reduction of combustion noise and instabilities a...

REDUCTION OF COMBUSTION NOISE AND INSTABILITIES

USING POROUS INERT MATERIAL WITH

A SWIRL-STABILIZED BURNER

by

DANIEL SEQUERA

A DISSERTATION

Submitted in partial fulfillment of the requirements for the degree of Doctor of Philosophy

in the Department of Mechanical Engineering in the Graduate School of

The University of Alabama

TUSCALOOSA, ALABAMA

2011

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ABSTRACT

Combustion instabilities represent a major problem during operation of power generation

systems that can lead to costly shutdown. Combustion instabilities are self excited large

amplitude pressure oscillations caused by the coupling of unsteady heat release and acoustic

modes of the combustor. These oscillations cause fluctuating mechanical loads and fluctuating

heat transfer that can result in catastrophic premature failure of components. Combustion noise, a

significant source of noise in gas turbines, can lead to combustion instabilities. Combustion noise

and instabilities are different phenomena; however, they both occur due to unsteady heat release

of turbulent flames that excites acoustic modes of the combustor. The instabilities self excite

when flame adds energy to the acoustic field at a faster rate than it can dissipate it. Swirl-

stabilized combustion and porous inert medium (PIM) combustion are two methods that have

extensively been used, although independently, for flame stabilization. In this study, the two

concepts are combined so that PIM serves as a passive device to mitigate combustion noise and

instabilities. A PIM insert is placed within the lean premixed, swirl-stabilized combustor to

affect the turbulent flow field reducing combustion noise. This study is the first step for eventual

implementation in liquid fuel systems. After presenting the concept, a numerical investigation of

the changes in the mean flow field caused by the PIM is presented. Changes in the flow field can

be beneficial for noise reduction by optimizing the geometric parameters of the PIM. Next,

atmospheric pressure experiments were conducted at low reactant inlet velocity (<10 m/s) and

low reactant inlet temperature (<120 °C) to investigate effect of PIM parameters on sound

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pressure level (SPL), and CO and NOx emissions. Surface and interior combustion modes were

identified and PIM geometric parameters were optimized. Next, a laboratory facility to conduct

experiments at high reactant inlet velocity, high inlet air temperature, and high pressure was

designed and developed. Results show that the porous insert substantially reduces combustion

noise for a range of operating conditions. Moreover, experiments show that the porous insert can

mitigate combustion instabilities without adversely affecting CO and NOx emissions.

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DEDICATION

This dissertation is dedicated to all my family, particularly to my parents, Yelitza and

Edgar, and my brothers, Axzel and Reinaldo.

v

LIST OF ABREVIATIONS AND SYMBOLS

A Model constant

AA Atomizing air

B Bias uncertainty �̃ Mean reaction progress variable

CD Turbulent length scale constant

C0 Model constant

C1 Model constant

CO Carbon monoxide

FFT Fast Fourier transform

FPGA Field-programmable gate array

HfC Hafnium Carbide

ID Inside diameter

k Turbulent kinetic energy

keff Effective thermal conductivity

Li ith acoustic energy loss process

LFE Laminar flow element

LPM Lean premixed

lpm Liters per minute

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lnpm Normal liters per minute

l t Turbulent length scale

n Number of products

NOx Nitrogen oxides

NG Natural gas

OD Outside diameter

PIM Porous inert medium

P Random uncertainty

Prms Root mean square of pressure

Pchamber Pressure inside enclosure

Pinlet Pressure at inlet

Pref Reference pressure

ppm Parts per million

p’ Combustor pressure oscillations

Q Combustion air flow rate

Qc Cooling air flow rate

q’ Heat addition oscillations

Re Reynolds number

RNG Renormalization group

RT Real time

Sc Mean reaction rate

Sct Turbulent Schmidt number

Si Momentum sink term

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SiC Silicon Carbide

slpm Standard liters per minute

SPL Sound pressure level

T Period of oscillations

Ti Inlet temperature

U Velocity

u’ RMS velocity

Ui Overall uncertainty

Ul Laminar flame speed

Ut Turbulent flame speed

V Combustor volume �� Velocity vector

Yi Mass fraction of product species i

Yi, eq Equilibrium mass fraction of product species i � Thermal diffusivity of unburnt mixture � Turbulence dissipation rate

Ф Equivalence ratio

µt Turbulent viscosity

µeff Effective viscosity � Density

ρu Density of unburnt mixture

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ACKNOWLEDGMENTS

I would like to take this opportunity to show my appreciation to everyone that directly and

indirectly had a contribution to make this dissertation possible.

First and foremost, I want to express my most sincere appreciation to my academic

advisor and friend, Dr. Ajay K. Agrawal, whose guidance throughout my studies made my

experience in Tuscaloosa one that I will always cherish. His technical knowledge and managerial

skills, patience and professional attitude towards any situation, even the most difficult ones,

taught me invaluable lessons and greatly influenced my professional growth. Moreover, his

ability to be demanding yet appreciative, his opportunistic advice, inside and outside the

academic environment, made my graduate studies absolutely enjoyable.

I want to gratefully thank all members of my committee, Dr. Baker, Dr. Olcmen, Dr.

Taylor and Dr. Wiest, excellent engineers I am privileged to have had as professors.

Next, I want to express my special gratitude to fellow graduate students I had the fortune

to go to class and share with. Heena, Pankaj, Ben, Troy, Tanisha, Justin, Lulin, Cristina, Cosmin,

and Vijay were always helpful, supportive and fun to enjoy conversation with over a cup of

coffee. I also want to express particular gratitude to Zach, whom I had the chance to work with in

the final stages of my studies. Completion of this investigation was only possible with his

unconditional and assertive support. I also want to express my gratitude to all staff members in

the Mechanical Engineering Department, whose dedication and professionalism make possible

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for students to successfully complete academic careers at UA. Lynn, Pamelia, Betsy, Lisa, Barry,

Ken, Jim, Sam, James, thank you all very much.

I also want to thank my good friends Paulo, Amanda, Jose, Troy, my cousin Miguel and

my brothers Axzel and Reinaldo for making my stay in Tuscaloosa unforgettable, for always

being supportive, encouraging, and fun to be around. Needless to say, I will be forever grateful to

my parents for all the support, guidance, help and unconditional love during my time in UA.

x

CONTENTS

ABSTRACT ...................................................................................................................... ii

DEDICATION ................................................................................................................. iv

LIST OF ABREVIATIONS AND SYMBOLS .................................................................. v

ACKNOWLEDGMENTS .............................................................................................. viii

LIST OF TABLES ......................................................................................................... xiv

LIST OF FIGURES ........................................................................................................ xvi

1. INTRODUCTION ....................................................................................................... 1

1.1 Background ............................................................................................................ 1

1.2 Overview ................................................................................................................ 4

2. NUMERICAL SIMULATIONS OF SWIRL STABILIZED COMBUSTION COUPLED WITH POROUS INERT MEDIUM .......................................................... 9

2.1 Background ............................................................................................................ 9

2.2 Physical Model ..................................................................................................... 11

2.2.1 Governing Equations ................................................................................... 11

2.2.2 Combustion Model ....................................................................................... 12

2.2.3 Boundary Conditions ................................................................................... 14

2.2.4 Model Validation ......................................................................................... 15

2.3 Results and Discussion ........................................................................................ 16

2.3.1 Non-Reacting Flow ...................................................................................... 16

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2.3.2 Reacting Flow ............................................................................................. 17

2.4 Conclusions .......................................................................................................... 19

3. NOISE REDUCTION IN SWIRL-STABILIZED COMBUSTOR COUPLED WITH PIM ..................................................................... 38

3.1 Background ......................................................................................................... 38

3.2 Experimental Setup .............................................................................................. 40

3.3 Results and Discussion ........................................................................................ 42

3.3.1 Effect of PIM Pore Density .......................................................................... 44

3.3.2 Effect of PIM Geometry ............................................................................... 46

3.3.3 Effect of Reactant Flow Rate ........................................................................ 47

3.3.4 CO and NOx Emissions ................................................................................ 49

3.3.5 Long Duration Experiments ......................................................................... 50

3.4 Conclusions .......................................................................................................... 51

4. DEVELOPMENT OF A FACILITY FOR HIGH FLOW RATE, HIGH INLET TEMPERATURE, AND HIGH PRESSURE COMBUSTION EXPERIMENTS ........ 78

4.1 Background ......................................................................................................... 78

4.2 Reactant Supply Systems ..................................................................................... 80

4.2.1 Air Lines ...................................................................................................... 80

4.2.2 Electric Heater ............................................................................................ 82

4.2.3 Fuel Line ..................................................................................................... 82

4.2.4 Product Exhaust Line .................................................................................. 83

4.3 Instruments and Data Acquisition System ............................................................ 84

4.4 Combustion Experimental Apparatus ................................................................... 87

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5. REDUCTION OF COMBUSTION NOISE AND INSTABILITIES WITH THE USE OF POROUS INERT MATERIAL .............................................. 120

5.1 Background ....................................................................................................... 120

5.2 Experimental Setup ............................................................................................ 122

5.3 Results and Discussion ...................................................................................... 126

5.3.1 Open Top Experiments ............................................................................... 130

a. Effect of Pore Density ......................................................................... 130

b. Effect of Flow Rate ............................................................................... 134

5.3.2 Restricted Top Experiments........................................................................ 137

a. Effect of PIM on Noise at P = 1 atm .................................................... 137

b. Effect of PIM on Noise at P = 2 atm ..................................................... 143

5.4 Conclusions ........................................................................................................ 145

6. CONCLUSIONS AND RECOMMENDATIONS .................................................... 220

6.1 Conclusions ....................................................................................................... 220

6.2 Recommendations .............................................................................................. 222

REFERENCES .............................................................................................................. 224

APPENDIX A COMBUSTION PERFORMANCE OF LIQUID BIO-FUELS IN A SWIRL STABILIZED BURNER ...................................... 229

APPENDIX B CALCULATION OF SWIRL NUMBER ............................................... 258

APPENDIX C CALCULATION OF AIR FLOW RATE IN LFE .................................. 260

APPENDIX D SAMPLE CALCULATIONS OF O2 AND CO2 CONCENTRATIONS .............................................................................. 263

APPENDIX E FLOW VELOCITY AND REYNOLDS NUMBER CALCULATIONS........................................................................................ 266

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APPENDIX F SOUND PRESSURE LEVEL CALCULATION SCRIPT ............................................................................................. 268

UNCERTAINTY ANALYSIS ....................................................................................... 274

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LIST OF TABLES

3.1 Summary of results, effect of pore density, Q = 300 slpm ..................................... 45

3.2 Summary of results, effect of geometry, Q = 300 slpm ........................................ 48

3.3 Summary of results, effect of flow rate ................................................................. 49

3.4 Summary of SPL for long-duration experiment .................................................... 51

4.1 Air supply line parts ............................................................................................. 81

4.2 Fuel supply line parts............................................................................................ 84

4.3 List of instruments for flow measurement ............................................................. 85

5.1 Effect of microphone location on SPL for open top experiments, Q = 1020 slpm, Ф = 0.7, Tinlet = 20 °C ................................................................ 128

5.2 Effect of microphone location on SPL for restricted top experiments, Q = 1020 slpm, Ф = 0.7, Tinlet = 20 °C, Qc = 990 slpm ........................................ 129

5.3 Summary of sound pressure levels for Q = 1020 slpm ........................................ 133

5.4 Summary of pressure measurements for Q = 1020 slpm...................................... 134

5.5 Summary of sound pressure levels for Q = 1400 slpm ........................................ 136

5.6 Summary of pressure measurements for Q = 1400 slpm...................................... 137

5.7 Summary of jet noise total SPL, Q = 1020 slpm, P = 1 atm ................................. 141

5.8 Summary of combustion noise total SPL, Q = 1020 slpm, P = 1 atm................... 142

5.9 Summary of pressure measurements for restricted top experiments, Q = 1020 slpm, P = 1 atm ........................................................ 143

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5.10 Summary of Combustion Noise total SPL, Q = 2040 slpm, P = 2 atm ................. 144

5.11 Summary of pressure measurements for restricted top experiments, Q = 2040 slpm, P = 2 atm ........................................................ 145

A.1 NREL biooil characteristics ................................................................................ 233

A.2 Experimental fuel blends (Vol%) ........................................................................ 233

A.3 Water contents in the fuel blend.......................................................................... 234

C.1 Calibration coefficients for air flow rate calculation............................................ 261

D.1 Summary of O2 and CO2 calculated and experimental results ............................. 265

E.1 Summary of flow velocity and Reynolds number calculations ............................ 267

G.1 Readings for air and fuel random uncertainty calculation, low pressure facility ........................................................................................... 277

G.2 Readings for air and fuel random uncertainty calculation, high pressure facility .......................................................................................... 278

G.3 Readings for pressure random uncertainty calculation, high pressure facility .......................................................................................... 279

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LIST OF FIGURES

1.1 Schematic diagram of swirl stabilization mechanism ............................................. 7

1.2 Proposed concepts .................................................................................................. 8

2.1 Schematic diagram of swirl stabilization mechanism ........................................... 21

2.2 Schematic of combustor with the swirler .............................................................. 22

2.3 Computational domain ......................................................................................... 23

2.4 Axial velocity profile at z = 20 mm, methane flame, Φ = 0.58 .............................. 23

2.5 Velocity vectors for non-reacting flow. (a) Experimental results (Wicksall, 2005), (b) Computed results................................................................. 24

2.6 Velocity vectors for reacting flow. (a) Experimental results (Wicksall, 2005), (b) Computed results................................................................. 25

2.7 Velocity vectors for non-reacting flow. (a) without PIM, (b) with PIM ................. 26

2.8 Axial velocity profiles at different axial locations for non-reacting flow: (a) z =10 mm, (b) 20 = mm, (c) 30 = mm ........................... 27

2.9 Swirl velocity profiles at different axial locations for non-reacting flow: (a) z =10 mm, (b) 20 = mm, (c) 30 = mm ........................... 28

2.10 Radial velocity profiles at different axial locations for non-reacting flow: (a) z =10 mm, (b) 20 = mm, (c) 30 = mm ........................... 29

2.11 Velocity vectors for reacting flow Ф = 0.58. (a) without PIM, (b) with PIM ......... 30

2.12 Axial velocity profiles at different axial locations for Ф = 0.58: (a) z =10 mm, (b) 20 = mm, (c) 30 = mm ........................................ 31

2.13 Swirl velocity profiles at different axial locations for Ф = 0.58: (a) z =10 mm, (b) 20 = mm, (c) 30 = mm ........................................ 32

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2.14 Radial velocity profiles at different axial locations for Ф = 0.58: (a) z =10 mm, (b) 20 = mm, (c) 30 = mm ........................................ 33

2.15 Velocity vectors for reacting flow Ф = 0.85. (a) without PIM, (b) with PIM ......... 34

2.16 Axial velocity profiles at different axial locations for Ф = 0.85: (a) z =10 mm, (b) 20 = mm, (c) 30 = mm ........................................ 35

2.17 Swirl velocity profiles at different axial locations for Ф = 0.85: (a) z =10 mm, (b) 20 = mm, (c) 30 = mm ........................................ 36

2.18 Radial velocity profiles at different axial locations for Ф = 0.85: (a) z =10 mm, (b) 20 = mm, (c) 30 = mm ........................................ 37

3.1 Schematic diagram of experimental setup ............................................................. 52

3.2 Photos of PIM inserts (a) PIM insert (b) combustor without PIM (c) combustor with two PIM pieces ...................................................................... 53

3.3 Description and schematic diagram of PIM configurations used in this study........ 54

3.4 Flame images, (a) without PIM (b) with PIM interior combustion (c) with PIM surface combustion .......................................................................... 55

3.5 Schematic diagram illustrating the PIM stabilization mechanism .......................... 56

3.6 One third octave band SPL for repeatability test ................................................... 57

3.7 Flame images for Q = 300 slpm, Ф = 0.7 (a) Configuration A (b) Configuration B (c) Configuration C (d) Configuration D (e) Configuration E (f) Configuration F (g) Configuration G (h) Configuration h (i) Configuration I ................................................................. 58

3.8 Flame images for Q = 300 slpm, Ф = 0.8 (a) Configuration A (b) Configuration B (c) Configuration C (d) Configuration D (e) Configuration E (f) Configuration F (g) Configuration G (h) Configuration h (i) Configuration I ................................................................. 59

3.9 Power spectra for Q = 300 slpm, Ф = 0.7 (a) Configuration A (b) Configuration B (c) Configuration C (d) Configuration D (e) Configuration E (f) Configuration F (g) Configuration G (h) Configuration h (i) Configuration I ................................................................. 60

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3.10 Power spectra for Q = 300 slpm, Ф = 0.8 (a) Configuration A (b) Configuration B (c) Configuration C (d) Configuration D (e) Configuration E (f) Configuration F (g) Configuration G (h) Configuration h (i) Configuration I ................................................................. 61

3.11 One third octave band SPL, effect of pore density, Q = 300 slpm (a) Ф = 0.7 (b) Ф = 0.8 ......................................................................................... 62

3.12 One third octave band SPL, effect of geometry, Q = 300 slpm (a) Ф = 0.7 (b) Ф = 0.8 ......................................................................................... 63

3.13 Flame images for Q = 300 slpm, Ф = 0.7 (a) Configuration A (b) Configuration D (c) Configuration G (d) Configuration I ................................ 64




3.17 Power spectra for Q = 600 slpm, Ф = 0.7 (a) Configuration A (no PIM) (b) Configuration D (divergent) (c) Configuration G (constant) (d) Configuration I (convergent) ........................... 68

3.18 Power spectra for Q = 600 slpm, Ф = 0.8 (a) Configuration A (no PIM) (b) Configuration D (divergent) (c) Configuration G (constant) (d) Configuration I (convergent) ........................... 69

3.19 One third octave band SPL, effect of reactants flow rate, Q = 300 slpm (a) Ф = 0.7 (b) Ф = 0.8 ................................................................... 70

3.20 One third octave band SPL, effect of reactants flow rate, Q = 600 slpm (a) Ф = 0.7 (b) Ф = 0.8 ................................................................... 71

3.21 CO and NOx emissions for Q = 300 slpm, Ф = 0.7, Ti = 100 °C (a) CO (b) NOx ................................................................................. 72



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3.25 One third octave band SPL, long duration test ...................................................... 76

3.26 CO and NOx emissions for Q = 600 slpm, Ф = 0.7, long duration test (a) CO (b) NOx ........................................................................ 77

4.1 General schematic of high pressure combustion laboratory ................................... 91

4.2 Layout of air flow control system ......................................................................... 92

4.3 Layout of high pressure combustion laboratory .................................................... 93

4.4 Layout of high pressure combustion laboratory .................................................... 94

4.5 (a) Photographic image of combustion air pre-heater (b) Schematic diagram of combustion air pre-heater ............................................. 95

4.6 Heater stand ......................................................................................................... 96

4.7 Fuel station ........................................................................................................... 97

4.8 Layout of fuel flow control system ....................................................................... 98

4.9 Exhaust side view ................................................................................................. 99

4.10 Exhaust overhead view ....................................................................................... 100

4.11 Exhaust front view.............................................................................................. 101

4.12 CompactRIO system ........................................................................................... 102

4.13 Sensor/controller and CompactRIO layout .......................................................... 103

4.14 Schematic of assembled experimental apparatus ................................................. 104

4.15 Exploded view of experimental apparatus ........................................................... 105

4.16 Photographic image of experimental apparatus ................................................... 106



4.19 Details of assembled plenum base ...................................................................... 109

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4.20 Details of support pipe/flange ............................................................................. 110

4.21 Details of plenum base ....................................................................................... 111

4.22 Details of enclosure ............................................................................................ 112

4.23 Details of faces of enclosure ............................................................................... 113

4.24 Details of cross section of enclosure ................................................................... 114

4.25 Details of windows on enclosure ........................................................................ 115

4.26 Details of ports on enclosure............................................................................... 116

4.27 Details of window covers ................................................................................... 117

4.28 Details of windows ............................................................................................. 118

4.29 Schematic diagram and photograph of sampling probe ....................................... 119

5.1 Schematic of experimental setup ........................................................................ 147

5.2 Photograph of fuel station ................................................................................... 148

5.3 Schematic of combustion chamber ..................................................................... 149

5.4 Swirler ............................................................................................................... 150

5.5 Schematic of experimental setup ........................................................................ 151

5.6 Schematic diagram of PIM ................................................................................. 152

5.7 Schematic diagram and photograph of sampling probe ....................................... 153

5.8 Schematic of PIM stabilization mechanism......................................................... 154

5.9 Microphone locations for open top experiments .................................................. 155

5.10 One third octave band SPL for repeatability test ................................................ 156

5.11 Effect of probe position on SPL for open top experiments, Q = 1020 slpm, Ф = 0.7, Tinlet = 20 °C ................................................................ 157

5.12 Microphone locations for restricted top experiments ........................................... 158

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5.13 Effect of microphone location on SPL for restricted top experiments, Q = 1020 slpm, Ф = 0.7, Tinlet = 20 °C, Qc = 990 slpm ........................................ 159

5.14 Power spectra for Q = 1020 slpm, Ф = 0.65, no PIM (a) Tinlet = 20 °C, (b) Tinlet = 130 °C, (c) Tinlet = 260 °C................................................................... 160



5.17 Power spectra for Q = 1020 slpm, Ф = 0.65, 18 ppcm PIM (a) Tinlet = 20 °C, (b) Tinlet = 130 °C, (c) Tinlet = 260 °C................................................................... 163






5.23 SPL in one third octave for Q = 1020 slpm, Ф = 0.65, (a) Tinlet = 20 °C, (b) Tinlet = 130 °C, (c) Tinlet = 260 °C................................................................... 169



5.26 CO emissions for Q = 1020 slpm, Ф = 0.65, (a) Tinlet = 20 °C, (b) Tinlet = 130 °C, (c) Tinlet = 260 °C................................................................... 172


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5.28 NOx emissions for Q = 1020 slpm, Ф = 0.65, (a) Tinlet = 20 °C, (b) Tinlet = 130 °C, (c) Tinlet = 260 °C................................................................... 174


5.30 Pressure drop measurements for open top experiments Q = 1020 slpm (a) no PIM (b) 18 ppcm PIM (c) 32 ppcm PIM .................................................. 176













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5.44 Pressure drop measurements for open top experiments Q = 1400 slpm (a) no PIM (b) 18 ppcm PIM .............................................................................. 190

5.45 Schematic diagram of nozzle for restricted flow experiments.............................. 191

5.46 Jet noise power spectra, no PIM, P = 1 atm, Ф = 0.75, Tinlet = 130 °C (a) sampling rate of 2000 Hz, (b) sampling rate of 4000 Hz ................................ 192

5.47 Combustion noise power spectra, no PIM, P = 1 atm, Ф = 0.75, Tinlet = 130 °C (a) Sampling rate of 2000 Hz, (b) Sampling rate of 4000 Hz ........ 193

5.48 Location of microphones for jet noise ................................................................. 194

5.49 Jet noise SPL in one third octave for no PIM, Q = 1020 slpm, Ф = 0.70, (a) Tinlet = 20 °C, (b) Tinlet = 130 °C, (c) Tinlet = 260 °C ........................ 195

5.50 Jet noise power spectra, no PIM, Q = 1020 slpm, Ф = 0.65 (a) Tinlet = 20 °C, (b) Tinlet = 130 °C, (c) Tinlet = 260 °C ....................................... 196

5.51 Jet noise power spectra, 18 ppcm PIM, Q = 1020 slpm, Ф = 0.65 (a) Tinlet = 20 °C, (b) Tinlet = 130 °C, (c) Tinlet = 260 °C ....................................... 197

5.52 Jet noise SPL in one third octave, Q = 1020 slpm, Ф = 0.65, P = 1 atm, (a) Tinlet = 20 °C, (b) Tinlet = 130 °C, (c) Tinlet = 260 °C ....................... 198

5.53 Combustion noise power spectra, no PIM, Q = 1020 slpm, Ф = 0.65 (a) Tinlet = 20 °C, (b) Tinlet = 130 °C, (c) Tinlet = 260 °C ....................................... 199

5.54 Combustion noise power spectra, 18 ppcm PIM, Q = 1020 slpm, Ф = 0.65 (a) Tinlet = 20 °C, (b) Tinlet = 130 °C, (c) Tinlet = 260 °C ....................................... 200

5.55 Combustion noise SPL in one third octave, Q = 1020 slpm, Ф = 0.65, P = 1 atm (a) Tinlet = 20 °C, (b) Tinlet = 130 °C, (c) Tinlet = 260 °C ........................ 201



5.58 Jet noise SPL in one third octave, Q = 1020 slpm, Ф = 0.70, P = 1 atm (a) Tinlet = 20 °C, (b) Tinlet = 130 °C, (c) Tinlet = 260 °C ........................ 204

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5.60 Combustion noise power spectra, 18 ppcm PIM, Q = 1020 slpm, Ф = 0.70 (a) Tinlet = 20 °C, (b) Tinlet = 130 °C, (c) Tinlet = 260 °C ......................... 206




5.64 Jet noise SPL in one third octave, Q = 1020 slpm, Ф = 0.75, P = 1 atm (a) Tinlet = 20 °C, (b) Tinlet = 130 °C, (c) Tinlet = 260 °C ........................ 210


5.66 Combustion noise power spectra, 18 ppcm PIM, Q = 1020 slpm, Ф = 0.75 (a) Tinlet = 20 °C, (b) Tinlet = 130 °C, (c) Tinlet = 260 °C ......................... 212


5.68 Pressure drop measurements for restricted top experiments, P = 1 atm (a) no PIM (b) 18 ppcm PIM .............................................................. 214

5.69 Schematic diagram of nozzle for restricted flow experiments.............................. 215




5.73 Pressure drop measurements for restricted top experiments, P = 2 atm (a) no PIM (b) 18 ppcm PIM .............................................................. 219

A.1 Schematic diagram of the experimental setup ..................................................... 242

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A.2 Schematic diagram (top) and photograph of the swirler (bottom) at the combustor inlet plane .................................................................................... 243

A.3 Injector details .................................................................................................... 244

A.4 Effect of atomizing air on flame images ............................................................. 245

A.5 Axial profiles of emissions for diesel, (a) NOx, (b) CO ...................................... 246

A.6 Axial profiles of emissions for biodiesel, (a) NOx, (b) CO ................................. 247

A.7 Axial profiles of emissions for biooil, (a) NOx, (b) CO ...................................... 248

A.8 Radial profiles of emissions for diesel, (a) NOx, (b) CO .................................... 249

A.9 Radial profiles of emissions for biodiesel, (a) NOx, (b) CO ............................... 250

A.10 Radial profiles of emissions for biooil, (a) NOx, (b) CO .................................... 251

A.11 Axial profiles of emissions for 15% AA, (a) NOx, (b) CO ................................. 252



A.14 Radial profiles of emissions for 15% AA, (a) NOx, (b) CO ................................ 255



1

CHAPTER 1

INTRODUCTION

Background

In recent years, considerable interest has been generated to develop fuel-flexible power

systems using advanced gas turbines to achieve high-efficiency with ultra low-emissions

(Richards, 2001). Renewable fuels produced from homegrown biomass are expected to

constitute a greater portion of the fuel feedstock in the near to mid-term. Increased production

and use of biofuels will not only benefit the environment but also contribute to the energy

security and economic growth. Further, liquid bio-fuels present an emerging opportunity for

power generating gas turbine applications. Thus, part of this study isolates the effects of fuel

composition and fluid dynamics on emissions from different liquid fuels in an atmospheric

pressure burner replicating typical features of a gas turbine combustor. The burner utilized a

commercial twin-fluid injector with primary air swirling around the injector. The fuels include

diesel, biodiesel, emulsified biooil, and diesel-biodiesel blends. For fixed volume flow rates of

fuel and air, experiments were conducted by varying the airflow split between the injector and

co-flow swirler. Results show that flow dynamics induced by inlet conditions, i.e., split ratio of

airflow, has a dramatic impact on combustion performance: fuel atomization is improved and

emissions are reduced by changing the flow structure of the flame. Details of the investigation of

liquid fuel combustion are presented in Appendix A. The remaining scope of this study is to

investigate effect of porous insert material on noise and instabilities in a combustor operated with

2

gaseous fuels as a first step to implement the concept in liquid-fuel operated power generation

systems.

