SPRAY CHARACTERISTICS AND COMBUSTION PERFORMANCE OF UNHEATED AND PREHEATED LIQUID BIOFUELS HEENA VINODKUMAR PANCHASARA A DISSERTATION

Transcription

1 SPRAY CHARACTERISTICS AND COMBUSTION PERFORMANCE OF UNHEATED AND PREHEATED LIQUID BIOFUELS by HEENA VINODKUMAR PANCHASARA 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 21

2 Copyright Heena Vinodkumar Panchasara 21 ALL RIGHTS RESERVED

3 ABSTRACT Recent increases in fuel costs, concerns for global warming, and limited supplies of fossil fuels have prompted wide spread research on renewable liquid biofuels produced domestically from agricultural feedstock. In the present research diesel, Vegetable Oil (VO), two types of biodiesel produced from VO and animal fat are investigated as potential fuels for gas turbines to generate power. Experiments are performed using a laboratory scale burner simulating gas turbine combustor operated at atmospheric pressure. A commercially available airblast (AB) atomizer is used to create the fuel spray. A parametric study of combustion performance (CO and NOx emissions) and spray characteristics (droplet diameter, drop size distribution, and mean and RMS axial velocities) is carried out by varying air to liquid mass ratio (ALR), and fuel inlet temperature in cold spray and spray flame with/without swirl air and without/with enclosure. The problems of high viscosity and poor volatility of VO (soybean oil) were addressed by using diesel-vo blends with up to 3% VO by volume. Gas chromatography/mass spectrometry, thermogravimetric analysis, and density, kinematic viscosity, surface tension and water content measurements are used to characterize the fuel properties. Characteristics of the resulting spray are measured using a laser sheet visualization system and a Phase Doppler Particle Analyzer system (PDPA). However, several operational and durability problems of using straight VO s for direct combustion occur because of their higher viscosity and low volatility compared to diesel fuel. The high kinematic viscosity of vegetable oil (VO) makes it unsuitable for direct combustion using conventional fuel preparation systems. Thus, we preheat the fuel to reduce its ii

4 kinematic viscosity and to improve fuel atomization. Measurements are obtained for fuel inlet temperature varying from 4 to C and for ALR varying from 2 to 4. Results show that an increase in the fuel inlet temperature decreases NOx and CO emissions, which can be attributed to improved fuel atomization resulting from decreased kinematic viscosity at higher fuel temperatures. Results also show a decrease in Sauter Mean Diameter (SMD) with an increase in VO temperature, regardless of the ALR at any given axial location in the spray. A significant difference in the distributions of mean and root mean square (RMS) axial velocity occurs with an increase in VO inlet temperature for a fixed ALR, presence of swirling air, and presence of flame. In general, the radial profiles show larger droplets distributed towards the edge of the spray and smaller droplets in the interior spray region. Higher VO inlet temperature and higher ALR produced a narrower spray with smaller diameter droplets and higher peak axial velocities. Swirling air flow and high temperatures in flames facilitate secondary breakup of larger droplets to significantly reduce the SMD. Finally the effect of enclosure is also studied since it represents a more realistic combustor design for any continuous flow system. The insulated enclosure eliminated the ambient air entrainment and minimized heat loss to the ambient air to create a fine spray flame with characteristics similar to those of an open flame. iii

5 DEDICATION This dissertation is dedicated to my beloved parents: Mr. Vinodkumar G. Panchasara and Mrs. Sadhana Vinod Panchasara. iv

6 ACKNOWLEDGMENTS Though only my name appears on the cover of this dissertation, many of my colleagues, friends and faculty members have encouraged and supported its production. I owe my gratitude to all those people who have made this dissertation possible and because of whom my graduate experience at UA has been one that I will cherish forever. At the outset, my sincere thanks and deepest gratitude is extended to my advisor, Dr. Ajay K. Agrawal. I have been amazingly fortunate to have an advisor who gave me the freedom to explore on my own and at the same time the guidance to recover when my steps faltered. Dr. Agrawal taught me how to question thoughts and express ideas. He has challenged me to think more critically about my work and made me a better graduate student. His patience, perseverance, practical and technical knowledge, organization skills and support helped me overcome many crisis situations and finish this dissertation. Without his continuous and untiring support, this work could not have seen its timely completion. I would take the opportunity to thank Southern Co., and U.S. Department of Energy (DOE) for supporting me during the course of my doctoral studies. I gratefully thank my committee members, Dr. Ashford, Dr. Baker, Dr. Daly and Dr. Lane, Dr. Taylor for their guidance and their invaluable support of my academic progress as well as dissertation. I wish to thank Dr. Woodbury for lending me the infra-red thermal imaging camera and spending time with me on data processing. I am also Department for helping me directly or indirectly in bringing this work to successful completion. v

7 I extend my sincere thanks to Jim Edmonds, James, Sam, Ken Dunn and Barry Johnson who have helped me to a great extent with their timely support in the machine shop as well as lab setting. I am extremely thankful to Ms. Lynn Hamric, Ms. Pamelia Bedingfield, Lisa Hinton and Ms. Betsy Singleton for their cooperation with the administrative affairs during my stay at U.A. My graduate studies would not have been the same without the social and academic challenges and diversions provided by all my colleagues-friends during my stay here. I am particularly thankful to Daniel Sequera, Ben Simmons, Troy Dent, Pankaj Kolhe, Tanisha booker, Vijaykant Sadasivuni and Lulin Jiang. We not only studied, relaxed, and traveled together, but they were always willing to provide me with any help and support in the lab as well as stylistic suggestions and substantive challenges to help me improve my presentation skills and clarify my arguments. Their efforts and friendship has really made a difference during the course of my graduate studies here. I also extend my gratitude to my friends Saskia Clayton,Aparajita Sengupta, Cehtna Maini, Indu Ankareddi,Vishal Warke, Krishna Chetri,Yogin Patel, Jayraj Sethji, Vishwas Gadhvi for being friends and for the support they have lent me during my stay in Tuscaloosa. My very special thanks to Yajuvendrasinh Parmar,Suhani Shah and Hardik Singh for their constant encouragement and support. Their unconditional love and friendship is invaluable. It s beyond words to express my gratitude towards my parents whom I owe everything I am today, Mrs. Sadhana V. Panchasara and Mr. Vinod G. Panchasara. Their unwavering faith and confidence in my abilities and in me is what has shaped me to be the person I am today. Thank you for everything. Without their love and support, encouragement and firm faith in me this would not have been possible. I also thank my brother Sumit Panchasara and my sister in law Ankita Panchasara for their love and support. vi

8 It is indeed a great pleasure to look back and take this opportunity to express my gratitude towards all the people for their help in my efforts which have really made a difference in my life. Thank you every one. vii

9 CONTENTS ABSTRACT... ii DEDICATION... iv ACKNOWLEDGMENTS...v LIST OF TABLES... xii LIST OF FIGURES... xiii 1. INTRODUCTION Background Combustion Fundamentals Liquid Fuel Combustion Objectives Overview PROPERTIES OF BIODIESEL AND DIESEL-VEGETABLE OIL BLENDS Background Fuel Property Measurements GC-MS Analysis Thermogravimetric Analysis (TGA) Density, Viscosity and Surface Tension Water Content Spray Characteristics...29 viii

10 2.3.1 Spray Angle Measurements Droplet Size Characteristics Conclusions EFFECT OF FUEL PREHEATING ON EMISSIONS FROM COMBUSTION OF VISCOUS BIOFUELS Background Experimental Set-Up Results and Discussion Visual Flame Images Emissions Profiles Conclusions CHARACTERISTICS OF PREHEATED NON-EVAPORATING BIO-OIL SPRAYS Background Experimental Set-Up Results and Discussion Spray Images SMD Contour Plots Axial Velocity Contour Plots Radial Profiles of SMD Droplet Diameter Distributions Conclusions CHARACTERISTICS OF PREHEATED BIO-OIL SPRAYS Background...97 ix

11 5.2 Experimental Set-Up and Procedure Combustion System Fuel Injection System Flare System Emissions Measurement System Phase Doppler Particle Analyzer (PDPA) Set-Up Traversing System Infrared (IR) Imaging Test Conditions Results and Discussion Effect of Swirling Air Flow on Open Cold Spray (a) Axial Velocity Contours (Mean and RMS) (b) SMD Contours (c) Transverse Profiles of Mean and RMS Axial Velocity (d) Transverse Profiles of SMD (e) Droplet Diameter Distribution Profiles Effect of Flame on Open Spray (a) Velocity Contours (Mean and RMS) (b) SMD Contours (c) Transverse Profiles of Mean and RMS Axial Velocity (d) Transverse Profiles of SMD (e) Droplet Diameter Distribution Profiles Effect of Enclosure on Spray Flame and Emissions x

12 (a) Transverse Profiles of Mean and RMS Axial Velocity (b) Transverse Profiles of SMD...13 (c) Droplet Diameter Distribution Profiles (d) Effect of Enclosed Spray on Combustion Emissions (d) Enclosure Exterior Surface Temperature Distributions Conclusions CONCLUSIONS AND RECOMMENDATIONS Conclusions Recommendations REFERENCES APPENDIX A APPENDIX B APPENDIX C APPENDIX D...27 APPENDIX E xi

13 LIST OF TABLES 2.1 List of Fuels used in the Study Results of GC-MS Analysis Physical Properties of Diesel, Biodiesel # 1, Biodiesel # 2 and Vegetable Oil Water Content in the Fuel Physical Properties of VO at 25 C Properties and Characteristics of the Insulating Material...13 xii

14 LIST OF FIGURES 1.1 (a) Airblast Injector Concept (b) Injector Details Working Principle of the Flow Blurring Injector Effect of Atomizing Airflow Rate on Visual Images of Diesel Flames: FB Injector (Top); AB Injector (Bottom) Axial Profiles of Emissions in Diesel Flames; (a) CO Concentration, (b) NOx Concentration. [Open symbols represent FB Injector, Closed Symbols represent AB Injector] Radial Profiles of Emissions in Diesel Flames; (a) CO Concentration, (b) NOx Concentration. [Open symbols represent FB Injector, Closed Symbols represent AB Injector] Axial Profiles of Emissions in Kerosene Flames; (a) CO Concentration, (b) NOx Concentration. [Open symbols represent FB Injector, Closed Symbols represent AB Injector] Radial Profiles of Emissions in Diesel Flames; (a) CO Concentration, (b) NOx Concentration. [Open symbols represent FB Injector, Closed Symbols represent AB Injector] Results of Thermogravimetric Analysis Kinematic Viscosity of Diesel and Biodiesel Fuels Kinematic Viscosity of Diesel-VO Blends Airblast Atomizer Details Spray Visualization Photographs. (a) ALR 2.52 BD-1, (b) ALR 4.35 BD-1, (c) ALR 3.92 VO Spray Cone Angle vs. Atomizing Airflow Rate SMD versus ALR. (a) Diesel, BD-1 and BD xiii

15 2.8 SMD versus ALR for Diesel, VO, and Diesel-VO Blends Schematic Diagram of the Experimental Setup; all dimensions are in cm Schematic Diagram (top) and Photograph of the Swirler (bottom) at the Combustor Inlet Plane; all dimensions are in cm Effect of Temperature on Kinematic Viscosity of Diesel and VO fuel Effect of ALR on Fuel Inlet Temperature VO Flame images at ALR 2.4 for different fuel temperatures (a) Unheated VO, (b) 58 o C, (c) 78 o C, (d) 99 o C, (e) 121 o C VO Flame images at ALR 4. for different fuel temperatures (a) Unheated VO, (b) 58 o C, (c) 78 o C, (d) 99 o C, (e) 121 o C [(a) and (b)] Axial profiles for CO and NOx for different fuel temperatures at ALR [(c) and (d)] Axial profiles for CO and NOx for different fuel temperatures at ALR [(e) and (f)] Axial profiles for CO and NOx for different fuel temperatures at ALR [(a) and (b)] Radial profiles for CO and NOx for different fuel temperatures at ALR [(c) and (d)] Radial profiles for CO and NOx for different fuel temperatures at ALR (e) and (f)] Radial profiles for CO and NOx for different fuel temperatures at ALR [(a) and (b)] Radial profiles for CO and NOx for ALR 2.4, ALR 3.5 and ALR 4. at T f = 58 o C...6 [(c) and (d)] Radial profiles for CO and NOx for ALR 2.4, ALR 3.5 and ALR 4. at T f = 78 o C...61 [(e) and (f)] Radial profiles for CO and NOx for ALR 2.4, ALR 3.5 and ALR 4. at T f = 99 o C [(g) and (h)] Radial profiles for CO and NOx for ALR 2.4, ALR 3.5 and ALR 4. at T f = 121 o C Kinematic Viscosity of Diesel and Vegetable Oil Schematic representation of the liquid break-up indicating geometry and different lengths...77 xiv

16 4.3 Schematic diagram of the experimental setup Air-Blast Injector Details Schematic of PDPA system Laser sheet spray images for VO at 4 C, 7 C and C at ALR 2.. (a) Spray Images at Exposure time ms. (b) Spray Images at Exposure time 4 ms Laser sheet spray images for VO at 4 C, 7 C and C at ALR 4.. (a) Spray Images at Exposure time ms. (b) Spray Images at Exposure time 4 ms Contours of Sauter Mean Diamter for (a) T = 4 C, ALR = 2., (b) T = 7 C, ALR (c) T = C, ALR = 2. and (d) T = 7 C, ALR = Contours of Axial Velocity for (a) T = 4 C, ALR = 2., (b) T = 7 C, ALR (c) T = C, ALR = 2. and (d) T = 7 C, ALR = Contours of Axial RMS Velocity for (a) T = 4 C, ALR = 2., (b) T = 7 C, ALR (c) T = C, ALR = 2. and (d) T = 7 C, ALR = Profiles of Sauter Mean Diameter for (a) Y =2 mm...89 (b) Y = 5 mm...9 (c) Y = 8 mm Profiles of Sauter Mean Diameter versus Fuel Viscosity for (a) Y =1 mm...92 (b) Y = 2 mm...93 (c) Y = 4 mm Profiles of Diameter Distribution for (a) T = 4 C, ALR = 2., (b) T = 7 C, ALR (c) T = C, ALR = 2. and (d) T = 7 C, ALR = Flame structure of an airblast atomizer Vaporization of a typical droplet in an idealized spray flame Schematic of the combustor experimental set-up...14 xv

17 5.4 Photographic view of the enclosure without and with insulation and schematic of the top view of the pentagonal enclosure Airblast Injector Details Schematic of Liquid Fuel Supply System Schematic of Gaseous Fuel Supply System Photographic representation of the Flare system (a) Emissions Analyzer; (b) Emissions Measurement Traversing system Schematic of the PDPA system Experimental set-up of a PDPA system mounted on a 3-way traversing system Plan view of the PDPA traversing mechanism in radial and axial coordinates Photographic view of the PDPA system integrated with the combustor assembly Photographic view of the Infrared Camera Mean axial velocity contour for cold spray without swirl Mean axial velocity contour for cold spray with swirl Mean axial velocity contour for cold spray with swirl at T f = C RMS axial velocity contour for cold spray without swirl RMS axial velocity contour for cold spray with swirl RMS axial velocity contour for cold spray with swirl at T f = C SMD contour for cold spray without swirling air SMD contour for cold spray with swirling air SMD contour for cold spray with swirling air at T f = C Transverse profiles of mean axial velocity and RMS axial velocity for cold spray with and without swirl at (a and b) Y = 5 mm...16 xvi

18 (c and d) Y = 1 mm (e and f) Y = 15 mm (g and h) Y = 2 mm (i and j) Y = 25 mm (k and l) Y = 3 mm (m and n) Y = 35 mm (o and p) Y = 4 mm (q and r) Y = 45 mm (s and t) Y = 5 mm (u and v) Y = 6 mm...17 (w and x) Y = 7 mm (y and z) Y = 75 mm Transverse profiles of mean axial velocity and RMS axial velocity for swirling air cold spray; unheated and VO at T f = C (a and b) Y = 5 mm (c and d) Y = 1 mm (e and f) Y = 15 mm (g and h) Y = 2 mm (i and j) Y = 25 mm (k and l) Y = 3 mm (m and n) Y = 35 mm (o and p) Y = 4 mm...18 (s and t) Y = 5 mm (u and v) Y = 6 mm xvii

19 (w and x) Y = 7 mm (y and z) Y = 75 mm Transverse profiles of SMD for cold spray with and without swirl at (a and b) Y = 5 mm and 1 mm (c and d) Y = 15 mm and 2 mm (e and f) Y = 25 mm and 3 mm (g and h) Y = 35 mm and 4 mm (i and j) Y = 45 mm and 5 mm (k and l) Y = 6 mm and 7 mm...19 (m) Y = 75mm Transverse profiles of SMD for swirling air cold spray; unheated and VO at T f = C (a and b) Y = 5 mm and 1 mm (c and d) Y = 15 mm and 2 mm (e and f) Y = 25 mm and 3 mm (g and h) Y = 35 mm and 4 mm (i and j) Y = 5 mm and 6 mm (k and l) Y = 7 mm and 75 mm Droplet Distribution Profile for cold spray with and without swirl at (a and b) Y = 5mm, X = mm and Y = 1 mm, X = mm c and d) Y = 4mm, X = mm and Y = 6 mm, X = mm Mean axial velocity contour for flame spray Mean axial velocity contour for flame spray for VO at C RMS axial velocity contour for flame spray...22 xviii

