the water fuel ratio, ψ, for a set of hydrogen

fractions x at p = 16 bar and (a) T = 1300 ◦C and (b) T = 900 ◦C. . . . . . 86

5.11 NOx emissions as a function of the water fuel ratio, ψ, for a set of pressures,

p, at T = 1300 ◦C and (a) 0 vol.% H2 and (b) 100 vol.% H2. . . . . . . . . . 87

5.12 NOx emissions as a function of the water fuel ratio, ψ, for a set of combustor

inlet temperatures, Tair, at p = 16 bar and T = 1300 ◦C and (a) 0 vol.% H2

and (b) 100 vol.% H2. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 88

5.13 Model response to varying model input parameters p, T , x, vrel, Tair for (a) the

relative stoichiometric air temperature, Tst/Tst,ref , and (b) the PSR 1 residence

time, τ/τref . The diagrams are comprised of five parameter variations, where

only single parameters are varied to show their individual influence on Tst and

all other operating conditions explicitly remain constant. The table assigns

the line styles to the parameters and indicates which parameters correspond

to each x-axis. The reference conditions are given in Tab. 4.3. . . . . . . . . 90

142 List of Figures

5.14 Response of the model NOx emissions, flame temperature, Tst, and residence

time, τ , to changes in the water fuel ratio, ψn, and hydrogen fuel content, x. 91

5.15 Sensitivity coefficients σ for NO formation in different hydrogen admixtures. 94

5.16 Sensitivity coefficients σ for NO formation with natural gas fuel and different

water injection levels. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 95

5.17 The H◦ net production rate for significant reactions at the reference parame-

ters for changing (a) hydrogen fraction of the fuel and (b) water injection in

the case of natural gas fuel. All other operating parameters p, T , vref , and

Tair are at reference conditions according to Tab. 4.3. . . . . . . . . . . . . . 97

5.18 The H◦ concentration at reference point conditions and selected variations of


and (b) shows results for water injection at reference point conditions (Tab. 4.3). 98

5.19 The O◦ net production rate for significant reactions at the reference parame-

ters for changing (a) hydrogen fraction of the fuel and (b) water injection in



5.20 The O◦ concentration at reference point conditions and selected variations of


and (b) shows results for water injection at reference point conditions (Tab. 4.3).101

5.21 The OH◦ net production rate for significant reactions at the reference param-




5.22 The OH◦ concentration at reference point conditions and selected variations

of p, T , vrel, and Tair: (a) shows results for varying hydrogen fractions in the

fuel and (b) shows results for water injection at reference point conditions

(Tab. 4.3). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 103

5.23 The HO2◦ net production rate for significant reactions at the reference pa-

rameters for changing (a) hydrogen fraction of the fuel and (b) water injection

in the case of natural gas fuel. All other operating parameters p, T , vref , and


5.24 The HO2◦ concentration at reference point conditions and selected variations



(Tab. 4.3). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 106

List of Figures 143

5.25 The CH◦ net production rate for significant reactions at the reference param-




5.26 The CH◦ concentration at reference point conditions and selected variations



(Tab. 4.3). . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 108

5.27 The NOx formation pathways for (a) combustor pressure p and (b) combustor

outlet temperature T . All other operating parameters p, T , x, v, Tair, and ψ

are at reference conditions as listed in Tab. 4.3. . . . . . . . . . . . . . . . . 111

5.28 The NOx formation pathways for (a) hydrogen fraction x and (b) relative air

inlet velocity vrel. All other operating parameters p, T , x, v, Tair, and ψ are

at reference conditions as listed in Tab. 4.3. . . . . . . . . . . . . . . . . . . 112

5.29 The NOx formation pathways for air inlet temperature, Tair. All other oper-

ating parameters p, T , x, vref , and ψ are at reference conditions as listed in

Tab. 4.3. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 114

5.30 The NOx formation pathways for different water fuel ratios for (a) natural gas

and (b) hydrogen fuel. All other operating parameters p, T , x, v, and Tair are

at reference conditions as listed in Tab. 4.3. . . . . . . . . . . . . . . . . . . 115

5.31 The distribution of the thermal, chemical and dilution effects of water on the

NO reduction for (a) reference conditions, (b) reduced pressure (p = 3 bar),

(c) reduced combustor outlet temperature (T=900 ◦C), (d) reduced air inlet

velocity (vrel = 0.6), and (e) reduced air inlet temperature (Tair = 400 ◦C). . 118

6.1 The correlation results of the residence time, τ . The conditions are given in

Tab. 4.3. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 124

6.2 The correlation results for the st. flame temperature, Tst. The conditions are

given in Tab. 4.3. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 126

6.3 A comparison of numerical and experimental and correlation results for NOx

emissions. The experimental conditions are given in Tab. 4.3. . . . . . . . . . 128

6.4 Comparison of numerical and experimental NOx results for a reference pres-

sure of 3 bar. The experimental conditions are given in Tab. 4.3. . . . . . . . 130

6.5 Comparison of numerical and experimental results with the correlation devel-

oped for 3 bar reference conditions. The experimental conditions are given in

Tab. 4.3. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 131

6.6 Comparison of correlation results with the experimental results of a second

combustor of similar design. . . . . . . . . . . . . . . . . . . . . . . . . . . . 133

List of Tables

2.1 Fuel properties of hydrogen, natural gas, and methane. . . . . . . . . . . . . 8

2.2 State-of-the-art gas turbine technology for hydrogen . . . . . . . . . . . . . . 12

4.1 Overview of high pressure combustion test rigs for gas turbine combustors. . 24

4.2 Summary of underlying measurement accuracies . . . . . . . . . . . . . . . . 31

4.3 Operational parameter ranges and reference point. . . . . . . . . . . . . . . . 38

5.1 Reaction mechanisms for numerical study. . . . . . . . . . . . . . . . . . . . 66

5.2 Input parameters and model results . . . . . . . . . . . . . . . . . . . . . . . 71

5.3 Flame volume, flame and total residence time, and stoichiometric temperature

at reference conditions* . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 74

5.4 Comparison of the underlying reaction schemes for NOx and CO2 emissions,

residual O2, and fuel mass flow rate. . . . . . . . . . . . . . . . . . . . . . . 78

6.1 Parameters and rRMSE of the correlation fitting for the residence time, τ . . 124

6.2 Parameters and rRMSE of the correlation fitting for the stoichiometic flame

temperature, Tst. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 126

6.3 Parameters and rRMSE of the NOx correlation for different data sets. . . . . 127

6.4 Results of correlations that fit τ and Tst for 3 bar reference conditions. . . . 131

A.1 Natural gas composition of November 2016 by Thyssengas (2016) . . . . . . 138

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