Lean premixed (LPM) combustion of hydrogen-rich syngas has emerged as a means to

effectively burn a variety of fuels while lowering emissions. LPM combustion can cause

autoignition, flame flashback, and/or combustion instabilities, which must be eliminated to

ensure reliable operation, structural rigidity, and acceptable NOx and CO emissions (Lieuwen,

2006; Jayasuria, 2006; Moriconi, 2005). Figure 1.1 illustrates the stabilization mechanism of a

typical swirl LPM combustor. Incoming reactants experience a sudden expansion of cross

sectional area, which creates the corner recirculation zone. This recirculation of hot products

provides energy to ignite incoming reactants. Also, high velocity gradient in the central region of

the combustor creates a central recirculation zone, which also provides energy to ignite incoming

reactants. These recirculation zones are highly turbulent, which results in high pressure

oscillations. On the other hand, heat release from the reaction zone is unsteady because of the

turbulent nature of the flame. Pressure oscillations and unsteady heat release excite the

surrounding acoustic field, generating combustion noise.

Combustion instabilities are a major challenge in today’s power generation systems.

Combustion instabilities are spontaneously excited by a feedback loop between combustion

oscillations and acoustic modes of the combustor, causing large pressure oscillations in the

combustor, large amplitude vibrations, increased heat transfer and thermal stresses on the

combustor walls, oscillatory mechanical loads and severe mechanical damage, thus, operation

down-time and costly repairs (Lieuwen, 2005). Combustion instabilities occur when unsteady

heat release from the combustion process adds energy to the acoustic field faster than it can

dissipate it via, for example, viscous dissipation and heat transfer. This is known as the

3

Rayleigh’s criterion (Rayleigh, 1945) and is expressed mathematically by the Rayleigh integral,

given as:

� ��, � ��, � �� ≥ ��, � �� (1.1)

Where:

p’(x,t) = combustor pressure oscillations

q’(x,t) = heat addition oscillations

V = combustor volume

T = period of the oscillations

Li = ith acoustic energy loss process

Rayleigh’s criterion is satisfied when flame adds energy to the acoustic field. Flame adds

energy to the acoustic field when phase between unsteady heat release from the combustion

process and pressure oscillations is less than 90 degrees. If phase between unsteady heat release

and pressure oscillations is greater than 90 degrees, then the combustion process removes energy

from the acoustic field. Thus, the Rayleigh integral states that combustion instabilities occur

when the magnitude of the driving force exceeds the damping process, i.e., net energy added to

the acoustic mode exceeds the dissipation mechanism. In this study, an experimental study of a

passive technique to mitigate combustion noise and instabilities is proposed for typical operating

conditions of a turbine engine.

Figure 1.2 shows two PIM configurations initially considered to implement the PIM

concept with a swirl-stabilized combustor (Agrawal, 2008). In Configuration 1, a PIM is placed

4

at the center of the combustor, presumably to affect the center flow recirculation region. In

Configuration 2, a circular ring is used to modify the flow in the corner and near wall regions of

the combustor. Preliminary experiments revealed that configuration 1 does not reduce

combustion noise. Configuration 2 was however identified as a promising concept for further

investigation (Agrawal, 2008). This concept differs fundamentally from the existing PIM

combustion literature dealing only with surface or submerged reactions throughout the

combustor (Howell, 1996, Trimis, 1996; Marbach, 2007, Waitz, 1998, Fernandez-Pello, 2002).

Instead, in this concept, the PIM is used synergistically to improve the performance of gaseous

flames produced in the swirl-stabilized combustor. Next, a brief description of the research

discussed in each chapter is presented.

1.2 Overview

1. First, an investigation of combustion performance of various alternative fuels is

presented. The fuels used in this study are diesel (commercial grade), biodiesel, biooil

and diesel-biodiesel blends. A commercial air blast injector was used to atomize the

liquid fuel. Total air supply is split two ways: atomizing air and combustion air.

Atomizing air is supplied to the air blast injector and used to atomize the fuel.

Combustion air is fed through a swirler to create a swirl-stabilized flame. Visual images,

CO and NOx emissions are presented in Appendix A for different split ratios with total

air supply kept constant (therefore overall equivalence ratio was constant) for different

fuels. Results show that for a given equivalence ratio flow effects have a significant

impact on NOx and CO emissions.

5

2. Chapter 2 presents a numerical investigation of the effect of porous inserts in the swirl

stabilized combustion chamber. The study reveals how strategically located porous

inserts combined with swirl stabilization mechanism, fundamentally changes the overall

combustion process, redistributing and redirecting reactants and products, eliminating

vortical structures resulting in a distributed flame front that can reduce sound pressure

levels.

3. Chapter 3 presents the experimental counterpart of the study discussed in chapter 2.

Premixed combustion of methane in a swirl-stabilized combustor is combined with

porous inert material to investigate the effect on combustion noise and emissions of CO

and NOx. Experiments are conducted with different PIM thicknesses, pore densities,

geometries, and equivalence ratios to investigate effect on sound pressure level. Tests are

conducted at relatively low reactant flow rates (up to Q = 600 slpm, Re = 10,000) and

inlet air temperatures of 100 and 120 °C. Different PIM combustion modes are identified

in this study.

4. Chapter 4 presents the development of a lab facility to conduct combustion experiments

at high reactant flow rates, high inlet air temperatures, and high operating pressure.

Facility design including details of air and fuel supply systems, instruments and data

acquisition system, and operational procedure are discussed. This chapter sets the stage

for experiments discussed in the next chapter

5. Chapter 5 presents experimental results for high reactant flow rates, high air inlet

temperatures, and high operating pressures, to closely simulate gas turbine operating

conditions. Experiments are conducted using annular diffuser-shaped porous inserts

identified as optimum PIM geometry in Chapter 3. Measurements of combustion noise

6

and jet noise are presented to identify the link between the two. Results show that porous

insert mitigates combustion noise and jet noise, and eliminate combustion instabilities,

when present.

6. Chapter 6 presents the concluding remarks of the investigation. Also recommendations

for future work are presented in this chapter.

7

Figure 1.1. Schematic diagram of swirl stabilization mechanism

Swirler

Central recirculation zone

Corner recirculation zone

Flame

Reactants

Combustor wall

Products

Dump plane

8

Figure 1.2. Proposed concepts

Configuration 1

Swirler

Reactants

Swirler

Reactants

Configuration 2

Combustor

Combustor

PIM

PIM

9

CHAPTER 2

NUMERICAL SIMULATIONS OF SWIRL STABILIZED COMBUSTION COUPLED WITH POROUS INERT MEDIUM

Background

Lean Premixed (LPM) combustion has proven to be an effective way to control the flame

temperature, therefore, avoiding the thermal NOx mechanism to play an important role in

combustion. LPM combustion is a simple concept that reduces NOx emissions without the need

for installing, maintaining and operating sophisticated post cleanup equipment. In recent years,

an extensive effort has been made to understand and mitigate problems associated with LPM

combustion systems. Much of the research has focused on natural gas (NG) fuel and hence, the

turbine installations of the past decade are mainly NG fueled (Richards, 2001). Swirling flows

are extensively used for flame stabilization (Gupta, 1984). Strong radial and axial pressure

gradients generated by the swirl flow induce axial recirculation zones. Additionally, corner

recirculation zones are generated by sudden expansion of cross-sectional area and existence of

bluff body. However, the precise flow structure depends on many factors, i.e., swirl injector

geometry, size of enclosure, particular exit velocity profiles, etc (Gupta, 1984). Recirculation

zones generate reversed hot flow of combustion products that ignite the incoming reactants,

providing a flame stabilization mechanism, as illustrated in Figure 2.1. Advanced gas turbines

for power generation utilize swirl-stabilized combustion systems operated in the LPM mode.

10

Porous inert media (PIM) has been utilized as another technique for flame stabilization,

capable of achieving ultra low NOx emissions (Marbach, 2005). Heat released by combustion is

transferred from the reaction zone to the PIM, which in turn radiates and convects heat upstream

to preheat the incoming reactants. The result is the capability to control flame stability and flame

temperature, lowering NOx emissions (Marbach, 2005).

Numerical models of swirl-stabilized combustion systems have been developed in the

past. Huang et al. (2003) illustrated instantaneous velocity field, instantaneous fluctuating

pressure field, and effect of increased air inlet temperature on temporal evolution of flame in a

swirl-stabilized burner. Stone and Menon (2002) used LES to investigate the effect of swirl and

equivalence ratio on flame dynamics. Grinstein et al. (2005) simulated a swirl combustor to

study the effect of combustor confinement on flowfield and flame evolution.

Past studies have investigated combustion performance of swirl-stabilized and PIM-

stabilized systems. However, those studies have utilized only one type of flame stabilization

technique, i.e., swirl stabilized or PIM stabilized. The objective of this investigation is to gain a

fundamental understanding of the changes in the flow structure induced by the presence of PIM

in the swirl-stabilized combustor. Results of the numerical simulation are compared to

experimental results for non-reacting and reacting flows without the PIM (Wicksall, 2005).

Several simplifying assumptions associated with combustion and turbulence models and

boundary conditions are made. Therefore, the computed results provide only a qualitative

assessment of the flow field.

11

2.2 Physical Model

Figure 2.2 shows a schematic diagram of the combustor with the swirler. The swirler has

six vanes positioned at 28° to the horizontal. The theoretical swirl number is 1.5, assuming that

the flow exits tangentially from the swirler vanes. The bulk axial inlet velocity is 10 m/s. The

combustor is an 8.1 cm inner diameter and 30.6 cm long quartz tube. Inlet velocity was specified

by radial, axial and swirl components. The combustor was modeled as 2D axisymetric geometry

with swirl, which assumes that there are no circumferential gradients in the flow. Figure 2.3

shows the computational domain with finer grids used near the inlet and PIM regions to resolve

the flow gradients.

2.2.1 Governing Equations

The flow field was computed from continuity and momentum equations in axial, radial

and circumferential directions. Turbulence was modeled using the RNG k – ε model. Cold flow

and reacting flow were modeled with and without the PIM. A simplified porous media model

was used in this study (Marbach, 2006). In this model, sink terms are added to the conservation

of momentum equations to approximate flow resistance associated with PIM. The sink term was

modeled using a power law correlation, with C0 and C1 determined experimentally (Marbach,

2005). An effective thermal conductivity (keff) was used to account for the solid and fluid

conductivities and the porosity of the porous media. The governing equations are:

Mass conservation equation: �� + �� = 0 (2.1)

12

Momentum conservation equations: �� + �� = − �� + �� + �� − 23�� + �� −��′��′�� + �� (2.2) �� = −�|�|� (2.3) !�� = �� (2.4)

Energy equation: �� " + �� #��(�" + )$ = �� !�� %�� + ��(&��)�� (2.5)

Where E is total energy, keff is the effective thermal conductivity and �&�� is the deviatoric

stress tensor.

2.2.2 Combustion Model

Combustion was modeled using turbulent premixed combustion model, based on the

work of Zimont et al. (2000, 1998, 1995). This model involves the solution of a transport

equation for the reaction progress variable. The closure of this equation is based on the definition

of the turbulent flame speed. The flame front propagation is modeled by solving for the density

weighted mean reaction progress variable, �̃:

∇ ∙ �� ̃ = ∇ ∙ ' �� ∇�̃(+ �� (2.6)

13

Where � ̃ is mean reaction progress variable, Sc the mean reaction rate (s-1) and Sct is turbulent

Schmidt number. The progress variable is defined as a normalized sum of the product species

mass fraction:

�̃ = ∑ *��∑ *�,�� (2.7)

Where n is number of products, Yi is mass fraction of product species i (CH4, O2 and N2), and

Yi,eq is mass fraction of product species i at chemical equilibrium (CO2, H2O, N2, O2). The value

of �̃ is defined as a boundary condition at all flow inlets. It is specified as either 0.0 (unburnt) or

1.0 (burnt). The mean reaction rate, Sc, is modeled as:

�� = ��+�|∇�̃| (2.8)

Where ρu is density of unburnt mixture, and Ut is turbulent flame speed. The closure of the

problem is based on the definition of turbulent flame speed (Zimont, 1998):

+� = ,(�′)��+��-�� (2.9)

Where A is model constant (0.52), u’ is RMS axial velocity (m/s), Ul = laminar flame speed

(m/s), α is thermal diffusivity of unburnt mixture (m2/s), and l t is turbulent length scale (m).

Laminar flame speed and thermal diffusivity of unburnt mixture are known constants (0.12 m/s

for Ф = 0.58, 0.35 m/s for Ф = 0.58 and 1.96x10-05 m2/s). The turbulence length scale, l t, is

computed from:

14

-� = �� ′ �� (2.10)

Where ε is the turbulence dissipation rate and CD is turbulent length scale constant (0.37).

Critical strain rate represents a measure of probability of flame stretching. Flame

stretching has an impact on mean turbulent heat release intensity and can result in flame blow-

off. Thus, critical strain rate indirectly represents a measure of probability of flame quenching. If

there is no flame stretching, the flame will be unquenched Critical strain rate was specified as

2000 s-1 (Wicksall, 2005)

2.2.3 Boundary Conditions

A numerical simulation was developed to model the effect of the PIM on the flow

structure of non-reacting and reacting swirling flows. The inlet boundary condition is a

simplified assumption based on experimental data (Chigier, 1964). Thus, conclusions must be

drawn carefully considering this limitation of the model. Swirling flow was modeled with

incoming flow entering at 28° angle, specifying velocity components (radial, axial and swirl).

Axial velocity was specified as 10 m/s which is also the measured bulk inlet velocity based on

the swirler flow cross-sectional area. Swirl and radial velocity components were specified as

linear profiles for each component (Gupta, 1984). At the inlet, turbulence intensity was specified

as 10% of the total kinetic energy and turbulent length scale was specified as 1.5 mm. The flow

enters the combustor at radial locations between 10 mm and 20 mm. The outlet boundary

condition is set to pressure outlet, to improve convergence if backflow occurs. Numerical

convergence was determined when all residuals reached values below 10-6

15

2.2.4 Model Validation

Computations were performed using four different grid sizes: 75 x 40, 100 x 60, 125 x 80

and 200 x 100. Figure 2.4 shows axial velocity profiles at the axial location of 20 mm for

different grid sizes for reacting flow case with Ф = 0.58. Since the results of 125 x 80 and 200 x

100 grids are nearly identical, 125 x 80 grid was used for all calculations to provide grid

independent solution. Figure 2.5 shows the computed and experimental (Wicksall, 2005) velocity

vectors for non-reacting flow with no porous media. Results show qualitative agreement between

computed and experimental velocity fields. Corner and central recirculation zones are seen in the

vector plots. Inlet flow enters approximately at 35° angle from the vertical, then turns

approximately an additional 10° as flow from the central recirculation zone re-attaches with the

inlet flow. Central and corner recirculation flows re-attach at similar locations for computed and

experimental results. Although simplifying assumptions were made in the computational model,

results qualitatively predict the main features of the flow.

Figure 2.6 shows the computed and experimental (Wicksall, 2005) velocity vectors for

combustion of methane at Ф = 0.58 without porous media. Similar to non-reacting flow,

computed and experimental results show qualitative agreement. Velocities are higher for reacting

case, as expected, due to decrease in density of the products. Main features of the measured flow

are replicated by computations: inlet flow at approximately 35° angle from the vertical,

additional 10° turning as central recirculation zone re-attaches; central and corner recirculation

zones with similar re-attachment locations.

16

2.3 Results and Discussion

Computations were performed for non-reacting and reacting flows. Methane flames of Ф

= 0.58 and 0.85 were modeled for the reacting flow. For each case, effect of PIM on the flow

structure was investigated. Comparison of computed and experimental results is presented.

Results include velocity vectors and radial profiles at different axial locations. Results for non-

reacting and reacting flows are presented in the following sections.

2.3.1 Non-reacting Flow

Figure 2.7 shows the computed velocity field in a 40 mm by 60 mm window. Flow

structures with no porous material and with porous material are presented for identical conditions

in Figures 2.7(a) and 2.7(b), respectively. Central and corner recirculation zones are present in

case of the flow field with no porous media. The corner recirculation zone results from the

sudden cross-sectional area expansion in the flow direction. Central recirculation zone extends

across much of the width of the domain. Larger flow velocities occur near the wall of the

enclosure. Presence of PIM in the enclosure dramatically changes the flow structure, as seen in

Figure 2.7(b). Corner recirculation zone disappears because flow is distributed within PIM.

Resistance to the flow introduced by PIM causes inlet flow to tilt vertically. Central recirculation

zone is narrow and more intense compared to its no PIM counterpart.

Next, Figure 2.8 shows axial velocity profiles at different axial locations (z) of the

domain to examine the evolution of the flow without and with PIM. Figure 2.8 shows different

locations of peak axial velocity for flow without and with PIM. Axial velocity peak without PIM

progresses toward the wall and becomes narrow at z = 30 mm. Near the wall, axial velocity is

negative indicating the corner recirculation zone. When PIM is present, axial velocity in the

17

porous region is positive but close to zero, indicating blockage of the flow created by the porous

insert. There is no evidence of a corner recirculation zone. Outside the porous region, axial

velocity peaks at around r = 16 mm. This peak location remains constant as the flow evolves in

the axial direction. Axial velocity near the center has a larger negative value, indicating a more

intense central recirculation zone compared to that without PIM.

Figure 2.9 shows the evolution of the swirl velocity without and with PIM. Without the

PIM, the location of peak swirl velocity progresses towards the combustor wall, similar to the

peak axial velocity. The peak value of the swirl velocity decreases in the axial direction. When

PIM is present, swirl velocity inside porous region is zero. Outside porous region, swirl velocity

remains approximately constant. Location of peak of swirl velocity remains constant at

approximately r = 16 mm, similar to the axial velocity. These results indicate that the PIM

produces a stronger swirl near the center region. Figure 2.10 compares the radial velocity profiles

without and with PIM. Without PIM, peak radial velocity decreases in the axial direction, and its

location progressively shifts towards the combustor wall. At z = 30 mm the radial velocity is

nearly zero. With PIM, the radial component of velocity is nearly zero at all axial locations.

Overall, results indicate that swirling effect induced by the swirl injector is intensified by the

porous insert. Furthermore, the corner recirculation zone is diminished and central recirculation

zone is also intensified.

2.3.2 Reacting Flow

Figure 2.11 shows computed velocity fields in a 40 mm by 60 mm window for reacting

flow without and with PIM. These results show the change in time-averaged flowfield caused by

PIM in the reaction zone. Similar to the non-reacting case, a corner recirculation zone exists

18

without porous insert, as seen in Figure 2.11(a). The central recirculation zone occupies a large

portion of the combustor width. Both corner and central recirculation zones provide a mechanism

for flame stabilization as hot products come in contact with incoming reactant flow, igniting the

mixture to sustain the flame. For case with PIM, Figure 2.11(b), corner recirculation zone is

eliminated by flow redistribution in the porous region. Central recirculation zone becomes more

intense. Presence of PIM changes the flow structure, although typical swirl-stabilization

mechanism is still present. A more intense central recirculation zone remains responsible for

igniting fresh reactant flow in the central region. Although the typical swirl-stabilization

mechanism is affected, the combined swirl-PIM system is also effective in stabilizing the flame.

This remark is consistent with experimental observations showing a stable flame.

Figure 2.12 shows axial velocity profiles at different axial location for reacting flow at Ф

= 0.58. Results indicate trends similar to those obtained for non-reacting flow. Figure 2.12 shows

the axial velocity without and with PIM. Without PIM, the axial velocity peak progresses

towards the wall where a section of negative axial velocity exists, indicating corner recirculation

zone. Presence of PIM causes the axial velocity peak to remain at nearly a constant radial

location. The axial velocity within the porous region is nearly zero. A region of negative axial

velocity with magnitude greater than the no PIM case evolves near the center of the combustor.

PIM restricts the radial extent of the flow recirculation region, which is stronger for the case with

PIM. Figure 2.13 shows the swirl component of the velocity for cases without and with PIM.

With no PIM, swirl velocity of similar magnitude extends across the radial direction. With PIM,

the peak swirl velocity is higher, and it remains approximately constant in the axial direction,

both in magnitude and location. Swirl velocity in the porous region is zero. Similar to the non-

reacting case, swirl effect does not diminish with the presence of PIM. Figure 2.14 shows the

19

radial velocity profiles without and with PIM. Without PIM, radial velocity decreases rapidly in

the axial direction. With PIM, radial velocity is approximately zero at all axial locations.

Reacting flow computations were also performed for Ф = 0.85. Figure 2.15 shows

velocity vectors in the combustion chamber. Results are similar to those for Ф = 0.58. Figures

2.16 to 2.18 show profiles of axial, swirl and radial velocity at different axial locations. Similar

trends to those obtained for reacting flow at Ф = 0.58 are observed. As expected, velocity

magnitudes in this case are greater because of the higher flame temperature. The overall flow

field without or with PIM is unaffected by an increase in the equivalence ratio.

2.4 Conclusions

Flow field in a LPM swirl-stabilized combustor integrated with porous media was

computed. Non-reacting flow and reacting flow were modeled without and with PIM. Flow

resistance associated with PIM was modeled by adding sink terms to the momentum

conservation equations. The sink term was modeled by a power law correlation with coefficients

determined experimentally. Methane flames were modeled for Ф = 0.58 and 0.85. Turbulence

was modeled using the RNG k – ε model. Inlet boundary conditions were simplified assumptions

based on experimental data. Combustion was modeled using a turbulent premixed combustion

model. The computed flow field was compared with experimentally obtained data for reacting

and non-reacting flows. Results show qualitative agreement for both reacting and non-reacting

flows. Change in the flow structure introduced by the porous insert was investigated next.

Results show that the porous insert significantly alters the flow structure. Porous insert

eliminates the corner recirculation zone, vertically orients the gaseous flame zone, intensifies the

central recirculation zone, maintains the swirling effect imparted by the swirl injector, and

20

creates a more uniform flow distribution at downstream locations. These unique features of the

present concept can improve the noise and instability performance of combustor as discussed

next. The flow field is similar for non-reacting and reacting cases, and it is not affected

significantly by an increase in the equivalence ratio.