20 5.32 RMS Axial Velocity contour for flame spray for VO at C SMD contour for flame spray SMD contour for flame spray for VO at C Transverse profiles of mean axial velocity and RMS axial velocity for flame spray and cold spray at (a and b) Y = 5 mm...26 (c and d) Y = 1 mm...27 (e and f) Y = 15 mm...28 (g and h) Y = 2 mm...29 (i and j) Y = 25 mm (k and l) Y = 3 mm (m and n) Y = 35 mm (o and p) Y = 4 mm Transverse profiles of mean axial velocity and RMS axial velocity for flame spray and cold spray for T f = C at (a and b) Y = 5 mm (c and d) Y = 1 mm (e and f) Y = 15 mm (g and h) Y = 2 mm (i and j) Y = 25 mm (k and l) Y = 3 mm (m and n) Y = 35 mm (o and p) Y = 4 mm Transverse profiles of SMD for flame spray and cold spray at (a and b) Y = 5 mm and 1 mm xix

21 (c and d) Y = 15 mm and 2 mm (e and f) Y = 25 mm and 3 mm (g and h) Y = 35 mm and 4 mm Transverse profiles of SMD for flame spray and cold spray for T f = C at (a and b) Y = 5 mm and 1 mm (c and d) Y = 15 mm and 2 mm (e and f) Y = 25 mm and 3 mm (g and h) Y = 35 mm and 4 mm Droplet Distribution Profile for flame spray and cold spray at (a and b) Y = 5mm, X = mm and Y = 1 mm, X = mm (c and d) Y = 4mm, X = mm and Y = 6 mm, X = mm Droplet Distribution Profile for flame spray and cold spray for T f = C at (a and b) Y = 5mm, X = mm and Y = 1 mm, X = mm (c and d) Y = 4mm, X = mm and Y = 6 mm, X = mm Transverse profiles of mean axial velocity and RMS axial velocity for enclosed flame for T f = C and T f = 15 C at (a and b) Y = 5 mm (c and d) Y = 1 mm (e and f) Y = 15 mm (g and h) Y = 2 mm (i and j) Y = 25 mm (k and l) Y = 3 mm (m and n) Y = 35 mm...24 xx

22 5.42 Transverse profiles of SMD for enclosed flame for T f = C and T f = 15 C at (a and b) Y = 5 mm and 1 mm (c and d) Y = 15 mm and 2 mm (e and f) Y = 25 mm and 3 mm (g and h) Y = 35 mm Droplet Distribution Profile for enclosed flame for T f = C and T f = 15 C at (a and b) Y = 5mm, X = mm and Y = 1 mm, X = mm (c and d) Y = 3mm, X = mm and Y = 35 mm, X = mm Transverse profiles of CO for enclosed VO flame at C and 15 C Transverse profiles of NOx for enclosed VO flame at C and 15 C IR Image of the enclosure surface temperature Profile plot for exterior surface temperature at different axial location on the enclosure xxi

23 CHAPTER 1 INTRODUCTION Background Global interest in clean power generation has driven continued improvement of power systems. Improvement efforts focus on reducing emissions, increasing efficiency and lowering costs without sacrificing reliability. Most continuous combustion systems operate on a single fuel, mostly gaseous fuel. However, fuel flexible power systems that can operate on both gaseous and liquid fuels are needed to minimize the operation costs. The ignition and combustion of liquid fuels plays a major role in diesel engines, gas turbines and industrial burners. Liquid fuels are usually burned as sprays of small liquid droplets, the droplets first evaporating to produce a cloud of vapor which then burns in the gas phase Combustion Fundamentals: One of the most important parameter used to characterize combustion is the fuel to air ratio (f/a), expressed either on a volume or mass basis. With precisely enough air to consume all of the fuel, theoretically, combustion is said to have a stoichiometric f/a ratio. Adding more air produces combustion that is fuel-lean, and operating on less air produces fuel-rich combustion. Because differing fuels have different stoichiometric fuel/air ratios, it is convenient to normalize the fuel/air ratio by the stoichiometric value, 1

24 leading to the well known equivalence ratio, Φ (f / a) φ = (f / a) stoich Combustion using different types of fuel is readily described as (i) Fuel-Lean if Φ < 1 (ii) Fuel-Rich if Φ > 1 Another important combustion parameter is the adiabatic flame temperature. In principle, the highest temperature would be attained at Φ = 1, because all the fuel and oxygen would be consumed. In practice, the effects of species dissociation and heat capacity shifts the peak temperature to slightly above stoichiometric Φ ~ Liquid Fuel Combustion Combustion of liquid fuels provides a significant fraction of current total energy supply. For efficient combustion to occur the fuel and air should be intimately mixed. Since combustion of solid fuels result in higher emissions like soot and nitric oxides (NOx), liquid and gaseous fuels are becoming more prevalent. Moreover the liquid fuels offer a greater flexibility since the fuel can be transported easily as compared to gaseous fuels requiring large scale stationary operation. Combustion of liquid fuels in diesel engines, spark ignition engines, gas turbines, rocket engines and industrial furnaces is dependent on effective atomization to increase the specific surface area of the fuel, thereby achieving high mixing rates and evaporation. The process of atomization is one in which a liquid jet or sheet is disintegrated by the kinetic energy 2

25 itself, or by being subjected to a high velocity gas. A spray serves as the heart of almost every type of liquid-fueled combustion system. Recent emphasis on fuel conservation and stringent environmental regulations has prompted increased research examining the processes involved in spray production. Ideally, to promote combustion with maximum efficiency and minimum harmful emissions, an injector should deliver a fuel spray that rapidly disperses and evaporates to yield a homogeneous mixture of gaseous fuel and air. The diffusion mode of burning droplets, which promotes the formation of soot and high nitric oxide (NOx) and carbon monoxide (CO) emissions (Lefebvre, 1989], is consequentially avoided. From a practical standpoint, the fuel injector should deliver good atomization over a range of fuel mass flow rates and incur low pressure drop. Fuel jets or sheets break up into sprays in various regimes that depend upon the influence and interaction of the liquid s kinetic energy, surface tension, viscosity, and surrounding air. The mechanism by which the liquid jet breaks up into a spray was first broached by Rayleigh who analyzed the stability of a low speed, inviscid jet (Lasheras & Hopfinger, 2). In what Lefebvre (1989) has defined as classical atomization, a liquid jet or sheet, subjected to destabilizing forces, breaks up after it leaves the injector. In the classical atomization regime, a liquid jet or sheet initially disintegrates into ligaments which in turn break up into droplets, whose size depends on the characteristic length scale of the liquid jet or sheet. If the Weber number, defined as the ratio of the jet s kinetic energy to its surface tension, is small, surface instabilities control the break-up. Some of the characteristics (Lefebvre,1989), has listed for an ideal injector include good atomization over a wide range of either steady or 3

26 transient fuel flow rates, unaffected by flow instabilities, low power requirements, scalability, resistant to blockages, and delivery of a fine spray. Lin and Reitz (1988) have provided a good review along with a stability analysis to predict the breakup of liquid jets. Reviews by McDonell and Samuelsen (1991), Radzan (1998),Mansour (25), and Nakamura et al (26) deal with fuel sprays optimized for lean combustion. Several approaches to atomize the fuel the have been developed in the past. For example, a high pressure gradient across a nozzle will ensure a large velocity differential at the jet s liquidair interface which, in turn leads to rapid break-up of the fuel jet. This simple arrangement of pressure atomizer can be problematic if the fuel flow rate is to extend over a wide range. The orifice size and pressure gradient must accommodate the largest anticipated fuel flow rate, and hence, the required pressure gradient cannot be maintained for low fuel flow rates. A second strategy for fuel jet breakup transfers energy from the surrounding air to the fuel jet through the use of air-assist or air blast. The air-assist injector employs a low volume flow of high velocity air to break up the fuel jet. The air-blast atomizer delivers a relatively larger volume flow of low velocity air to both break up the fuel jet and to disperse the resulting spray to the combustion zone. For the air-assist and air-blast atomization procedures, the production of either small volumes of high pressure air or large volumes of low pressure air can be expensive in terms of auxiliary power requirements. Prefilming the fuel represents a third type of atomization process in which the liquid fuel leaves the injector surface as a sheet. Subsequently, the thin liquid sheet is subjected to air-blast on both sides to affect a fine droplet spray. The above three atomization concepts have been combined into wide varieties of fuel injectors designed to suit their particular usage. A fourth process known as effervescent atomization injects high pressure air directly into the liquid fuel before injection. Effervescent atomization, as described by Sovani et al (21), 4

27 typically produces a spray with fine, non-uniform, droplets; however, the resulting spray angle is small and air pumping costs are high since the air must be pressurized to the fuel s pumping level. Recently, Ganan-Calvo reported the discovery of a simple, reproducible atomizer configuration, the so called Flow blurring injector, with gas-liquid interactions that yield in high atomization efficiency. For a specified liquid flow rate and total energy input, Ganan-Calvo claims that the FB atomizer creates about five to fifty times more droplet surface area than any other pneumatic atomizer of the plain-jet airblast type. The basic operating principle of the FB atomizer can be seen in Figure 1. Conceptually, in a FB atomizer the air is forced through a small gap between the exit of fuel tube with inside diameter of d and a coaxial orifice in a perpendicular plate of the same diameter as d. The nozzle s wall is tapered at the outlet, and the sharpened edge of the orifice plate is the same diameter as the inner diameter of the nozzle (d). When the axial distance between the nozzle exit and orifice place (H) is small (i.e., H/d <.25), some of the gas flowing into the lateral cylindrical passageway (LCP) between the nozzle exit and orifice plane is forced upstream a short distance into the liquid nozzle. Flow recirculation and vigorous turbulent mixing of the liquid and the gas occur in this short nozzle length result in a fine spray forming downstream of the orifice plate. The FB atomizer operates on the same atomization principle as the effervescent process but overcomes some of the effervescent atomizer difficulties. A study was conducted by Panchasara (29), and Simmons (28), to demonstrate the behavior of a FB and an AB atomizer in a swirl stabilized combustion system with diesel and kerosene as fuels. The FB injector concept has never been demonstrated in combustion systems. 5

28 Some of the results from the study are presented in Figures 1.3, 1.4, 1.5 and 1.6. The performance of these atomizers is shown in terms of emission results and qualitative flame images at a given fuel and air flow rates. Figure 1.4 presents the NOx and CO emission profiles along the axis of the combustor in diesel flames for both injectors. The axial distance (z) is these profiles is measured from the injector exit plane; thus z = 45 cm refers to the combustor exit plane Figure 1.4 (a) shows that the CO concentrations tend to increase and then decrease in the axial direction. Initially, the CO is produced in the flame during the fuel breakdown process and it is subsequently oxidized in the reaction and post-reaction zones. An increase in the atomizing air flow rate decreases the CO emissions since the premixed mode of combustion becomes more prevalent. The NOx concentrations presented in figure 1.4 (b) are nearly constant in the axial direction. Evidently, most of the NOx is formed in the reaction zone within z = 12 cm. Emissions measurements were also taken at the combustor exit plane to characterize variations in the transverse flow direction. Figure 1.5 shows that the NOx and CO concentrations are nearly independent of the radial coordinate at the combustor exit plane, although flow asymmetry is evident at low atomizing airflow rates because of the turbulent fluctuations in the resulting diffusion flame. Results show that the flow mixing within the combustor forms a homogeneous product gas mixture at the combustor exit, especially at higher atomizing airflow rate. Figure 1.5 shows that the CO and NOx emissions decrease with increasing atomizing air flow rate. Emissions profiles in the axial and radial directions for the kerosene flames presented, respectively, in figures 1.6 and 1.7, show the same trends as those observed previously in diesel flames: (i) the CO emissions initially increase and then decrease in the axial direction, (ii) much of the NOx is formed in a short reaction zone near the injector exit, (iii) the CO profiles at the combustor exit display asymmetry for low atomizing airflow rates but 6

29 they are uniform at higher atomizing airflow rates when combustion occurs in the premixed mode, (iv) the NOx emission profiles are nearly uniform at the combustor exit plane, (v) both CO and NOx emissions decrease with increase in the atomizing airflow rate. Combustion of liquid fuels is largely dependent on effective ways of atomization. Fuel atomization plays an important role in combustion of liquid biofuels as well. Economic and political factors make supplies of petroleum uncertain and have given rise to tremendous price escalation, for developing countries where oil imports are imperative. Moreover, the depleting resources of petroleum-based fuels, the increasing threat to the environment from NOx, CO, hydrocarbon emissions and global warming, have given momentum to the search for alternative fuels which are renewable. Current research on biofuels mostly focuses on their use in internal combustion engine applications, but they do show promise in power generation sector in gas turbines and process burners. As fuel candidates for power generation are mushrooming worldwide; there is presently a surge of interest around the liquid biofuels. Among these biofuels available today, biodiesel fuels are the natural candidates due to their compatibility with gasoil and their increasing use in the transportation sector as well as other areas of combustion such as gas turbines for power generation. Among the many different types of alternative fuels available, edible and non-edilble oils and their derivatives can be promising options. They are renewable as the carbon released by burning of vegetable oils is used when the oil crops undergo photosynthesis. In United States, the focus of vegetable oils research is on the most abundant oil-producing crop, the soybean oil. Vegetable oils have long hydrocarbon chains with trace amounts of triglycerides, that makes them highly viscous with low volatility and hence they show difficulty in atomizing compared to 7

30 diesel fuels. Different methods are used to reduce the kinematic viscosity and improve volatility of VO fuel. One of the methods to improve the physical properties of original VO is by the process of transesterification. This chemical process however requires additional energy, time, and cost penalty to prepare the fuel for acceptable atomization. Second method is to preheat the fuel to reduce the kinematic viscosity and thereby making it possible to atomize in combustion systems. Although fuel properties are important, it has been shown experimentally that the fuel atomization process is critical for a turbine combustor to minimize the emissions for a given fuel. Previous studies on various diesel, bio-oils and biodiesel by [Sequera,et.al., 27.] have found that by tailoring the atomization process, the emissions of CO and NOx can be decreased by a factor of 2 to Objectives The present research focuses on detailed spray characteristics and combustion performance of bio-oils. At first, the fuel properties are measured to study the physical properties of biodiesel fuels, soybean oil and diesel fuels. Then the combustion performance is studied using viscous bio-oils followed with detailed spray characteristics at different fuel inlet temperature in the cold flow. Later part of the research involves the study of flame spray characteristics of these viscous bio-oils and their effect on combustion emissions. Such measurements in preheated viscous bio-oils are non-existent at this time. Experiments are conducted to measure spray characteristics as well as provide an insight into CO and NOx emissions using a commercial airblast atomizer. 8

31 1.5. Overview 1. In Chapter 2 experiments are reported to study characteristics of alternative fuels for gas turbine combustion. The fuels used in this study are biodiesels produced from vegetable (soybean) oil and animal fat (chicken), and blends of diesel-vegetable oil. Measurements are obtained to characterize the physical and chemical properties of fuel by Gas Chromatography and Mass Spectroscopy (GC-MS) analysis, Thermogravimetric Analysis (TGA) analysis, density, kinematic viscosity, surface tension and water content. Spray is characterized by spray angle measurements and drop size predictions of SMD from an air blast atomizer. Measurements are used to compare the combustion performance of bio-oils to conventional diesel fuel emissions. This part of the study provided motivation to demonstrate the combustion performance and spray characteristics of viscous bio-oils by improving the fluid mechanics of combustion process as compared to chemistry of the fuel. 2. Chapter 3 provides the emissions measurements for viscous biofuels combustion to study the effect of fuel inlet temperature on emissions. A method to preheat the vegetable oil prior to injection in the combustor is developed and implemented. Visual flame images and emissions of CO and NOx showing the effect of fuel inlet temperature are presented. 3. Chapter 4 examines the effect of preheated soybean oil using an airblast atomizer. The non-reacting spray is created at ambient conditions of temperature and pressure. Both qualitative and quantitative measurements of the spray are presented. Laser sheet visualization system is used to present the qualitative spray images and a Phase Doppler Particle Analyzer (PDPA) system is used to present the detailed quantitative droplet size 9

32 data in the spray. Such measurements of the spray using bio-oils are nonexistent at this time. 4. Chapter 5 presents the study of flame spray characteristics of soybean oil using an airblast atomizer. The measurements are obtained to present the (i) effect of swirl on spray, (ii) effect of flame on spay, (iii) effect of enclosure on spray and emissions of CO and NOx. The results of this study is expected to provide some insights into understanding the effect of droplet diameter and droplet diameter distribution in a cold spray to demonstrate the effect of co flow swirling air. For the flame spray the results of spray characteristics obtained show the effect on pollutant emissions since they are affected by the fuel spray and atomization characteristics and the fuel properties and hence improve the combustion performance. The measurements are taken at various locations in the spray (near and far field of the spray) axially and radially. 5. Finally, in chapter 6, important conclusions from this study are presented and recommendations for future studies are provided. 1

33 (a) Air Fuel (b) Orifice Disc Distributor Nozzle Body Screw Pin O-ring Seal Figure 1.1 (a) Air blast Injector Concept (b) Injector Details 11

34 Spray d Atomizing air H d Fuel Figure 1.2 Working Principle of the Flow Blurring Injector 12

35 % 15% 2% 25% Figure 1.3. Effect of Atomizing Airflow Rate on Visual Images of Kerosene Flames: FB Injector (Top), AB Injector (Bottom) % 15% 2% 25% 13

36 (a) 1% AA-FB 18 15% AA -FB 18 2% AA - FB 16 15% AA - AB 16 2% AA -AB CO (ppm) Axial Distance (cm) (b) NOx (ppm) 1% AA - FB % AA - FB 125 2% AA - FB 15%AA - AB 2% AA - AB Axial Distance (cm) Figure 1.4. Axial Profiles of Emissions in Diesel Flames; (a) CO Concentration, (b) NOx Concentration. [Open symbols represent FB Injector, Closed Symbols represent AB Injector]. 14

37 (a) 18 1% AA- FB 15% AA- FB % AA- FB 15% AA- AB % AA- AB 14 CO (ppm) Radial Distance (cm) (b) NOx (ppm) 125 1% AA- FB % AA- FB 2% AA- FB 15% AA- AB 2% AA- AB Radial Distance (cm) Figure 1.5. Radial Profiles of Emissions in Diesel Flames; (a) CO Concentration, (b) NOx Concentration. [Open symbols represent FB Injector, Closed Symbols represent AB Injector]. 15

38 (a) 1% AA- FB 15% AA- FB 2% AA- FB 1% AA- AB 15% AA- AB 1 3 2% AA- AB 1 3 CO (ppm) Axial Distance (cm) (b) NOx (ppm) % AA- FB 15% AA- FB 8 2% AA- FB 1% AA- AB 8 15% AA- AB 6 2% AA- AB Axial Distacne (cm) Figure 1.6. Axial Profiles of Emissions in Kerosene Flames; (a) CO Concentration, (b) NOx Concentration. [Open symbols represent FB Injector, Closed Symbols represent AB Injector]. 16

39 (a) 1% AA - FB 15% AA - FB 2% AA- FB 1% AA- AB 15% AA- AB 1 3 2% AA- AB 1 3 CO (ppm) Radial Distance (cm) (b) NOx (ppm) 1% AA- FB % AA- FB 125 2% AA- FB 1% AA- AB 15% AA- AB 2% AA- AB Radial Distance (cm) Figure 1.7. Radial Profiles of Emissions in Diesel Flames; (a) CO Concentration, (b) NOx Concentration. [Open symbols represent FB Injector, Closed Symbols represent AB Injector]. 17

40 CHAPTER 2 PROPERTIES OF BIODIESEL AND DIESEL-VEGETABLE OIL BLENDS Background In recent years, two important factors have influenced the world s energy market. First, the increased demand for petroleum fuels has steadily raised their price to impose economic burden on importing nations. Second, the concerns for global warming have increased scrutiny on emissions of greenhouse gases such as carbon dioxide (CO 2 ) formed by the combustion of fossil fuels. The depletion of conventional fossil fuels is also a cause for concern. The above factors have prompted research on alternative energy sources. Biofuels derived from agricultural products (i.e., biomass) offer the potential for a sustainable energy future. Unlike nonrenewable fossil fuels, the CO 2 emitted from the combustion of biofuels is recycled during cultivation of the biomass. Biofuels can increase energy security since the biomass feedstock is domestic, diffuse, and hence, secure. However, for commercial viability, the land usage and fossil energy input for biofuel production must be reduced through process integration/optimization and/or specialty energy crops such as microalgae (Chisti, 27). Ethanol and biodiesel are the two commonly available liquid biofuels in the US. These fuels are currently mixed with gasoline or diesel fuel in small amounts (up to 1% for ethanol and up to 2% for biodiesel) to reduce the demand for fossil fuels for vehicular transportation. Accordingly, several studies have been performed to evaluate performance of these fuels for automotive internal combustion engines (Agrawal, 27; Monyem, et.al, 21 and Rakopoulos, 26). 18

41 Although market demand currently favors the use of biodiesel mainly for vehicles, the liquid biofuels could also serve as an alternative fuel for the power generation industry, for example, by replacing the fuel in land based diesel or gas turbine engines. Biodiesel is obtained from oils such as vegetable oil (VO) or animal fats through a process known as trans-esterification. In this process, the vegetable oil or animal fat (triglyceride) reacts with an alcohol in the presence of a catalyst such as sodium or potassium hydroxide to produce biodiesel and glycerol (Demirbas, 25). The most common form of biodiesel in the United States is made from soybean oil (vegetable oil) and methanol and it is known as soy methyl ester (SME) or methyl soyate. The trans-esterification process improves the physical properties of the original VO, and, in particular, decreases the viscosity to improve fuel atomization in the combustion system. This process however incurs a cost penalty and hence, the price of biodiesel exceeds that of the VO use d to produce it. Furthermore, glycerol, co-produced with biodiesel, is often a waste product since it is yet to find a commercial market. Thus, the direct use of VO in a gas turbine to generate power can offer economic benefits, especially if the problem of fuel atomization associated with high viscosity can be alleviated by blending VO with diesel or biodiesel fuels. The continuous flow operation of gas turbines offers greater flexibility, unlike the unsteady processes of internal combustion engines imposing constraints on fuel chemistry in terms of octane and/or cetane numbers. Field tests in recent years have demonstrated the technical feasibility of gas turbine engine operation on fuels such as bio-oil, ethanol, and biodiesel (Lupandin, et. al, 25, Chiang, et. al, 27, Moliere, et. al, 27, and Bolszo, et.al, 27). However, these tests have produced limited data on emissions of nitric oxides (NOx) and carbon monoxide (CO). Detailed NOx and CO emissions measurements in an atmospheric 19

42 pressure burner simulating typical features of a gas turbine combustor were reported by Sequera et al. (27) for flames of diesel, biodiesel, emulsified biooil, and diesel-biodiesel fuel blends. (Sequera et al. 27) found that the NOx and CO emissions are determined mainly by the fuel atomization process. Although fuel properties are important, the atomization process itself could be tailored to minimize the emissions for a given fuel. The objective of this part of the study is to document physical properties of biodiesel and diesel-vo blends. Biodiesels produced from vegetable oil (soybean) and animal fat (chicken) are considered in this study. In case of VO, the fuel must be heated to reduce its kinematic viscosity so that the pressure drop in the fuel supply line is reasonable, atomization occurs with fine droplets, and combustion emissions are acceptable. To circumvent the need for heating the VO, and thus, to avoid hardware changes to the fuel supply system, diesel-vo blends are investigated. The emissions measurements to demonstrate the combustion performance of different fuels are reported by Panchasara, et.al, (29). Table 2.1 lists the fuels used for the experiments. The physical and chemical properties of fuel are characterized by gas chromatography and mass spectrometry (GC-MS), thermogravimetric (TGA) analysis, and kinematic viscosity, surface tension, density, and water content measurements. Spray characteristics are determined by spray angle measurements and predictions of the Sauter mean diameter from an air blast atomizer. The following sections report the fuel property measurements and the spray characteristics followed by the conclusions. 2

43 Table 2.1. List of Fuels used in the Study Fuel Description Diesel Commercial grade # 2 Biodiesel # 1 (BD-1) Biodiesel # 2 (BD-2) Vegetable Oil (VO) Methyl soyate made from soybean oil Methyl soyate made from chicken fat % refined soybean oil 9-1 Blend 9% Diesel 1% VO 8-2 Blend 8% Diesel 2% VO 7-3 Blend 7% Diesel 3% VO 2.2. Fuel property measurements It is important to characterize chemical and physical properties of various fuels because they were obtained from different sources. First, the fuel composition was determined using the GC-MS technique. The GC-MS analysis of biodiesel fuels (BD-1 and BD-2) was used to determine the fuel s low heating value (LHV) and an equivalent chemical formula. The TGA analysis was performed to characterize the fuel volatility. Fuel kinematic viscosity and surface tension measurements were measured to determine an acceptable operational range and to compute the Sauter mean diameter. 21

44 Table 2.2. Results of GC-MS Analysis Fatty Acid Formula %Mass %Mass Molecular Gas Heat of LHV Methyl Esters BD-1 BD-2 Weight Formation kj/kg KJ/kmol (Turns,S.R.) Lenoleic Methyl C 19 H 34 O , 37,634 ester Oleic Methyl C 19 H 36 O ,25 37,756 Ester Palmitic Methyl C 17 H 34 O ,5 37,284 Ester Steartic Methyl C 19 H 38 O ,4 38,314 Ester Tetradecanoic C 15 H 3 O ,3 36,526 Methyl Ester 9-Hexadecanoic C 17 H 32 O , 37,184 Methyl Ester 22

45 GC-MS Analysis The GC-MS analysis technique combines gas chromatography and mass spectrometry to determine % mass (or % volume) of species present in a chemical (by gas chromatography) together with the elemental composition of each species (by mass spectrometry). The analysis was carried out at UA s GC-MS Facility using HP 689 GC connected with a Micro mass AutoSpec-Ultima TM NT Mass spectrometer. GC was performed with column DB-1, inlet temperature of 225 C; split ratio of 3, oven temperature of C, heating rate of 1 C/min and hold time of 25 min. The GC column did not produce any peaks for the VO sample because of the low column temperature. Table 2.2 presents the analysis results for the two biodiesel fuels. In BD-1, linoleic methyl ester is the major component, while that in BD-2 is oleic methyl ester. The last two esters, i.e., tetradecanoic methyl ester and 9-hexadecenoic methyl ester were found in small amounts in BD-2, but they were not present in BD-1. Table 2.2 also lists the enthalpy of formation of each component taken from ( chemistry).this information was used to compute fuel component s low heating value (LHV) assuming complete combustion. Evidently, the LHV of different biodiesel components is within 2% of each other. Based on these results, the equivalent chemical formulas for BD-1 and BD-2 are, C H O 2 and C H 35.1 O 2, respectively Thermogravimetric Analysis (TGA) Thermogravimetric analysis is an experimental approach to determine a material s thermal stability and/or its fraction of volatile components by monitoring the weight change as a specimen is heated. A volatile sample loses mass with an increase in temperature and after a certain temperature it loses maximum of its mass and what remains is the residue mass. The 23

46 volatility has a significant effect on fuel vaporization which in turn affects fuel-air mixing and combustion processes. For example, poor vaporization can results in fuel droplets burning in the diffusion mode. The resulting high flame temperatures are known to produce high concentrations of NOx, CO, and soot emissions (Turns, S.R., 2). The fuel samples (8 to 17 mg) were characterized using TA 295 analyzer from TA instruments. The maximum temperature was set at 6 C, and testing was conducted from the ambient temperature to 6 C with a heating rate of 5 C/min, in an atmosphere of air supplied at the rate of 9 cm 3 /min. Before starting the experiment the sample loading pan was flame cleaned and tarred to zero. The microbalance was calibrated with standard weights varying from 1 mg to mg. Figure 2.1 shows the TGA profiles for diesel, BD-1, BD-2, and VO. Results show that diesel is the most volatile fuel among all the fuels listed in Fig Diesel evaporated rapidly when heated to 8 C and it completely vaporized at approximately 18 C. Both biodiesel fuels show similar volatility characteristics; negligible mass loss when heated below C, rapid mass loss (or vaporization) at temperatures above 15 C, and nearly complete vaporization approximately at 225 C. The vegetable oil is the least volatile fuel, showing negligible mass loss at 225 C. In this case, much of the mass loss occurs at temperatures between 25 C and 5 C. Clearly, VO will require significantly higher thermal feedback from the flame to the fuel droplets to fully pre-vaporize the fuel, and thus, to create a homogenous fuel-air mixture prior to combustion. Results in Fig. 2.1 explain the choice of diesel-vo fuel blends; blending a high volatility fuel (diesel) with a low volatility fuel (VO) would improve fuel vaporization, which can result in lower combustion emissions. Figure 2.1 includes TGA data for BD-2 from (Cayli, 27). Results show an excellent agreement between present results and data reported in the literature. Minor differences are attributed to the 24

47 presence of an oxidizing environment in this study (as opposed to an inert gas), which tends to increase the oxidation of the fuel, especially at higher temperatures Density, Viscosity, and Surface Tension Table 2.3 lists the measured physical properties of various fuels together with the relevant measurement uncertainties. The density measurements were performed gravimetrically at 25 C using volumetric glassware. Three test runs were taken to ensure the repeatability and accuracy of the data. Diesel is a lighter fuel compared to both biodiesels, while the density of the VO is the highest. Density measurements are important to determine the volumetric flow rate of the fuel to obtain a specified heat release rate in the combustor. Table 2.3 also lists the LHV of fuels on a volumetric basis. The LHV of the two biodiesel fuels was determined from component analysis presented in Table 2.2, while data from existing literature were used for the remaining fuels. For a given heat release rate, biodiesel fuels will require nearly 13% higher volume flow rate compared to the diesel fuel. Kinematic viscosity is an important physical property affecting pressure drop in the fuel line as well as the fuel atomization in the combustor. High fuel viscosity can result in excessive pressure drop and produce spray with large droplets, which deteriorates the combustion performance. The kinematic viscosity was measured by Cannon-Fenske Type 2 viscometer (Cannon Instrument Co., State College, PA, USA) with viscosity range of 2- mm 2 /sec (centi-stokes or cst). The viscometer was immersed in a water bath connected to a temperature controlled heat exchanger. The temperature was controlled using a Fisher Scientific Isotemp 116S recirculating water bath (Pittsburgh, PA, USA). The kinematic viscosity was measured for temperature range from 2 C to 7 C. Table 2.3 lists the viscosity values at 25 C, while the detailed profiles are shown in Figs 2.2 and

48 Table 2.3 and Figure 2.2 show that the kinematic viscosity of biodiesels is 45-55% higher compared to that of diesel. The kinematic viscosity of both biodiesels is nearly the same over the entire temperature range. The higher fuel viscosity will require higher pumping power for biodiesel compared to diesel to overcome the pressure drop in the fuel supply line. Furthermore, the high fuel viscosity of biodiesel can also negatively impact fuel atomization, an effect that will be quantified later on by droplet size predictions. Table 2.3. Physical properties of Diesel, Biodiesel # 1, Biodiesel # 2 and Vegetable Oil Property Fuel Diesel BD-1 BD-2 VO Mol. Weight kg/kmol Density at 25 C, 925. ± kg/m ± ± ± Viscosity at 25 C, ± mm ± ± ±.16 /s.22 Surface Tension at 25 C mn/m 28.2 ± ± ± ±.6 37 LHV, kj/kg (Rakopoul os,26) LHV, MJ/m 3 37,198 (Turns, S.R.,2) 33,442 32,689 34,225 (Rakopoul os,26) Table 2.3 and Figure 2.3 show that the kinematic viscosity of vegetable oil is higher than that of diesel by more than an order of magnitude. Fuels with such high viscosity do not produce acceptable atomization using existing fuel injectors, which was also verified experimentally. Hence, blends of VO with diesel were prepared to reduce the fuel viscosity. Figure 2.3 shows the 26

49 measured and computed kinematic viscosity of the 7-3 diesel-vo blend (by volume). The viscosity of the blend was calculated from Eq. 1 for a mixture (Ejim, C.E, 27): n lnν = x lnν Eq. 1 i= 1 i i Where, x i is the volume fraction of the species i, and ν i is the kinematic viscosity of the species i. The excellent agreement between measured and computed values for the 7-3 diesel- VO blend in Figure 2.3 verifies application of Eq. 1. Thus, Eq. 1 was used to predict the kinematic viscosity of 8-2 and 9-1 diesel-vo blends as shown in Figure 2.3. Results show that the diesel viscosity at 25 C matches the viscosity of 9-1 blend at 4 C, 8-2 blend at 5 C and 7-3 blend at 6 C. At 25 C, the biodiesel viscosity is nearly the same as that of the 8-2 diesel-vo blend. Further, vegetable oil will require temperature above 8 C to replicate the room temperature viscosity of diesel. For a given temperature, the viscosity of 7-3 blend is more than twice of that of the diesel. Hence, this blend was considered as the upper operating limit for VO content in the diesel fuel. The blends remained miscible for all volume ratios. Accordingly, the present combustor system was operated with a variety of fuels in the viscosity range of 3.5 mm 2 /s 1. mm 2 /s at room temperature. A Kruss K-12 model Processor Tensiometer with a Platinum ring was used to measure the surface tension of the fuel samples at 25 C. Three test runs were taken for each fuel to ensure the repeatability of the results. The surface tension was also calculated theoretically from the surface tension data of individual methyl esters listed in Table-2.3. The surface tension (σ m ) for the fuel blends was evaluated from Eq. 2 (Ejim, et.al, 27) n σ m = yi σ i i= 1 ( ) 1/ 4 4 Eq. 2 27

50 Here, y i is the mass fraction of species i, and σ i is the surface tension of species i. For biodiesel fuels, measured values of surface tension were within.5% of those computed using Eq. 2 and published surface tension data for esters listed in Table 2.3 (Ejim, et.al, 27) Water Content The water content of the fuels was measured to observe the percent volume of water present in the fuels. The water content of each fuel (as received) was determined experimentally using a volumetric Aquastar Karl Fischer titrator (EM Science, Gibbstown, NJ, USA) with Composite 5 solution as the titrant and anhydrous methanol as the solvent. Each sample was at least.1 g and triplicate measurements were performed on each sample at 25 C. The equilibrium water content for each sample was determined by first contacting equal volumes of fuel and deionized water for 24 hours followed by titration of the fuel sample using the above method. Deionized water was obtained from a commercial deionizer (Culligan, Northbrook, IL, USA). Table 2.4 summarizes results from the water content analysis. Table 2.4. Water Content in the Fuel Fuel % Mass of Water in % Mass of Water in Fuel Equilibrated Fuel Diesel.6.6 BD BD VO