21

Figure 2.1. Schematic diagram of swirl stabilization mechanism

Swirler

Central recirculation zone

Corner recirculation zone

Flame

Reactants

Combustor wall

Products

Dump plane

22

Figure 2.2. Schematic of combustor with the swirler

8.1

Dimensions in cm

Inlet Swirler

30.6

4.0

23

Figure 2.3. Computational domain

Figure 2.4.Axial velocity profile at z = 20 mm, methane flame, Φ = 0.58

0 100 200 300

010203040

Inlet Outflow

Symmetry plane / axial location

Radial location (mm)

Axi

alve

loci

ty(m

/s)

40 30 20 10 0-4

-2

0

2

4

6

8

10

75x40100x60125x80200x100

Combustor wall

24

Figure 2.5. Velocity vectors for non-reacting flow. (a) Experimental results (Wicksall, 2005), (b)

Computed results


Axi

allo

catio

n(m

m)

25 20 15 10 5

5

10

15

20

25

(a) (b)

25

Figure 2.6. Velocity vectors for reacting flow. (a) Experimental results (Wicksall, 2005), (b)

Computed results


Axi

allo

catio

n(m

m)

25 20 15 10 5

5

10

15

20

25

(a) (b)

26

Figure 2.7. Velocity vectors for non-reacting flow. (a) without PIM, (b) with PIM

(a) (b)


Axi

allo

catio

n(m

m)

40 30 20 10 00

20

40

60m/s: -4 -2 0 2 4 6 8 10


Axi

allo

catio

n(m

m)

40 30 20 10 00

20

40

60m/s: -8 -4 -2 0 2 4 6 10

27

Figure 2.8. Axial velocity profiles at different axial locations for non-reacting flow: (a) z = 10

mm, (b) z = 20 mm, (c) z = 30 mm


Axi

alve

loci

ty(m

/s)

40 30 20 10 0-15

-10

-5

0

5

10

15

No PIMPIM


Axi

alve

loci

ty(m

/s)

40 30 20 10 0-15

-10

-5

0

5

10

15

No PIMPIM


Axi

alve

loci

ty(m

/s)

40 30 20 10 0-15

-10

-5

0

5

10

15

No PIMPIM

(a)

(b)

(c)

28

Figure 2.9. Swirl velocity profiles at different axial locations for non-reacting flow: (a) z = 10

mm, (b) z = 20 mm, (c) z = 30 mm


Sw

irlv

elo

city

(m/s

)

40 30 20 10 0-15

-10

-5

0

5

10

15

No PIMPIM


Sw

irlve

loci

ty(m

/s)

40 30 20 10 0-15

-10

-5

0

5

10

15

No PIMPIM


Sw

irlv

elo

city

(m/s

)

40 30 20 10 0-15

-10

-5

0

5

10

15

No PIMPIM

(a)

(b)

(c)

29

Figure 2.10. Radial velocity profiles at different axial locations for non-reacting flow: (a) z = 10

mm, (b) z = 20 mm, (c) z = 30 mm


Rad

ialv

elo

city

(m/s

)

40 30 20 10 0-15

-10

-5

0

5

10

15

No PIMPIM


Rad

ialv

eloc

ity(m

/s)

40 30 20 10 0-15

-10

-5

0

5

10

15

No PIMPIM


Rad

ialv

eloc

ity(m

/s)

40 30 20 10 0-15

-10

-5

0

5

10

15

No PIMPIM

(a)

(b)

(c)

30

Figure 2.11. Velocity vectors for reacting flow Ф = 0.58. (a) without PIM, (b) with PIM

(a) (b) Radial location (mm)

Axi

allo

catio

n(m

m)

40 30 20 10 0

20

40

60m/s: -8 -6 -2 0 2 6 10 12


Axi

allo

catio

n(m

m)

40 30 20 10 00

20

40

60m/s: -4 -2 0 2 4 6 8 10

31

Figure 2.12.Axial velocity profiles at different axial locations for Ф = 0.58: (a) z = 10 mm, (b) z

= 20 mm, (c) z = 30 mm


Axi

alve

loci

ty(m

/s)

40 30 20 10 0-15

-10

-5

0

5

10

15

No PIMPIM


Axi

alve

loci

ty(m

/s)

40 30 20 10 0-15

-10

-5

0

5

10

15

No PIMPIM


Axi

alve

loci

ty(m

/s)

40 30 20 10 0-15

-10

-5

0

5

10

15

No PIMPIM

(a)

(c)

(b)

32

Figure 2.13. Swirl velocity profiles at different axial locations for Ф = 0.58: (a) z = 10 mm, (b) z

= 20 mm, (c) z = 30 mm


Sw

irlve

loci

ty(m

/s)

40 30 20 10 0-15

-10

-5

0

5

10

15

No PIMPIM


Sw

irlve

loci

ty(m

/s)

40 30 20 10 0-15

-10

-5

0

5

10

15

No PIMPIM


Sw

irlv

elo

city

(m/s

)

40 30 20 10 0-15

-10

-5

0

5

10

15

No PIMPIM

(a)

(c)

(b)

33

Figure 2.14. Radial velocity profiles at different axial locations for Ф = 0.58: (a) z = 10 mm, (b)

z = 20 mm, (c) z = 30 mm


Rad

ialv

elo

city

(m/s

)

40 30 20 10 0-15

-10

-5

0

5

10

15

No PIMPIM


Rad

ialv

elo

city

(m/s

)

40 30 20 10 0-15

-10

-5

0

5

10

15

No PIMPIM


Rad

ialv

eloc

ity(m

/s)

40 30 20 10 0-15

-10

-5

0

5

10

15

No PIMPIM

(a)

(c)

(b)

34

Figure 2.15. Velocity vectors for reacting flow Ф = 0.85. (a) without PIM, (b) with PIM


Axi

allo

catio

n(m

m)

40 30 20 10 00

20

40

60


Axi

allo

catio

n(m

m)

40 30 20 10 00

20

40

60m/s: -6 -4 -2 0 4 8 10 14 m/s: -3 -2 0 4 8 10 12 14

35

Figure 2.16. Axial velocity profiles at different axial locations for Ф = 0.85: (a) z = 10 mm, (b) z

= 20 mm, (c) z = 30 mm

Radial position (mm)

Axi

alve

loci

ty(m

/s)

40 30 20 10 0-15

-10

-5

0

5

10

15

PIMNo PIM


Axi

alve

loci

ty(m

/s)

40 30 20 10 0-15

-10

-5

0

5

10

15

PIMNo PIM


Axi

alve

loci

ty(m

/s)

40 30 20 10 0-15

-10

-5

0

5

10

15

PIMNo PIM

(a)

(c)

(b)

36

Figure 2.17. Swirl velocity profiles at different axial locations for Ф = 0.85: (a) z = 10 mm, (b) z

= 20 mm, (c) z = 30 mm


Sw

irlve

loci

ty(m

/s)

40 30 20 10 0-15

-10

-5

0

5

10

15

No PIMPIM


Sw

irlv

elo

city

(m/s

)

40 30 20 10 0-15

-10

-5

0

5

10

15

No PIMPIM


Sw

irlv

elo

city

(m/s

)

40 30 20 10 0-15

-10

-5

0

5

10

15

No PIMPIM

(a)

(c)

(b)

37

Figure 2.18. Radial velocity profiles at different axial locations for Ф = 0.85: (a) z = 10 mm, (b)

z = 20 mm, (c) z = 30 mm


Rad

ialv

elo

city

(m/s

)

40 30 20 10 0-15

-10

-5

0

5

10

15

No PIMPIM


Rad

ialv

elo

city

(m/s

)

40 30 20 10 0-15

-10

-5

0

5

10

15

No PIMPIM


Rad

ialv

elo

city

(m/s

)

40 30 20 10 0-15

-10

-5

0

5

10

15

No PIMPIM

(a)

(c)

(b)

38

CHAPTER 3

NOISE REDUCTION IN SWIRL-STABILIZED COMBUSTOR COUPLED WITH PIM

Background

In recent years, noise emission has become increasingly important to industry and

society. The combustion process is a common source of noise production in gas turbines,

internal combustion engines, industrial burners, and commercial furnaces. Heat release in the

reacting mixture causes dilatation, which produces pressure pulsations that propagate outside the

flame zone as sound waves. Most practical combustion systems operate with the working fluid

in turbulent motion with embedded reaction zones, which alter the noise production mechanism.

In addition to the direct noise produced in the reaction zone, thermal non-uniformities arising

from the combustor can generate indirect noise in downstream components. The topic of this

study is the direct combustion noise, which is often the dominant component of the total sound

power.

The early research on combustion noise is summarized by Putnum (1976) and Strahle

(1978), who report analytical models and empirical data base for sound power level as a function

of burner geometry, flow rate, fuel type, equivalence ratio, etc. In recent years, the research

focus has shifted to LPM combustion systems, driven by the need to comply with the

increasingly stringent emissions regulations. However, typical LPM combustion systems are

prone not only to combustion noise but also to combustion instability characterized by coherent,

fixed frequency feedback oscillations. Combustion noise and instability are distinct outcomes,

39

yet they both arise from the same source, i.e., heat release fluctuations in a turbulent flow with

multiple length and time scales. In recent years, several experimental and computational studies

have helped in the understanding of the noise generation mechanisms (Schwarz, 2009; Rajaram,

2006; Flemming, 2005; Choi, 2005; Hirsch, 2007; Tiribuzi, 2008; Duchaine, 2009; Lieuwen,

2005; Lee, 2003; Richards, 2003; Noiray, 2009). In particular, the advent of computational fluid

dynamics (CFD) has afforded the opportunity to analyze noise generation mechanism by

incorporating detailed physics of turbulent combustion with the acoustics. These studies have

identified passive and active methods to control combustion noise and instabilities.

The effectiveness of an active combustion control system depends upon actuation,

sensing, and control algorithms, among other factors. In spite of the significant progress in these

areas, complete reliability of active combustion control is still a major concern since an

unexpected event can destroy the system within a fraction of second. Thus, in this study, a

passive technique involving the use of PIM to suppress combustion noise and/or instability is

investigated. The PIM is an open-cell foam structure alloyed with HfC/SiC to protect the foam

material from high temperature oxidation by creating refractory surface oxides that offer

nominally 600°C higher use temperature than SiC alone. The vortex shedding mechanism of

combustion noise and/or instability is curtailed on source by inserting a properly designed porous

structure to constrain the recirculation regions in the swirl-stabilized combustor. The porous

structure is intended to limit and/or disintegrate vortical structures in the flame to produce a more

homogeneous flow field with limited regions of flow reversal. Previous experiments revealed

that introducing a porous insert is a promising concept for further investigation (Agrawal, 2008).

In this chapter, the experimental set up to acquire noise and pollutant emissions data for a range

40

of geometric and operating conditions is described. Then, results and discussions are presented

followed by the conclusions of the study.

3.2 Experimental Setup

The test apparatus is shown schematically in Figure 3.1. The combustion chamber

operated at atmospheric pressure is a 30.0 cm long, 8.0 cm ID quartz cylinder to enclose to

flame. Heated air enters the system through a plenum filled with marbles to breakdown the large

vortical structures. The air passes through a swirler into the mixing section, where the gaseous

methane is supplied. Air and fuel premix in the mixing region before entering the dump plane of

the combustion chamber through a swirler. The swirler has six vanes positioned at 28o to the

horizontal, and it results in theoretical swirl number of 1.5 (see Appendix B for calculation). The

inlet Reynolds number based on the equivalent diameter ranged from 5,000 to 10,000. The

combustor is back-side cooled by natural convection. A compressed storage tank supplies air

that passes through a pressure regulator, water traps, flow control valve, and an in-line electrical

heater before reaching the experiment. Methane fuel is also supplied from a rack of storage

tanks. Air flow rate is measured by a laminar flow elements (LFE) calibrated for 0 to 1000 liters

per minute (lpm) of air. The LFE for fuel flow rate measurements is calibrated for 0 to 100 lpm

of methane. The flow rate measured by the LFE is corrected for temperature and pressure as

specified by the manufacturer.

Sound pressure data are collected using a Brüel & Kjær microphone probe (Model 4189)

located 28 cm from the edge of the combustor exit plane. A total number of 10000 samples are

acquired in 5 sec at 2000 Hz. The measured voltage signal is converted to pressure fluctuation

data using the probe sensitivity of 44.3 mV/Pa. An FFT analysis is performed to obtain the

41

power spectrum. A Matlab script was written to compute the sound pressure level (SPL) in

decibels (dB) given as (Bussman, 2001):

�.� = 10 ∗ log� �.��.�� (3.1)

Where Pref = 20 µPa. Total SPL was calculated by:

�.�� = 10 ∗ log� /�10.�∗� !�� 0 (3.2)

Where SPLi is the SPL at each frequency level, n is the number of frequency bands. One-third

octave frequency bands were used, from bands 13 to 29, with centers at 20 Hz and 800 Hz

respectively (Bussman, 2001).

Emissions measurements are taken by continuously sampling the product gas by a quartz

probe (OD = 7.0 mm) attached to a three-way manual traversing system. The upstream tip of the

probe was tapered to 1 mm ID to quench reactions inside the probe. The gas sample passed

through an ice bath and water traps to remove moisture upstream of the gas analyzers. The dry

sample passed through electrochemical analyzers to measure the concentrations of CO and NOx

in ppm. The analyzer also measures oxygen and carbon dioxide concentrations, which are used

to cross-check the equivalence ratio obtained from the measured fuel and air flow rates. The

uncorrected emissions data on dry basis are reported with measurement uncertainty of ±2 ppm.

42


In this study, porous inserts of different pore sizes and inside diameters were used. Figure

3.2 shows a photograph of a PIM insert and photographs of the combustor without and with the

PIM insert. Figure 3.3 shows diagrams of the PIM inserts configurations used in this study. All

inserts were 2.5 cm thick, 8.0 cm OD, and had porosity of 85%. The inside diameters were 3.8

cm, 4.4 cm and 5.0 cm, and the pore densities were 4 pores per cm (ppcm), 8 ppcm, and 18

ppcm. For each experiment, two pieces of PIM were stacked together labeled, for example, as

D38-P4 + D44-P4. This configuration pertains to a PIM piece with 3.8 cm inside diameter and

pore density of 4 ppcm followed by (in the direction of the flow) another PIM piece with 4.4 cm

inside diameter and pore density of 4 ppcm. Experiments were conducted at equivalence ratios

(Ф) of 0.7 and 0.8 with airflow rate (Q) of (1) 300 slpm at inlet temperature (Ti) of 100˚C, and

(2) 600 slpm at inlet temperature of 120˚C.

Figure 3.4 shows flames images for Ф=0.7, Q=300 slpm, and Ti=100˚C. Without the

porous insert, the image in Figure 3.4(a) depicts a blue flame typical of LPM combustion. The

image in Figure 3.4(b) pertains to configuration D38-P4+D44-P4. The orange glow in this

image indicates combustion stabilized within the PIM. A confined blue gaseous flame is also

visible (though barely) downstream of the PIM. The image in Figure 3.4(c) reveals a

fundamentally different combustion mode for configuration D38-P18+D38-P18; small blue

flamelets are stabilized on the surface of the porous insert while majority of the reactants burn in

the confined, swirl-stabilized gaseous flame region. The interior combustion mode for the

configuration in Figure 3.4(b) is detrimental for material strength, and it also produced higher

sound pressure levels.

43

Porous insert in the combustor alters the flow field and flame structure by restricting and

re-directing reactants and products inside the combustion chamber. Elimination of the corner

recirculation zone also fundamentally changes the stabilization mechanism of the flame.

Reactant flow exiting the swirl injector is divided into the center (core) region and PIM region.

The reactant flow rate in the PIM would depend upon the flow resistance offered by the PIM.

The reactant flow in the center region results in the typical swirl-stabilized flame, albeit of higher

swirl intensity because of the constrained volume of the free flame. Combustion products from

the free flame would enter the porous insert through the inner surface and mix with the reactants

introduced upstream, as illustrated in Figure 3.5. Combustion products would also transfer heat

to the porous insert, and further to the reactants flowing through the PIM, which leads to interior

or surface combustion depending upon the resulting flame speed and porous insert geometry.

This fundamental mechanism is the basis for all cases studied with PIM inserts. Interior

combustion occurs when the reactants ignite and sustain a flame within the PIM. In this case, a

balance is achieved between: (1) energy of unburned reactants flowing in the PIM, (2) energy of

products from the free flame penetrating the PIM, and (3) heat transfer between free flame and

the PIM. This balance can be expected to depend upon the PIM pore density, PIM geometry,

reactants velocity. Surface combustion occurs if reactants do not ignite within the PIM. Instead,

the flame is established on the downstream surface of the PIM.

Interior and surface combustion modes excite the acoustic field differently, thus resulting

in different noise levels. Interior combustion is accompanied with intense thermal radiation from

the PIM, and it generally produced higher noise levels. The surface combustion produced blue

flamelets similar to the free flame, and this mode of combustion generally mitigated the noise

produced in the free flame without the porous insert. Surface combustion mode is hypothesized

44

to reduce the noise levels for the following reasons: (1) the heat release rate in the free flame is

reduced because some of the reactants also burn downstream of the PIM surface as a distributed

flame, (2) PIM suppresses the pressure fluctuations in the adjacent free flame, and (3) the

vortical structures produced in the corner recirculation zone and shear regions are virtually

eliminated.

3.3.1 Effect of PIM Pore Density

Figure 3.6 shows five data sets for Q = 300 slpm, no PIM and Ф = 0.7. These data were

taken independently to demonstrate repeatability of noise measurements. Sound pressure levels

overlap one another, indicating measurements are repeatable. Table 3.1 presents a summary of

total SPL measured for the baseline case (no PIM) and each configuration with PIM. Results are

presented at Φ = 0.7 and 0.8. Each result presented is the average of 5 independent

measurements. Figure 3.7 shows flame images for Q = 300 slpm at Ф = 0.7 for baseline case

without and with PIM. Figure 3.8 shows flame images for Q = 300 slpm at Ф = 0.8. Porous insert

of 4 ppcm resulted in interior combustion for both Φ (Figures 3.7(b) and 3.8(b)). Compared to

baseline case, total SPL decreased by 2.9 dB for Φ = 0.7, and increased by 3.2 dB for Φ = 0.8.

Porous insert of 8 ppcm also resulted in interior combustion for both Φ (Figures 3.7(c) and

3.8(c)). For Φ = 0.7, the total SPL decreased by 6.8 dB. For Φ = 0.8, noise level increased by 1.0

dB. Porous insert of 18 ppcm resulted in surface combustion for Φ = 0.7 and reduction in total

SPL by 7.1 dB. For Φ = 0.8, interior combustion was confined to a narrow downstream region of

the PIM insert (Figure 3.8(d)) and the total SPL decreased by 4.0 dB. Results show that the

highest pore density of 18 ppcm is most effective in mitigating the combustion noise. Note that

45

this case resulted in surface combustion (or interior combustion in a narrow downstream region).

Thus, achieving PIM surface combustion mode is important to reduce the total SPLs.

Figure 3.9 presents the power spectra for all configuration at Ф = 0.7. Figures 3.9(a),

3.9(b) and 3.9(g) show a peak at a frequency of about 250 Hz typical of combustion instability.

These cases pertain to the highest SPLs. For the remaining cases, the power spectra are

broadband with no distinct peak, thus, combustion instability without PIM was mitigated by

PIM. Figure 3.10 shows power spectra for all configurations at Φ = 0.8. Figure 3.10(b) shows a

peak at frequency of 450 Hz, and Figures 3.10(c) and 3.10(h) show peaks at frequency of 700

Hz. Figures 3.10(d) and 3.10(g) show peaks at approximately 250 Hz, and power spectra in

Figures 3.10 (e) and 3.10(f) are broadband. Figure 3.11 shows SPL in one third octave band for

all PIM configurations. Results show that the pore density of 18 ppcm is most effective in

mitigating the combustion noise at higher frequencies.

Table 3.1

Summary of results, effect of pore density, Q = 300 slpm

Configuration PIM Ф = 0.7 Ф = 0.8 Pore density

(ppcm)

A None 103.0 dB 103.9 dB N/A

B D38-P4 + D44-P4 100.1 dB 107.1 dB 4

C D38-P8 + D44-P8 96.2 dB 104.9 dB 8

D D38-P18 + D44-P18 95.4 dB 99.8 dB 18

46

3.3.2 Effect of PIM Geometry

By changing PIM dimensions with a fixed pore density, the flow structure can be affected

to favor or avoid the interior combustion. Thus, PIM geometry can be expected to affect the total

SPLs. Experiments were conducted with 18 ppcm porous inserts of different inside diameters.

The PIM geometries studied are: constant, increasing, and decreasing inside diameter in the flow

direction. Table 3.2 lists all cases studied and a summary of the total SPLs.

Figure 3.12 shows the SPL in one third octave band for Q = 300 slpm and different PIM

geometries. PIM configurations D, E and F resulted in very similar total SPLs; 6.0 dB reduction

at Φ = 0.7, and 4.0 dB reduction at Φ = 0.8 compared to the baseline case with no PIM. Note that

these cases had no indication of interior combustion. For these cases, the inside diameter (3.8

cm) of the upstream porous insert is the same as the outside diameter of the swirl injector at the

dump plane. Similar to Configuration F, the Configuration G uses constant inside diameter

porous rings. However, the total SPL decreased only by 2.2 dB at Φ = 0.7 and increased by 1.9

dB at Φ = 0.8. This result is attributed to the increased volume of the free flame region, which

changes the stabilization mechanism by approaching conditions similar to the baseline case, i.e.,

corner recirculation zone accompanied with vortical structures in the shear layer of the flame. In

this case, interior combustion was observed on the inner surface of the porous insert. PIM

configuration H reduced the total SPL by 6.8 dB at Φ = 0.7. However, at Φ = 0.8, the total SPL

increased by 3.9 dB. The downstream confinement of the free flame increases entrainment of

products into the porous insert, which promotes interior combustion mode to increase noise

levels.

47

3.3.3 Effect of Reactant Flow Rate

Experiments were conducted at a higher air flow rate of Q=600 slpm with Ti=120˚C. In

this case, none of the PIM flames at Ф=0.7 experienced interior combustion indicating

sufficiently high flow velocity through the pores. In spite of the high reactant flow rate,

submerged combustion still occurred at Ф = 0.8. However, tests indicate a dramatic reduction in

combustion noise at the higher flow rate. That is, the PIM is very effective in redistributing the

flow in the flame region to suppress heat release fluctuations, for example, originating from the

turbulent vortical structures in the recirculation zones. Figures 3.13 to 3.16 show flame images

for cases with PIM, Q = 300 and 600 slpm at Ф = 0.7 and 0.8. Figure 3.17 shows the power

spectra for configurations A (no PIM), D (divergent), G (constant) and I (convergent) for Q =

600 slpm at Ф = 0.7. Configuration A without PIM shows peaks at frequencies between 350 and

700 Hz. All configurations with PIM resulted in broadband power spectra. Thus, dominant peaks

present without PIM were mitigated with all PIM geometries. Figure 3.18 shows power spectra

for Q = 600 slpm at Ф = 0.8. Several peaks occur without the PIM (Figure 3.18(a)), which is an

indication of combustion instabilities. Figure 3.18(b) shows that the spectral peaks are virtually

eliminated with the use of divergent PIM in the combustor. Total SPL at this high flow rate are

also summarized in Table 3.3.

Figure 3.19 shows SPL in one third octave band for Q = 300 slpm, Ф = 0.7 and 0.8.

Figure 3.20 shows SPL in one third octave band for Q = 600 slpm, Ф = 0.7 and 0.8. For Ф=0.7,

all porous inserts were effective in reducing the noise/instability, with typical reductions in total

SPL of 13 to 14 dB, particularly at higher frequencies. At the higher equivalence ratio, only the

divergent configuration was effective in reducing the total SPL by 13 dB, also by mitigating

power at higher frequencies. Configurations G (constant) and I (convergent) were either

48

ineffective or marginally effective because of the increased propensity for interior combustion.

These results show that the porous insert can effectively mitigate combustion instabilities when

interior combustion can be avoided.

Table 3.2

Summary of results, effect of geometry, Q = 300 slpm

Configuration PIM Ф = 0.7 Ф = 0.8 Inside wall

A None 103.0 dB 103.9 dB N/A

D D38-P18 + D44-P18 95.4 dB 99.8 dB Divergent

E D38-P18 + D50-P18 95.4 dB 99.9 dB

F D38-P18 + D38-P18 95.7 dB 99.8 dB Constant

G D50-P18 + D50-P18 96.1 dB 104.8 dB

H D50-P18 + D38-P18 96.2 dB 106.8 dB Convergent

I D44-P18 + D38-P18 96.2 dB 106.8 dB

49

Table 3.3

Summary of results, effect of flow rate

Configuration Air Flow Rate 300 slpm 600 slpm

PIM / Ф 0.7 0.8 0.7 0.8

A None 103.0 dB 103.9 dB 118.9 dB 120.5 dB

D D38-P18 + D44-P18 95.4 dB 99.8 dB 105.9 dB 107.1 dB

G D50-P18 + D50-P18 96.1 dB 104.8 dB 104.7 dB 120.4 dB

I D44-P18 + D38-P18 96.2 dB 106.8 dB 105.8 dB 115.1 dB

3.3.4 CO and NOx Emissions

Figures 3.21 and 3.22 present the CO and NOx emissions measured for Q=300 slpm, and

Ti=100˚C at Ф=0.7 and 0.8. For Ф=0.7, the CO concentrations for the divergent configuration

are slightly higher and the NOx emissions are comparable for all configurations. For Ф=0.8, all

cases result in comparable CO emissions, but NOx emissions are the highest for the constant area

insert. Interestingly, the NOx emissions for the divergent PIM are comparable to the case

without PIM. For all cases, emissions profiles are nearly flat in the radial direction, indicating

good spatial uniformity of combustion. Overall, results show that the porous insert does not

have an adverse effect on CO and NOx emissions.

Figures 3.23 and 3.24 show the CO and NOx emissions measured at the combustor exit

plane for Q = 600 slpm, Ti=100˚C at Ф=0.7 and 0.8.. For Ф=0.7, Figure 3.23(a) shows that the

CO emissions without the PIM vary from 25 to 30 ppm. With PIM, the CO emissions decrease

50

significantly to below 10 ppm for all porous inserts. The same trend is also observed at Ф=0.8

(Figure 3.24(a)), where the CO emissions with PIM insert are nearly one-third of those without

the PIM. Figures 3.23(b) and 3.24(b) show that the NOx emissions increase slightly with the

PIM insert. Overall, the results are very encouraging and suggest that significant reductions in

noise and pollutant emissions are feasible by a judicial choice of PIM configuration.

3.3.5 Long Duration Experiments

To ensure durability of the porous material used in this investigation, a long-duration test

was conducted, simulating a realistic gas turbine scenario. Test was run continuously during a 4-

hour period. A divergent PIM configuration was used for Q = 600 slpm and Ф = 0.7 because of

promising results obtained in previous section. The objective was to test material endurance over

several hours of operation, to identify damage if any, and its effect on noise, CO and NOx

emissions. Every 30 minutes, CO and NOx emissions data were collected at the center point of

the exit plane and noise data were collected at the combustor exit plane. Results show a steady

flame throughout the duration of the test as reflected by SPL and emissions data. Figure 3.25

shows a plot of SPL in one third octave band taken at 30 minute intervals. Profiles nearly overlap

with each other, indicating no major changes during the experiment. Table 3.4 summarizes the

total SPL throughout the test, and shows that it is nearly constant. Figure 3.26 shows CO and

NOx emissions measured at 30 minute intervals. CO and NOx concentrations are nearly constant

throughout the experiment. Again, these results indicate reliable operation of combustor with

PIM for several hours.

51

Table 3.4

Summary of SPL for long-duration experiment

Time (min) 0 30 60 90 120 150 180 210 240

Total SPL (dB) 106.9 105.7 106.3 105.3 104.2 106.1 105.5 106.1 105.7

3.4 Conclusions

In this study, a novel concept to integrate PIM with swirl-stabilized LPM combustion is

proposed to passively reduce combustion noise. Experimental study shows that the swirl-

stabilized combustion is supplemented with submerged or surface combustion in the presence of

the porous insert. Equivalence ratio, reactant flow rate, and parameters such as PIM pore size and

ID determine the combustion mode. Surface combustion is a desirable mode, while submerged

combustion must be avoided to achieve low SPLs. Results show that a divergent porous insert

with pore density of 18 ppcm can reduce the total SPL by up to 14 dB, depending upon the

reactant flow rate, without adversely affecting the NOx and CO emissions. Furthermore, PIM

insert mitigates combustion instability present without PIM. No noticeable change in noise and

emissions measurements and material properties was observed for long duration tests over a

period of 4 hours. This remarkable performance can be achieved with minimal changes to the

combustor hardware. Thus, the concept presented here is attractive for passive control of

combustion noise and instabilities.