51 2.3. Spray characteristics Spray Angle Measurements The spray angle measurements were done on the experimental set-up shown in figure 2.4. The test set-up consists of combustor assembly and injector assembly. The details of the combustor assembly will be explained in Chapter 3. Primarily the combustor assembly consists of a plenum, mixing chamber and a swirler. The injector system runs through the plenum and the mixing chamber. An O-ring within a sleeve is located at the bottom of the plenum to prevent air leakage. A commercial air-blast atomizer (Delavan Siphon type SNA nozzle) with its details shown schematically and photographically in Fig. 2.4 was used for the experiments. This commercial version creates a swirling flow of atomizing air to breakdown the fuel jet as the fluids exit the orifice plate. The liquid fuel was supplied by a peristaltic pump with reported calibration error of ±.25% of the flow rate reading ranging from 12 ml/min to 13 ml/min in steps of 2 ml/min. Viton tubes were used to prevent degradation of the fuel lines. A 25µm filter was placed in the fuel supply line to prevent dirt and foreign particles from entering into the injector. The flow rate of injection air was varied for ALR of 2.52 and The fuel flow rate was kept constant at 12 mlpm. The spray images were taken for open cold spray conditions as shown in figure 2.5 for BD-1, BD-2 and VO. Images were taken with a regular digital camera for different ALR s. The spray cone angle was calculated manually based on the slope of best curve fit line on the edge of the spray from the pixel coordinate information obtained using image editing software as shown in figure 2.6. The Figure 2.5 shows photographs of BD-1 spray for different atomizing airflow rates. A photograph of the VO spray is also shown for comparison. Visually, biodiesel produced a fine spray, while the VO spray contained large droplets around the outer edge. These large droplets give the appearance of a larger cone angle with VO, a 29

52 problem that was addressed during data analysis using intensity profile to compute the cone angle. Figure 2.6 shows similar spray cone angle for diesel and biodiesel fuels. The cone angle for VO is smaller by about 5 degrees. For all cases, the cone angle increased with increasing atomizing air flow rate, signifying increased penetration of fuel into the combustor Droplet Size Characteristics Detailed characterization of the fuel spray requires measurements of drop size distributions, for example, using phase Doppler particle analyzer and/or laser diffraction techniques. As the first step, we performed droplet size calculations using the correlation of Rizk and Lefebvre, (1984), for the Sauter Mean Diameter (SMD) from an air-blast injector: SMD d o =.48 σ 2 ρ U d A R o ALR μ L σρ d L o ALR Eq. 3 with, SMD: Sauter Mean diameter, μm d o : σ : μ L: ρ A : Liquid discharge orifice diameter, m Surface tension, N/m Liquid fuel viscosity, m 2 /sec Density of air, kg/m3 ρ L : Liquid fuel density, kg/m 3 U R : Relative velocity of co flowing ALR : Air to liquid mass flow ratio Figure 2.7(a) presents the Sauter mean diameter versus the air to liquid mass flow ratio (ALR) for diesel and biodiesel. Note that 15% AA pertains to ALR of 2.65 and 2.52, 3

53 respectively, for diesel and biodiesel fuels. The corresponding values for 25% AA are 4.35 and Figure 2.7(a) shows that increasing the atomizing airflow rate (for a fixed fuel flow rate) improves atomization by decreasing the SMD of the droplets produced. For a given ALR, the biodiesel droplets are larger than diesel droplets by about 2 μm. The SMD of the two biodiesel fuels is nearly the same for all ALR values. Figure 2.7(b) presents the SMD for diesel, VO, and diesel-vo blends. For VO, the 15% AA and 25% AA correspond to ALR of 2.4 and 3.92, respectively. For the entire range of ALR, the SMD of VO is two to three times higher than that of the diesel. However, the SMD of diesel-vo blends is only moderately higher than that of the diesel. Interestingly, the 7 3 diesel-vo blend produces nearly the same droplet diameter at different ALR values as the two biodiesels (see Fig. 2.7(a)). The results in Fig. 2.7 have provided a quantitative assessment of atomization behavior of different fuels, but an accurate knowledge of the droplet size requires direct measurements in the spray as presented in later chapters Conclusions Fuel properties and spray characteristics were measured for diesel, two types of biodiesel, and diesel-vo blends. The composition of biodiesel fuels was determined using the GC-MS technique, which also allowed us to compute the low heating value of the fuel. The main constituent of VO biodiesel was linoleic methyl ester while that for animal fat biodiesel was oleic methyl ester. In spite of the compositional differences, the volatility, kinematic viscosity, and surface tension properties of the two biodiesel fuels were similar. The kinematic viscosity of biodiesel fuels was about 5% higher than that of diesel. The kinematic viscosity of VO was more than 1 times that of diesel, and hence, VO would require blending with a low viscosity fuel. Diesel-VO blends with up to 3% VO (by volume) were used in this study to achieve an acceptable range of fuel viscosity; about twice that of the diesel and equal to that of biodiesel. 31

54 The TGA analysis indicated diesel to be the most volatile, while VO was the least volatile fuel. The Sauter mean diameter in biodiesel spray was estimated to be only slightly larger than that in diesel spray. The estimated SMD of biodiesel and 7-3 diesel-vo blends was nearly the same. 32

55 Diesel BD-1 BD-2 Oleic Methyl Ester [Cayli, [Ref. 13] VO 12 %Mass Temperature ( o C) Figure 2.1. Results of Thermo gravimetric Analysis 33

56 Kinematic Viscosity (mm 2 /s ) Diesel BD-1 8 BD Temperature ( o C) Figure 2.2. Kinematic Viscosity of Diesel and Biodiesel Fuels 34

57 Kinematic Viscosity, (mm 2 /s) 1 2 Diesel VO 7-3 blend ( Exp) 7-3 blend (Eq.1) 8-2 blend 9-1 blend Temperature o C Figure 2.3. Kinematic Viscosity of Diesel VO blends 35

58 Figure 2.4 Air Blast Atomizer Details 36

59 BD-1 BD-1 VO ALR 2.52 ALR 2.52 ALR 4.35 Figure 2.5.Spray Visualization Photographs 37

60 Spray Angle (Degrees) Diesel 4 BD-1 BD-2 4 VO %AA Figure 2.6 Spray Cone Angle vs Atomizing Airflow Rate 38

61 Sauter Mean Diameter (μm) Diesel 35 BD-1 BD AirtoLiquidMassRatio(ALR) Figure 2.7. SMD versus ALR for Diesel, BD-1 and BD-2. 39

62 Sauter Mean Diameter (μm) Diesel VO 7-3 Diesel-VO 8-2 Diesel-VO Diesel-VO AirtoLiquidMassRatio(ALR) 4 5 Figure 2.8. SMD versus ALR for Diesel, VO, and Diesel-VO Blends 4

63 CHAPTER 3 EFFECT OF FUEL PREHEATING ON EMISSIONS FROM COMBUSTION OF VISCOUS BIOFUELS Background Growing concern over effects of greenhouse gases on the environment has revived interest in alternate fuels to replace the currently prevalent fossil fuels. Vegetable oils (VOs) and their derivatives are potential biofuel alternative that have received significant attention in recent years. The most common of all the vegetable oils in the U.S. is the soybean oil. Results from Chapter 2 have shown that it is difficult to atomize VO because longer hydrocarbon chain lengths with trace amounts of triglycerides results in high kinematic viscosity and low volatility of VO compared to diesel fuel. At room temperature, the VO kinematic viscosity is about 2 times higher than that of diesel fuel [Chapter 2]. Poor volatility of VO also makes it difficult for the fuel droplets to fully pre-vaporize before combustion. One of the methods to improve the physical properties and fuel atomization of VO for efficient combustion is the process of trans-esterification, which produces soy methyl ester (SME) or biodiesel. This chemical process however incurs a cost penalty, and hence, the price of biodiesel is higher than of VO. Furthermore, glycerol, a byproduct of the process remains to be a waste product. The second method is to preheat the VO to reduce its kinematic viscosity, and thus, make it possible to atomize the fuel. Much of the previous research with heated VO has been conducted in diesel engines [Simmons, 28; and Bari, 27]. Heated VO could also be used in stationary systems such as power generating gas turbines and furnaces. 41

64 Thus, this research seeks to investigate how heating the VO affects combustion emissions. A commercially available air blast (AB) atomizer is used to atomize the fuel (soybean oil) for experiments conducted in a swirl-stabilized burner operated at atmospheric pressure. Details of the experiment set-up and some preliminary results of visual flame images are presented in the following sections Experimental Set-Up The test apparatus is shown schematically in Fig The test set-up consists of combustor assembly and injector assembly. The injector assembly and the details of the injector are explained in Chapter 2 and figure 2.4. The modification here is with the addition of fuel heater to preheat the liquid fuel before atomization and modification of the fuel flow path to overcome the heat losses. The combustion airflow path includes a plenum filled with marbles to breakdown the large vortical structures, a swirler in the mixing section, and another swirler at the combustor inlet (see Fig. 3.2) to enhance fuel-air mixing and improve flame stability. The bulk inlet velocity of the combustion air is 2.7 m/s, which results in a Reynolds number varying from 315 based on the equivalent diameter of the injector. An air-cooled quartz glass column of inside diameter 8. cm and length 46 cm is used to enclose the flame. Air from a compressor passed through a pressure regulator, a dehumidifier, and water traps to remove any moisture present. Then, the air was split into combustion air supply and atomizing air supply lines. The air flow rates were measured by the laminar flow elements (LFE). The flow rates measured by the LFE s were corrected for temperature and pressure as specified by the manufacturer. The liquid fuel supplied by a peristaltic pump passes through a pulse dampener, a fuel filter and an electric fuel heater (Infinity Fluids Corporations, CRES-ILB- 12/24 inline water/liquid heater). The heater uses a Proportional Derivative Controller (PID) 42

65 control unit to control the fuel temperature within the accuracy of ±.5 o C, measured at the heater outlet by a K-type thermocouple. Heated fuel enters fuel injector through 5 cm long tube of 5 mm ID to minimize the heat loss. The fuel temperature at the injector inlet was measured by a K- Type thermocouple. The experiment was started by supplying gaseous methane and then, igniting the methane-air reactant mixture 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 rate of total air (combustion + atomizing) was constant at 15 standard lpm. Experiments were conducted for fixed volume flow rate of fuel at 12 mlpm which is measured before preheating of the fuel. The atomizing airflow rate was varied to obtain ALR of 2.4, 3.5 and 4.. The equivalence ratio during the experiments was kept constant at Φ =.78 for the fixed volume flow rate of fuel and air. The time required for the fuel to flow through the system and reach the atomizer was about 3 minutes. Another -4 minutes were required before the fuel temperature reached the set value. The product gas was sampled continuously by a quartz probe (OD = 7. 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 probe was traversed in the axial direction at the center of the combustor and in the radial direction at the combustor exit plane. The gas sample passed through an ice bath and water traps to remove moisture upstream of the gas analyzers. The dry sample was sent through the electrochemical analyzers to measure the concentrations of CO and NOx in ppm. The analyzer also measured oxygen and CO 2 concentrations, which were used to cross-check the equivalence ratio computed from the measured fuel and air flow rates. The uncorrected emissions data on dry basis are reported with uncertainty of +/-2 ppm. 43

66 Combustion performance is characterized by measuring CO and NOx emission profiles within the combustor. Experiments are conducted using an AB atomizer for a fixed fuel flow rate and the total air flow rate which will be discussed in the experimental set-up section. The total air flow rate is split between the atomizing air and primary combustion air while air to fuel mass ratio is varied through the injector to document the fluid dynamics effects on emissions. Results and discussion section include the visual flame images for different fuel inlet temperatures as increased on heater in the steps of 3 C at different atomizing air ratios, and axial and radial emissions profiles Results and Discussion Table 3.1 lists the key physical properties of VO taken from chapter 2. Figure 3.2 shows the kinematic viscosity of VO, which is 2 times higher than that of diesel at room temperature. Figure 3.3 shows that VO should be heated to about 8 C to replicate the room temperature kinematic viscosity of diesel. In this study, the fuel temperature at the heater outlet (T h ) was varied from 7 to 16 C in steps of 3 C. It resulted in fuel temperature (T f ) at the injector inlet from 57 C to 122 C, depending upon the ALR. Figure 3.4 illustrates this relationship between the heater exit and injector inlet temperatures. For a fixed heater exit temperature, the fuel inlet temperature increased by about.5 to 1.5 o C as the ALR increased from 2.4 to 4.. Fuel inlet temperature for ALR = 3.5 is chosen to explain the results. Table 3.1. Physical properties of VO at 25 C Density Kinematic Viscosity kg/m 3 mm 2 /sec Surface Tension mn/m LHV MJ/m ± ± ±.6 34,225 44

67 Visual Flame Images Figures 3.5 and 3.6 show the photographic flame images at ALR of 2.4 and 4., respectively, taken by a digital camera to obtain a qualitative understanding of the flame characteristics. For ALR of 2.4, Figure 3.5 depicts a short blue premixed region near the injector exit followed by a large yellow-orange diffusion flame zone. As the fuel inlet temperature increases, the flame appears to be shorter and the blue premixed zone extends farther downstream. At this low ALR the fuel droplets burn in diffusion mode attributing to the distinct yellow flame since they don t fully vaporize and premix with air before entering the reaction zone. Further, poor atomization at low ALR produces large droplets that tended to accumulate on the combustor wall and gradually formed a ring around the interior of the glass enclosure as seen in the photograph. The ring can still be observed at elevated fuel temperatures since the droplet accumulation started during the initial start-up when the fuel did not yet reach the set temperature. Accumulated droplets subsequently pyrolized to form a distinct black ring around the combustor wall. At a higher ALR of 4., as shown in Fig. 3.6, the flames show a distinct blue color typical of premixed mode of combustion. These images suggest that most of the fuel droplets prevaporize and premix with air before the combustion. The soot ring around the interior of the glass enclosure resulting from the burn-out of the fuel accumulated during the startup was still present. Large droplets were observed to reach the combustor wall, although their number decreased significantly with increasing fuel temperature. The flames exhibited stable operation for all test conditions Emissions Profiles Figure 3.7 shows the effect of ALR and fuel temperature on axial profiles of CO and NOx emissions. The axial distance (z) is measured from the atomizer exit plane, thus z = 45 cm 45

68 refers to the combustor exit plane. Figure 3.7 (a) shows the decreasing trend for CO emissions in the flow direction for all conditions. This result signifies that CO is produced in the flame initially during the fuel decomposition and is oxidized in the downstream zone. The CO emissions are highest for the fuel inlet temperature T f = 58 C and lowest for T f = 121 C for all conditions shown in Fig. 8. The CO concentrations decrease with increase in ALR (see Figs 3.7c and 3.7e) signifying that an increase in the atomizing air flow rate makes the premixed mode of combustion more prevalent. The corresponding NOx profiles are shown in Figs 3.7(b), 3.7(d), and 3.7(f). Figure 3.7(b) for ALR = 2.4 shows that the NOx concentration decreases with an increase in the fuel inlet temperature. Most of the NOx is formed within the reaction zone, i.e., z < 2 cm. Figure 3.7(d) and 3.7(e) show that increasing the ALR decreases the NOx emissions for all fuel temperatures. For a given ALR, an increase in temperature shows a decrease in NOx emissions in the range of 2-25 ppm while for a fixed fuel temperature an increase in ALR from 2.4 to 4. decreases the NOx emissions by a factor of up to 15. Figure 3.8 shows the radial emissions profiles taken at the combustor exit plane. Figure 3.8 (a) shows a decrease of about 1 ppm in CO concentration for each 2 C increase in the fuel temperature. Data for the lowest fuel temperature show slight asymmetry in the flame, attributed to larger turbulent fluctuations in the flame. The flow asymmetry reduces at higher ALR of 3.5 and ALR 4. (Fig. 3.8 (c) and (e)). At the highest ALR, the CO concentrations are small and tend to overlap each other for different fuel inlet temperatures. Overall, the CO emissions were in the range of 2 to 1 ppm for all fuel temperatures at all ALR conditions except for the worst condition of low fuel temperature and low atomizing air flow rate, i.e., T f = 58 C and ALR = 2.4. Figure 3.8 (b), (d) and (f) show the radial NOx profiles for ALR of 2.4, 3.5 and 4. respectively. 46