52

Figure3.1. Schematic diagram of experimental setup

Combustion chamber

PIM

Swirler

Premix section

Fuel inlet

Plenum Air inlet

Products

Air heater

53

Figure 3.2. Photos of PIM inserts (a) PIM insert (b) combustor without PIM (c) combustor with

two PIM pieces

(a)

(b) (c)

54

Figure 3.3. Description and schematic diagram of PIM configurations used in this study

Configuration I Pore density: 18 ppcm IDs: 4.4, 3.8 cm

Configuration A Pore density: None IDs: None

Configuration B Pore density: 4 ppcm IDs: 3.8, 4.4 cm

Configuration C Pore density: 8 ppcm IDs: 3.8, 4.4 cm

Configuration H Pore density: 18 ppcm IDs: 5.0, 3.8 cm

Configuration G Pore density: 18 ppcm IDs: 5.0, 5.0 cm

Configuration F Pore density: 18 ppcm IDs: 3.8, 3.8 cm

Configuration E Pore density: 18 ppcm IDs: 3.8, 5.0 cm

Configuration D Pore density: 18 ppcm IDs: 3.8, 4.4 cm

None (Baseline)

55

Figure 3.4. Flame images, (a) without PIM (b) with PIM interior combustion (c) with PIM

surface combustion

(a) (b) (c)

56

Figure 3.5. Schematic diagram illustrating the PIM stabilization mechanism

Surface flame

Reactants

Products

PIM

Swirler

Reactants

Core region flame

57

Figure 3.6. One third octave band SPL for repeatability test

Frequency (Hz)

SP

L(d

B)

0

0

200

200

400

400

600

600

800

800

1000

1000

40 40

60 60

80 80

100 100

120 120

140 140

58

Figure 3.7. Flame images for Q = 300 slpm, Ф = 0.7 (a) Configuration A (b) Configuration B (c)

Configuration C (d) Configuration D (e) Configuration E (f) Configuration F (g) Configuration

G (h) Configuration H (i) Configuration I

(a) (b) (c)

(d) (e) (f)

(g) (h) (i)

59

Figure 3.8. Flame images for Q = 300 slpm, Ф = 0.8 (a) Configuration A (b) Configuration B (c)

Configuration C (d) Configuration D (e) Configuration E (f) Configuration F (g) Configuration

G (h) Configuration H (i) Configuration I

(a) (b) (c)

(d) (e) (f)

(g) (h) (i)

60

Figure 3.9. Power spectra for Q = 300 slpm, Ф = 0.70 (a) Configuration A (b) Configuration B

(c) Configuration C (d) Configuration D (e) Configuration E (f) Configuration F (g)

Configuration G (h) Configuration H (i) Configuration I

(a) (b) (c)

(d) (e) (f)

(g) (h) (i)

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

0.1

0.2

0.3

0.4

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

0.1

0.2

0.3

0.4

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

0.1

0.2

0.3

0.4

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

0.1

0.2

0.3

0.4

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

0.1

0.2

0.3

0.4

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

0.1

0.2

0.3

0.4

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

0.1

0.2

0.3

0.4

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

0.1

0.2

0.3

0.4

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

0.1

0.2

0.3

0.4

61

Figure 3.10. Power spectra for Q = 300 slpm, Ф = 0.80 (a) Configuration A (b) Configuration B

(c) Configuration C (d) Configuration D (e) Configuration E (f) Configuration F (g)

Configuration G (h) Configuration H (i) Configuration I

(a) (b) (c)

(d) (e) (f)

(g) (h) (i)

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

0.1

0.2

0.3

0.4

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

0.1

0.2

0.3

0.4

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

0.1

0.2

0.3

0.4

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

0.1

0.2

0.3

0.4

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

0.1

0.2

0.3

0.4

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

0.1

0.2

0.3

0.4

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

0.1

0.2

0.3

0.4

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

0.1

0.2

0.3

0.4

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

0.1

0.2

0.3

0.4

62

Figure 3.11. One third octave band SPL, effect of pore density, Q = 300 slpm (a) Ф = 0.70 (b) Ф

= 0.8

(a)

(b)

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100050

60

70

80

90

100

110

120

130

No PIM4 ppcm8 ppcm18 ppcm

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100050

60

70

80

90

100

110

120

130

No PIM4 ppcm8 ppcm18 ppcm

63

Figure 3.12. One third octave band SPL, effect of geometry, Q = 300 slpm (a) Ф = 0.70 (b) Ф =

0.8

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100050

60

70

80

90

100

110

120

130

No PIMDivergent 1 (D)Divergent 2 (E)Constant 1 (F)Constant 2 (G)Convergent 1 (H)Convergent 2 (I)

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100050

60

70

80

90

100

110

120

130

No PIMDivergent 1 (D)Divergent 2 (E)Constant 1 (F)Constant 2 (G)Convergent 1 (H)Convergent 2 (I)

(a)

(b)

64

Figure 3.13. Flame images for Q = 300 slpm, Ф = 0.7 (a) Configuration A (b) Configuration D

(c) Configuration G (d) Configuration I

(a)

(b)

(c)

(d)

65



(a)

(b)

(c)

(d)

66



(a)

(b)

(c)

(d)

67



(a)

(b)

(c)

(d)

68

Figure 3.17. Power spectra for Q = 600 slpm, Ф = 0.70 (a) Configuration A (no PIM) (b)

Configuration D (divergent) (c) Configuration G (constant) (d) Configuration I (convergent)

(a) (b)

(c) (d)

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

0.2

0.4

0.6

0.8

1

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

0.2

0.4

0.6

0.8

1

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

0.2

0.4

0.6

0.8

1

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

0.2

0.4

0.6

0.8

1

69

Figure 3.18. Power spectra for Q = 600 slpm, Ф = 0.80 (a) Configuration A (no PIM) (b)

Configuration D (divergent) (c) Configuration G (constant) (d) Configuration I (convergent)

(a) (b)

(c) (d)

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

0.2

0.4

0.6

0.8

1

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

0.2

0.4

0.6

0.8

1

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

0.2

0.4

0.6

0.8

1

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

0.2

0.4

0.6

0.8

1

70

Figure 3.19. One third octave band SPL, effect of reactants flow rate, Q = 300 slpm (a) Ф = 0.7

(b) Ф = 0.8

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100050

60

70

80

90

100

110

120

130

No PIMDivergentConstantConvergent

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100050

60

70

80

90

100

110

120

130


(a)

(b)

71

Figure 3.20. One third octave band SPL, effect of reactants flow rate, Q = 600 slpm (a) Ф = 0.7

(b) Ф = 0.8

(a)

(b)

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100050

60

70

80

90

100

110

120

130


Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100050

60

70

80

90

100

110

120

130


72

Figure 3.21. CO and NOx emissions for Q = 300 slpm, Ф = 0.7, Ti = 100˚C (a) CO (b) NOx

(a)

(b)

Transverse distance (cm)

CO

(ppm

)

-4 -2 0 2 40

5

10

15

20No PIMDivergentConstantConvergent


NO

x(p

pm

)

-4 -2 0 2 40

5

10

15

20

25


73


(a)

(b)


CO

(pp

m)

-4 -2 0 2 40

5

10

15

20

25

30

No PIMConstantConvergentDivergent


NO

x(p

pm

)

-4 -2 0 2 410

20

30

40

50

60

70


74


(a)

(b)


NO

x(p

pm

)

-4 -2 0 2 40

5

10

15

20

25

30



CO

(ppm

)

-4 -2 0 2 40

5

10

15

20

25

30

35


75


(a)

(b)


CO

(pp

m)

-4 -2 0 2 40

20

40

60

80



NO

x(p

pm

)

-4 -2 0 2 40

20

40

60

80No PIMConstantConvergentDivergent

76

Figure 3.25. One third octave band SPL, long duration test

XXXXXX

XXX X

X

X X X X

X

X

*****

***

* **

* * **

*

*

Frequency (Hz)

SP

L(d

B)

0

0

200

200

400

400

600

600

800

800

1000

1000

40 40

60 60

80 80

100 100

120 120

140 140

77

Figure 3.26. CO and NOx emissions for Q = 600 slpm, Ф = 0.7, long duration test (a) CO (b)

NOx

Time (min)

CO

(ppm

)

0

0

60

60

120

120

180

180

240

240

0 0

2 2

4 4

6 6

8 8

10 10

12 12

Time (min)

NO

x(p

pm)

0

0

60

60

120

120

180

180

240

240

0 0

10 10

20 20

30 30

40 40

50 50

60 60

(a)

(b)

78

CHAPTER 4

DEVELOPMENT OF A FACILITY FOR HIGH FLOW RATE, HIGH INLET TEMPERATURE, AND HIGH PRESSURE

COMBUSTION EXPERIMENTS

Background

Combustion research has gained significant importance in recent years. One reason for

this increased importance is strict emissions regulations. Many studies of combustion science are

conducted under atmospheric pressure because of its simplicity of operation. These studies have

made great contribution to the field by understanding fundamental aspects of combustion: flame

propagation, flame instabilities, chemistry, acoustic behavior, etc. Conversely, gas turbine

systems for land-based power generation operate at high pressure conditions. Although low

pressure experiments have made significant contributions to the field, a closer insight to

combustion processes used in typical applications requires studies at high pressures. Such

experiments also require high reactant flow rates and high inlet air temperatures.

The University of Alabama is doing its part to become more environmentally friendly by

conducting research on combustion systems that reduce pollutant emissions. However, the

existing facilities for combustion experiments at UA can operate only at atmospheric pressure.

Thus, there is a need to develop a laboratory facility capable of safely operating high pressure

combustion experiments. This would integrate air and fuel supply systems into a high pressure

and high temperature experimental apparatus which mimics the conditions of a gas turbine. The

air and fuel supply systems should be able to operate at high pressures, be easy to connect and

79

disconnect, and offer flexibility for modifications and upgrades. These conditions will be used to

simulate turbine engine conditions and allow measurements of noise and pollutant emissions for

a range of operating conditions. To achieve this objective, an effective method to supply desired

fuel and air flows must be designed and installed. Flow rate will need to be measured,

interpreted, and used to obtain the desired air-fuel ratio. The combustion air line must include a

heater to simulate temperature at compressor discharge in a gas turbine engine. The combustion

products exiting the experimental apparatus will also need to be safely routed out of the

laboratory.

A key requirement of the new combustion laboratory is that there need to be two

combustion test rigs: one for experiments conducted at atmospheric pressure and one for

experiments conducted at pressures higher than atmospheric. In turn, the high pressure side of

the combustion laboratory must be operated remotely, that is, air and fuel flow rates must be

controlled from a computer located safely away from the experimental apparatus. Combustion

air needs to be preheated by a heater prior to entering the combustion testing apparatus. Air and

fuel supply lines, control equipment (pressure regulator, valves, etc.) and heater need to be

placed at safe locations to avoid tripping hazard, while allowing access for flow control and

adjustments. The general layout of the facility is shown in Figure 4.1. The facility allows

experiments to be performed at pressures ranging from 1 to 10 atm and inlet air temperatures up

to 800 K. The facility consists of three major components: (1) supply and exhaust systems, (2)

instruments and data acquisition, (3) combustion experimental apparatus.

80

4.2 Reactant Supply Systems

4.2.1 Air Lines

The general layout of the facility consists of air and fuel lines routed to the test area in the

designated laboratory space. Air line enters the lab area directly from a dedicated compressor-

tank system, followed by a pressure regulator and filter. A flow control valve is used to regulate

flow rate of air. The flow control system is a FLOWSERVE Kammer Series 5 (model 0350E3-P)

electric actuator and Kammer Total Flow globe valve. The main air supply line then turns back

up a column and runs overhead to the combustion chambers, as seen in Figure 4.2. The

combustion laboratory layout is depicted in Figures 4.3 and 4.4. Next, air supply line is split into

a high pressure system and a low pressure system. Immediately after the split, each line is

equipped with a Nibco carbon steel shutoff ball valve. The high pressure system air line further

splits into two branches: one for combustion air, and another for cooling air. A Nibco Class 150

Bronze Globe valve is placed in each of these lines to manually control flow split. Each of the air

lines is routed to a laminar flow element (LFE) used for flow measurement. Details of the LFE

are provided in instruments and data acquisition section. Combustion air is routed to an electric

air heater, and then to the experimental apparatus. Downstream of the LFE cooling air connects

to the experimental apparatus through a 4-way manifold to distribute the cooling air flow in the

combustion chamber. Table 4.1 shows a list of parts used to assemble the air supply lines.

81

Table 4.1

Air supply line parts

Quantity Description Material Supplier

50 ft 2” schedule 80 Carbon steel Consolidated pipe

16 ft 1” schedule 80 Carbon steel Consolidated pipe

4 ft 1” schedule 80 Stainless steel Consolidated pipe

9 2” 90° elbow schedule 80 Carbon steel Consolidated pipe



2 2” tee schedule 80 Carbon steel Consolidated pipe

2 2” x 1” reducer schedule 80 Carbon steel Consolidated pipe

1 2” x 1” reducer schedule 80 Stainless steel Consolidated pipe

2 6” x 2” reducer schedule 80 Stainless steel Consolidated pipe

1 2” tee 300# schedule 80 Carbon steel Consolidated pipe

1 1” tee 300# schedule 80 Stainless steel Consolidated pipe

10 2” ANSI 300# flange Carbon steel Consolidated pipe

4 2” pipe unions Carbon steel McMaster-Carr

6 1” ANSI 300# flange Stainless steel Consolidated pipe

2 6” ANSI 300# flange Stainless steel Consolidated pipe

2 2” ¼ turn ball valve Carbon steel Nibco

1 2” globe valve Bronze Nibco

1 Air pressure regulator - Ingersoll Rand

82

4.2.2 Electric Heater

An Osram Sylvania 72kW 480 volt 6 inch flanged in-line heater, displayed in Figure 4.5,

is used to heat the air entering the combustion experimental apparatus. Preheating the reactant

air simulates gas turbine engine conditions by mimicking hot air exiting the compression stage.

The heater is a compact, robust industrial electric heat source capable of heating air at high

pressure (150 psi) to 1400 °F. It is supported by two 6 inch 300 lb ANSI Stainless Steel Flanges.

The heater is mounted vertically on the floor near the experimental apparatus. This provides an

interface between the incoming air line and the heated air line connecting to the combustion

chamber. A steel stand was designed and built to support the weight of the heater, as seen in

Figure 4.6. The stand was built from two inch eleven gauge steel square tubing and one quarter

inch steel plate. A gasket is used between the stand and the heater to dampen any vibrations. A

2”x6” 304 stainless steel flanged diffuser is used upstream of the heater to connect the incoming

air line (2” diameter) to the heater. Combustion inlet pipe diameter is 1 inch, thus, a 6”x1”

reducer is used to route air from heater to the combustor inlet. Air temperature inside the heater

is measured by thermocouples connected to an independent controller to obtain the desired air

temperature. Type K (chromel/alumel) thermocouple wire is used to make this connection.

Electrical power to the heater is controlled by an Avatar A3P power controller. An E5CN digital

temperature controller is used to set the temperature of the heater. The control panel is mounted

on the wall away from the combustion chamber.

4.2.3 Fuel Line

Fuel is supplied by a fuel station located outside the building as shown in Figure 4.7. City

NG from the building line is compressed to a rack of ten 50 liter tanks kept at 3000 psi using a

83

Fuelmaker Corp. compressor. NG is supplied to the experimental apparatus from these

pressurized tanks. Fuel line is routed to a pressure regulator to adjust the fuel supply pressure

desired for a specific experimental condition. Then, similar to air line, fuel supply line splits into

two branches: one to high pressure system and one to low pressure system, as shown in Figure

4.1. Fuel line is connected to a solenoid valve (model number) with an electrical cutoff switch

for safety. Fuel is then routed to the experimental apparatus. The fuel line in the high pressure

system has a diameter of 0.5 inches. The fuel line diameter in the low pressure system is 3/8

inches. A Swagelok, stainless steel, ball valve is installed on each line to manually shut off the

fuel in case of an emergency. Fuel rate is controlled and measured by a Bronkhorst mass flow

controller calibrated for 0 to 465 normal liters per minute (lnpm). Figure 4.8 shows the fuel line

routed through the column. The Bronkhorst Controller is attached to the NI CompactRIO data

acquisition and control system using a RS-232 serial connection. Data acquired are processed by

the LabVIEW software to monitor and control the fuel flow rate. The fuel line runs overhead and

then branches into the high pressure system and the low pressure system. The fuel is injected into

the combustion air line where mixing occurs upstream of the pressurized chamber. Table 4.2 lists

the parts used in the fuel supply lines.

4.2.4 Product Exhaust Line

During experiments, hot gases produced from combustion must be safely routed to

outside the building. The layout of the exhaust system is shown in Figures 4.9, 4.10 and 4.11.

(top view and side views) The exhaust system rests atop of high and low pressure testing

apparatus. The main exhaust duct is a stainless steel 12 inch pipe to handle high temperature

combustion products and is connected to an exhaust fan, located on the roof the building.

84

Table 4.2

Fuel supply line parts

Quantity Description Material Supplier

2 ½” ¼ turn ball valve Stainless steel Swagelok

33 ft ½” tube Stainless steel Swagelok

30 ft 3/8” tube Stainless steel Swagelok

1 ½” by 3/8” reducer Stainless steel Swagelok

3 ½” compression fittings Stainless steel Swagelok

3 3/8” compression fittings Stainless steel Swagelok

1 Gas pressure regulator Stainless steel Swagelok

4.3 Instruments and Data Acquisition System

Split of total air flow in combustion and cooling air lines was controlled by a manual

globe valve on each line, as explained above. Air flow rate on each line was measured by a LFE.

The pressure drop across the LFE and absolute pressure in the line are measured to calculate the

air flow rate using the calibration curve provided by the manufacturer. Both LFE’s are

MERIAM, model 50MW20, calibrated for air flow rate of 0-1400 lpm. Outflow from the

differential and absolute pressure transducers is digitized by a LabVIEW based data acquisition

system. An example showing flow rate calculation from these measurements is presented in

Appendix C. Fuel flow rate was controlled and measured by a Bronkhorst EL-FLOW mass flow

controller calibrated for 0 – 465 lnpm of methane. A RS-232 computer interface provided by the

manufacturer was used to control this instrument. A list of instruments used to measure fuel and

air flow rates is given in Table 4.3.

85

Table 4.3

List of instruments for flow measurement

Description Model Range Accuracy*

Laminar flow element

(LFE)

MERIAM

50MW20 1400 lpm 0.72% reading

K-type thermocouple Omega

KQSS-14G-10 1250 °C 0.75% reading

Absolute pressure sensor Omega

MMA150 150 psi 0.20% reading

Differential pressure sensor Omega

MMDDU10WC 10 inH2O 0.03% reading

Mass flow controller Bronkhorst

EL-FLOW 465 lnpm of CH4

0.5% reading +

0.1% full scale

*Specified by manufacturer

CompactRIO system from National Instruments is used for data acquisition and control.

The CompactRIO system offers a balance of versatility, reliability, and compact design. The

CompactRIO data acquisition system, displayed in Figure 4.12, consists of a real-time controller,

a chassis, and four modules. The real-time controller, which acts as a stand-alone PC, has a 533

MHz processor, 2 GB storage, and 256 MB RAM. Most of the software program runs on this

controller. The reconfigurable chassis holds up to eight modules and houses the field-

programmable gate array (FPGA) chip. Because of its reliability, most of the signal conditioning

86

is done on the FPGA chip. The current configuration utilizes four modules: a 4-channel 20 mA

current output, an 8-channel 20 mA current input, a 32-channel 10 V voltage input, and a 16-

channel thermocouple input. The CompactRIO system connects to the operator PC via Ethernet

from the real-time controller.

The flow control system utilizes LabVIEW 8.6 software. LabVIEW interfaces with the

sensors and controllers via the CompactRIO data acquisition system. A layout of the sensors,

controllers, and CompactRIO system is depicted in Figure 4.13. As discussed previously,

cooling and combustion air lines each houses and LFE, which in turn is equipped with an

absolute pressure transducer, a differential pressure transducer, and K-type thermocouple. The

LabVIEW interface is programmed to interpret the sensor input and then generate a signal for the

flow control valve to adjust the total air flow rate. The manual control valve on each line is used

to fix the ratio of combustion air to cooling air, eliminating the need for separate actuated valves.

The combustion air line also includes an electric heater, which is controlled by a separate power

controller supplied by the manufacturer. The fuel line contains a mass flow controller, which is

controlled by the software provided by the manufacturer.

Programming the system using LabVIEW is a three step process that involves the FPGA,

the real-time controller (RT) and the PC. The first step is programming the FPGA chip

embedded in the chassis of CompactRIO system. It determines the initial conditioning of the

signal, the sampling rate and further routing of the signal. The FPGA chip is a reliable part of the

CompactRIO system because when a new program is sent to the chip, it reconfigures its gates to

embed the program into hardware. The next programming step is the real-time controller, which

can act as a stand-alone PC without a display. By embedding it on the real-time controller, the

program has a dedicated processor along with its own memory and storage allowing it to run

87

independent of the PC. This functionality reduces the risk of hardware crash and speeds up the

entire process. Most of the calculations are conducted on the real-time controller. It receives the

voltage or current signal, scales it to the sensors’ calibration curve to produce useful parameters

and uses these parameters to calculate the mass flow rate of each air line. The final step is to

program the PC itself. Since most of the program runs on the CompactRIO system, the PC

program is primarily used to monitor the system, displaying all of the properties measured and

providing a link to input values of the variables to the controllers. This is done by displaying

indicators for the shared variables already defined.

4.4 Combustion Experimental Apparatus

The combustion experimental apparatus consists of supply of preheated air, supply of

fuel, a premixing section, a burst disk section, a pressurized compartment where the combustion

liner is located, a concentric reducer and a nozzle exit. Figure 4.14 shows a schematic of the

assembled experimental apparatus, and Figure 4.15 shows an exploded view of the experimental

apparatus. Air supply line is a 1 inch schedule 80 stainless steel pipe attached to the exit of the

electric heater. Combustion air lines between electrical heater and combustion the combustion

chamber are insulated to reduce heat loss, as shown in Figure 4.16. Combustion air supply line

connects to a flanged tee, where it splits into a burst disk section and the premixing section. The

burst disk is a safety feature of the system. In the event of a sudden increase in pressure in the

combustion chamber, the burst disk would break rerouting incoming reactants. High pressure in

the chamber would force reactants to the atmospheric pressure exit away from the reaction zone

in the combustion chamber. The nominal burst pressure of the disk is 150 psi. Thus, if this

pressure is reached in the combustion system, disk will burst to release the pressure.

88

In normal operation, combustion air enters the premixing section, where fuel is injected.

The premixing section is a 30 inch long pipe located upstream of the combustion chamber. Fuel

enters the premixing section through a side hose. A section of porous material of 4 ppcm is

inserted in the premixing section to improve mixing of fuel and air. The premixing section

allows reactants to mix prior to entering the reaction zone. Thus, a homogeneous mixture with

nearly a constant equivalence ratio is obtained to avoid localized rich and/or lean mixture regions

that may cause auto ignition and/or flame blow off. Reactants in the premixer pass through a

swirler prior to entering the combustion chamber. In the premixer section, wall static pressure

and temperature are measured upstream of the swirler. The measured temperature pertains to

inlet temperature. Wall pressure measurement is used to determine the pressure drop across the

swirler/combustor. The swirling reactant flow enters the combustion chamber.

The combustion chamber consists of a quartz cylinder placed on the dump plane, as seen

in Figure 4.17. The quartz cylinder is 30 cm (12 inches) long by 7.0 cm (2.75 inches) inside

diameter. The quartz cylinder is secured by a holder plate on the downstream edge of the quartz.

The holder plate is secured mechanically by threaded rods screwed onto the dump plane, holding

the quartz against it (refer to figures). The combustion chamber is surrounded by a pressurized

compartment from now on referred to as enclosure. The enclosure is a stainless steel 15 inch

outside diameter cylinder with 3 inch thick walls. The enclosure is 60 cm (24 inches) long. The

enclosure has two rectangular windows for optical access to the combustion chamber and a total

of 12 access ports on opposite sides for instrumentation. The ports are ½ inch NPT threaded

holes to mount various probes. Windows can be covered with 3.8 cm (1.5 inch) thick quartz

glass designed to withstand chamber pressure. Quartz windows are held in place by a stainless

steel window frame. On the upstream side, the enclosure seals with a plate referred to as plenum

89

base. Plenum base and enclosure have bolt patterns of an 8 inch 300-lb flange for assembly. Seal

is provided by high temperature resistant gasket.

Enclosure and plenum base assembly is designed so that dump plane of combustor is

horizontally aligned with the bottom edge of the windows, thus, dump plane is raised from the

plenum base. This design serves two purposes: (1) it provides maximum optical access to the

combustion chamber and (2) it exposes dump plane flange directly to incoming cooling air.

Cooling air then impinges on dump plane flange to enhance cooling. Four ¾ inch hoses are

connected on the outside face of the plenum base to supply cooling air, as seen in Figure 4.18.

On the downstream side, enclosure is connected to a concentric reducer, with 8 inch 300 lb bolt

pattern. Downstream of the concentric reducer, a 3 inch 300 lb flange is located for nozzle

assembly. A converging nozzle is machined to choke the flow, thus building pressure in the

enclosure. Specifications of the components of the experimental apparatus explained above are

shown in Figures 4.19 to 4.28.

A sampling probe is used to measure combustion emissions. The sampling probe is

divided into two sections: an L-shaped section made of quartz with tapered tip for reaction

quenching, and a stainless steel straight section. Figure 4.29 shows an illustration and photograph

of the sampling probe. Straight stainless steel section runs through the wall of enclosure. Tip of

the L-shaped section is exposed to combustion products. These two parts connect with the use of

a stainless steel Swagelok nut-ferrule connection without exposing fitting to high temperatures.

The assembled two-piece probe seals to the wall of the enclosure with the use of a Swagelok nut-

ferrule fitting. This fitting is manually loosened to traverse probe, then tightened again.

Access to the combustor, for example, to place a porous insert in the combustor, is

facilitated by removing the concentric reducer. Thus the combustor is accessed vertically from

90

above. Removal of windows allows access from the sides of the enclosure, however it is not

recommended to remove windows on a regular basis. Once reducer is dismounted, quartz

holding mechanism is disengaged from the dump plane. Thus, the quartz cylinder combustion

chamber can be removed.

91

Figure 4.1. General Schematic of high pressure combustion laboratory

92

Figure 4.2. Layout of air flow control system

Dimension in inches

93

Figure 4.3. Layout of high pressure combustion laboratory

Dimension in inches

94

Figure 4.4. Layout of high pressure combustion laboratory

Dimension in inches

95

Figure 4.5. (a) Photographic image of combustion air pre-heater (b) Schematic diagram of

combustion air pre-heater

(a)

(b)

Flow

Dimensions in inches [mm]

96

Figure 4.6. Heater stand

Holes for securing heater to stand

Holes for securing stand to floor

97

Figure 4.7 Fuel station

Fuel tanks (10)

Portable fuel tanks

Compressor

98

Figure 4.8. Layout of fuel flow control system

Dimensions in inches

99

Figure 4.9. Exhaust side view


Exhaust pipe

Combustor

100

Figure 4.10. Exhaust overhead view


Low pressure combustor

High pressure combustor

Exhaust pipe

101

Figure 4.11. Exhaust front view


High pressure combustor

102

Figure 4.12. CompactRIO system

Real -time controller

Current in

Current out

Temperature in

Voltage in Chassis with 4 extra slots

103

Air/Fuel

Mix

Cooling

Air

Compressed

Air In

Electronic

Flow Control

Valve

Mass-flow

Controller

Manual

Control ValvesAbs. Pressure

Transducer

Diff. Pressure

Transducers

Type K

Thermocouples

Current

Out

Temp.

In

Current

In

Natural

Gas In

RS232Serial

Electric Heater

Digital Heater Control

Ethernet

To PC

CompactRIO DAQ System

Figure 4.13. Sensor/controller and CompactRIO Layout

104

Figure 4.14. Schematic of assembled experimental apparatus

Concentric reducer

Window cover

Window

Access ports

Premix section

Fuel inlet

Air inlet Burst disk section

105

Figure 4.15. Exploded view of experimental apparatus

Window cover

Combustion chamber

Dump plane

Plenum base

Premix section

106

Figure 4.16. Photographic image of experimental apparatus

Enclosure

Window cover

Access ports

Exhaust pipe

Electric heater

Nozzle

Air inlet Burst disk section

107


Electric heater

Solenoid valve

Fuel line

Fuel inlet

Concentric reducer

Nozzle

Combustion chamber

Window

Dump plane

108


Fuel inlet

Cooling air hoses

Premix section

Window cover

Enclosure

109

Figure 4.19. Details of assembled plenum base

110

Figure 4.20. Details of support pipe/flange

111

Figure 4.21. Details of plenum base

112

Figure 4.22. Details of enclosure

113

Figure 4.23. Details of faces of enclosure

114

Figure 4.24. Details of cross section of enclosure

115

Figure 4.25. Details of windows on enclosure

116

Figure 4.26. Details of ports on enclosure

117

Figure 4.27. Details of window covers

118

Figure 4.28. Details of windows

119

Figure 4.29 Schematic diagram and photograph of sampling probe

Stainless steel section Quartz section

Union

To gas analyzer

Fittings

To gas analyzer

Fitting (connects to wall of enclosure) Union

Gas sample

120

CHAPTER 5

REDUCTION OF COMBUSTION NOISE AND INSTABILITIES WITH THE USE OF POROUS INERT MATERIAL

Background

The stabilization mechanism of a swirl-stabilized combustor consists of inducing a

tangential or swirl velocity prior to reactants entering the combustion chamber. A sudden

expansion of the cross sectional area in the direction of the flow takes place as reactants enter the

combustion chamber. This sudden expansion of the swirling flow creates corner and central

recirculation zones, which in turn allow heat transfer from products to reactants, providing

ignition energy to incoming reactants. Corner recirculation zone is presumably the main source

of noise in swirl-stabilized flame, due to high turbulent fluctuations dominant in this region. A

numerical study presented in Chapter 2 revealed that use of porous insert in the corner

recirculation region redistributes the flow, mitigating highly turbulent structures that generate

noise.

Combustion instability is the result of the coupling of pressure waves generated by

unsteady heat release rate with those generated by oscillating pressure excitation of air

surrounding reaction zone. Typically active methods involving costly equipment that are difficult

to operate are used to control noise and instabilities. This investigation presents a passive

technique, requiring no control during operation, to reduce combustion noise. Furthermore,

combustion noise and combustion instabilities are different outcomes of the same fundamental

121

problem. Thus, reduction in combustion noise is a step forward to mitigate combustion

instabilities. The combustion process in gas turbines for power generation occurs at an elevated

pressure to improve thermal efficiency of the system. Previous experiments presented in Chapter

3 were conducted at atmospheric pressure, low reactant flow rates, and low air inlet

temperatures. Although these conditions are not typical of gas turbine combustors, the use of

PIM as a passive technique indicated significant reduction in combustion noise, and

consequently in combustion instabilities. In this chapter, experiments are conducted at conditions

replicating gas turbine combustor operation, with reactant inlet temperature of up to 260 °C, and

average reactant inlet velocity of up to 76 m/s, at fuel lean conditions with equivalence ratio

varying from 0.65 to 0.75. Corresponding, the reactant inlet Reynolds numbers vary from 2x104

to 11x104(see Appendix E for details).

The present approach is similar to that presented in Chapter 3, i.e., a passive technique

that combines the features of a swirl stabilized flame with those of introducing a PIM in the

combustor to reduce combustion noise and instability. An experimental investigation simulating

conditions of a gas turbine engine, combining use of swirl flow and PIM to reduce combustion

noise and instabilities is lacking in the literature. Mechanical damage in gas turbine engines

caused by pressure fluctuations, driven by combustion instabilities, has been a recurring

problem. Down time and repairs associated with these problems are costly, and damages can be

catastrophic. This study finds relevance since the technique presented can significantly mitigate

combustion noise and instabilities at source. Specifically, PIM changes the velocity profiles and

turbulent structures inside the combustion chamber favorably, in such a way that turbulent

fluctuations are mitigated to eliminate dominant frequencies that produce noise and instabilities.

122

5.2 Experimental Setup

The experimental apparatus described in chapter 4 is shown schematically in Figure 5.1.

A compressor and dryer assembly supplied dry air at 200 psi. A ball valve and pressure regulator

were used to reduce pressure down to 60 psi. A control valve (Kammer model 0350E3-P) was

used to regulate the total air flow rate. Air supply was split into two lines: combustion air and

cooling air lines. Combustion air is mixed with fuel and participates in the combustion reaction.

Cooling air maintains the apparatus at a safe temperature and it does not participate in

combustion. Downstream of the split, globe valves on each air line were used to control the flow

rate in each line. Downstream of each valve, combustion and cooling air flows were measured

independently with identical LFE calibrated for 0 to 1400 lpm. Pressure drop across the LFE was

measured by a differential pressure transducer. An absolute pressure transducer was used to

measure the absolute pressure of air passing through the LFE. A k-type thermocouple was used

to measure the temperature of the air passing through the LFE. The flow rate measured by the

LFE is corrected for temperature and pressure as specified by the manufacturer (see Appendix C

for details).

Next, cooling air was routed directly to combustion apparatus. Flow of cooling air

surrounds the combustion chamber, inside an enclosure. Combustion air was routed through an

electrical heater (Osram Sylvania 72 KW model number 073377) used for preheating prior to the

air entering combustion apparatus. Downstream of the electric heater, pipes were insulated to

reduce heat loss. Combustion air then entered the premix section, where fuel is injected. The

premix section is a 60 cm (24 inches) long, 2.5 cm (1 inch) stainless steel schedule 80 pipe with

ID = 2.44 cm (0.96 inches) and OD = 3.35 cm (1.32 inches). Fuel and air mixed prior to entering

the combustion chamber.

123

Fuel is supplied by a fuel station located outside the building. City NG from the building

line is compressed to 3000 psi and stored in a rack of ten 50 liter tanks. Figure 5.2 shows a

photograph of the fuel station. NG is supplied to the experimental apparatus from these

interconnected pressurized tanks. A pressure regulator is used to control supply pressure down to

100 psi. Fuel flow rate is controlled by a Bronkhorst mass flow controller calibrated for 0 to 465

normal liters per minute (lnpm). For safety, a solenoid valve with an electrical cutoff switch is

placed in the fuel line. Fuel is then routed to the premixer section, where it mixes with

combustion air.

In the premixer section, a swirler resides 2.5 cm (1 inch) upstream of the dump plane as

depicted in Figure 5.3. The swirler had six vanes positioned at 49° to the horizontal, as depicted

in Figure 5.4. The swirl number (S) is 0.9 (see Appendix B for calculation) Reactants enter the

combustion chamber, a 30 cm (12 inches) long quartz tube with inside diameter of 7.3 cm (2.9

inches), where reaction occurs. The combustor is back-side cooled by the flow of cooling air

around it. Pressure drop across the swirl injector/combustor is measured by absolute pressure

transducers mounted at two locations: (1) on the premix section upstream of the injector and (2)

on the wall of the enclosure, as depicted in Figure 5.5. The pressure data are reported with

measurement uncertainty of ±0.2 KPa (see Appendix G). The porous inserts used in this study

were supplied by Ultramet. Two pore densities were used: 18 pores per cm (ppcm) and 32 ppcm.

All pieces used were of the same dimension and were placed on the dump plane of the

combustor. PIM inserts are 5 cm tall, 7 cm outside diameter (OD). Previous testing demonstrated

that a diverging shape of the inside wall provides best performance in terms of noise reduction.

Thus, porous inserts for this investigation were fabricated with tapered inside wall, as shown

124

schematically in Figure 5.6. Porous insert was kept in place by press fitting the PIM into the

quartz chamber using a temperature resistant carbon foil, provided by Ultramet.

The product gas was sampled continuously by a quartz probe (OD = 5.0 mm) mounted on

the enclosure and traversed in the radial direction across the combustor. The probe was mounted

in two steps: (1) straight stainless steel section of the probe was connected to a fitting on wall of

enclosure on designated location (see Figure 5.5); (2) L-shape quartz section was connected to

straight stainless steel section using a union and fitting. The upstream tip of the probe was

tapered to 1 mm ID to quench reactions inside the probe. Figure 5.7 shows a diagram and

photograph of the sampling probe. The gas sample passed through an ice bath and water traps to

remove moisture prior to entering the gas analyzers. The dry sample was routed through

electrochemical analyzers to measure concentrations of CO and NOx in ppm. The analyzer also

measured oxygen and carbon dioxide concentrations used to cross-check the equivalence ratio

obtained from the measured fuel and air flow rates. The gas analyzer manufacturer reports a

measurement uncertainty of ±2 ppm on dry basis. Flame was ignited with the use of a high-

voltage spark generator. Poles of the generator were exposed to reactant flow and then generator

was activated to create a strong spark between the poles to ignite reactants. A low reactant flow

rate was used during ignition to avoid sudden pressure increase that could damage the sampling

probe.

For open top experiments, sound pressure data were measured outside the enclosure,

aligned with the exit plane at different locations. For restricted top experiments, sound pressure

data were measured at two different locations simultaneously, targeting two different sources of

noise: (1) within the enclosure, hereon referred to as combustion noise, (2) outside the enclosure,

hereon referred to as jet noise. Combustion noise data were acquired using a Kistler water-cooled

125

pressure sensor (model 601B1) with sensitivity of 1.132 pC/psi. A signal conditioner (model

Kistler 5051) converts the sensor output to voltage signal, digitized by the LabVIEW data

acquisition system. Pressure sensor is mounted on the wall of the enclosure, as illustrated in

Figure 5.5. Sensor probe for combustion noise measurements is exposed to pressure oscillations

resulting from coupling (or cancelling) of acoustic waves with waves driven by unsteady heat

release rate. Combustion noise sensor is also exposed to cooling air flow, thus to the turbulent

fluctuations associated with it. Jet noise is measured outside of the enclosure, thus, the probe is

exposed to pressure fluctuations induced by the high velocity jet at the nozzle exit. Jet noise

pressure data were collected using a Brüel & Kjær condenser microphone probe (Model 4189-A-

021) located at different positions outside the enclosure (sound pressure data for open top

experiments were acquired with the same instrument). The measured voltage signal is converted

to pressure fluctuation data using the calibrated probe sensitivity of 45.8 mV/Pa. Measurements

were acquired at sampling rates of 2000 and 4000 Hz. Signals from both instruments are

processed using FFT analysis to obtain the sound power spectra. A Matlab script (see Appendix

F for details) was written to compute the total sound pressure level (SPL) in decibels (dB), given

by (Bussman, 2001):

�� = 10 ∗ log��

Where Pref = 20 µPa. Total SPL was calculated by:

�� = 10 ∗ log��∑ 10�.�∗��

(5.1)

(5.2)

126

Where SPLi is the SPL at each frequency band, and n is the number of frequency bands. One-

third octave frequency bands were used, from bands 13 to 29 for sampling rate of 2000 Hz and

from 13 to 32 for sampling rate of 4000 Hz.


Previously conducted experiments (Chapter 4) indicated a reduction of combustion noise

with the use of porous inserts in the reaction zone of a swirl stabilized flame. Reduction in noise

was achieved by re-distributing the flow of reactants and products within the combustion

chamber. Presence of porous material offers resistance to the incoming flow of reactants, thus, a

limited amount of flow penetrates the porous insert to eliminate the corner recirculation zone.

Combustion noise can be generated by the turbulent fluctuations dominant in the corner

recirculation zone. Thus, it follows that elimination of the corner recirculation zone would result

in reduced sound pressure levels. The flame is sustained and confined within the core region of

the porous insert. Products of this gas flame penetrate the porous insert through the inside wall

and mix with the reactants flowing through the PIM. The mixture of reactants and products exits

the porous insert, and ignites to sustain flamelets at the downstream surface of PIM, as illustrated

in Figure 5.8. Noise reduction can be expected to depend upon PIM pore density, and location

and geometry of the porous insert. PIM must target highly turbulent regions within the

combustor to be effective. Preliminary experiments indicated two types of free flame: (1)

anchored within the premixer tube downstream of the swirler, (2) stabilized within the quartz

combustor. Flame anchoring within the premixer tube was undesired because flame propagation

in turbulent recirculation zone downstream of swirler resulted in very high SPL. Increased

reactant flow rate resulted in flame stabilized within the combustor, which reduced the SPL.

127

Previous experiments (Chapter 3) were conducted at relatively low reactant inlet velocities (10

m/s). In this section, experiments were conducted at high reactant inlet velocities, closely

simulating inlet conditions of gas turbine engines. Experiments were conducted with combustion

air flow rates (Q) of 1020, 1400 and 2040 standard liters per minute (slpm), which yield bulk

inlet axial velocities of 30 to 76 m/s at atmospheric pressure (see Appendix E). Experiments

were conducted with no porous insert as the baseline case and then with tapered inserts of pore

density of 18 and 32 ppcm. Tests were conducted with open top (i.e., without the nozzle section

downstream of the combustion chamber) and with restricted top, whereby a nozzle section was

placed downstream of the enclosure.

Sound pressure measurements were taken to examine noise trends at different operating

conditions. First, with open top, five sets of data were taken to demonstrate repeatability of

measurements for Q = 1020 slpm, Ф = 0.7 and Tinlet = 20 °C and no supply of cooling air (Qc).

Microphone probe was aligned with the exit plane of the enclosure, as shown schematically in

Figure 5.9. Data were taken at distance (d) of 25 cm from axis of combustor. Figure 5.10 shows

that one third octave band SPL overlap one another, indicating measurements are repeatable.

Next, SPL was measured at 5 radial locations (d). The combustor was operated at Ф = 0.7 with

no supply of cooling air and no porous insert, Q = 1020 slpm and Tinlet = 20 °C. Table 5.1

provides a summary of the total SPL measured for each case. Figure 5.11 shows the measured

SPL spectra in one-third octave band. SPL spectra show similar trends for each case, while total

SPL decreases with increase in distance (d) of the measurement point. This result is expected due

to atmospheric attenuation of sound. Next, with restricted top, SPL was measured at 16 different

points surrounding the jet as shown schematically in Figure 5.12. Experiments were conducted

for Q = 1020 slpm, Ф = 0.7 with a porous insert of 18 ppcm and Tinlet = 20 °C. Cooling air flow

128

rate was kept at Qc = 990 slpm. Table 5.2 shows a summary of results, and Figure 5.13 shows the

SPL spectra in one third octave band for all positions. Results show that SPL increases slightly in

the flow direction. In the radial direction, SPL decreases with increased distance from the center

of the jet. For all cases, the profiles show same trend, with identical frequencies at peak power.

Table 5.1

Effect of microphone location on SPL for open top experiments, Q = 1020 slpm, Ф = 0.7, Tinlet =

20 °C

Microphone

position (d) (cm)

Total sound

pressure level (dB)

26 118.0

34 116.2

49 111.6

79 110.4

139 104.1

129

Table 5.2

Effect of microphone location on SPL for restricted top experiments, Q = 1020 slpm, Ф = 0.7,

Tinlet = 20 °C, Qc = 990 slpm

Microphone position #

(see Figure 5.12)

Total sound pressure

level (dB)

11 103.1

21 103.8

31 104.6

41 105.3

12 101.6

22 102.2

32 103.0

42 103.8

13 100.9

23 101.4

33 102.3

43 102.9

14 99.7

24 100.2

34 100.6

44 101.0

130

5.3.1 Open Top Experiments

(a) Effect of Pore Density

Experiments were conducted with open top enclosure for two combustion air flow rates

Q =1020 and 1400 slpm, pertaining to mass flow rates of 1.30 and 1.80 Kg/min, respectively.

Tests were conducted at three equivalence ratios (Ф): 0.65, 0.70 and 0.75 and three combustion

air inlet temperatures of Tinlet = 20, 130 and 260 °C. Experiments were conducted without porous

insert as the baseline case and with 18 ppcm and 32 ppcm porous inserts to investigate the effect

of pore size. Sound pressure measurements were taken at the exit plane of enclosure at d = 25

cm, as depicted in Figure 5.9. Figure 5.14 presents the power spectra without porous insert for Q

= 1020 slpm, Ф = 0.65, and different air inlet temperatures. Power spectra are presented in this

study to recognize dominant frequencies, if any. Thus, data are presented in linear scale, as

opposed to log scale involving dB. For Tinlet = 20 and 130 °C, power spectra appears broadband,

yet, at Tinlet = 130 °C, an instability is observed at 525 Hz. For Tinlet = 260 °C, there is dominant

frequency at 750 Hz signifying combustion instability, i.e., heat release oscillations are coupled

with acoustic oscillations. Observation during experiments also indicated instability: combustion

roar was noticeably louder with high pitch. Figure 5.15 shows the power spectra at Ф = 0.7 for

all air inlet temperatures. Results for Tinlet = 20 °C show broadband power spectra with a minor

peak centered around 400 Hz. However, for Tinlet = 130 °C the instability is dominant at 525 Hz.

At Tinlet = 260 °C, power spectra is broadband and no instability is present. Figure 5.16 shows

power spectra for Ф = 0.75. Instability is present for Tinlet = 130 °C only, but no distinct peak is

observed at other inlet air temperatures. These results show that changes in the equivalence ratio

and/or inlet air temperatures can manifest instabilities in the combustor.

131

Figure 5.17 shows power spectra for experiments conducted with porous insert of 18

ppcm, Q = 1020 slpm and all inlet air temperatures. Results show significant reduction in SPL

for all inlet air temperatures and, most importantly, instabilities present without porous insert

were completely mitigated. Flame under these conditions was confined within the inside walls of

the porous insert, with combustion also occurring on the downstream surface of the porous

insert. Thus, No combustion within the PIM (interior combustion) was present. Figures 5.18 and

5.19 show results with porous insert for Ф = 0.70 and 0.75 respectively. Results are similar to

those for Ф = 0.65, i.e., the dominant peak frequency indicative of combustion instability, was

eliminated by the use of porous insert. Total SPL was also reduced with use of PIM, and

reduction was particularly large for cases with instability in the baseline case as summarized in

Table 5.3. Results indicate SPL reduction of 4 to 9 dB. Figure 5.20 to 5.22 show power spectra

for cases studied with 32 ppcm porous insert. Again, porous insert eliminates combustion

instabilities and reduces total SPL, also summarized in Table 5.3. Use of 32 ppcm porous insert

also reduces SPL by up to 9 dB, principally for cases when combustion instability is present with

no porous insert. SPL obtained with 18 ppcm and 32 ppcm are comparable, thus, the effect of

pore size is not significant. Porous inserts of both pore densities are effective in reducing SPL.

Figure 5.23 shows SPL in dB versus one third octave bands for cases with Ф = 0.65 and Tinlet =

20, 130 and 260 °C. Porous insert mitigates fluctuations at higher frequencies: 400 Hz and

above. Both 18 and 32 ppcm porous inserts resulted in SPL reduction; however, difference

between them is not significant as discussed above. Figures 5.24 and 5.25 depict similar results

for Ф = 0.70 and 0.75, respectively. Clearly, porous insert reduces SPL by mitigating combustion

instabilities.

132

Figure 5.26 shows CO emissions for Ф = 0.65, Q = 1020 slpm, and all inlet air

temperatures. Results show that CO emissions without and with porous insert are within

uncertainty of measurements. Thus, PIM does not affect CO concentrations. Figure 5.27 shows

CO emissions for Ф = 0.75. Again, CO concentrations without and with PIM are within the

uncertainty of measurements. Higher CO concentration on one side of the combustor in this case

is likely caused by imperfect mixing of reactants entering the combustion chamber. Figures 5.28

and 5.29 present NOx concentrations for Ф = 0.65 and 0.75 respectively, with air Tinlet = 20, 130

and 260 °C. Similar to CO emissions, porous insert does not affect NOx formation. Thus NOx

concentrations remain within measurement uncertainties for cases without and with PIM.

Increased reactant inlet temperature increases thermal NOx production, which results in higher

NOx concentrations.

Pressure drop across the swirl injector and combustor was measured for all cases. Pinlet

pertains to pressure upstream of the swirl injector, and Pchamber refers to pressure in the enclosure,

near the combustor exit. Figure 5.5 illustrates probe locations for Pinlet and Pchamber measurements.

Table 5.4 presents a summary of pressure measurements (absolute pressure) for all cases studied

for Q = 1020 slpm. Pressure drop across the injector increases with increase in inlet temperature,

as seen in Figure 5.30. At Tinlet = 260 °C, pressure drop of 15.5 KPa is the largest registered

value. Porous inserts of 18 and 32 ppcm have negligible effect on the pressure drop. This is an

important result indicating that the porous insert does not incur a pressure loss penalty.

133

Table 5.3

Summary of sound pressure levels for Q=1020 slpm

Tinlet (°C) Ф Qc (slpm)

Total sound pressure level (dB)

No PIM 18 ppcm PIM 32 ppcm PIM

20

0.65 990 117.0 110.2 111.4

0.70 990 120.0 111.7 111.0

0.75 990 120.6 114.9 113.2

130

0.65 1160 120.5 113.9 114.7

0.70 1160 125.8 116.4 116.4

0.75 1160 124.3 118.3 116.2

260

0.65 1350 122.1 117.4 118.2

0.70 1350 119.0 119.0 118.2

0.75 1350 118.8 119.8 118.8

134

Table 5.4

Summary of pressure measurements for Q = 1020 slpm

Tinlet (°C)

Ф

No PIM 18 ppcm PIM 32 ppcm PIM

Pinlet (KPa)

Pchamber (KPa)

∆P (KPa)

Pinlet (KPa)

Pchamber (KPa)

∆P (KPa)

Pinlet (KPa)

Pchamber (KPa)

∆P (KPa)

20

0.65 102.8 100.6 2.2 102.7 101.3 1.4 100.6 100.0 0.6

0.70 102.8 100.6 2.2 103.4 101.3 2.1 100.6 100.0 0.6

0.75 102.8 100.0 2.8 103.4 101.3 2.1 101.3 99.3 2.0

130

0.65 111.7 100.6 11.1 110.3 101.3 9.0 107.5 100.6 6.9

0.70 111.0 100.6 10.4 110.3 102.0 8.3 107.5 100.6 6.9

0.75 111.7 102.0 9.7 111.7 101.3 10.4 108.9 100.6 8.3

260

0.65 112.4 101.3 11.1 115.8 102.0 13.8 113.1 102.0 11.1

0.70 111.7 101.3 10.4 117.2 102.0 15.2 115.1 100.3 14.8

0.75 113.8 102.0 11.8 117.2 102.0 15.2 115.8 100.3 15.5

(b) Effect of Flow Rate

Next, effect of PIM was investigated for reactant flow rate, Q = 1400 slpm. This high

flow rate results in average reactant inlet axial velocities of 30 to 52 m/s. Tests were conducted

without porous insert and with porous insert of 18 ppcm only, because of the similarity between

18 and 32 ppcm PIM results discussed previously. Figure 5.31 shows power spectra for Ф = 0.65

of all reactant inlet air temperatures for tests without porous insert. Tinlet = 130 °C resulted in

highest SPL because of a dominant frequency at approximately 650 Hz. This dominant

135

frequency was also present for cases with Ф = 0.70 and 0.75, as seen in Figures 5.32 and 5.33

respectively. For cases with Tinlet = 20 and 260 °C, power spectra was broadband with no distinct

peaks. Thus, the total SPL for Tinlet = 130 °C was highest for all Ф. Similar to cases with lower

reactant flow rates, porous inserts resulted in significant reduction of SPL and combustion

instability. Figure 5.34 shows power spectra with 18 ppcm PIM. Results show that for all Tinlet,

power spectra were broadband. Most importantly, dominant frequency present at Tinlet = 130 °C

indicative of combustion instability was completely mitigated. Similarly, Figure 5.35 shows that

dominant frequency of combustion instability at Ф = 0.7 was mitigated by porous insert. Figure

5.36 (c) shows a peak at 400 Hz with the porous insert. In this case, localized interior combustion

was present at the downstream surface of the porous insert. This finding is consistent with

previous results in Chapter 3 indicating that PIM interior combustion is not desirable. Despite an

increase in reactant velocity by 40%, PIM remains effective in reducing SPL and mitigating

combustion instability.

Length of the porous insert is an important geometric parameter. Preliminary

experiments, not shown here, demonstrated that PIM with shorter length (2.5 cm) had small

effect in reducing SPL. With a longer porous insert, a larger percentage of products of the

confined free flame penetrate the PIM through the inner wall. Thus, heat is transferred and

distributed more effectively by PIM. Table 5.5 presents a summary of total SPL obtained for all

cases with Q = 1400 slpm. Results indicate that, similar to Q = 1020 slpm, PIM reduces the total

SPL and mitigates combustion instability. However, case with PIM and Tinlet = 260 ° C resulted

in higher SPL than that without PIM because of localized interior combustion. Figures 5.37 to

5.39 compare SPL without and with PIM. Results indicate a favorable trend with the use of PIM,

particularly at frequencies higher than 400 Hz. Figures 5.40 to 5.43 show CO and NOx emissions

136

for Q = 1400 slpm. Similar to Q = 1020 slpm, CO emissions were not affected by the PIM.

Further, CO emissions were nearly independent of the reactant inlet temperature. NOx emissions

presented in Figure 5.43 show an increase with inlet air temperature as expected. However, NOx

formation was not affected by the PIM.

Table 5.6 presents a summary of pressure drop data measured across the swirl injector

and combustor for Q = 1400 slpm. Similar to cases with Q = 1020 slpm, pressure drop increased

with inlet air temperature and it was highest for Tinlet = 260 °C, as seen in Figure 5.44. Flow

resistance by the porous insert is minor compared to the overall pressure drop in the swirler and

combustor.