69 The NOx emissions decrease significantly with an increase in the fuel temperature. The NOx emissions also decrease with an increase in the atomizing air flow rate. Figure 3.9 shows the same results plotted for a fixed fuel temperature, but with varying ALR. Figure 3.9 (a) and (b) show the radial profiles CO and NOx concentrations for T f = 58 C. Results show a decreasing trend with an increase in ALR. Both CO and NOx emissions are nearly uniform at the combustor exit, except for ALR = 2.4, which results in greater asymmetry. NOx concentrations show a large effect of ALR for a given fuel temperature. At T f = 58 C, the NOx concentration decreased by a factor of 6 as ALR increased from 2.4 to Conclusions In this study, the effect of fuel temperature on NOx and CO emissions was studied using high viscosity vegetable oil (soybean oil) as the fuel of interest. Experiments were conducted in an atmospheric pressure, swirl stabilized burner. Results show that both CO and NOx concentrations decrease with an increase in the fuel temperature. For a given ALR, the CO emissions decreased by 1 to 2 ppm for each 2 C rise in the fuel temperature. The decrease in NOx emissions was much greater, especially at higher fuel temperatures whereby a 2 C rise in fuel temperature decreased NOx emissions by up to 2 ppm. For a given fuel inlet temperature, increasing ALR from 2. to 4. showed a decrease in CO emissions by a factor of 2 to 3 and a decrease in NOx emissions by factor of 6 to 15. The visual images were consistent with the emission measurements, but they also reveal large wetting of the combustor wall by unburnt droplets at all test conditions, although its extent decreased significantly with an increase in the fuel inlet temperature and/or ALR. 47

70 Figure 3.1. Schematic Diagram of the Experimental Setup. 48

71 Figure3. 2. Schematic Diagram of Swirler 49

72 VO 6 Diesel 5 5 Viscosity ( mm 2 /s) Temperature ( o C) Figure 3.3 Effect of Temperature on Kinematic Viscosity of Diesel and VO fuel. 5

73 Measured Fuel Inlet Temperature ( o C) o C 11 o C 13 o C o C ALR Figure 3.4. Effect of ALR on Fuel Inlet Temperature 51

74 (a) (b) (c) (d) (e) Figure 3.5 VO Flame Images at ALR 2.4 for different fuel temperatures (a) Unheated VO, (b) 58 o C, (c) 78 o C, (d) 99 o C, (e) 121 o C 52

75 (a) (b) (c) (d) (e) Figure 3.6. VO Flame Images at ALR 4. for different fuel temperatures (a) Unheated VO, (b) 58 o C, (c) 78 o C, (d) 99 o C, (e) 121 o C 53

76 (a T f =58 o C 6 78 o C 99 o C o C CO (ppm) ALR Axial Distance (cm) (b) NOx (ppm) 12 T f =58 o C o C 99 o C 121 o C ALR Axial Distance (cm) Figure 3.7. Axial Profiles for CO and NOx for different fuel temperatures at different ALR s [(a) and (b)] ALR

77 (c) T f =58 o C 6 78 o C 99 o C o C CO (ppm) ALR (d) Axial Distance (cm) T f =58 o C o C o C 121 o C NOx (ppm) ALR Axial Distance (cm) Figure 3.7. Axial Profiles for CO and NOx for different fuel temperatures at different ALR s [(c) and (d)] ALR

78 (e) Unheated VO T f =58 o C 6 78 o C 99 o C o C CO (ppm) ALR (f) Axial Distance (cm) NOx (ppm) 14 Unheated VO 14 T f =58 o C o C 99 o C o C ALR Axial Distance (cm) Figure 3.7. Axial Profiles for CO and NOx for different fuel temperatures at different ALR s [(e) and (f)] ALR 4. 56

79 (a) T f = 58 o C o C o C o C CO (ppm) ALR (b) Radial Distance (cm) NOx (ppm) 12 T f =58 o C o C 99 o C 121 o C ALR Radial Distance (cm) Figure 3.8. Radial Profiles for CO and NOx for different fuel temperatures at different ALR s [(a) and (b)] ALR

80 (c) CO (ppm) T f =58 o C o C o C o C 14 ALR Radial Distance (cm) (d) T f =58 o C o C o C ALR o C NOx (ppm) Radial Distance (cm) Figure 3.8. Radial Profiles for CO and NOx for different fuel temperatures at different ALR s [(a) and (b)] ALR

81 (e) CO (ppm) 18 Unheated VO T f = 58 o C 78 o C o C 14 ALR o C (f) Radial Distance (cm) NOx (ppm) Unheated VO 12 T f =58 o C o C 99 o C 121 o C 8 8 ALR Radial Distance (cm) Figure 3.8 Radial Profiles for CO and NOx for different fuel temperatures at different ALR s [(a) and (b)] ALR

82 (a) ALR ALR ALR CO (ppm) T f =58 o C (b) Radial Distance (cm) ALR ALR ALR NOX (ppm) T f =58 o C Radial Distance (cm) Figure 3.9. Radial Profiles for CO and NOx for ALR 2.4, ALR 3.5 and ALR 4. at different fuel inlet temperatures. [(a) and (b)] T f = 58 o C. 6

83 (c) ALR 2.4 ALR ALR CO (ppm) T f =78 o C (d) Radial Distance (cm) ALR ALR ALR NOx (ppm) T f =78 o C Radial Distance (cm) Figure 3.9. Radial Profiles for CO and NOx for ALR 2.4, ALR 3.5 and ALR 4. at different fuel inlet temperatures. [(c) and (d)] T f = 78 o C. 61

84 (e) ALR ALR ALR T f =99 o C 14 CO (ppm) (f) Radial Distance (cm) NOx (ppm) ALR ALR 3.5 ALR T f =99 o C Radial Distance (cm) Figure 3.9. Radial Profiles for CO and NOx for ALR 2.4, ALR 3.5 and ALR 4. at different fuel inlet temperatures. [(e) and (f)] T f = 99 o C. 62

85 (g) CO (ppm) 14 ALR T ALR 3.5 f =121 o C ALR (h Radial Distance (cm) NOx (ppm) T f = 121 o C ALR ALR 3.5 ALR Radial Distance (cm) Figure 3.9. Radial Profiles for CO and NOx for ALR 2.4, ALR 3.5 and ALR 4. at different fuel inlet temperatures. [(g) and (h)] T f = 121 o C 63

86 CHAPTER 4 CHARACTERISTICS OF PREHEATED NON-EVAPORATING BIO-OIL SPRAYS Background In chapter 2, we have investigated biodiesel fuels and diesel-vo blends as potential gas turbine fuels to generate power. The fuel properties were characterized by gas chromatographmass spectrometer, thermogravimetric analysis, and density, kinematic viscosity, surface tension and water content measurements. The combustion performance (in terms of carbon monoxide or CO and nitric oxides or NOx emissions) of different fuels was compared experimentally in an atmospheric pressure burner using an air-blast (AB) injector and swirling primary air around it [Panchasara, 29]. Results show that the atomization and fuel-air mixing processes have a major impact on combustion emissions for all fuels. For example, an increase in the atomizing airflow rate decreased CO emissions by a factor of up to 5 and NOx emissions by a factor of up to 1. Most importantly, the 7-3 diesel-vo fuel blend was equivalent to biodiesel fuels, in terms of NOx and CO emissions from the combustor. One of the key parameters affecting atomization is the kinematic viscosity of the fuel (Lefebvre, 1989; Lasheras & Hopfinger, 2; Faeth, et.al, 1995; Babinsky & Sojka, 22). Figure 4.1 shows that the kinematic viscosity of VO is higher than that of diesel by more than an order of magnitude which has been shown in chapter 2. The diesel kinematic viscosity at 25 C matches the viscosity of 9-1 diesel-vo blend at 4 C, 8-2 blend at 5 C and 7-3 blend 64

87 at 6 C. At 25 C, the biodiesel viscosity is nearly the same as that of the 8-2 diesel-vo blend. For a given temperature, the kinematic viscosity of 7-3 diesel-vo blend is similar to that of biodiesel and about twice that of the diesel. Thus, emissions from 7-2 diesel-vo blend were comparable to those from biodiesel (Panchasara, et.al. 29). An alternative to biodiesel and diesel-vo blends is direct combustion of source oils, which can result in significant economic and environmental benefits. It will however require preheating of the oil to reduce its kinematic viscosity so that the pressure drop in the fuel supply line is reasonable, atomization occurs with fine droplets, and diffusion mode of combustion is avoided to reduce pollutants of PM, NOx, CO, etc. In Chapter 3, we have shown that increasing VO inlet temperature decreased both CO and NOx concentrations in the flame (Faeth, 1987). For a given air to liquid mass ratio (ALR), the CO emissions decreased by 1 to 2 ppm for each 2 C rise in the fuel temperature. The decrease in NOx emissions was much greater, especially at higher fuel temperatures whereby a 2 C rise in fuel temperature decreased NOx emissions by up to 2 ppm. Spray characteristics such as drop diameter, drop size distribution, cone angle and penetration determine fuel-air mixing in the combustor, and hence, pollutant formation, life, durability, and efficiency of the gas turbine engine. Spray is produced by break-up of liquid fuel into small droplets through the atomization process. Small droplets produce larger surface area, thereby reducing the liquid vaporization time and improving fuel-air mixing. As stated by Lefebvre, efficient combustion requires optimal droplet size distribution within the spray, for a range of operating conditions to include droplets both large enough to penetrate into the combustion chamber and small enough to pre-vaporize within the short residence time in the flame region. 65

88 Past studies have presented the structure and break-up properties of sprays (Faeth,et.al.,1995;Babinsky & Sojka,22), different modeling techniques of drop size distributions (Simmons, et.al, 29) and processes of mixing, transport and combustion in sprays (Faeth,et.al., 1995). Figure 4.2 depicts the geometry, instabilities, wavelength and the break up process in a spray of liquid jet injected into a high velocity annular coaxial gas stream (Bari, et.al. 22). Small jet diameters and low Reynolds numbers demonstrate a spray with Rayleigh instability. At higher Reynolds numbers, the jet is subjected to waviness caused by aerodynamic effects, a regime called non-axisymmetric Rayleigh break-up. At even higher Reynolds numbers, droplets are stripped off by wind stress at the liquid/gas interface and atomization results from short wavelength shear instability (Lasheras & Hopfinger,2).Related research on preheated oils has focused mainly on injector performance and emissions characteristics of diesel engines operated on cottonseed and palm oils (Bari, et.al.,22;nwafor,23). These studies showed that the reduction in kinematic viscosity resulting from fuel preheating improves the combustion and emissions performance of the engine. The objective of the current study is to examine the spray characteristics of preheated soybean oil atomized by an AB injector. Such spray measurements of preheated VO are nonexistent at this time. The non-reacting spray is created at ambient conditions of temperature and pressure. First a laser sheet visualization system is used to obtain the qualitative spray images. Next, a two component Phase Doppler Particle Analyzer (PDPA) system is used to obtain detailed quantitative measurements of droplet diameter and gas phase velocities. Details of the experimental set-up, results and discussions, and conclusions are presented in the following sections. 66

89 4.2. Experimental set-up Figure 4.3 shows a schematic of the experimental set-up consisting of fuel (VO) and air supply systems, fuel heater, AB injector, and spray collection system. The VO from the reservoir is pumped by a high pressure peristaltic metering pump (Cole Parmer Model EW ) with an uncertainty of ±.25% of the reading. The pressure oscillations introduced by the pump are eliminated by a pulse dampener introduced in the flow path to achieve a steady flow. Next, an electric inline liquid heater (CRES-ILB-12/24) is used to preheat the VO upstream of the AB injector. The heater uses a Proportional Derivative Controller (PID) unit to control the VO temperature within accuracy of ±.5 C, measured at the heater outlet by a K-type thermocouple. Heated VO enters the injector through an insulated 5 cm long tube of 3 mm ID to minimize the heat loss. The VO temperature at the injector inlet is measured using a K-type thermocouple. Air for atomization is supplied by an air compressor and is measured using a mass flow meter with - liters per minute range. The spray collection system includes a funnel, filters, collector, and an exhaust fan. The VO in the spray is collected by the funnel and removed from the test area by the suction fan located downstream. The resulting flow passes through a filter element where the VO is recovered prior to exhausting out nearly oil-free air to the atmosphere. Figure 4 shows the detailed view of the commercially available AB injector (Delavan Model 369-2). In this injector, atomization occurs by the shear interaction between gas and liquid phases. The liquid supply tube of.3 mm inside diameter is surrounded by a swirling stream of atomizing air to break up the liquid jet into fine droplets. A laser sheet visualization system was used for qualitative visualization of the spray. The laser sheet was created by a 2 mw diode-pumped solid state laser producing 532 nm green 67

90 laser beam. A combination of cylindrical and spherical lenses from Edmund Optics was used to form a laser sheet at the mid-plane of the spray. Uncoated cylindrical lens of 12.5mm diameter x 12.5mm focal length and an uncoated Plano convex lens of 25 mm diameter x 75 mm focal length used created a laser sheet of about 1mm thick Spray images were taken by a digital camera with exposure times of 4µs and µs was mounted on a fixed tripod to ensure consistent field of view for different operating conditions. Quantitative drop size measurements were obtained using 2D Phase Doppler Particle Analyzer (PDPA). PDPA is a point sampling device based on the light scattering interferometery. The laser beams from the transmitter probe intersect to form a sample measurement volume. Principally when a particle or drop passes through the beam intersection region, the scattered light forms a fringe pattern. Since the droplet is moving, the fringe pattern sweeps past the receiver aperture at the Doppler difference frequency, which is directly proportional to the drop velocity. The spatial frequency of the fringe pattern is inversely proportional to the drop diameter. The phase shift between the Doppler burst signals from different detectors is proportional to the diameter of the spherical shaped droplet. There is no calibration required for the PDPA method since the drop size and the droplet velocity are dependant only on laser wavelength and optical configuration. The schematic of the 2D PDPA system is as shown in Figure 4.5. The laser beam from a 2-W water cooled argon-ion laser is separated into a pair of nm green beams and a pair of 488 nm blue beams using a beam separator assembly. One beam of each pair is shifted by a 4 MHz Bragg cell. Next, each beam is focused onto a fiber optic cable to deliver the beams to a 25 mm focal length PDPA transmitter. The PDPA receiver is set at an angle of 135 degrees from the transmitter to collect the refracted light intensities from the spray. The detected signal is 68

91 acquired by a data acquisition system, and analyzed using the TSI Flow Sizer software to obtain mean and root mean square (RMS) velocities, Sauter mean diameter (SMD), and drop size distribution data. Experiments were conducted for fixed VO flow rate of 12 liters per minute (lpm). The atomizing air flow rate through the injector was varied to obtain ALR of 2. and 4.. Measurements were obtained by moving the AB injector using a three-way traversing system, while the PDPA system was stationary. The injector was traversed to acquire radial profiles at axial planes between Y = 1 mm and mm, in.5 mm intervals. The center of the spray pertained to the peak location of the axial velocity profile. Data rates of up to 55 khz were obtained towards the centre of the spray. The data rate and mean axial velocity both decreased to nearly zero as the detection volume reached the outer edge of the spray Results and Discussion Spray Images The qualitative spray images for VO at 4 C, 7 C and C are shown in Figure 4.6 for ALR = 2. and in Figure 7 for ALR = 4.. For each case, images were obtained for µs (top) and 4µs (bottom) exposure times, to capture steady and transient features of the spray. Figure 4.6 shows a decreasing trend of spray angle with increasing VO temperature. This result can be explained by an increasing number of larger droplets reaching farther away from the center of the spray for lower VO temperatures. Because of the high momentum, large droplets also tend to move farther downstream in the spray. The bottom images at a smaller exposure time reveal similar characteristics, but they also reveal the waviness at the edge of the spray. The observations of Figure 4.6 are replicated in Figure 7 for ALR = 4.. An increase in ALR has 69

92 similar effect as an increase in VO inlet temperature, i.e., the spray becomes narrower, shorter, and denser with an increase in the VO inlet temperature and/or ALR SMD Contour Plots Point-wise measurements acquired at different axial and radial locations were used to construct contour plots for different operating conditions. Figure 4.8 shows the contour plots of SMD for T = 4 C, 7 C and C at ALR 2., and T = 7 C at ALR = 4.. For the lowest VO inlet temperature of 4 C, the contour plot in Figure 4.8 (a) can be divided into three regions: the smallest droplets of 5 to 15 µm reside in the center core region, mid-size droplets up to 4 µm are found in the shear layer region, both near and farther away from the injector exit. The largest droplets with diameter exceeding 4 µm are observed at the outer edge of the spray, both in the near and far downstream locations. Evidently, the high momentum of large droplets is responsible for carrying them away from the injector exit. At this low VO inlet temperature, the maximum SMD is about 6 µm, and the SMD range is correspondingly wider. Figure 4.8 (b) shows the contour plot of SMD for VO at T = 7 C. Results show similar trends; smallest droplets in the core region and largest droplets at the edge of the spray. The maximum SMD value for this case is however around 4 µm, which is significantly lower than that of 6 µm for T = 4C. Figure 4.8 (c) shows the same trend in the SMD for VO at T = C. The maximum SMD for this case is still around 42 µm. Results show that an increase in VO temperature improves atomization by decreasing the SMD, although benefits diminish at higher VO temperatures. The effect of ALR on SMD is observed in Figure 4.8(d) showing SMD contour plot for T = 7 C and ALR = 4.. For this high ALR, the maximum SMD is around 34 µm and the spray cone angle is smaller than that for lower ALR. Results show that increases in both the VO temperature and ALR improve atomization. Since increasing VO temperature 7