Table 5.5

Summary of sound pressure levels for Q=1400 slpm

Tinlet (°C) Ф Qc (slpm) Total sound pressure level (dB)

No PIM PIM (18 ppcm)

20

0.65 1300 118.0 113.6

0.70 1300 120.4 115.6

0.75 1300 120.8 116.6

130

0.65 1550 124.3 117.8

0.70 1550 125.6 119.0

0.75 1550 125.4 120.5

260

0.65 1750 119.6 119.7

0.70 1750 119.1 121.3

0.75 1750 119.2 122.8

137

Table 5.6

Summary of pressure measurements for Q = 1400 slpm

Tinlet (°C)

Ф No PIM 18 ppcm PIM

Pinlet (KPa)

Pchamber (KPa)

∆P (KPa)

Pinlet (KPa)

Pchamber (KPa)

∆P (KPa)

20

0.65 105.5 102.0 3.5 105.5 101.3 4.2

0.70 106.1 101.3 4.8 105.5 100.6 4.9

0.75 105.5 102.0 3.5 105.5 100.6 4.9

130

0.65 113.0 102.0 11.0 111.7 101.3 10.4

0.70 113.7 101.3 12.4 113.1 101.3 11.8

0.75 113.0 102.0 11.0 112.4 102.0 10.4

260

0.65 118.6 102.0 16.6 117.2 101.3 15.9

0.70 118.6 101.3 17.3 119.3 102.0 17.3

0.75 117.9 102.0 15.9 119.3 102.0 17.3

5.3.2 Restricted Top Experiments

(a) Effect of PIM on Noise at P = 1 atm

Experiments were conducted with the reducer and exit nozzle installed to independently

investigate effect of porous insert on combustion noise and jet noise. Tests were conducted with

a 7.6 cm by 3.8 cm tapered exit nozzle. Figure 5.45 shows a schematic diagram of the nozzle

used for these experiments. This size nozzle did not build pressure inside the enclosure during

testing, thus tests pertain nearly to atmospheric pressure. Sound pressure data were measured

138

simultaneously at two locations: (1) outside the enclosure to obtain SPL of the high velocity jet,

or jet noise; and (2) on the wall of the enclosure, to obtain SPL in the reaction zone, or

combustion noise. First, sound pressure data were collected at two different sampling rates to

identify possible frequency aliasing. Figure 5.46 shows the power spectra of jet noise at 2000

and 4000 Hz sampling rates for the baseline case of Ф = 0.70 and Tinlet = 130 °C. Figures 5.46 (a)

and (b) both show a frequency peak at 600 Hz, indicating lack of frequency aliasing by sampling

rate. Figure 5.47 shows the power spectra for combustion noise data at 2000 and 4000 Hz

collected by the pressure sensor. Figure 5.47 (a) shows a peak frequency at 725 Hz, and 5.47 (b)

shows a peak at 1250 Hz, suggesting signal aliasing. Thus, subsequent measurements for both jet

noise and combustion noise were taken at 4000 Hz. Sound pressure data at jet exit were collected

at three different locations along the axis of the jet, as depicted in Figure 5.48. Figure 5.49

presents SPL versus one-third octave bands for Ф = 0.70 and Tinlet = 20, 130 and 260 °C. For

Tinlet = 20 °C, SPL is broadband for all microphone locations. Case with Tinlet = 130 °C shows a

dominant peak at 600 Hz. Case with Tinlet = 260 °C shows peaks at 800 Hz and 1600 Hz. In

general, microphone at location 3, farthest from the nozzle exit, registers highest total SPL for all

cases. Nonetheless, power spectra shows similar trends at each probe location. Thus, subsequent

results are presented only for data collected at location 1, i.e., for probe aligned with the jet exit

plane.

Figures 5.50 and 5.51 present power spectra at jet exit without porous insert and with 18

ppcm insert, respectively, for Ф = 0.65 and Tinlet = 20, 130 and 260 °C. As observed with

experiments in the previous section, PIM insert resulted in reduction of noise and mitigation of

combustion instability (notice change in vertical scale). Use of porous insert affects sound

pressure in the reaction zone, which in turn affects noise generated at the nozzle exit. This is the

139

result of the acoustic attenuation of pressure waves generated at the reaction zone by the

presence of PIM. Figure 5.52 compares SPL without and with porous insert for Ф = 0.65. Results

show that PIM is effective in reducing SPL, particularly at frequencies higher than 500 Hz.

Figures 5.53 and 5.54 present the combustion noise power spectra without and with PIM

respectively. As observed before, instability present without PIM, observed in Figure 5.53 (c) at

1250 Hz, was mitigated with use of PIM. Power spectra shown in Figure 5.54 have low power

for all cases. Next, Figure 5.55 presents combustion noise without and with porous insert for Ф =

0.65 and Tinlet = 20, 130 and 260 °C. Similar to jet noise results, use of PIM reduces combustion

noise and mitigates combustion instabilities in the reaction zone.

Next, jet noise and combustion noise data are presented for Ф = 0.70, without and with

porous insert in Figures 5.56 and 5.57, respectively. Figure 5.56 (b) shows a dominant peak at

550 Hz without porous insert, which is mitigated with use of PIM, as shown in Figure 5.57 (b).

Figure 5.58 compares SPL without and with PIM for Ф = 0.70. Case without PIM results in high

SPL at high frequencies. These high peaks are suppressed with the use of PIM. Next, similar to

jet noise, combustion noise power spectra are presented in Figures 5.59 and 5.60. For these

cases, noise is mainly broadband, except a peak present at 1250 Hz for Tinlet = 130 °C for no PIM

case. As observed before, this peak is eliminated by use of PIM. Figure 5.61 compares SPL for

Ф = 0.7 without and with PIM. As before, PIM reduces SPL by eliminating combustion

instability and attenuating broadband noise. Figures 5.62 to 5.67 present results for Ф = 0.75 for

jet noise and combustion noise. Again, results are consistent with previous observations at lower

equivalence ratios. A summary of test results is presented in Table 5.7 for jet noise, and Table

5.8 for combustion noise.

140

Similar to previous experiments, pressure drop across swirler and combustor was also

measured. Table 5.9 presents a summary of pressure drop data without and with porous insert.

Increase in inlet air temperature increases the pressure drop, which is highest at 12.4 KPa for

Tinlet = 260 °C, as seen in Figure 5.68. As observed for previous cases, PIM has negligible effect

on pressure drop across the injector/combustor.

Porous insert in the combustor reduces SPL and eliminates dominant frequencies of

fluctuation, caused by combustion instability. This result was consistently observed for all cases

operated at atmospheric pressure. Results show that porous insert reduces total SPL by

approximately 1-2 dB for cases with broadband noise. Most importantly, combustion instability

is also mitigated by PIM with SPL reduction of nearly 10 dB. In summary, when combustion

instability is present, PIM eliminates it or mitigates it, and when the combustion instability is not

present, PIM reduces broadband combustion noise.

141

Table 5.7

Summary of jet noise total SPL, Q = 1020 slpm, P = 1 atm


Tinlet (°C) Ф Qc (slpm) Position No PIM 18 ppcm PIM

20

0.65 990

1 104.5 102.8

2 106.0 104.4

3 106.7 105.1

0.70 990

1 106.6 104.1

2 108.3 105.5

3 109.0 106.0

0.75 990

1 108.2 105.9

2 109.8 107.7

3 110.4 108.5

130

0.65 1160

1 111.6 105.1

2 113.3 106.6

3 113.5 107.5

0.70 1160

1 118.7 106.7

2 121.0 108.7

3 122.7 109.1

0.75 1160

1 125.0 109.1

2 125.2 111.2

3 124.1 112.3

260

0.65 1350

1 116.6 108.2

2 117.0 110.1

3 117.2 111.1

0.70 1350

1 112.6 110.4

2 115.1 112.3

3 115.4 113.3

0.75 1350

1 114.4 111.6

2 117.4 113.6

3 117.5 114.3

142

Table 5.8

Summary of combustion noise total SPL, Q = 1020 slpm, P = 1 atm


Tinlet (°C) Ф Qc (slpm) No PIM 18 ppcm PIM

20

0.65 990 145.6 145.0

0.70 990 145.7 144.8

0.75 990 146.0 144.9

130

0.65 1160 146.7 145.0

0.70 1160 148.0 145.1

0.75 1160 157.5 145.5

260

0.65 1350 149.3 145.4

0.70 1350 147.4 145.9

0.75 1350 147.6 146.3

143

Table 5.9

Summary of pressure measurements for restricted top experiments, Q = 1020 slpm, P = 1 atm

Tinlet (°C)


Pinlet (KPa)

Pchamber (KPa)

∆P (KPa)

Pinlet (KPa)

Pchamber (KPa)

∆P (KPa)

20

0.65 104.1 102.7 1.4 104.8 103.4 1.4

0.70 104.8 103.4 1.4 104.8 104.1 0.7

0.75 104.1 103.4 0.7 104.8 104.1 0.7

130

0.65 109.6 104.8 4.8 114.4 104.8 9.6

0.70 111.0 103.4 7.6 113.1 105.5 7.6

0.75 111.7 104.1 7.6 113.7 105.5 8.2

260

0.65 116.5 104.8 11.7 116.5 106.2 10.3

0.70 117.2 105.5 11.7 117.9 106.8 11.1

0.75 118.6 106.2 12.4 117.2 106.2 11.0

(b) Effect of PIM on Noise at Elevated Pressure

In this section, experiments were conducted with a tapered nozzle of 3.8 cm by 1.9 cm,

illustrated in Figure 5.69. Previous experiments conducted with cold flow through this nozzle

resulted in high pressure inside the enclosure. Tests were conducted with Q = 2040 slpm, Ф =

0.65, 0.70 and 0.75 and enclosure pressure of approximately 2 atm. Figures 5.70 to 5.72 show

the SPL in one third octave bands for combustion noise. Cases without porous insert have

dominant frequencies above 500 Hz. Similar to previous results, PIM effectively reduces these

high frequency fluctuations. Table 5.10 shows a summary of obtained results for elevated

pressure tests. Results indicate that PIM is also beneficial at elevated pressures, however, total

SPL reduction is limited to 1-2 dB, thus, decrease in total SPL is not as dramatic. Based on

144

previous observations, flame is presumably partly submerged within the PIM, however, direct

flame observation was not feasible in the present setup. Table 5.11 shows a summary of pressure

measurements, which are also illustrated in Figure 5.73. Highest pressure drop of 9.0 KPa,

pertains to Tinlet = 130 °C. At Tinlet = 260 °C, pressure sensor exceeded operating temperature and

failed to generate signal.

Table 5.10

Summary of combustion noise total SPL, Q = 2040 slpm, P = 2 atm


Tinlet (°C) Ф Qc (slpm) No PIM 18 ppcm PIM

20

0.65 1230 148.9 146.6

0.70 1230 148.8 148.1

0.75 1230 149.4 152.5

130

0.65 1700 148.8 147.7

0.70 1700 148.5 147.0

0.75 1700 148.2 146.2

260

0.65 1900 148.3 146.8

0.70 1900 149.8 146.7

0.75 1900 150.3 146.6

145

Table 5.11

Summary of pressure measurements for restricted top experiments, Q = 2040 slpm, P = 2 atm

Tinlet (°C)


Pinlet (KPa)

Pchamber (KPa)

∆P (KPa) Pinlet

(KPa) Pchamber (KPa)

∆P (KPa)

20

0.65 236.5 234.4 2.1 209.6 208.2 1.4

0.70 238.6 237.2 1.4 212.4 210.3 2.1

0.75 243.4 241.3 2.1 212.4 210.5 1.9

130

0.65 249.6 243.4 6.2 228.9 221.3 7.6

0.70 252.4 245.5 6.9 233.1 224.1 9.0

0.75 258.6 250.3 8.3 234.4 225.5 8.9

260

0.65 249.6 - - 242.0 233.7 8.3

0.70 254.4 - - 243.4 235.1 8.3

0.75 257.2 - - 246.8 - -

5.4 Conclusions

Experiments were conducted at high inlet air flow rates, and high air inlet temperature to

more closely simulate operating conditions of a gas turbine engine. Average axial velocity at

combustor inlet for conditions ranged from 30 to 76 m/s. Inlet air temperatures were 20, 130 and

260 °C and equivalence ratios were 0.65, 0.70 and 0.75. Effect of pore density was investigated,

using 18 and 32 ppcm porous inserts. Jet noise and combustion noise were measured

independently. Tests were conducted at chamber pressure P = 1 and 2 atm. Pressure drop across

swirl injector and combustor was also measured. Product gas was sampled to measure CO and

NOx concentrations. Results demonstrated that porous insert reduce total SPL for all cases.

146

Furthermore, combustion instability present without porous insert was mitigated by use of PIM

in the reaction zone. Porous media redistributes incoming flow and attenuates turbulent flow

fluctuations, eliminates corner recirculation zone, and effectively attenuates sound pressure

levels. Porous media attenuates heat release fluctuations associated with turbulent nature of

flame. This result was obtained only with combustion occurring on the surface of the porous

insert. PIM interior combustion tends to increase SPL, and thus, it is undesirable. For flow

conditions investigated, flame stabilized within core region and on surface of PIM for both pore

densities. Difference between PIM of 18 and 32 ppcm is not significant. Highest pressure drop

registered pertained highest inlet temperature. PIM did not affect CO or NOx emissions for cases

studied.

147

Figure 5.1. Schematic of experimental setup

Cooling air inlet

Reducer

Nozzle

Enclosure

Combustion chamber

PIM

Swirler

Premix section

Fuel inlet

Preheated air inlet

Products

Dump plane

148

Figure 5.2. Photograph of fuel station

149

Figure 5.3. Schematic of combustion chamber

Quartz combustion chamber

PIM

Swirler

Premix section

2.5 cm

Reactants

Cooling air Cooling air

Products

Dump plane

150

Figure 5.4. Swirler

Flow

Flow

1.18

48.9°

R.02

R.26 R.30

0.50 .25 0.40

0.04

0.04

0.21 Ø0.96

Ø0.54

151

Figure 5.5. Schematic of experimental setup

Cooling air inlet

Reducer

Nozzle

Enclosure

Swirler

Fuel inlet

Preheated air inlet

Products

Port for Pinlet measurement

Port for Tinlet measurement

Port for Pchamber measurement

Port for emissions measurement

27 cm

152

Figure 5.6. Schematic diagram of PIM

5

3.75

2.5

7

Direction of the flow

Dimensions in cm

153

Figure 5.7. Schematic diagram and photograph of sampling probe

Stainless steel section Quartz section

Union

To gas analyzer

Fittings

To gas analyzer

Fitting (connects to wall of enclosure) Union

Gas sample

154

Figure 5.8. Schematic of PIM stabilization mechanism

Surface flame

Reactants

Products

PIM

Swirler

Reactants

Core region flame

155

Figure 5.9. Microphone locations for open top experiments

d=80

d=25

d=35

d=50

d=140

Dimensions in cm

Cooling air inlet

Enclosure

Fuel inlet

Preheated air inlet

156

Figure 5.10. One third octave band SPL for repeatability test

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100060

70

80

90

100

110

120

130

157

Figure 5.11. Effect of probe position on SPL for open top experiments, Q = 1020 slpm, Ф = 0.7,

Tinlet = 20 °C

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100050

60

70

80

90

100

110

120

130

d = 25 cmd = 35 cmd = 50 cmd = 80 cmd = 140 cm

158

Figure 5.12. Microphone locations for restricted top experiments

11

2.5

2.5

2.5

21

31

12

2.5

41

2.5 2.5

22

32

42

13

23

33

43

14

24

34

44

Dimensions in cm

25

Cooling air inlet

Reducer

Nozzle

Enclosure

Fuel inlet

Preheated air inlet

159

Figure 5.13. Effect of microphone location on SPL for restricted top experiments, Q = 1020

slpm, Ф = 0.7, Tinlet = 20 °C, Qc = 990 slpm

X

XXXX

XXX

XX

XX

X

X X

X

X

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100065

70

75

80

85

90

95

100

105

Position 11Position 21Position 31Position 41

XX

XXX

X

XXXXX

X

X

X

X X

X

X

Frequency (Hz)S

PL

(dB

)0 200 400 600 800 100065

70

75

80

85

90

95

100

105


X

XXXXX

XXXX

X X

X

XX X

X

X

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100065

70

75

80

85

90

95

100

105


XXXXXX

XXXX

X X

X

X

X X

X

X

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100065

70

75

80

85

90

95

100

105


X

160

Figure 5.14. Power spectra for Q = 1020 slpm, Ф = 0.65, no PIM (a) Tinlet = 20 °C, (b) Tinlet =

130 °C, (c) Tinlet = 260 °C

(a)

(b)

(c)

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

161


130 °C, (c) Tinlet = 260°C

(b)

(c)

(a)

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

162


130 °C, (c) Tinlet = 260 °C

(b)

(c)

(a)

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

163

Figure 5.17. Power spectra for Q = 1020 slpm, Ф = 0.65, 18ppcm PIM (a) Tinlet = 20 °C, (b) Tinlet

= 130 °C, (c) Tinlet = 260 °C

(b)

(c)

(a)

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

164

Figure 5.18. Power spectra for Q = 1020 slpm, Ф = 0.70, 18 ppcm PIM (a) Tinlet = 20 °C, (b)

Tinlet = 130 °C, (c) Tinlet = 260 °C

(b)

(c)

(a)

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

165



(b)

(c)

(a)

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

166



(b)

(c)

(a)

Frequency (Hz)

Po

wer

(A.U

.)0 200 400 600 800 1000

0

5

10

15

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

5

10

15

167



(b)

(c)

(a)

Frequency (Hz)

Po

wer

(A.U

.)0 200 400 600 800 1000

0

5

10

15

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

5

10

15

168



(b)

(c)

(a)

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

169

Figure 5.23. SPL in one third octave for Q = 1020 slpm, Ф = 0.65, (a) Tinlet = 20 °C, (b) Tinlet =

130 °C, (c) Tinlet = 260 °C

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100060

70

80

90

100

110

120

130

No PIM18 ppcm PIM32 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100060

70

80

90

100

110

120

130


Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100060

70

80

90

100

110

120

130


(b)

(c)

(a)

170


130 °C, (c) Tinlet = 260 °C

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100060

70

80

90

100

110

120

130


Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100060

70

80

90

100

110

120

130


Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100060

70

80

90

100

110

120

130


(b)

(c)

(a)

171


130 °C, (c) Tinlet = 260 °C

(b)

(c)

(a)

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100060

70

80

90

100

110

120

130


Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100060

70

80

90

100

110

120

130


Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100060

70

80

90

100

110

120

130


172

Figure 5.26. CO emissions for Q = 1020 slpm, Ф = 0.65, (a) Tinlet = 20 °C, (b) Tinlet = 130 °C, (c)

Tinlet = 260 °C


CO

(pp

m)

-3 -2 -1 0 1 2 30

10

20

30

40

50



CO

(pp

m)

-3 -2 -1 0 1 2 30

10

20

30

40

50



CO

(pp

m)

-3 -2 -1 0 1 2 30

10

20

30

40

50


(b)

(c)

(a)

173


Tinlet = 260 °C


CO

(pp

m)

-3 -2 -1 0 1 2 30

10

20

30

40

50



CO

(pp

m)

-3 -2 -1 0 1 2 30

10

20

30

40

50



CO

(pp

m)

-3 -2 -1 0 1 2 30

10

20

30

40

50


(b)

(c)

(a)

174

Figure 5.28. NOx emissions for Q = 1020 slpm, Ф = 0.65, (a) Tinlet = 20 °C, (b) Tinlet = 130 °C,

(c) Tinlet = 260 °C


NO

x(p

pm

)

-3 -2 -1 0 1 2 30

20

40

60

80

100



NO

x(p

pm

)

-3 -2 -1 0 1 2 30

20

40

60

80

100



NO

x(p

pm

)

-3 -2 -1 0 1 2 30

20

40

60

80

100


(b)

(c)

(a)

175




NO

x(p

pm

)-3 -2 -1 0 1 2 3

0

20

40

60

80

100



NO

x(p

pm

)

-3 -2 -1 0 1 2 30

20

40

60

80

100



NO

x(p

pm

)

-3 -2 -1 0 1 2 30

20

40

60

80

100


(b)

(c)

(a)

176

Figure 5.30. Pressure drop measurements for open top experiments Q = 1020 slpm (a) no PIM

(b) 18 ppcm PIM (c) 32 ppcm PIM

(b)

(c)

(a)

Φ

∆P

(KP

a)

0.65 0.7 0.750

5

10

15

20

25

Tinlet = 20 oCTinlet = 130 oCTinlet = 260 oC

Φ

∆P

(KP

a)

0.65 0.7 0.750

5

10

15

20

25


Φ

∆P

(KP

a)

0.65 0.7 0.750

5

10

15

20

25


177

Figure 5.31. Power spectra for Q = 1400 slpm, Ф = 0.65, no PIM, (a) Tinlet = 20 °C, (b) Tinlet =

130 °C, (c) Tinlet = 260 °C

(b)

(c)

(a)

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

178


130 °C, (c) Tinlet = 260°C

(b)

(c)

(a)

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

179


130 °C, (c) Tinlet = 260 °C

(b)

(c)

(a)

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

180

Figure 5.34. Power spectra for Q = 1400 slpm, Ф = 0.65, 18 ppcm PIM, (a) Tinlet = 20 °C, (b)


(b)

(c)

(a)

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

181


Tinlet = 130 °C, (c) Tinlet = 260°C

(b)

(c)

(a)

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

182



(b)

(c)

(a)

Frequency (Hz)

Po

wer

(A.U

.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

Frequency (Hz)

Po

we

r(A

.U.)

0 200 400 600 800 10000

5

10

15

183


130 °C, (c) Tinlet = 260 °C

(b)

(c)

(a)

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100060

70

80

90

100

110

120

130

No PIM18 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100060

70

80

90

100

110

120

130

No PIM18 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100060

70

80

90

100

110

120

130

No PIM18 ppcm PIM

184


130 °C, (c) Tinlet = 260 °C

(b)

(c)

(a)

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100060

70

80

90

100

110

120

130

No PIM18 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100060

70

80

90

100

110

120

130

No PIM18 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100060

70

80

90

100

110

120

130

No PIM18 ppcm PIM

185


130 °C, (c) Tinlet = 260 °C

(b)

(c)

(a)

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100060

70

80

90

100

110

120

130

No PIM18 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100060

70

80

90

100

110

120

130

No PIM18 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 200 400 600 800 100060

70

80

90

100

110

120

130

No PIM18 ppcm PIM

186


Tinlet = 250 °C


CO

(pp

m)

-3 -2 -1 0 1 2 30

10

20

30

40

50

60

70

No PIM18 ppcm PIM


CO

(pp

m)

-3 -2 -1 0 1 2 30

10

20

30

40

50

60

70

No PIM18 ppcm PIM


CO

(pp

m)

-3 -2 -1 0 1 2 30

10

20

30

40

50

60

70

No PIM18 ppcm PIM

(b)

(c)

(a)

187


Tinlet = 250 °C


CO

(pp

m)

-3 -2 -1 0 1 2 30

10

20

30

40

50

60

70

No PIM18 ppcm PIM


CO

(pp

m)

-3 -2 -1 0 1 2 30

10

20

30

40

50

60

70

No PIM18 ppcm PIM


CO

(pp

m)

-3 -2 -1 0 1 2 30

10

20

30

40

50

60

70

No PIM18 ppcm PIM

(b)

(c)

(a)

188



(b)

(c)

(a)


NO

x(p

pm

)

-3 -2 -1 0 1 2 30

20

40

60

80

100

No PIM18 ppcm PIM


NO

x(p

pm

)

-3 -2 -1 0 1 2 30

20

40

60

80

100

No PIM18 ppcm PIM


NO

x(p

pm

)

-3 -2 -1 0 1 2 30

20

40

60

80

100

No PIM18 ppcm PIM

189



(b)

(c)

(a)


NO

x(p

pm

)

-3 -2 -1 0 1 2 30

20

40

60

80

100

120

No PIM18 ppcm PIM


NO

x(p

pm

)

-3 -2 -1 0 1 2 30

20

40

60

80

100

120

No PIM18 ppcm PIM


NO

x(p

pm

)

-3 -2 -1 0 1 2 30

20

40

60

80

100

120

No PIM18 ppcm PIM

190

Figure 5.44. Pressure drop measurements for open top experiments Q = 1400 slpm (a) no PIM

(b) 18 ppcm PIM

(b)

(a)

Φ

∆P

(KP

a)

0.65 0.7 0.750

5

10

15

20

25


Φ

∆P

(KP

a)

0.65 0.7 0.750

5

10

15

20

25


191

Figure 5.45. Schematic diagram of nozzle for restricted flow experiments

Dimensions in cm

Direction of flow

7.6

3.8

2.9

192

Figure 5.46. Jet noise power spectra, no PIM, P = 1 atm, Ф = 0.75, Tinlet = 130 °C (a) sampling

rate of 2000 Hz, (b) sampling rate of 4000 Hz

(b)

(a)

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

20

40

60

80

100

120

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

20

40

60

80

100

120

193

Figure 5.47. Combustion noise power spectra, no PIM, P = 1 atm, Ф = 0.75, Tinlet = 130 °C (a)

sampling rate of 2000 Hz, (b) sampling rate of 4000 Hz

(b)

(a)

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

20000

40000

60000

Frequency (Hz)

Pow

er(A

.U.)

0 500 1000 1500 20000

20000

40000

60000

194

Figure 5.48. Location of microphones for jet noise

Dimensions in cm

25

Cooling air inlet

Reducer

Nozzle

Enclosure

Fuel inlet

Preheated air inlet

5

5 2

1

3

195

Figure 5.49. Jet noise SPL in one third octave for no PIM, Q = 1020 slpm, Ф = 0.70 (a) Tinlet =

20 °C, (b) Tinlet = 130 °C, (c) Tinlet = 260 °C

(b)

(c)

(a)

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 200060

70

80

90

100

110

120

130

No PIM, location 1No PIM, location 2No PIM, location 3

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 200060

70

80

90

100

110

120

130


Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 200060

70

80

90

100

110

120

130


196

Figure 5.50. Jet noise power spectra, no PIM, Q = 1020 slpm, Ф = 0.65 (a) Tinlet = 20 °C, (b)

Tinlet = 130 °C, (c) Tinlet = 260°C

(b)

(c)

(a)

Frequency (Hz)

Po

we

r(A

.U.)

0 500 1000 1500 20000

1

2

3

4

5

Frequency (Hz)

Po

we

r(A

.U.)

0 500 1000 1500 20000

1

2

3

4

5

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

1

2

3

4

5

197

Figure 5.51. Jet noise power spectra, 18 ppcm PIM, Q = 1020 slpm, Ф = 0.65 (a) Tinlet = 20 °C,

(b) Tinlet = 130 °C, (c) Tinlet = 260 °C

Frequency (Hz)

Po

we

r(A

.U.)

0 500 1000 1500 20000

0.2

0.4

0.6

0.8

1

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

0.2

0.4

0.6

0.8

1

(b)

(c)

(a)

Frequency (Hz)

Po

we

r(A

.U.)

0 500 1000 1500 20000

0.2

0.4

0.6

0.8

1

198

Figure 5.52. Jet noise SPL in one third octave, Q = 1020 slpm, Ф = 0.65, P = 1 atm, (a) Tinlet =

20 °C, (b) Tinlet = 130 °C, (c) Tinlet = 260 °C

(b)

(c)

(a)

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 200060

70

80

90

100

110

120

130

No PIM18 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 200060

70

80

90

100

110

120

130

No PIM18 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 200060

70

80

90

100

110

120

130

No PIM18 ppcm PIM

199

Figure 5.53. Combustion noise power spectra, no PIM, Q = 1020 slpm, Ф = 0.65 (a) Tinlet = 20

°C, (b) Tinlet = 130 °C, (c) Tinlet = 260 °C

(b)

(c)

(a)

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

1000

2000

3000

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

1000

2000

3000

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

1000

2000

3000

200

Figure 5.54. Combustion noise power spectra, 18 ppcm PIM, Q = 1020 slpm, Ф = 0.65 (a) Tinlet

= 20 °C, (b) Tinlet = 130 °C, (c) Tinlet = 260 °C

(b)

(c)

(a)

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

1000

2000

3000

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

1000

2000

3000

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

1000

2000

3000

201

Figure 5.55. Combustion noise SPL in one third octave, Q = 1020 slpm, Ф = 0.65, P = 1 atm (a)

Tinlet = 20 °C, (b) Tinlet = 130 °C, (c) Tinlet= 260 °C

(b)

(c)

(a)

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 2000110

120

130

140

150

160

No PIM18 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 2000110

120

130

140

150

160

No PIM18 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 2000110

120

130

140

150

160

No PIM18 ppcm PIM

202



(b)

(c)

(a)

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

1

2

3

4

5

Frequency (Hz)

Po

we

r(A

.U.)

0 500 1000 1500 20000

1

2

3

4

5

Frequency (Hz)

Po

we

r(A

.U.)

0 500 1000 1500 20000

5

10

15

203



(b)

(c)

(a)

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

0.2

0.4

0.6

0.8

1

Frequency (Hz)

Po

we

r(A

.U.)

0 500 1000 1500 20000

0.2

0.4

0.6

0.8

1

Frequency (Hz)

Po

we

r(A

.U.)

0 500 1000 1500 20000

0.2

0.4

0.6

0.8

1

204

Figure 5.58. Jet noise SPL in one third octave, Q = 1020 slpm, Ф = 0.70, P = 1 atm (a) Tinlet = 20


(b)

(c)

(a)

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 200060

70

80

90

100

110

120

130

No PIM18 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 200060

70

80

90

100

110

120

130

No PIM, location 118 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 200060

70

80

90

100

110

120

130

No PIM18 ppcm PIM

205

Figure 5.59. Combustion noise power spectra, no PIM, Q = 1020 slpm, Ф = 0.70, (a) Tinlet = 20


(b)

(c)

(a)

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

1000

2000

3000

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

1000

2000

3000

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

1000

2000

3000

206



(b)

(c)

(a)

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

1000

2000

3000

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

1000

2000

3000

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

1000

2000

3000

207


Tinlet = 20 °C, (b) Tinlet = 130 °C, (c) Tinlet = 260 °C

(b)

(c)

(a)

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 2000110

120

130

140

150

160

No PIM18 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 2000110

120

130

140

150

160

No PIM18 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 2000110

120

130

140

150

160

No PIM18 ppcm PIM

208



Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

1

2

3

4

5

Frequency (Hz)

Po

we

r(A

.U.)

0 500 1000 1500 20000

20

40

60

80

100

120

(b)

(c)

(a)

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

1

2

3

4

5

209



(b)

(c)

(a)

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

0.2

0.4

0.6

0.8

1

Frequency (Hz)

Po

we

r(A

.U.)

0 500 1000 1500 20000

0.2

0.4

0.6

0.8

1

Frequency (Hz)

Po

we

r(A

.U.)

0 500 1000 1500 20000

0.2

0.4

0.6

0.8

1

210

Figure 5.64. Jet noise SPL in one third octave, Q = 1020 slpm, Ф = 0.75, P = 1 atm (a) Tinlet = 20

°C, (b) Tinlet=130 °C, (c) Tinlet=260 °C

(b)

(c)

(a)

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 200060

70

80

90

100

110

120

130

No PIM18 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 200060

70

80

90

100

110

120

130

No PIM18 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 200060

70

80

90

100

110

120

130

No PIM18 ppcm PIM

211

Figure 5.65. Combustion noise power spectra, no PIM, Q = 1020 slpm, Ф = 0.75 (a) Tinlet = 20


(b)

(c)

(a)

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

80000

160000

240000

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

1000

2000

3000

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

1000

2000

3000

212



(b)

(c)

(a)

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

1000

2000

3000

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

1000

2000

3000

Frequency (Hz)

Po

wer

(A.U

.)

0 500 1000 1500 20000

1000

2000

3000

213



(b)

(c)

(a)

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 2000110

120

130

140

150

160

No PIM18 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 2000110

120

130

140

150

160

No PIM18 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 2000110

120

130

140

150

160

No PIM18 ppcm PIM

214

Figure 5.68. Pressure drop measurements for restricted top experiments, P = 1 atm (a) no PIM

(b) 18 ppcm PIM

Φ

∆P

(KP

a)

0.65 0.7 0.750

5

10

15

20

25


Φ

∆P

(KP

a)

0.65 0.7 0.750

5

10

15

20

25


(b)

(a)

215

Figure 5.69. Schematic diagram of nozzle for restricted flow experiments

Dimensions in cm

Direction of flow

3.8

1.9

2.9

216



(b)

(c)

(a)

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 2000110

120

130

140

150

160

No PIM18 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 2000110

120

130

140

150

160

No PIM18 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 2000110

120

130

140

150

160

No PIM18 ppcm PIM

217



(b)

(c)

(a)

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 2000110

120

130

140

150

160

No PIM18 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 2000110

120

130

140

150

160

No PIM18 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 2000110

120

130

140

150

160

No PIM18 ppcm PIM

218



(b)

(c)

(a)

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 2000110

120

130

140

150

160

No PIM18 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 2000110

120

130

140

150

160

No PIM18 ppcm PIM

Frequency (Hz)

SP

L(d

B)

0 500 1000 1500 2000110

120

130

140

150

160

No PIM18 ppcm PIM

219

Figure 5.73. Pressure drop measurements for restricted top experiments, P = 2 atm (a) no PIM

(b) 18 ppcm PIM

Φ

∆P

(KP

a)

0.65 0.7 0.750

5

10

15

20

25

Tinlet = 20 oCTinlet = 130 oC

Φ

∆P

(KP

a)

0.65 0.7 0.750

5

10

15

20

25


(b)

(a)

220

CHAPTER 6

CONCLUSIONS AND RECOMMENDATIONS

6.1 Conclusions

In the present study, swirl stabilization mechanism is combined with porous inert media

for lean premixed combustion with the goal to reduce combustion generated noise without

affecting CO and NOx emissions. Gaseous fuel is used as a first step to eventually implement the

concept in a liquid fuel combustor. The approach involves passive control of the flow structures

that generate noise in combustion chamber. This approach can also mitigate combustion

instabilities that self-excite within the combustor. First, time-averaged numerical simulation of

non-reacting and reacting flows is performed to gain insight into the flow structure without and

with the porous insert. Next, an experimental investigation was conducted to determine impact of

PIM geometric parameters (such as ID, pore size) on sound pressure levels. Initially, experiments

were conducted at atmospheric pressure, and relatively low reactant flow rates and low inlet air

temperatures. Results were used to identify combustion modes and PIM configurations that are

most beneficial for noise reduction. Next, a laboratory facility was designed and developed to

conduct combustion experiments at high reactant flow rate, high inlet air temperature, and high

operating pressures. Details of the design, installation and operation of this laboratory facility are

presented. The facility was utilized for the next series of experiments to more closely simulate

typical gas turbines operating conditions, i.e., high reactant flow rate, high inlet air temperature

221

and high pressure. PIM configurations optimized in previous study were utilized. Following are

the main conclusions:

• Numerical model shows qualitative agreement with experimentally obtained data for

reacting and non-reacting flows. The study shows that flow inside combustion chamber is

redistributed: porous insert eliminates the corner recirculation zone, vertically orients the

gaseous flame zone, intensifies the central recirculation zone, maintains the swirling

effect imparted by the swirl injector, and creates a more uniform flow distribution at

downstream locations. These features can be expected to improve the noise and

instability performance of combustor.

• Experimental study identified interior and surface modes of swirl-stabilized combustion

with porous insert. Equivalence ratio, reactant flow rate, and PIM geometric parameters

such as ID and pore size determine the combustion mode. Surface combustion mode is

desirable, while interior combustion must be avoided to achieve low SPLs. Lower pore

density (<18 ppcm) resulted in undesired interior combustion mode and increased SPL. A

divergent porous insert with pore density of 18 ppcm was found to reduce the total SPL

by up to 14 dB. NOx and CO emissions were not adversely affected by the porous insert.

• A new laboratory facility for combustion experiments at high reactant flow rate, high

inlet air temperature, and high operating pressure was developed. A combustor

experimental apparatus was designed and fabricated to operate at pressure up to 10 atm.

The combustor designed consisted of different size nozzle to choke the flow downstream

of the combustion chamber. The experimental apparatus was demonstrated to operate

with air flow rate up to 3 kg/min, Tinlet = 260 °C and pressure of up to 4 atm.

222

• Experiments were conducted in the new facility at a higher operating pressure, high

reactant flow rate, and high inlet air temperature, to more closely simulate gas turbine

engine operating condition. Porous inserts of 18 and 32 ppcm were used. Jet noise and

combustion noise were measured independently. Pressure drop across swirl injector and

combustor was also measured. Results show that porous insert reduces both jet noise and

combustion noise. Furthermore, combustion instability present without porous insert was

mitigated by use of PIM. Porous insert effectively attenuated heat release fluctuations

associated with turbulent nature of flame. This result was obtained only with combustion

occurring on the surface of the porous insert. Difference between PIM of 18 and 32 ppcm

is not significant. Cases with no instabilities also showed a reduction in the total SPL

with use of porous insert. Highest pressure drop registered pertained highest inlet

temperature, and PIM did not affect CO or NOx emissions or pressure drop across the

combustor.

6.2 Recommendations

Recommendations for future scope and improvements of current work are listed below:

• A time dependent numerical model is recommended. Flow field, pressure field and heat

release fluctuations are of particular interest. A numerical study with porous insert of

diffuser shaped inside wall is recommended. Model can be extended to predict acoustic

behavior of the combustor.

• A study of swirl number effect on SPL. Swirl number can affect turbulence intensity in

corner recirculation zone, which in turn affects pressure fluctuations. Also, location of

swirler with respect to dump plane impacts total SPL by creating recirculation zones

223

within the premixer tube. Thus, an investigation of swirler parameters, such as swirl

number and location within the premixer, is recommended.

• Experiments with longer porous inserts (in the axial direction) at elevated flow rate. It

was observed that high reactant velocities tend to stretch reaction zone in the axial

direction of the combustor. Thus a longer porous insert could enhance heat transfer and

acoustic dissipation inside combustion chamber for these cases.

• An experimental investigation that combines PIM with liquid fuel combustion under

optimum atomization operating conditions is recommended.

• A more durable and flexible traversing system/emission probe to facilitate mounting and

emission measurements. Quartz holding mechanism that is more durable and reliable, and

a safe means to easily remove/mount combustor reducer apparatus.

• For safety, automation of an ignition system that can be remotely activated is

recommended.

• An automated control of combustion/cooling air split by merging air and fuel control

software interfaces is recommended.

• Redirect cooling air jets inside enclosure to improve cooling of windows.

224

REFERENCES

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B. F. Magnussen and B. H. Hjertager, (1976). On mathematical models of turbulent combustion

with special emphasis on soot formation and combustion. Proceedings of the Combustion Institute, vol. 16, pp. 719-729.

Bussman, W., and Jayakaran, J., (2001). Noise. The John Zink Combustion Handbook, CRC

Press, New York. Bertoli, C., D’Alessio, J., Del Giacomo, N., Lazzaro, M., Massoli, P., and Moccia, V., (2000).

Running Light-Duty DI Diesel Engines with Wood Pyrolysis Oil. SAE Paper 200-01-2975.

Broda, J. C., Seo, S., Santoro, R. J., Shirhattikar, G., and Yang, V., (1998). An Experimental

Study of Combustion Dynamics of a Premixed Swirl Injector. Proceedings of the Combustion Institute, Pittsburgh, PA, vol. 27, pp. 1849-1856.

Cheng, R.K., Yegian, D.T., Miyasato, M.M., Samuelsen, G.S., Pellizzari, R., and Benson, C.,

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APPENDIX A

COMBUSTION PERFORMANCE OF LIQUID BIO-FUELS IN A SWIRL-STABILIZED BURNER

Background

Unlike non-renewable fossil fuels, the CO2 emitted from the combustion of biofuels is

recycled in the environment through future plant growth. A recent study projects that with

relatively modest changes in land use and agricultural and forestry practices, an annual supply of

1.36 billion dry tons of biomass could be available for large-scale biofuel production in the

United States by the mid-21st century, while still meeting demands for forestry products, food

and fiber (De La Torre, 2003)

Biomass can be converted into solid, gaseous or liquid fuel depending upon the

conversion processes and economic factors. Several current biomass technologies are reviewed

in (Demirbas, 2004). Co-firing of biomass with coal is among the most cost effective

approaches (Damstedt, 2007). Since combustion of solid fuels results in higher emissions (e.g.,

soot and nitric oxides or NOx), liquid and/or gaseous fuels produced from biomass are likely to

become increasingly prevalent in the near term. Biomass can be gasified in air-blown or oxygen-

blown gasifier, followed by the cleanup of the product gas (known as synthetic gas or syngas)

containing carbon monoxide (CO) and hydrogen (H2) as the primary reactants. The biomass

syngas can serve as the fuel to generate power in a high-efficiency combined cycle power plant.

The combined cycle power plant integrated with the gasification system can operate

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synergistically with biomass and/or coal as feedstock, since the syngas produced from

gasification of either of these sources (biomass or coal) contains the same reactive ingredients.

In contrast to biomass syngas requiring large-scale stationary operation, liquid biofuels

offer greater flexibility since the fuel can be transported easily. Liquid biofuels offer the

prospects of distributed generation, whereby the power is produced closer to the source without

hauling the bulky biomass to a distant central location. At present, ethanol and biodiesel are the

two commonly used liquid biofuels. However, the feedstocks for these fuels compete with the

food-chain crops. For example, virtually all of the ethanol produced in the US comes from corn

and biodiesel is produced by the transesterification of vegetable oils such as soybean oil. Thus, a

long-term strategy will require liquid fuels from biomass feedstocks such as wood and energy

crops that do not interfere with the food chain.

The biomass must undergo gasification or pyrolysis to produce the liquid fuel. Starting

with the biomass syngas, the well-known Fischer-Tropsch (FT) process can be used to produce

liquid biofuels of desired composition and physical characteristics. Although the FT process is

attractive to produce liquid biofuels for vehicular transportation (William, 2002), the fuels

produced by pyrolysis of biomass can be economic alternatives for power generation

applications. Pyrolysis is the thermal destruction of organic material in the absence of or limited

supply of oxygen (Damstedt, 2007). In fast pyrolysis, the thermal decomposition occurs at

moderate temperatures with a high heat transfer rate to the biomass particles and a short hot

vapor residence time in the reaction zone (Oasma, 1999; Czernik, 2004; Mohan, 2006). The

main product is the pyrolysis oil, also known as biooil, which is usually a dark-brown free-

flowing liquid with a distinctive smoky odor. Biooils have been successfully tested in diesel

engines and gas turbines (Strenziok, 2001; Bertoli, 2000; Lupandin, 2006), although

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modifications to the fuel handling system can incur unacceptable financial cost. Thus, a near-

term strategy would be to emulsify biooil using fuels compatible with the fuel handling

equipment. Ikura et al. (2003) produced biooil emulsified with the diesel fuel. The cost for

producing emulsions with zero stratification increased with increasing amounts of biooil in the

diesel fuel.

The literature review shows few studies on combustion performance of liquid bio-fuels

for gas turbine applications. Thus, the primary objective of this study is to isolate the effects of

fuel composition and fluid dynamics on emissions from different liquid fuels in an atmospheric

pressure burner replicating typical features of a gas turbine combustor. The burner utilized a

commercial twin-fluid injector with primary air swirling around the injector. The fuels include

diesel, biodiesel, emulsified biooil, and diesel-biodiesel blends. Biooil emulsions were produced

using blends of biodiesel and diesel to increase the amount of biomass derived fuel in the final

product. The emulsified biooil produced in this manner is expected to require minimal

modifications to the fuel handling system. For fixed volume flow rates of fuel and air,

experiments were conducted by varying the airflow split between the injector and co-flow

swirler. Results include visual flame images, and axial and radial profiles of NOx and CO

concentrations at different operating conditions. In the following sections, the fuel preparation

steps and experimental setup details are outlined followed by results and discussions.

A.2 Fuel Preparation

As mentioned above, the fuels in this study included diesel, biodiesel, emulsified biooil,

and diesel-biodiesel blends. The diesel fuel used was of a commercial grade (No. 2 diesel fuel)

purchased from a local filling station. The biodiesel was supplied by Alabama Biodiesel

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Corporation (Moundville, AL) and it was a soybean oil methyl ester (SME). Pyrolysis oil, also

known as, biooil (from hardwood) was provided by the National Renewable Energy Labortory

(NREL; Boulder, CO). It was a hot-vapor filtered biooil with very low ash content. The biooil

was approximately one-year old, and was not phase separated or treated with any viscosity

reducing agents. The NREL biooil composition is summarized below in Table A.1.

The emulsified biooil was formulated by mixing diesel, biooil and ‘surfactants’. The

‘surfactants’ used in this study was a blend of biodiesel, 2-ethyl-1-hexanol (an alcohol), and n-

octylamine. The latter two chemicals were purchased from Aldrich (Milwaukee, WI). The

emulsified fuel was produced by mixing diesel, biodiesel, biooil, alcohol, and amine using a high

shear (10,000 rpm) Oster blender (Boca Raton, FL). Mixing was done at room temperature and

pressure until an emulsified liquid was obtained in about 2 minutes. In this study, two types of

biodiesel (SME and 90% ethyl oleate hereafter referred to as SEE) were mixed with diesel by

gentle stirring to form diesel-biodiesel fuel blends. These blends are completely miscible over

all concentration ranges. Table A.2 summarizes the fuels used in this study.

The water content of each fuel was determined by a volumetric Aquastar Karl-Fischer

titrator (EM Science, Gibbstown, NJ) with Composite 5 solution as the titrant and anhydrous

methanol as the solvent. All measurements were made in triplicate and at 25oC. The water

content in the fuel blends is summarized below in Table A.3.

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Table A.1

NREL Biooil Characteristics

Biooil Diesel* Biodiesel†

Moisture content (wt. %) 20.0

Ash (wt. %) 0.018

Elemental composition (wt. %)

Carbon 45.6 86.7 78.61

Hydrogen 7.6 13.5 11.99

Nitrogen 0.05 0.04 2.1‡

Sulfur 0.02 0.03 <0.004

Oxygen 46.8 9.38

* From Chiaramonti et al, 2006 † From Ikura et al, 2003. ‡ ppm.

Table A.2

Experimental Fuel Blends (Vol%)

Fuel Diesel Biodiesel

(SME+SEE)

Biooil Alcohol/

Amine

Diesel 100

Biodiesel (100 + 0)

Biooil 45 (30 + 0) 15 8/2

SOME 80 (20 + 0)

SOEE 80 (0 + 20)

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Table A.3

Water Contents in the Fuel Blend

Fuel % w/w ± 3σ**

Diesel 0.021 ± .006

Biodiesel 0.10 ± .01

Biooil 0.27 ± .04

SOME 0.03 ± .01

SOEE 0.017 ± .004

BioOil (as received)† 22.9 ± 1.4

†Water content most likely increased during storage.

A.3 Experimental Setup

The test apparatus shown schematically in Figure A.1 consists of the combustor assembly

and the injector assembly. The primary air enters the system through a plenum filled with

marbles to breakdown the large vortical structures. The air passes through a swirler into the

mixing section, where the gaseous fuel is supplied during the startup. The reactant(s) enter the

combustor through a swirler used to improve the fuel-air mixing. Figure A.2 shows a schematic

diagram and a photograph of the combustor inlet section with the swirler.

The swirler had six vanes positioned at 28o to the horizontal. The theoretical swirl

number was 1.5, assuming that the flow exited tangentially from the swirler vanes. The bulk

axial inlet velocity of the primary air was 1.9 to 2.1 m/s, which resulted in Reynolds number

varying from 5960 to 6750. The liquid fuel is supplied from an injector with separate concentric

inlets for fuel and atomization air. The injector system runs through the plenum and the mixing

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chamber. An O-ring within a sleeve is provided at the bottom of plenum to prevent any leakage.

The injector is a commercial air-blast atomizer (Delavan Siphon type SNA nozzle) and it creates

a swirling flow of atomizing air to breakdown the fuel jet. The injector details are shown in

Figure A.3. The combustor itself is a 8.0 cm ID and 46 cm long quartz tube. The combustor is

back-side cooled by natural convection.

The liquid fuel is supplied by a peristaltic pump with the range of flows rates from 2

ml/min to 130 ml/min in steps of 2 ml/min. The reported calibration error of the pump is +/-

0.25% of the flow rate reading. Viton tubes were used prevent any degradation of the fuel lines.

A 25 micron filter was used to prevent dirt and other foreign particles from clogging the injector.

The primary and atomizing air is supplied by an air compressor. The air passes through a

pressure regulator and a water trap to remove the moisture. Then, the air is split into primary air

supply and atomizing air supply lines. The primary air flow rate is measured by a laminar flow

element calibrated for 0 to 1000 liters per minute (lpm) of air. The pressure drop across the LFE

is measured by a differential pressure transducer. An absolute pressure transducer is used to

measure the pressure of air passing through the LFE. The flow rate measured by the LFE is

corrected for temperature and pressure as specified by the manufacturer. The atomizing air is

measured by calibrated mass flow meter.

The product gas was sampled continuously by a quartz probe (OD = 7.0 mm) attached to

a three-way manual traversing system. The upstream tip of the probe was tapered to 1 mm ID to

quench reactions inside the probe. The sample passed through an ice bath and water traps to

remove moisture upstream of the gas analyzers. The dry sample passed through electrochemical

analyzers to measure the concentrations of CO and NOx in ppm. The analyzer also measured

oxygen and carbon dioxide concentrations, which were used to cross-check the equivalence ratio

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obtained from the measured fuel and air flow rates. The uncorrected emissions data on dry basis

are reported with measurement uncertainty of +/- 2 ppm.

The experiment was started by supplying the gaseous methane and then, igniting the

methane-air reactant mixtures in the combustor. Next, the liquid fuel flow rate was gradually

increased to attain the desired value, while the methane flow rate was slowly decreased to zero.

In this study, the volume flow rates of total air (primary + atomizing) and fuel were kept

constant, respectively, at 150 standard lpm and 12 ml/min. It could result in small variations in

the amount of heat-released since the heating value of fuels is different. The combustion

performance is strongly dependent upon the spray characteristics determined by the atomizing air

flow rate. Initial experiments indicated yellow, sooty flames dominated by the diffusion mode of

combustion for atomizing air flow rates below 10% of the total air. Thus, experiments focused

on the premixed combustion mode with strong fuel-air premixing prompted by fine droplets

formed with large atomizing air flow rates. Accordingly, the experiments were conducted by

varying the percentage of the atomizing air (AA) from 15% to 25% of the total air. Since the

overall air-fuel ratio is constant, the effects of atomizing air on combustion emissions can be

ascertained from these measurements.

A.4 Results and Discussion

A.4.1 Visual Flame Images

Direct photographs of flame were taken by a digital camera to obtain qualitative

understanding of the flame characteristics. These photographs are reproduced in Figure A.4 for

diesel, biodiesel, and biooil flames. For each case, the flame images are shown for atomizing air

(AA) of 15%, 20%, and 25% of the total air. All flames in Fig. A.4 show a distinctive blue color

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typical of the premixed combustion. In contrast, combustion in the diffusion mode produces

yellow, sooty flames, because the fuel droplets do not vaporize and premix with air before

reactions take place. Results in Figure A.4 suggest that the injector is producing spray with fine

droplets that pre-vaporize and premix with air to form the reactant mixture prior to the

combustion. This is also true for the emulsified biooil, since it contained a relatively small

amount of biooil (15% by volume) that is otherwise difficult to pre-vaporize. Excellent

atomization is attributed to the large atomizing air used in this study. Note that the pressure drop

associated with atomizing air flow in the injector could increase the operating cost. However,

the operating conditions in this study provide a consistent basis to compare different fuels and

they also point towards the need to optimize the injector design to cost-effectively reduce

emissions of soot, NOx, CO, and unburned hydrocarbons.

Figure A.4(a) shows that the width and height of the diesel flame increases with increase

in the atomizing air. For example, the flame is short and intense for 15% AA compared to that

for 25% AA. In case of 15% AA, the fuel droplets pre-vaporize to form reactant mixture of

higher equivalence ratio, which would burn at an elevated flame temperature. With increase in

the atomizing air, the local equivalence ratio of the reactant mixture decreases and hence, the

reactions occur at a lower temperature. Note that the temperature of the homogenized products

would be the same for both cases since the overall air-fuel ratio is constant. Thus, the observed

differences occur because of the local inhomogeneities in the flow field. Clearly, the flow

structure has profound impact on flame characteristics as indicated by Figure A.4(a). Figures

A.4(b)-(c) show that the effects of atomizing air on biodiesel and biooil flames is similar to that

for diesel flames. The images of biooil flames in Figure A.4(c) reveal green tint in the post-

combustion zone, whose origin is unknown at present.

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A.4.2 Effect of Atomizing Air on NOx and CO Emissions

Figures A.5 to A.7 present NOx and CO emissions profiles along the axis of the

combustor for different fuels used. The axial distance (z) in these profiles is measured from the

combustor inlet plane; thus z = 45 cm refers to the combustor exit plane. For diesel flames,

Figure A.5 (a) shows that the NOx concentration is nearly constant in the axial direction.

Evidently, all of the NOx is formed in a short reaction zone within z = 12 cm. Figure A.5 shows

a significant decrease in the NOx concentrations as the atomizing air is increased from 15% to

20% of the total air. Further increase in atomizing air (to 25%) results in only a modest decrease

in the NOx concentrations. This effect is related to the equivalence ratio (and hence, the reaction

zone temperature) of the reactant mixture produced for different atomizing air flow rates, as

discussed above. Higher atomizing air produces a leaner reactant mixture that burns at a lower

flame temperature to produce lower NOx concentrations.

For diesel flames, Figure A.5(b) shows that the CO concentrations increase and then

decreases in the axial direction. Initially, the CO is produced during the fuel breakdown and it is

subsequently oxidized in the reaction zone. The increase in the atomizing air tends to decrease

the CO emissions since the reactions occur at a lower flame temperature as explained previously.

The axial profiles of NOx and CO concentrations in biodiesel and biooil flames in Figures A.6

and A.7 reveal the same general trends: (i) the NOx emissions are formed within z = 12 cm and

NOx concentration is independent of the axial distance for z > 12 cm, (ii) the NOx concentration

decreases significantly with increase in atomizing air from 15% to 20% of the total air, but

marginally for increase in atomizing air from 20% to 25% of the total air, (iii) the CO

concentrations initially increase and then decrease in the axial direction, and (iv) the CO

concentrations decrease with increase in the atomizing air.