93 requires thermal energy input and increasing ALR increases the pressure drop penalty, a proper balance between the two parameters is important to achieve the optimum injector performance. Pressure drop measurements for the AB injector used in this study are given by Simmons et al Axial Velocity Contour Plots Figure 4.9 shows the contour plots of axial velocity at aforementioned VO inlet temperatures and ALR values of 2. and 4.. Figure 4.9 (a) for T = 4 C and ALR = 2. shows that the axial velocity peaks in the center of the spray, with velocity magnitude diminishing towards the edge of the spray. Moreover, the axial velocity decreases in the axial direction indicating the loss of momentum (or kinetic energy) of the droplets. For ALR = 2., contour plots in Figure 4.9(a)-(c) show that the peak axial velocity increases with increasing VO inlet temperature, evidently because of the heating of the atomizing air by VO. For T = 7 C, an increase in ALR from 2. to 4. changes the peak axial velocity near the injector exit from 4 m/s to about 95 m/s. At the higher ALR, the radial extent of the flow field decreases indicating a narrow spray. Figure 4.1 shows the contour plots of RMS axial velocity for conditions pertaining to those in Figures 4.8 and 4.9. Figures 4.1(a)-(c) show that an increase in VO temperature increases the RMS axial velocity in the near field. For example, the maximum RMS axial velocity is about 7 m/s, 14 m/s and 15 m/s, respectively, for VO inlet temperature of 4 C, 7 C, and C. Clearly, the higher axial velocity resulting from heating of the atomizing air by the VO is also producing higher turbulence labels, which would tend to reduce the droplet diameter as observed in Figure 4.8. These results show that reduction in kinematic viscosity associated with heated VO, increase in mean axial velocity of the atomizing air because of the heat transfer 71

94 from the heated VO, and increase in turbulent fluctuations, possibly due to high mean velocity, are complimentary effects in reducing the SMD of heated VO. Figure 4.1 (d) shows the contour of RMS axial velocity for T = 7 C and ALR = 4.. In this case, the RMS axial velocity is significantly higher, with the peak value reaching up to 32 m/s. Again, this increase in RMS axial velocity is a major contributing factor in reducing the SMD at the higher ALR as shown in Figure 4.8(d) Radial Profiles of SMD Figure 4.11(a) shows that the VO inlet temperature and/or ALR have a minor effect on SMD in the center region of the spray in the near field, i.e., Y = 2 mm. For all cases, the minimum SMD at the center of the spray is nearly the same, i.e., 8 µm. Away from the center, the radial profile is steeper at the lower temperature, which signifies a spray with a wider range of droplet diameters. The VO inlet temperature and ALR have the greatest effect in the outer region of the spray. The effects of VO temperature and ALR are more obvious at downstream locations of z = 5 and 8 mm, as shown in Figures 4.11(b and c). Increase in ALR has the most dramatic effect; the radial profile for ALR = 4. is much flatter compared to that for ALR = 2.. Increase in VO inlet temperature results in smaller droplets in the outer regions of the spray. Thus, the minimum SMD in the spray is independent of the VO temperature or ALR. However, increasing VO temperature or ALR decreases the radial spread of the spray as well as the SMD range. Results in Figure 4.11 are rearranged in Figure 4.12 to illustrate how VO inlet temperature, and hence, fuel kinematic viscosity affects SMD at different locations in the spray. Results are plotted in terms of SMD versus kinematic viscosity for the three test cases with ALR = 2.. Figure 4.12(a) shows that near the injector exit, Y = 15 mm, for all radial locations, the 72

95 SMD increases with an increase in the fuel kinematic viscosity. However, the SMD increase is minor near the center (r = and 5 mm) and much greater away from the center (r = 8 and 12 mm). Similar results are obtained at Y = 2 mm and Y = 3 mm as shown in Figures 4.12(b and c) Droplet Diameter Distributions Figure 4.13 (a, b and c) show the droplet diameter distributions in the spray for VO at 4 C, 7 C and C at ALR = 2. and Figure 4.13 (d) shows the distribution profile for VO at 7 C and ALR = 4.. The distribution profiles are taken at the edge of the spray at Y = 2 mm, r = 1 mm and Y = 8 mm, r = 25 Figure 4.13(b and c) show the same trend of droplet diameter distribution profiles for VO at 7 C and C and ALR = 2.. For a given ALR and given axial location in the spray, higher VO inlet temperature shows a lower range of droplet diameters with a narrower distribution profile. At Y = 2 mm, the largest droplet diameter is about 12 µm for VO at 7 C whereas for VO at C the largest droplet diameter is about 8 µm (see Figures 4.13 b-c). At Y = 8 mm, the largest droplet diameter is about 225 µm for VO at 7 C whereas for VO at C the largest droplet diameter is about 13 µm. The larger percentage of smaller droplets at higher VO inlet temperatures result in the decrease in the SMD as observed in Figure 4.8. Figure 13(d) shows the droplet distribution profile for VO at 7 C and ALR 4.. Profiles show a higher percentage of larger droplets in the near field, Y = 2 mm, compared to those Y = 8mm. As shown in Figure 4.13 (a), for VO at 4 C, the droplet diameter varies widely at Y = 2mm and Y = 8 mm. The count of larger droplets increases at the downstream location of Y = 8 mm. For example, the largest droplet diameter in the near field (Y = 2 mm) is about 15 µm whereas at Y = 8 mm, the largest drop diameter has increased to about 25 µm. The wider drop 73

96 diameter distribution in the far field is attributed to the larger droplets of high momentum migrating towards the periphery of the spray. Figure 13(b and c) show the same trend of droplet distribution profiles for VO at 7 C and C for ALR = 2.. For a given ALR and given axial location in the spray, higher VO inlet temperature results in a lower range of droplet diameters with a narrower distribution profile. At Y = 2 mm, the largest diameter is about 12 µm for VO at 7 C whereas for VO at C the largest droplet diameter is of about 8 µm (see Figures 4.13 b-c). =At Y = 8 mm, the largest droplet diameter is about 225 µm for VO at 7 C whereas for VO at C the largest droplet diameter is about 13 µm. The larger percentage of smaller droplets at higher VO inlet temperatures result in the decrease in the SMD as observed in Figure 8. Figure 4.13(d) shows the droplet distribution profile for VO at 7 C and ALR = 4.. Profiles show a higher percentage of larger droplets in the near field, Y = 2 mm, compared to those at Y = 8 mm. The largest drop diameter for this case is about 18 µm and 8 µm, respectively, at Y = 2 mm and 8 mm Conclusions In this chapter, measurements of SMD, axial velocity, axial RMS velocity and droplet distribution are presented to explain the effect of operating conditions in a non-evaporating spray of VO created by an AB injector. Laser sheet visualization system was used to acquire the spray images. For ALR = 2.5 and VO inlet temperature of 4 to C, the SMD of droplets in the spray ranged from 35 to 6 µm. The SMD decreased down to 3 µm for a higher ALR of 4.. The SMD of the VO spray decreased as the VO inlet temperature increased, indicating improved atomization. Because of the high momentum, the larger droplets migrated towards the edge of the spray 74

97 The number of larger droplets increased for the lower VO temperature, which led to an increase in the SMD. The spray angle decreased with an increase in VO inlet temperatures and ALR, thereby, improving the spray quality in terms of decreased droplet diameter. The mean axial velocity was the highest at the center of the spray and it decreased towards the edge of the spray and at downstream locations. The RMS axial velocity increased with an increase in VO inlet temperature and/or ALR. Experiments presented in the next chapter will focus on drop size measurements in VO spray flames to delineate the effects of heat release rate on fuel vaporization. The spray characteristics are expected to be much different because of the thermal feedback from the flame. 75

98 Diesel Vegetable Oil 6 Kinematic Viscosity mm 2 /s Temperature o C Figure 4.1 Kinematic Viscosity of Diesel and Vegetable Oil. 76

99 Figure 4.2 Schematic representation of the liquid break-up indicating geometry and different lengths. 77

100 X Y Figure 4.3 Schematic diagram of the experimental setup 78

101 Orifice disc Swirler vanes Fuel inlet Fuel supply Tube Figure 4.4 Air-Blast Injector Details 79

102 X Y Figure 4.5 Schematic of PDPA system 8

103 (a) (b) Figure 4.6 Laser sheet spray images for VO at 4 C, 7 C and C at ALR 2.. (a) Spray Images at Exposure time ms. (b) Spray Images at Exposure time 4 ms. 81

104 (a) (b) Figure 4.7 Laser sheet spray images for VO at 4 C, 7 C and C at ALR 4.. (a) Spray Images at Exposure time ms. (b) Spray Images at Exposure time 4 ms. 82

105 Axial location (mm) SMD (microns) T=4 o C, ALR = Radial location (mm) Axial locaiton (mm) T=7 o C, ALR = SMD (microns) Radial location (mm) Figure 4.8. Contours of Sauter Mean Diamter for (a) T = 4 C, ALR = 2., (b) T = 7 C, ALR

106 Axial location (mm) T=C, ALR = 2. SMD (microns) Radial location (mm) Axial locaiton (mm) T=7 o C, ALR = SMD (microns) Radial location (mm) Figure 4.8. Contours of Sauter Mean Diamter for (c) T = C, ALR = 2., (d) T = 7 C, ALR

107 (a) Axial Location (mm) T=4C, ALR = 2. Axial Velocity (m/s) Axial Location (mm) Radial Location (mm) (b) T=7C, 9 9 ALR = 2. 8 Axial Velocity (m/s) Radial Location (mm) Figure 4.9. Contours of Axial Velocity for (a) T = 4 C, ALR = 2., (b) T = 7 C, ALR

108 Axial Location (mm) Axial Location (mm) (c) Radial Location (mm) (d) T = C, ALR = 2. Axial Velocity (m/s) Radial Location (mm) Figure 4.9. Contours of Axial Velocity for (c) T = C, ALR = 2. and (d) T = 7 C, ALR = T=7C, 9. 9 ALR = Axial Velocity (m/s)

109 Axial Locaiton (mm) (a) Radial Locaiton (mm) T=4C, ALR = 2. Axial RMS Velocity Axial Location (mm) (b) Axial RMS Velocity T=7C, ALR = 2. Radial Location (mm) Figure 4.1.Contours of Axial RMS Velocity for (a) T = 4 C, ALR = 2., (b) T = 7 C, ALR

110 Axial Location (mm) Axial Location (mm) (c) T=C, ALR = Axial RMS Velocity Radial Location (mm) (d) T=7C, ALR = Axial RMS Velocity Radial Location (mm) Figure 4.1.Contours of Axial RMS Velocity for (c) T = C, ALR = 2. and (d) T = 7 C, ALR = 4. 88

111 Y=2mm 6 T=4C 6 T=7 T=C 5 T=7C,ALR=4. 5 SMD (microns) Radial Distance (mm) Figure 4.11.Profiles of Sauter Mean Diameter for (a) Y =2 mm 89

112 Y=5mm 6 T=4C 6 T=7C T=C 5 T = 7 C, ALR=4. 5 SMD (microns) Radial Distance (mm) Figure 4.11.Profiles of Sauter Mean Diameter for (b) Y =5 mm 9

113 SMD (microns) Y=8mm 6 T=4C 6 T=7C 5 T=C T = 7 C_ALR_ Radial Distance (mm) Figure 4.11.Profiles of Sauter Mean Diameter for (c) Y =8 mm 91

114 Y=15mm 3 3 SMD (microns) 25 r=mm 25 r=5mm 2 r=8mm 2 r=12 mm Viscosity (mm 2 /s) Figure Profiles of Sauter Mean Diameter versus Fuel Viscosity for (a) Y =15 mm. 92

115 Y= 2 mm 3 3 SMD (microns) r=mm r=5mm 2 r=8mm 15 r=12mm Viscosity (mm 2 /s) Figure 4.12.Profiles of Sauter Mean Diameter versus Fuel Viscosity for (b) Y = 2 mm 93

116 SMD (microns) Y=3mm 35 r=mm 35 r=5mm 3 r=8mm r=12mm Viscosity (mm 2 /s) Figure Profiles of Sauter Mean Diameter versus Fuel Viscosity for (c) Y = 3 mm 94

117 T=4C, ALR = Y= 2mm, r= 1mm Y = 8mm, r = 25mm Diameter Count Diameter (microns) T=C, ALR = Y = 2 mm, r =1 mm Y = 8 mm, r =25 mm Diameter Count Diameter (microns) 2 25 Figure 4.13.Profiles of Diameter Distribution for (a) T = 4 C, ALR = 2., (b) T = C, ALR 2. 95

118 T=7C, ALR = Y =2 mm, r=1 mm Y =8 mm, r=25 mm Diameter Count Diameter(microns) Diameter Count T=7C, ALR = Y = 2mm, r = 1 mm Y = 8mm, r = 25 mm Diameter (microns) 2 25 Figure 4.13.Profiles of Diameter Distribution for (c) T = 7 C, ALR = 2., (b) T = 7 C, ALR 4. 96

119 CHAPTER 5 CHARACTERISTICS OF PREHEATED BIO-OIL SPRAYS Background Global atmosphere pollution has become a serious problem for today. The emissions from the combustion of fossil fuels contribute a notable part to this pollution. Environmental care together with the limited stock and growing prices of fossil fuels has given alternative fuels the potential to supplant a significant portion of fuel for combustion applications such as gas turbine engines and IC engines. Given a wide spread of different biofuels available for combustion applications, the present study concentrates on atomization spray characteristics of vegetable oils. Vegetable oils have energy density, cetane number, heat of vaporization and stoichiometric air/fuel ratio comparable to diesel fuel. Different techniques have been employed so far to improve on the physical properties of bio-oils and thus opening the doors to clean combustion. The high kinematic viscosity has an adverse effect on the combustion of vegetable oils, posing problems in the associated fuel supply line and injector system, as discussed in chapters 2 and 3. Some well-known techniques to deal with high kinematic viscosity levels of neat vegetable oils include dilution, pyrolysis, micro-emulsion and trans-esterification. These techniques however, require additional energy input to improve the physical properties of the fuel. Preheating of the fuel is also one of the ways that reduces the kinematic viscosity to improve the atomization. Preheating is employed in the present study to investigate the spray characteristics in a non evaporating spray as well as flame spray. 97

120 The atomization and subsequent propagation of the fuel droplets, their vaporization and combustion are the most important processes concerning the formation of pollutants with the use of liquid fuels. For example in diesel engines, gas turbine engines, and oil burners, the combustion rate of fuel is controlled by effective vaporization of the fuel. The liquid fuel atomization rate has a strong influence on vaporization rates because the total surface area of the fuel is increased greatly by the atomization process. The fundamental mechanisms of atomization have been under extensive experimental and theoretical study for more than a century (Liu & Reitz, 1993). Still, one of the major thrusts in worldwide combustion research has been to gain insight into the physics of liquid fuel combustion in the primary zone of the combustor (Kneer, et.al.1993). Traditionally, the fundamental basis for the atomization process relies on either injecting the fuel under relatively high pressure into a relatively slow moving gas or subjecting the fuel to a high velocity air blast or a combination of two atomization mechanisms. The fuel injector assumes an important role in the combustor by providing some degree of fuel/air mixing close to the atomizer. Almost all combustor performance characteristics are strongly governed by the quality of the spray produced by the fuel atomizer. Calculation of the evaporation and reaction rates in the combustor involves the evaluation of parameters such as mean spray angle, range of drop sizes in the spray, and drop trajectory. It is a well known fact that the achievement of efficient fuel atomization and rapid evaporation in addition to mixing with combustion air can have a significant effect on such parameters as emissions, exhaust gas temperature profiles, and pattern factor. There are numerous studies on liquid breakup in the literature. As the relative velocity between the drop and gas increases, the drop break up regimes such as bag break up, shear 98

121 breakup and catastrophic break up are encountered. The jet break up as summarized by Liu & Reitz (1993) comprises of four liquid break up regimes based on breakup drop size and unbroken jet length. Rayleigh breakup regime; first-wind induced break up regime; second break up regime and the atomization break up regimes which are encountered as the jet velocities increases. At lower jet velocities, the growth of the small disturbances on the liquid surface due to the interaction with the surrounding gas is believed to be the dominant reason for the liquid break up. As the jet velocity is increased and the aerodynamic effect becomes more prominent, the breakup mechanisms become increasingly complex. Air blast atomization is an attractive strategy for liquid fuel breakup in gas turbines. Principally, the air blast atomizer functions by employing the kinetic energy of a flowing airstream to break the fuel jet or sheet into ligaments and then drops. Air blast atomizers offer great advantages over pressure atomizers, since a finer spray can be produced with a lower fuel pumping pressure. Air blast atomization maximizes the interaction between the air and liquid flows by, taking advantage of high velocity airflow to produce fine drops in a well distributed spray. The atomizing air also serves as emissions reduction strategy improving the fuel air mixing in the combustor, to reduce soot particulates, CO and NOx emissions. The performance of a given spray combustion system depends not only on the fuel droplet size distribution but also on the spray spatial distribution and the interaction of the droplets with the gas turbulence that involves a physical mechanism that is yet to be well understood. For this reason spatially and temporally resolved information such as mean droplet size, droplet size distribution, mean drop velocity rms velocities needs to be studied to understand most favorable spray conditions for optimal combustion performance. 99

122 Physical properties such as kinematic viscosity, surface tension and volatility are the key parameters that affect the process of fuel atomization and evaporation. The liquid kinematic viscosity affects not only the drop size distribution of the spray but also the fuel injector pressure drop. An increase in kinematic viscosity lowers the Reynolds number, hindering the development of any natural instabilities in the fuel jet or sheet, which help to further disintegrate the drops. These combined effects delay any further disintegration thus increasing the droplet sizes in the spray. Many alternative fuels are expected to have high kinematic viscosity which makes them difficult to atomize well and thus affecting the combustion efficiency. A comprehensive study on turbulent diffusion flames using intrusive probing techniques was made by Onuma and Ogasawara, (1975). They suggested that the spray flame structure is similar to that of a gaseous diffusion flame in turbulent flow. Chigier,(1974) used a non intrusive detection technique to measure the Sauter Mean diameter (SMD), drop velocity, and number density of air assisted spray and spray flames. A series of experimental and numerical studies of air assisted sprays and spray flames have been made by McDonell, & Samuelsen, (1991). Their observations concluded that the presence of fuel drops and reactions alters the structure of the gas-phase turbulence and that local clustering of drops exists for both non-reacting and reacting cases. A large portion of the experimental research in liquid fuel combustion is focused on pressure atomization mainly in the diesel engine applications. Relatively few studies have been reported on air blast atomization and their potential optimum strategies in alternative fuel combustion. Moreover, very little attention has been given to the evaporation characteristics of the air blast atomized sprays of alternative fuels. Detailed studies on the characteristics of spray flames are necessary to mitigate environment problems and enhance the performance and efficiency of liquid bio fuel combustion systems.