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Emissions measurements were also taken at the combustor exit plane to identify

unmixedness in the radial direction. For diesel flames, Figure A.8 shows that the NOx and CO

concentrations are nearly constant at the combustor exit plane. These results indicate that

sufficient flow mixing has taken place within the combustor to form a homogeneous product gas

mixture at the exit plane. Figure A.8 shows that NOx and CO concentrations at the combustor

exit plane decrease with increase in the atomizing air. The radial profiles of NOx and CO

concentrations at combustor exit plane for biodiesel and biooil flames in Figures A.9 and A.10

show the same general trends: (i) the NOx emissions are constant in the radial direction, (ii) the

NOx concentrations decrease with increase in the atomizing air, (iii) The CO concentrations are

independent of the radial coordinate except for 15% atomizing air in the biooil flame, where a

parabolic profile is observed, and (iv) the CO emissions decrease with increase in atomizing air.

Overall, results show that the fluid mechanics associated with the atomization process has

significant effect on the flame structure and emissions, and that different fuels respond similarly

to the flow-induced effects of the injector. Clearly, emissions are dependent not only on the fuel

properties but also upon the flame characteristics determined by the flow processes. Next, the

fuel effects are isolated by comparing the NOx and CO emissions for different fuels using the

same volume flow rates of fuel, atomizing air, and total air.

A.4.3 Fuel Effects on NOx and CO Emissions

Figures A.11to A.13 show axial profiles of NOx and CO emissions for different fuels.

Profiles in Figure A.13 (a) show that with 25% atomizing air the NOx emissions are highest for

the biooil and lowest for the biodiesel. Among the three remaining fuels (diesel, SOME, and

SOEE), the NOx emissions are highest for the diesel fuel. The high NOx emissions with

emulsified biooil are likely caused by the nitrogen present in the n-octylamine used as

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“surfactants.” Thus, alternate surfactants must be considered in the future to reduce NOx

emission from the fuel-bound nitrogen. Figure A.13 (b) shows that the CO emissions are

generally higher for diesel, lowest for biodiesel, and similar for the remaining fuels (biooil,

SOME, and SOEE).

Results show that both NOx and CO emissions are lowest for biodiesel with 25%

atomizing air. The biooil CO emissions are low, and by choosing an alternative to the nitrogen-

containing surfactant to eliminate the fuel-bound nitrogen, the biooil NOx emissions could be

reduced further. The CO and NOx emissions for diesel-biodiesel fuel blends (SOME and SOEE)

are generally lower than those for the diesel fuel. Thus, biodiesel, emulsified biooil, and bio-

diesel blends could provide emissions performance superior to the diesel fuel. Results for 20%

and 15% atomizing air show the same general trends in Figures A.11 and A.12. The only

exception is that the NOx emissions for biodiesel with 15% atomizing air are higher than those

for the diesel fuel. This result suggests that the emissions performance must be optimized by

tailoring the injector design for a given fuel. The emissions measurements at the combustor exit

plane are shown in Figures A.14 to A.16 for different fuels. Results are consistent with the

previous observations, i.e., the biodiesel produced lowest CO and NOx emissions, biooil NOx

emissions are the highest, and the CO emissions are highest for the diesel fuel.

A.5 Conclusions

In this study, NOx and CO emissions from diesel, biodiesel, emulsified biooil, and diesel-

biodiesel fuel blends were measured in an atmospheric pressure burner simulating typical

features of a gas turbine. The emulsified biooil was made by blending biooil with surfactants

containing biodiesel, alcohol, and amine. In general, the biodiesel flames produced the least

241

amounts of NOx and CO concentrations. Diesel flames produced higher CO emissions

compared to the other fuels. The high NOx emissions from biooil flames were attributed to the

nitrogen-containing surfactant. Both NOx and CO emissions were affected significantly by the

fraction of the total air used for atomization; emissions decreased with increase in the atomizing

air. Results show that even though the fuel properties are important, the flow effects can

dominate the NOx and CO emissions. For a given fuel, the emissions can be minimized by

properly tailoring the injector design and the associated combustion processes.

242

Figure A.1. Schematic diagram of the experimental setup

All dimensions in cm

243

Figure A.2. Schematic diagram (top) and photograph of the swirler (bottom) at the combustor

inlet plane

All dimensions in cm

244

Figure A.3. Injector details

245

Figure A.4. Effect of atomizing air on flame images

-4 -2 0 2 4-4 -2 0 2 4-4 -2 0 2 40

5

10

15

20

25

-4 -2 0 2 40

5

10

15

20

25

-4 -2 0 2 4 -4 -2 0 2 4

-4 -2 0 2 4-4 -2 0 2 4-4 -2 0 2 40

5

10

15

20

25

b. Biodiesel

a. Diesel

c. Biooil

15% AA 20% AA 25% AA

246

Figure A.5. Axial profiles of emissions for diesel, (a) NOx, (b) CO

Axial Distance (cm)

CO

conc

entr

atio

n(p

pm)

0

0

10

10

20

20

30

30

40

40

50

50

0 0

50 50

100 100

150 150

200 200

250 250

300 300

350 350

400 400

15% AA20% AA25% AA

(b)

Axial Distance (cm)

NO

xco

nce

ntr

atio

n(p

pm)

0

0

10

10

20

20

30

30

40

40

50

50

0 0

20 20

40 40

60 60

80 80

15% AA20% AA25% AA

(a)

247

Figure A.6. Axial profiles of emissions for biodiesel, (a) NOx, (b) CO

Axial Distance (cm)

CO

conc

entr

atio

n(p

pm

)

0

0

10

10

20

20

30

30

40

40

50

50

0 0

10 10

20 20

30 30

40 40

50 50

60 60

15% AA20% AA25% AA

(b)

Axial Distance (cm)

NO

xco

nce

ntr

atio

n(p

pm)

0

0

10

10

20

20

30

30

40

40

50

50

0 0

20 20

40 40

60 60

80 80

100 100

15% AA20% AA25% AA

(a)

248

Figure A.7. Axial profiles of emissions for biooil, (a) NOx, (b) CO

Axial Distance (cm)

NO

xco

nce

ntr

atio

n(p

pm)

0

0

10

10

20

20

30

30

40

40

50

50

0 0

40 40

80 80

120 120

160 160

200 200

15% AA20% AA25% AA

(a)

Axial Distance (cm)

CO

conc

entr

atio

n(p

pm)

0

0

10

10

20

20

30

30

40

40

50

50

0 0

30 30

60 60

90 90

120 120

150 150

15% AA20% AA25% AA

(b)

249

Figure A.8. Radial profiles of emissions for diesel, (a) NOx, (b) CO

Radial Distance (cm)

NO

xco

ncen

trat

ion

(ppm

)

-4

-4

-2

-2

0

0

2

2

4

4

0 0

20 20

40 40

60 60

80 80

15% AA20% AA25% AA

(a)


CO

conc

entr

atio

n(p

pm)

-4

-4

-2

-2

0

0

2

2

4

4

0 0

5 5

10 10

15 15

20 20

25 25

30 30

15% AA20% AA25% AA

(b)

250

Figure A.9. Radial profiles of emissions for biodiesel, (a) NOx, (b) CO


CO

conc

entr

atio

n(p

pm)

-4

-4

-2

-2

0

0

2

2

4

4

0 0

30 30

60 60

90 90

120 120

150 150

15% AA20% AA25% AA

(b)


NO

xco

ncen

trat

ion

(ppm

)

-4

-4

-2

-2

0

0

2

2

4

4

0 0

20 20

40 40

60 60

80 80

100 100

15% AA20% AA25% AA

(a)

251

Figure A.10. Radial profiles of emissions for biooil, (a) NOx, (b) CO


NO

xco

ncen

trat

ion

(ppm

)

-4

-4

-2

-2

0

0

2

2

4

4

0 0

40 40

80 80

120 120

160 160

200 200

15% AA20% AA25% AA

(a)


CO

conc

entr

atio

n(p

pm)

-4

-4

-2

-2

0

0

2

2

4

4

0 0

10 10

20 20

30 30

40 40

50 50

15% AA20% AA25% AA

(b)

252

Figure A.11. Axial profiles of emissions for 15% AA, (a) NOx, (b) CO

∗∗∗∗∗

∗

Axial Distance (cm)

NO

xco

ncen

trat

ion

(ppm

)

0

0

10

10

20

20

30

30

40

40

50

50

0 0

30 30

60 60

90 90

120 120

150 150

180 180

DieselBiodieselBiooilSOMESOEE

∗

(a)

∗∗

∗∗∗

∗

Axial Distance (cm)

CO

conc

entr

atio

n(p

pm)

0

0

10

10

20

20

30

30

40

40

50

50

0 0

50 50

100 100

150 150

200 200

250 250

300 300


∗

(b)

253


∗∗∗∗∗∗

Axial Distance (cm)

CO

conc

entr

atio

n(p

pm)

0

0

10

10

20

20

30

30

40

40

50

50

0 0

50 50

100 100

150 150

200 200

250 250

300 300

350 350

400 400


∗

(b)

∗∗∗∗∗∗

Axial Distance (cm)

NO

xco

ncen

trat

ion

(ppm

)

0

0

10

10

20

20

30

30

40

40

50

50

0 0

30 30

60 60

90 90

120 120

150 150

180 180


∗(a)

254


∗∗∗∗∗∗

Axial Distance (cm)

NO

xco

nce

ntr

atio

n(p

pm)

0

0

10

10

20

20

30

30

40

40

50

50

0 0

30 30

60 60

90 90

120 120

150 150

180 180


∗(a)

∗∗

∗

∗

∗∗

Axial Distance (cm)

CO

con

cen

trat

ion

(ppm

)

0

0

10

10

20

20

30

30

40

40

50

50

0 0

20 20

40 40

60 60

80 80

100 100


∗

(b)

255

Figure A.14. Radial profiles of emissions for 15% AA, (a) NOx, (b) CO

∗∗∗∗∗∗∗∗∗


NO

xco

ncen

trat

ion

(pp

m)

-4

-4

-2

-2

0

0

2

2

4

4

0 0

40 40

80 80

120 120

160 160

200 200


∗

(a)

∗∗

∗∗∗∗∗∗∗


CO

con

cen

trat

ion

(pp

m)

-4

-4

-2

-2

0

0

2

2

4

4

0 0

30 30

60 60

90 90

120 120

150 150


∗

(b)

256


∗∗∗∗∗∗∗∗∗


NO

xco

nce

ntr

atio

n(p

pm)

-4

-4

-2

-2

0

0

2

2

4

4

0 0

40 40

80 80

120 120

160 160

200 200


∗

(a)

∗∗∗∗∗∗∗∗∗


CO

con

cen

trat

ion

(ppm

)

-4

-4

-2

-2

0

0

2

2

4

4

0 0

20 20

40 40

60 60

80 80


∗

(b)

257


∗∗∗∗∗∗∗∗∗


CO

con

cen

trat

ion

(pp

m)

-4

-4

-2

-2

0

0

2

2

4

4

0 0

5 5

10 10

15 15

20 20

25 25

30 30


∗(b)

∗∗∗∗∗∗∗∗∗


NO

xco

ncen

trat

ion

(pp

m)

-4

-4

-2

-2

0

0

2

2

4

4

0 0

40 40

80 80

120 120

160 160

200 200


∗(a)

258

APPENDIX B

CALCULATION OF SWIRL NUMBER

The swirl number S, is defined as (Johnson, 2005):

� = 23 tan � 1− ��1− �� (B.1) � = �� (B.2)

Rc

Ri

Swirl blades

Center body

Tube

α

Flow

259

Thus, for swirler in Chapter 3,

Ri = 2.0 cm

Rc = 1.0 cm

α = 62°

S = 1.5

And for swirler in Chapter 5,

Ri = 1.3 cm

Rc = 0.7 cm

α = 48.9°

S = 0.9

260

APPENDIX C

CALCULATION OF AIR FLOW RATE IN LFE

To determine the air flow rate using the LFE, differential pressure, absolute pressure and

temperature must be measured, as specified by the manufacturer. The equation is of the form:

� = �� ∗ ∆� + � ∗ ∆�� ∗ �� (C. 1) Where:

Q = actual volumetric flow rate, CFM

B, C = flow coefficients derived from calibration

∆P = differential pressure, inH2O

µstd = viscosity of flowing gas at 20 °C, micropoise

µf = viscosity of flowing gas at flowing temperature, micropoise

The actual volumetric flow rate in liters per minute, LPM, is given by: �� = � ∗ 28.317 (C. 2) (1ft3 = 28.317 L)

Calibration coefficients are presented in Table C.1 for LFE used for air flow rate measurements

261

Table C.1

Calibration coefficients for air flow rate calculation

Coefficients Value

B 5.58597E+00

C -4.51946E-02

The viscosity in micropoise (µP) is given by Sutherland’s formula, as follows:

ST

bT

+=

23

µ

Where T is temperature in Kelvin (K) and for air:

21

58.14K

Pb

µ=

S = 110.4 K

Density in grams/liter (g/l) is given by the ideal gas law, as follows:

RT

Pabs=ρ

Where Pabs is measured absolute pressure in Pascal (Pa), R is the specific gas constant and T is

temperature in K. For air:

KKg

JR

.98.286=

Air flow rate in standard liters per minute (slpm) is given by:

SLPMLPMSLPM QQ

ρρ

×=

(C.3)

(C.4)

(C.5)

(C. 6)

(C.7)

(C.8)

262

QLPM = volumetric flow rate, LPM

=ρ air density, g/l

Where SLPMρ represents density of air at Standard Temperature and Pressure (STP). STP is

defined as 0 ºC and 1 atm. For air:

SLPMρ =1.276 g/l

Equivalence ratio (Ф) for 100% CH4 combustion is calculated as follows:

AF

AFst=φ

Where AFst is the mass-based stoichiometric Air-Fuel rate. For combustion of CH4 with air:

AFst = 17.11

AF is the actual mass-based Air-Fuel ratio and is given by:

( )( )

44 CH

air

CH

air

Q

Q

m

mAF

⋅

⋅==

ρρ

&

&

(C.9)

(C.10)

263

APPENDIX D

SAMPLE CALCULATIONS OF O2 AND CO2 CONCENTRATIONS

Oxygen (O2) and carbon dioxide (CO2) concentrations are determined from the

equilibrium reaction of each test condition. This result is used to cross-check the equivalence

ratio obtained from the measured fuel and air fuel flow rates. In this section, a sample calculation

of O2 and CO2 concentrations is presented. The procedure is to mass-balance the equilibrium

chemical equation of the case in question, assuming complete combustion, (or no dissociation),

thus, no CO, NOx or other minor species are considered in products containing only CO2, H2O,

N2 and O2. After balancing the equation, the concentration of O2 or CO2 is equal to the number

of moles of O2 or CO2 divided by the total number of moles in the dry product gas. The moles of

H2O must be subtracted from the total moles in the product gas because the sample is dried prior

to entering the gas analyzers. Next, a sample calculation is shown.

The CH4 – air stoichiometric reaction is represented by:

222224 52.72)76.3(2 NCOOHNOCH ++→++

Thus, for a given equivalence ratio Ф < 1, the lean chemical reaction is represented by:

(D.1)

264

2222224 2252.7

2)76.3(2

ONCOOHNOCH

−+++→++

φφφ

For example, for Ф = 0.8, the equation representing the chemical reaction is given by:

2222224 28.0

2

8.0

52.72)76.3(

8.0

2ONCOOHNOCH

−+++→++

Accordingly, the concentrations of O2 and CO2 in the products are given by:

222

22%

nOnNnCO

nOO

++=

222

22%

nOnNnCO

nCOCO

++=

where n is the number of moles of each species.

Substituting:

%6.41002

8.0

2

8.0

52.71

28.0

2

% 2 =×

−++

−=O

%2.91002

8.0

2

8.0

52.71

1% 2 =×

−++=CO

(D.2)

(D.3)

(D.4)

(D.5)

(D.6)

(D.7)

265

The O2 and CO2 concentrations were calculated using measured air and fuel flow rates

for each test case. These calculated values were compared with the O2 and CO2 concentrations

measured experimentally by the gas analyzer at the center of the combustor. Table D.1 shows a

summary of the results.

Table D.1

Summary of O2 and CO2 calculated and experimental results

Q (slpm) Ф

Calculated

Based on air and fuel

flow rates

Experimental

Measured by gas

analyzer

%O2 %CO2 %O2 %CO2

1020 0.65 7.9 7.3 7.0 8.0

0.75 5.7 8.6 4.6 9.5

1400 0.65 7.9 7.3 7.3 7.8

0.75 5.7 8.6 5.0 9.2

266

APPENDIX E

FLOW VELOCITY AND REYNOLDS NUMBER CALCULATIONS

Average Flow velocity (Uo) and Reynolds Number (Re) at the combustor inlet are

calculated using the injector diameter (D) as characteristic length. Uo and Re are calculated as

follows: �� = ��

Where:

Uo = Average Flow velocity, m/sec �� = Air mass flow rate, Kg/sec

ρ = Air density, Kg/m3

A = cross-section area of injector, m2, given by � = �� = ��4

Where:

D = Injector diameter = 0.0254 m

The viscosity (µ) in micropoise (µP) is given by Sutherland’s formula, as follows:

ST

bT

+=

23

µ

(E.2)

(E.1)

(E.3)

(E.4)

267

Where T is temperature in Kelvin (K) and for air:

21

58.14K

Pb

µ=

S = 110.4 K

µρ DU o=Re

For test conditions of this study, Uo and Re results are summarized in Table E.1.

Table E.1

Summary of flow velocity and Reynolds number calculations �� (Kg/min)

Q (slpm)

Tinlet (°C)

Pinlet (psi)

ρ (Kg/m3) µ (Pa.s) Uo (m/s) Re

1.3 1020

20 14.8 1.21 1.81x10-5 35.3 50876

130 15.9 0.95 2.30 x10-5 45.1 31353

260 17.0 0.77 2.79 x10-5 55.8 18588

1.8 1400

20 15.3 1.25 1.81 x10-5 47.2 62998

130 16.4 0.98 2.30 x10-5 60.6 36248

260 17.2 0.77 2.79 x10-5 76.4 22638

2.6 2040

20 34.2 2.81 1.81 x10-5 30.4 116613

130 36.2 2.16 2.30 x10-5 39.6 60635

260 36.2 1.63 2.79 x10-5 52.4 42026

(E.5)

(E.6)

(E.7)

268

APPENDIX F

SOUND PRESSURE LEVEL CALCULATION SCRIPT

This Matlab script is used to calculate total sound pressure level (dB) and A-scale (dBA)

the one-third octave sound pressure signal. An FFT of the time domain pressure signal is done in

LabVIEW. Then the total SPL is calculated using the script presented next. The total SPL is

based on a reference pressure of 20 µPa.

%Determine power per one-third octave frequency band

j=1;

for i=91:113

band1(i)=Prms(i); %generating array

end

P(j)=sum(band1); %total power in frequency band and generating SPL array

j=j+1;

for i=113:141


end

P(j)=sum(band2); %total power in frequency band

j=j+1;

for i=141:178

269


end


j=j+1;

for i=178:226


end


j=j+1;

for i=226:281


end


j=j+1;

for i=281:356


end


j=j+1;

for i=356:451


end


270

j=j+1;

for i=451:561


end


j=j+1;

for i=561:701


end


j=j+1;

for i=701:901


end


j=j+1;

for i=901:1121


end


j=j+1;

for i=1121:1401


271

end


j=j+1;

for i=1401:1776


end


j=j+1;

for i=1776:2251


end


j=j+1;

for i=2251:2801


end


j=j+1;

for i=2801:3551


end


j=j+1;

272

for i=3551:4501


end


j=j+1;

for i=4501:5616


end


j=j+1;

for i=5616:7066


end


j=j+1;

for i=7066:8896


end


%CALCULATION OF SOUND PRESSURE LEVEL PER ONE-THIRD OCTAVE BAND

Freq_band=[20 25 31.5 40 50 63 80 100 125 160 200 250 315 400 500 630 800 1000 1250

1600]; %Centers of frequency bands

273

SPL=10*log10(P/(4E-10)); %SPL per 1/3 octave band

%CALCULATION OF SOUND PRESSURE LEVEL PER FREQUENCY

Freq=0:0.2:1999.8; %Frequency range, by frequency resolution

dB_per_freq=10*log10(Prms/4E-10); %SPL at each frequency

%CALCULATION OF TOTAL dB

Rel_power=10.^(SPL./10); %Converting dB to relative power to sum dB levels

Sum_Rel_power=sum(Rel_power);

Total_dB=10*log10(Sum_Rel_power);

%CALCULATION OF dBA AND TOTAL dBA

%A-weighting correction factors:

A_factor=[50.50 44.70 39.40 34.60 30.20 26.20 22.50 19.10 16.10 13.40 10.90 8.60 6.60 4.80

3.20 1.90 0.80 0.0 -0.60 -1.00];

SPL_dBA=SPL-A_factor; %Sound pressure level in dBA

Rel_power_dBA=10.^(SPL_dBA./10); %Converting dB to relative power to sum dB levels

Sum_Rel_power_dBA=sum(Rel_power_dBA);

Total_dBA=10*log10(Sum_Rel_power_dBA);

274

APPENDIX G

UNCERTAINTY ANALYSIS

The uncertainty analysis calculation detailed in this section is consistent with the

procedure outlined by Coleman and Steele (1999). The systematic and random errors are

considered for each measured variable and the total uncertainty is given by both errors

propagated in the experimental result. The general data reduction equation is given by:

...),( 3,21 XXXrr = (G.1)

where r is the experimental result determined by Xi measured variables with 95% confidence.

The overall uncertainty of the result Ur is based on the systematic error Br and respective,

random error Pr of the result.

222rrr PBU += (G.2)

The expressions for Br and Pr are given by the systematic and random uncertainty of the

variables Xi, respectively, assuming no correlated precision uncertainties.

∑

∂∂

= 2

2

2

iXi

r BX

rB (G.3)

275

∑

∂∂

= 2

2

2

iXi

r PX

rP (G.4)

The systematic or bias uncertainty BXi is given by each instrument calibration error. The

random uncertainty or precision limit PXi is given by multiple readings of variable Xi and is

calculated as:

N

tStSP X

XX i== (G.5)

where t is the tabulated distribution with 95 % confidence for N-1 degrees of freedom, N is the

number of measurements (readings) and SX is the standard deviation of the sample set. The

standard deviation of the sample population is defined by:

2/12)(

1

1

−−

= ∑ XXN

S ix (G.6)

where the mean of the sample of the N readings is:

∑= iXN

X1

(G.7)

Overall uncertainty Ur is calculated for measured air and fuel flow rates with LFE,

followed by propagated uncertainty in equivalence ratio for experiments presented in Chapter 3.

Next, overall uncertainty is calculated for air and fuel flow measurements, and propagated

276

uncertainty in equivalence ratio for experiments in Chapter 5. Next, overall uncertainties for jet

noise and combustion noise SPL are calculated.

G.1 Flow Measurements, Low Pressure Facility

The uncertainty analysis in this section pertains to flow measurements of air and fuel

presented in Chapter 3. The uncertainty analysis is performed for an air flow rate of 150 slpm

and a fuel flow rate of 10.9 slpm. From Equation G.2, the total uncertainties in air and fuel flow

measurements are:

222AAA PBU += (G.8)

222FFF PBU += (G.9)

Both air and fuel flow rates are measured with a LFE. The bias errors for air BA and fuel

BF are given by manufacturer as 0.5 % of the measured values, thus BA = 0.75 slpm and BF =

0.05 slpm. Table G.1 shows the set of data collected to determine the random uncertainty PA in

the air flow rate and PF in the fuel flow rate, along with the mean values and standard deviations

according to equations G.6 and G.7.

277

Table G.1

Readings for air and fuel random uncertainty calculation, low pressure facility

N Air measurement (slpm) Fuel Measurement (slpm)

1 149.86 10.81

2 149.63 10.84

3 149.93 10.81

4 149.68 10.82

5 149.83 10.83

6 149.42 10.83

Mean Value 149.73 10.82

Standard Deviation 0.19 0.01

Replacing t = 2.571 for five degrees of freedom, the air flow rate random uncertainty is

PA = 0.48 slpm and the fuel flow rate random uncertainty is PF = 0.03. Substituting into

Equations G.8 and G.9 gives: UA = 0.89 slpm and UF = 0.06 slpm.

The uncertainties in measurements of air in LFE, and CH4 in LFE are propagated to

obtain the overall uncertainty in equivalence ratio. The general data reduction equation for

combustion of CH4 is:

A

F52.9=φ

From Equations G.2, G. 3 and G.4, the total uncertainty in equivalence ratio is:

( ) ( )22

22

2FA U

FU

AU

∂

∂+

∂

∂=

φφφ (G.12)

Where 0046.052.9

2−=−=

∂∂

A

F

A

φ and 0635.0

52.9==

∂∂

AF

φ

(G.11)

278

The uncertainties in air and CH4 measurements are respectively UA = ±0.89 and UF =

±0.06. Substituting in Equation G.12 allows:

0056.0±=φU

G.2 Flow Measurements, High Pressure Facility

The uncertainty analysis in this section pertains to flow measurements of air with LFE,

and fuel with mass flow controller, presented in Chapter 5. The uncertainty analysis is performed

for an air flow rate of 1020 slpm and a fuel flow rate of 69.6 slpm. The bias error for air BA is

given by manufacturer as 0.72% of measured value, thus BA = 7.3 slpm. Bias error for fuel BF is

given by manufacturer as 0.5 % of measured value plus 0.1% of full scale, BF = 0.8 slpm. Table

G.2 lists values for random uncertainty calculation.

Table G.2

Readings for air and fuel random uncertainty calculation, high pressure facility

N Air measurement (slpm) Fuel Measurement (slpm)

1 1020 69.5

2 1012 69.3

3 1025 69.7

4 1030 69.6

5 1041 69.6

6 1033 69.5

Mean Value 1027 68.8

Standard Deviation 10.2 0.14

Next, similar to calculations above, it follows that PA = 10.7 slpm, PF = 0.14 slpm. Thus,

UA = ±12.9 and UF = ±0.8, which allows:

279

UØ = ±0.01.

G.3 Pressure Measurements

The uncertainty analysis in this section pertains to pressure measurements across

injector/combustor presented in Chapter 5. The uncertainty analysis is performed for a pressure

of 101.3 KPa. The bias error BP is given by manufacturer as 0.2% of measured value, thus BP =

0.2 KPa. Table G.3 lists values for random uncertainty calculation.

Table G.3

Readings for pressure random uncertainty calculation, high pressure facility

N Air measurement (slpm)

1 100.6

2 100.6

3 101.3

4 102.0

5 101.3

6 101.3

Mean Value 101.2

Standard Deviation 0.53

Similar to calculations above, it follows that PP = 0.6 KPa. Thus,

UP = ±0.2 KPa

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