123 The present work seeks to experimentally investigate the spray characteristics of the fuel droplets in a non-evaporating as well as flame spray conditions using the Phase Doppler Particle Analyzer. An air blast atomizer is selected for the present investigation to generate the spray. The photographic image and the key structural features of a flame from a typical air blast atomizer are shown in figure 5.1. Larger drops are distributed mostly on the outer edge of the spray in region C and the smaller drops are located in the centre region A. These larger drops on the edge are affected by the presence of the flame. Due to slower evaporation rate of the large drops, a blue lean reaction sheath is formed inside the spray boundary as seen in the photographic image. The main reaction zone is a mixture of fuel vapor and air that is sufficient for combustion. In the central core of the spray, as seen in region A the fuel vapor concentration is too rich to allow the chemical reactions to take place. Above the spray flame, the smaller droplets produce tiny blue streak flames as well as the larger drops that burn incompletely produces a yellow orange streak flame flying in all directions (regions B and C ). Mostly on the outer periphery of the spray flame, single drop burning with self sustained envelope flame is observed. The swirling air flow improves mixing and creates a homogenous combustible mixture and hence stabilizes the liquid fuel flame. This is shown in figure 5.1 by the region D which is a methane air mixture flame. The swirling combustion air makes the liquid sheet unstable and hence helps any further disintegration in to smaller droplets that follow the flow. Figure 5.2 shows the droplet vaporization in spray flames. The spray consists of initial cool and hot zones. Within the cool zone, heat transfer is restricted to radiation from the flame front to the droplet surface. In the hot zone, heat transfer takes place both by radiation from the flame front and by turbulent convection. The droplet diameter reduces by the d 2 law: -d(d 2 )/dt = λ 11

124 Where λ is the evaporation constant for forced convection. The objective of the present work is to investigate the effects of combustion on spray and spray flames of bio-oil for which little experimental data are available. Experiments are conducted using unheated VO as well as heated VO at C. Measurements are obtained for both in open flame and an enclosed flame to simulate realistic gas turbine conditions. The mean axial and RMS velocities, SMD and drop size distribution data are acquired. The primary focus is placed on liquid fuel spray characteristics, and their effects on emissions. It can be envisioned that smaller droplets in the spray would lead to premixed combustion and hence lower emissions. The inferences from this study would aid in designing future liquid fuel combustors. Hereafter the details of the experimental set-up, results and discussion, and conclusions are presented in this chapter Experimental Set-up and Procedure The experimental set-up is shown schematically in Fig It consists of the combustion, fuel injection, and flare systems as discussed below. Phase Doppler Particle Analyzer (PDPA) mounted on a 3-way traversing system was used to acquire quantitative spray measurements in the cold flow as well as flame. In addition, emissions data were acquired by a continuous sampling probe and an infrared camera was used to record the combustor wall temperature Combustion System Combustion system was housed within a test chamber of dimensions 63.5 cm by 63.5 cm by m. The primary air enters the combustion 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 reactants or combustion-air 12

125 enter the combustor through a swirler section shown schematically in Fig. 5.3.The swirler is used to enhance fuel-air mixing and it also helps to stabilize the flame. The swirler has six vanes positioned at 28 to the horizontal to produce swirl number of about 1.5. The bulk axial inlet velocity of the primary air is 1.9 to 2.1 m/s, which resulted in Reynolds number varying from 596 to 675. The combustor is enclosed within a 15 cm inside diameter, 46 cm long pentagonal enclosure, with two sides of quartz glass and remaining three sides of metal plates. Table 5.1 Properties and Characteristics of the Insulating Material [ Nominal Composition, Wt% Buster Blanket Al 2 O 3 97 SiO 2 3 Organics <3 Trace Elements <.5 Melting Point, ºC (ºF) 238 (37) Maximum Use Temperature, ºC (ºF) 16 (2912) W/mK (BTU/hr ft 2 o F/inch) at 16 o C (2912 o F).476 (3.3) Bulk Density, g/cc <4 In order to have optical access and to be able to measure the spray data, a flat sided pentagonal enclosure was designed and fabricated in house. The photographic view of the enclosure with out and with insulation and the schematic of the top view of the pentagonal enclosure are shown in figure 5.4. Transparent quartz glass windows of 8. cm by 46. cm by 3mm thick, grounded and polished to optical quality were used for the enclosure to allow optical 13

126 access through the enclosure. The enclosure was insulated with 4 layers of insulation to minimize the heat loss from the flame. Alumina Mat type Buster Blanket insulation (1 standard thickness) from Zircar with properties listed in Table 5.1 was used to insulate the enclosure. Buster blanket is a flexible, Hi-Alpha fiber insulation material. It is needled into a durable blanket with the addition of organic fiber reinforcement with temperature up to 16oC. These polycrystalline fiber blankets offer higher temperature capability, less shrinkage and greater chemical resistance than standard alumina-silica blankets. The insulation helped retaining heat inside the combustion area and hence improves flame stability. For optical access a small window of 1.16 cm by 1.16 cm was cut on the insulation on two glass windows at the bottom of the enclosure as observed in the photographic image of figure 5.4. Air for combustion and atomization was supplied by an air compressor. The air passed through a pressure regulator, and a dehumidifier and water traps to remove the moisture. Then, the air was split into combustion air supply and atomizing air supply lines. The combustion air flow rate was measured by a laminar flow element (LFE) with reported calibration error of ±5 lpm. The pressure drop across the LFE was measured by a differential pressure transducer. An absolute pressure transducer was used to measure air pressure in the LFE. The atomizing airflow rate was measured by calibrated mass flow controller from Sierra (Model 81S-M., Mass-Track Flow Controller, 15- SLPM), with measurement uncertainty of ±.5 lpm Fuel Injection System The fuel injector system runs through the plenum and the mixing chamber, as shown in figure 5.3. An O-ring within a sleeve is located at the bottom of the plenum to prevent air leakage. A commercial air-blast atomizer (Delavan Siphon type SNA.2 nozzle) with its details shown schematically and photographically in Fig. 5.5 was used for the experiments. The 14

127 injector creates a swirling flow of atomizing air to breakdown the fuel jet as the fluids exit the injector. The inside diameter of the fuel supply tube is about.4 mm. The liquid fuel supplied by a peristaltic pump passes through a pulse dampener, a fuel filter and an electric fuel heater (Infinity Fluids Corporations, CRES-ILB-12/24 inline water/liquid heater) as shown in figure 5.6. The heater uses a Proportional Derivative Controller (PID) control unit to control the fuel temperature within accuracy of ±.5 o C, measured at the heater outlet by a K-type thermocouple. Heated fuel enters fuel injector through 5 cm long tube of 5 mm ID to minimize the heat loss. The fuel temperature at the injector inlet was measured by a K- Type thermocouple Flare system To burn the fuel droplets before they enter the exhaust system was a major challenge while conducting cold flow measurements. Thus a flare system that could successfully combust the liquid fuel upstream of the exhaust duct was designed in house. Several different approaches were attempted to create a flare system to burn most of the fuel droplets upstream of the exhaust duct while also avoiding bellowing of the spray in the operating area. Figure 5.8 shows the photographic view of the final flare system used with open cold spray. Spray measurements in the cold flow were acquired with and without swirling flow of combustion air. The flare system consists of 3 diffusion torches with co flowing methane and air streams as shown in figure 5.8. The flare system was located at about 6.96 cm downstream of the open cold spray. The methane flow rate through the flare system was kept to about 4 SLPM while the air flow rate was about 3 SLPM. As seen from the photographic image, the open cold spray is encompassed in the long flame from the flare creating a combustible mixture with most of the fuel droplets burning upstream of the exhaust duct as shown in figure 5.8. Sufficient air was entrained to burn the fuel 15

128 droplets using the co-flow air. The flares ensured that the fuel spray mixture passed through a sufficiently long flame so that enough residence time and high flame temperature would ensure proper droplet burning of the spray upstream of the exhaust duct Emissions Measurement System The product gas was sampled continuously by a quartz probe (OD = 7. 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 probe was traversed in the axial direction at the center of the combustor and in the radial direction at the combustor exit plane. The analyzer consist of a built-in sample pump which draws the flue gases through the stainless steel probe, pre-cooler, condensate trap and filter, flow meter, internal secondary filter, Teflon liquid blocker and the sample collecting sensors as the basic built in functions. Water traps mounted upstream of the emission analyzer prevent the sensors from the contamination. The dry sample was sent through the electrochemical analyzers to measure the concentrations of CO and NOx in ppm. As shown in figure 5.9, a NOVA gas analyzer was used to measure the concentrations of CO and NOx in the exhaust gas sample. The emission concentrations are measured using electrochemical sensors. Electrochemical fuel cell type sensors produce a small electrical output proportional to the gas being detected. The output signal is amplified and displayed on LCD digital panel meters. The emission analyzer measures the concentrations of CO in the range of to 2 ppm and NOx in the range of to ppm. The analyzer also measured oxygen and carbon dioxide concentrations, which were used to cross-check the equivalence ratio computed from the measured fuel and air flow rates. The uncorrected emissions data on dry basis are reported with uncertainty of +/-2 ppm. 16

129 5.2.5 Phase Doppler Particle Analyzer (PDPA) Set-up A 2D Phase Doppler Particle Analyzer (PDPA) was used for flow velocity and drop size measurements. PDPA is a point sampling device based on the light scattering interferometery principle. The laser beams from the transmitter probe intersect to form a sample measurement volume. Principally when a particle or drop passes through the beam intersection region, the scattered light forms a fringe pattern. Since the droplet is moving, the fringe pattern sweeps past the receiver aperture at the Doppler difference frequency, which is directly proportional to the drop velocity. The droplet size is measured by the phase shift of the light encoded in the spatial variation of fringes reaching two detectors after travelling paths of different lengths through the droplets. The phase shift is measured by the two detectors, each focused at a spatially distinct portion of the receiver lens. The spatial frequency of the fringe pattern is inversely proportional to the drop diameter. The phase shift between the Doppler burst signals from different detectors is proportional to the diameter of the spherical shaped droplet. There is no calibration required for the PDPA method since the drop size and the droplet velocity are dependant only on laser wavelength and optical configuration. The schematic of the 2D PDPA system is shown in Figure 5.1. The laser beam from a 2- W water cooled argon-ion laser is separated into a pair of nm green beams and a pair of 488 nm blue beams using a beam separator assembly. One beam of each pair is shifted by a 4 MHz Bragg cell. Next, each beam is focused onto a fiber optic cable to deliver the beams to a 75 mm focal length PDPA transmitter. In order to be able to take flame spray measurements using PDPA, the lenses were replaced to the focal length of 75 mm (29 inches) for both transmitter and receiver probes so that the optical probes are kept away from the flame. Precisions Achromatic Doublet lenses from Newport Optics were used. The transmitter lens, of 17

130 5.8 mm diameter, 75 mm focal length (Model No. PAC86) and the receiver lens of 76.2 mm diameter, 75 mm focal length (Model No. PAC46) are mounted on the respective probes. The PDPA receiver is set at an angle of 144 degrees from the transmitter to collect the refracted light intensities from the spray. The detected signal is acquired by a data acquisition system, and analyzed using the TSI Flow Sizer software to obtain mean and root mean square (RMS) velocities, Sauter mean diameter (SMD), and drop size distribution data. The photographic view of the PDPA system integrated with the combustion system is shown in figure Traversing System Measurements were obtained by moving the PDPA system using a three-way traversing system, while the combustor was kept stationary. The PDPA system was traversed to acquire radial profiles at axial planes between Y = 5 mm and 75 mm, in 5mm intervals for the cold spray as well as from Y = 5 mm and 4 mm, in 5 mm intervals for the flame spray measurements. The 3-way traversing system (Model No. MT1--M2-31), from Velmex Inc., was used to mount the PDPA system. The schematic of the system is shown in figure The plan view of the PDPA traversing mechanism in radial and axial direction is shown in figure Each (x, y and z) axis was traversed by the single shaft stepper motors Model NO. PK296-3AA-A6-3/8. The Y-axis was traversed to measure the radial velocity and the Z-axis measured the axial velocity. The traversing system was bolt mounted on a carbon steel table of 8. cm by 8. cm and 82 cm high. The transmitter and the receiver probes were mounted on the rail attached to the Z axis of the traversing system. To obtain a good signal quality, the angle between the transmitter and receiver was maintained at 144. To obtain this angle, both the probes were set at 18 with respective to the optical rail attached to the Z axis of the traversing system carrying the transmitter and receiver. To simplify the measurements in radial directions, we aligned the 18

131 transmitter direction with respect to one of the traversing system axis, i.e. the X axis. This demands the optical rail to be set at precisely the same angle with respect to X axis. This arrangement of setting the optical rail at 18 facilitates to take radial location measurements along Y axis. The center of the spray pertained to the peak location of the axial velocity profile. Data rates of up to 3-4 khz were obtained towards the centre of the spray. The data rate and mean axial velocity both decreased to nearly zero as the detection volume reached the outer edge of the spray Infrared (IR) Imaging An infra red camera was used to measure the exterior surface temperature of the insulated enclosure as shown in the photographic image in figure The camera used was form Mikron Infrared, Inc, Model no. 72v. The IR imaging camera (MikroScan 72 V) is a non-contact, high sensitivity infrared radiometer. It measures the infrared radiation emitted by the target surface and converts this radiation into a two-dimensional image related to the temperature distribution at the target surface. The IR camera senses thermal energy that is emitted from the target object. Through use of the software (MikroSpec), temperature variations over the area included in the field of view can be displayed. The total radiant emission of a blackbody is given by Stefan Boltzman equation W = at 4. The temperature of the blackbody can be obtained directly from the radiant energy of the blackbody by this equation. For normal objects, the right side of above equation multiplied by the emissivity. The data processing software (MikroSpec) allows the use r to view the temperature data of one or more points at selected locations anywhere within the field of view. The camera was mounted on a regular tripod. Two different temperature ranges were used to 19

132 take IR images at different location on the enclosure. The ranges used were to 5 C and to 2 C. The default emissivity value of.85 was used Test Conditions Experiments were conducted for fixed VO flow rate of 12 liters per minute (lpm). The atomizing air flow rate through the injector was kept constant at about 25 LPM which corresponds to an ALR of 2.5. The combustion air flow rate was kept to about 125 LPM to maintain a total air flow rate of 15 LPM throughout the experiments. As a first step the cold open spray experiments were done with and without the swirling air to study the effect of swirling air on the spray. Then the open flame spray data were taken which required swirling air flow with small amounts of methane fuel to stabilize the flame. Experiments were conducted with unheated and heated (to C) VO. The PDPA system was traversed radially and axially to acquire flow and drop size measurements at different locations in the open cold spray as well as open flame. Next measurements were acquired in spray flame within an insulated enclosure. Again, the VO was supplied either at room temperature or at C. The experiment was started by supplying gaseous methane and then, igniting the methane-air reactant mixture in the combustor. Next, the liquid fuel flow rate was gradually increased in steps of 2 mlpm to attain the desired value of 12 mlpm, while the methane flow rate was slowly decreased to zero. In this study, the volume flow rate of total air (combustion + atomizing) was constant at 15 standard lpm. Experiments were conducted for fixed volume flow rate of fuel at 12 mlpm. The time required for the fuel to flow through the system and reach the atomizer was about 3 minutes. Another -4 minutes were required before the fuel temperature reached the set value. 11

133 Once a stable flame was achieved, the liquid fuel flow rate was gradually increased while the methane flow rate was gradually reduced. Initially the methane fuel was supplied at 18 lpm for about 1.5 hours and then gradually reduced to about 3.88 lpm. This procedure took about 2.5 hours before a stable VO flame at 12 mlpm was achieved without any condensation in the optical measurement area. The equivalence ratio (φ) was maintained constant at throughout the experiment. The equivalence ratio in case of open flame was about 1.2 which did not account for the ambient air entrainment. The emissions measurements were taken for the spray flame along with the spray measurements using a 2 way velmex traversing system shown in figure 5.9 (b). 5.3 Results and Discussion Experiments were conducted to measure the spray characteristics to demonstrate the (i) effects of swirling air flow on open cold spray, (ii) effects of flame on open spray, and (iii) effects of enclosure on spray flame and emissions. The effect of fuel inlet temperature is also studied for unheated and preheated VO at C. The following sections will discuss important observations from the current study. Results are shown in the form of contour plots and profiles of axial velocity, RMS velocity, SMD, and droplet diameter distribution. For each case, the contour plots will be discussed first followed by the profiles plots Effect of Swirling Air Flow on Open Cold Spray (a) Axial Velocity Contours (Mean and RMS) Figures 5.15 and 5.16 shows the contour plots of axial velocity for cold spray with and without swirl, respectively. The plots were constructed using measurements taken for different axial and radial locations in the spray. The axial location was varied from 5 mm to 75 mm from the injector exit in steps of 5 mm interval. Transverse measurements were taken in steps of 1 mm 111

134 across the spray. As mentioned earlier, the experimental system was kept stationary while the PDPA system was traversed radially and axially. In general, for the contour plots shown in both the cases, the mean axial velocity peaks in the centre of the spray, with the velocity magnitude diminishing towards the edge of the spray. The peak axial velocities are seen to be slightly higher in the near field axial locations from 5 mm to 2 mm in the cold spray with swirl as compared to the cold spray without swirl. Farther downstream in the mid field (Y = 2 mm to 4 mm) and far field (Y = 45 mm to 75 mm) locations in the spray, the peak axial velocities for cold spray without swirl are higher as compared to the cold spray with swirl. For both cases as the flow field diverges axially downstream, the velocity decreases with the loss of kinetic energy of the droplets. For the cold spray without swirl, the axial velocity in the midfield and far field locations is about 12 m/s and 8 m/s higher as compared to cold spray with swirl. Figure 5.16 reveals a well-pronounced flow recirculation zone near the outer edge of the spray between the radial locations of 1 mm and 4 mm.note that the near field region of the spray, the values of axial velocity are negative. The presence of swirl widens the spread of the spray and increases the velocity near the injector exit while drastically reducing the velocity magnitude farther downstream compared to cold spray without swirl. Near the injector exit, the effect of swirl is to create high axial velocity owing to the increased airflow rate. However, farther downstream, the swirling air flow introduces additional turbulent mixing to widen the radial extent of the jet spray with sufficient reduction in axial velocity as compared to cold spray without swirl. Figure 5.17 shows the mean axial velocity contour plot for cold spray with swirling air flow for preheated VO at C. As expected, the axial velocities show an increase in magnitudes compared to the unheated case. This increase in the axial velocity results from the 112

135 heating of the atomizing air by heated VO. For preheated VO, the peak velocities at any given axial location in the spray are higher by about 1-15 m/s as compared to the unheated VO case in figure 5.8. With increase in fuel inlet temperature, the radial extent of the spray also decreases as compared to the unheated case. Higher axial momentum decreases the radial diffusion of the jet spray, resulting in a narrower spray in case of preheated fuel. Figure 5.18, 5.19 and 5.2 shows the contour plots for RMS axial velocity for no swirl flow, swirl flow and cold spray with swirl flow for VO at C, respectively. As shown in figure 5.15, for cold spray without swirl flow, the peak RMS axial velocity value is observed to be of 18 m/s while for cold spray with swirl the peak RMS velocity in figure 5.19 is around 22 m/s. Both the plots in figures 5.17 and 5.19 show a central dip and peaks on either side of the jet due indicating shear interactions between the jet and surrounding gaseous phase in the near injector exit locations. The peak in the RMS axial velocity for cold spray with no swirl flow is observed at X = 5 mm for Y = 5 mm to 1 mm while for cold spray with swirl flow the RMS peaks at X = 1 mm for Y = 5 mm to 2 mm. Farther downstream for both the cases without and with swirl flow, the RMS velocities show more uniform distribution compared to the near injector plane locations. The RMS velocities trail off trend on the outside edge of the spray. The higher magnitude of RMS velocity in cold spray with swirl flow indicates improved inherent turbulent mixing due to the surrounding swirling air flow as compared to no swirl flow in figure Figure 5.21 shows the RMS contour for cold spray with swirl for VO at C. Higher fuel inlet temperature shows a slight increase in RMS axial velocity compared to the unheated cold spray. The increase in the RMS axial velocity results from the heated fuel raising the temperature of the atomizing air to produce higher turbulence levels leading into smaller diameter droplets. The peak RMS value is about 25 m/s. 113

136 (b) SMD Contours The effective method of expressing the quality of the atomization is the mean diameter. Probably, the most common form of mean droplet size is the Sauter Mean Diameter (SMD), which has the physical interpretation as the diameter of the drop having the same volume/surface area ratio as the entire spray. Figure 5.21 and 5.22 shows the contour plots constructed from point-wise measurements acquired at different axial and radial locations. Figure 5.21 and 5.22 shows respectively the contour plots of SMD for open cold spray with and without co-flow swirling air. As seen from the plot the spray consists of the droplets with wide range of SMD values of 2 microns to 8 microns. The plot shows the distribution of the droplets in the various regions of the spray in the mid field (Y = 2 mm to 4 mm) and far field (Y = 45 mm to 75 mm) locations. The larger droplets of 55 microns and above diameter reside in the centre core region near the injector exit (Y= 5 mm to 15 mm) as well as at the outer edge (X = ± 2 mm) of the spray for Y = 5 mm to 75 mm. The presence of larger droplets in centre of the spray can be attributed due to higher axial momentum and weaker mixing in the jet core region compared to the shear layer region of the spray, i.e. X = ± 5 to15 mm. Some larger droplets are observed towards the spray edge due to their higher momentum. At radial locations of X = ±1 to 25 mm and Y = 3 mm to 75 mm. As observer from the plot, the smaller and intermediate size droplets reside in the shear layer (X = ± 5-15 mm) both near and far away from the injector exit. Droplets that escape from the centre liquid core undergo further breakup which leads to formation of smaller droplets. However, farther downstream, more uniform and homogenous distribution of the droplet size is observed in the radial direction. For all the axial locations of the spray, the SMD values were observed to range from 2 microns to 45 microns. 114

137 Figure 5.22, shows the SMD contour for cold spray with swirling air flow and unheated VO. The SMD ranges from 2 to 6 microns. The radial spread in this case is seen to be about 3 mm as compared to 15 mm in case of no swirling air as shown in figure As observed from the plot the larger droplets are located near the jet centre, while the droplets diameter decreases away from the jet centre. Droplet diameter in the near field location of the spray, i.e. Y = 5 mm to 15 mm, ranges from 5 to 6 microns. For downstream spray locations the droplet diameter gradually decreases and farther downstream of the spray at Y = 6 mm to 75 mm the SMD decreases to about 3 to 35 microns. The droplet diameter in the shear layer, i.e. X = 1 to 2 mm, is in the range of 2 to 35 microns. It can be seen from figures 5.21 and 5.22; that the SMD is smaller for spray with swirling air compared to spray with no swirling air. The swirling air promotes secondary droplet breakup mechanism resulting into smaller SMD in the spray. Hence, the droplets escaping from the central jet core undergo further disintegration into smaller droplets ranging from 2 microns to 35 microns, due to the co flowing swirling air. This process takes place for X = 15 mm to 45 mm for almost all the axial locations in the spray. The jet spreads wider in case of swirling air compared to without swirling air. Thus swirling air flow widens the spray and it also decreases the droplet diameter because of the secondary disintegration process. Figure 5.23 shows the SMD contour plot for cold spray with swirling air for VO preheated to C. Results show similar general trends as discussed above for figures 5.21 and Larger droplets in the central core region and smaller droplets in the shear layer region near and far field locations in the spray. The maximum SMD for this case however is around 4 µm which is significantly lower than the maximum SMD of 6 µm for the unheated VO cold spray. The difference in the results is attributed to the improved atomization due to increase in VO 115

138 temperature. Increase in VO temperature, improves the thermo physical properties of the VO, e.g. reducing the kinematic viscosity and surface tension which make it easier to break-up the droplets producing finer spray. Increasing the VO temperature however requires thermal energy of about.53 % of the VO energy content (LHV = 37 kj/kg) to achieve the benefits of improved spray characteristics. (c)transverse Profiles of Mean and RMS Axial Velocity Figure 5.24 (a- z) shows transverse profiles of mean axial velocity for cold spray with and without swirling air and unheated VO. The trend showing the effect of swirling air on cold spray is explained by figure 5.24 (a) at one axial location and for the rest of the axial locations the same trends follow. Figure 5.24 (a) shows a dome shaped velocity profile for Y = 5 mm, for cold spray with and without swirl air flow. The mean axial velocity is highest at the center and it decreases towards the periphery of the spray. The peak axial velocity is similar for both the cases. But the general trend also observed in other profiles of figure 5.24 is that in the near field region, the cold spray with swirling air flow shows marginally higher axial velocity, which decreases at farther downstream locations. The peak axial velocity near the injector exit region of the spray with swirling air is about ~5-8 m/s higher compared to non swirling air spray. From figures 5.24 (a, b, c and d), we can observe a well pronounced recirculation zone in case of swirling air flow, characterized by negative velocities on either side of the spray centre. The recirculation flow improves mixing of the fuel and air to further improve atomization. Swirling air flow is also the cause of the further disintegration the larger droplets. The swirling air flow tends to increase the radial spread of the spray compared to that for the non swirling air flow. The recirculation zone can be identified from the dotted line in the plot and it can be seen at X = 1mm to 4mm for Y = 5 mm to 2 mm. 116

139 Further downstream, the peak axial velocity decreases for each case with non swirl air flow resulting in higher velocities compared to the swirl air flow case. With increased axial distance, (Y = 3 mm to 75 mm), peak axial velocities show an increase for the non swirl flow compared to the swirl flow. This effect is attributed to the intense mixing and radial spread of the spray with swirling airflow. While the non-swirling air spray maintains the high axial momentum because of poor mixing with the ambient air. Figure 5.25 shows the effect of fuel temperature on cold spray. The figure 5.25 (a) shows the axial velocity at Y = 5 mm, which shows a slight increase in peak axial velocity for VO at C mostly in the region near the injector exit in the spray. The peak velocities of about 45 m/s, 34 m/s and 33 m/s are observed at respectively Y =5 mm, 1 mm and 15 mm in the near field region of the spray. A similar trend of mean axial velocity profile is observed for both the cases at all axial locations in the spray. Only a slight increase in mean axial velocity with heated VO is observed in the near and far field regions of the spray. The effect of fuel inlet temperature is more pronounced to reduce the droplet diameter which can be seen from the SMD profile plots. For all of the aforementioned test conditions, RMS axial velocities are shown in figure 5.24 (a-z). The ratio of RMS over the mean axial velocity represents the turbulence intensity. Figure 5.24 (b), shows the RMS axial velocity at Y = 5 mm for cold spray with and without swirl. In general the profiles for both the cases show a dip in the centre of the spray, which is also observed at Y = 1 mm, 15 mm and 2 mm. This result is attributed to the intense turbulent fluctuations in the shear region of the spray where the respective gradients reach the highest values to dominate the turbulence generation rate. Further downstream the turbulence level begins to decay owing to the diffusion and reduction in kinetic energy of the droplets. In 117

140 particular, turbulent kinetic energy attenuation is significant in the shear layer because of the additional dissipation introduced by the presence of the droplets. As seen in the figure 5.24 (a), the magnitude of the RMS axial velocity values for the swirl air flow as well as non swirl flow is almost overlapping. This result is consistent with the higher axial mean velocities as discussed previously in the near field regions of the spray. A local maximum is seen at radial location X = 5 mm, indicating that the faster and smaller drops have higher RMS axial velocity values. For Y = 3 mm to 75 mm, the dip in the RMS axial velocity is no longer observed and the profiles are more uniform indicating shear layer merging with the centerline at these downstream locations in the spray. The RMS axial velocity values for Y = 3 mm, 35 mm, 4 mm are about 1 m/s, 8 m/s and 7 m/s for both the cases. Further downstream at Y = 6 mm, 7 mm and 75 mm, the RMS axial velocity of about 3 to 5 m/s are similar for swirl air flow and non swirl air flow respectively. (d)transverse Profiles of SMD Figure 5.26 shows the radial profiles of SMD at different axial locations in the spray to demonstrate the effect of swirl air flow on the cold spray. Figure 5.26 (a) shows the SMD profile for cold spray with and without swirl air flow at of Y = 5 mm. In general, the SMD values are seen to be higher at the centerline with a diminishing trend of smaller droplets towards the edge of the spray. For the near field location in the spray i.e. Y = 5mm, SMD values at the centre is around 75 µm. Moving away from the centre, the droplet diameter decreases. A slight difference is observed in the SMD values for cold spray with and without swirling air, the latter showing smaller droplets. The profile for cold spray with swirl is tapered radially at X = 2 mm, since the swirling co flow air tends to spread the spray and further disintegrate the droplets to form smaller drops. The radial extent of the spray is wider for swirl air flow case compared to the without 118

141 swirl case, i.e. 55 mm for swirl case and 3 mm for non swirl spray. SMD values at Y = 5 mm range from 25 µm to 72µm for cold spray with swirl air flow, while that of cold flow without swirl air flow ranges from 4 µm to 8 µm. Similar trend is observed for other near field locations in the spray i.e. at Y = 1 mm, 15 mm and 2 mm. The maximum SMD is nearly the same at Y = 1 mm, 15 mm and 2 mm with and without swirl air flow. However, the magnitude of the SMD decreases in the axial direction. The noticeable difference in the droplet diameter is observed from Y = 3 mm to 75 mm in both the cases. The co flow swirling spray decreases the SMD by about 25 µm as compared to the non swirling spray. For co flow swirling spray there is a decrease of about 25 microns in SMD for Y = 3 mm to 75 mm. In the same region the SMD for the non swirling spray decreases from 57 microns to 45 microns. Thus the swirl air flow in the cold spray decreases the droplet diameter which would improve atomization and combustion. Figure 5.27 shows the effect of fuel inlet temperature on radial profiles for SMD in the spray. With the preheated VO at C, the maximum SMD decreases to 4 microns form 7 microns at Y = 5 mm. The range of SMD at this axial location is about 17 microns as compared to the wider range of 38 microns in case of unheated VO spray. The radial extent with the smaller droplets is greater in case of preheated VO spray compared to the unheated VO spray. Increase in fuel inlet temperature produces a spray with droplet diameter ranging from 18 microns to 4 microns; whereas the droplet diameter range in the unheated fuel spray is from 26 microns to 7 microns. Evidently the effect of fuel inlet temperature is seen at almost all axial locations in the spray ranging from near field to far field locations.the higher fuel inlet temperature decreases the fuel kinematic viscosity and surface tension, which aids in improving the atomization characteristics to produce finer droplets. Increased fuel inlet temperature 119

142 decreases the SMD and its range as well as the radial spread of the spray as seen from the mean and RMS axial velocity profiles. (e) Droplet Diameter Distribution Profiles To better understand the spray structure, figures 5.28 compares the droplet size distribution profiles at the centre of the injector exit. Figure 5.28 (a) shows the droplet size distribution profiles at the Y = 5mm, and X = mm. The SMD is 64 µm and 76 µm, respectively for cold spray with swirling air and without swirling air. As seen from the plot, cold flow with swirling air shows a higher percentage of smaller diameter droplets in the range of < 5 µm as compared to the non swirling spray. The largest droplet diameter is about 2 µm for both swirling and non swirling spray, although very few larger diameter droplets are observed in case of the swirling spray. For swirling spray, the SMD distribution shows most of the droplets with diameter of about µm or less while for non swirling spray the distribution spread out with a higher percentage of larger diameter droplets. The greater number of smaller diameter resulted in the smaller SMD in case of swirling spray as compared to non swirling spray. Figure 5.28 (b) shows the distribution profiles at Y = 4 mm, X = mm which is considered here as the mid field location in the spray. Close to 9% of the droplets in case of swirling spray are smaller than µm. The largest diameter in non swirling spray is about 15 µm. A significant difference in the distribution count is seen for droplets of diameter < 35µm, with swirling flow showing higher percentage compared to the non swirling flow. The presence of these larger droplets in case of non swirling spray is the cause of higher SMD of about 46 µm compared to that of 36 µm in case of swirling spray. Comparatively, the near field region of the spray shows greater number of larger droplets than at the axial locations away from the injector exit. 12

143 Figure 5.28 (c) shows the distribution profile for the swirling and non swirling spray at Y = 6 mm, X = mm, i.e. at a far field location in the spray. The profiles for both the cases are very similar, still showing greater number of larger droplets in case of non swirling spray. The highest droplet diameter in case of non swirling spray is 15 µm as compared to that of 8 µm for swirling flow. The presence of these larger droplets dominates the SMD in case of non swirling spray showing higher overall SMD s compared to the cold spray with swirling air flow. The SMD s are 36 µm and 44 µm for swirling spray and non swirling spray, respectively. As explained earlier, the droplet distribution is narrower with greater number of smaller drops at far field regions of the spray than that observed at the near field locations, which is consistent with the of higher SMD in the near the injector exit at the centre of the spray and smaller SMD father downstream. Close to the injector exit the spray is dense and due to the presence of various droplet sizes the distribution becomes rather wider displaying discrete maximum that that father downstream centre of the spray Effect of Flame on Open Spray (a) Velocity Contours (Mean and RMS) Figure 5.29 and 5.3 shows respectively the mean axial velocity contour for spray flame of unheated and preheated VO. The measurements were acquired for Y = 5 mm to 4 mm in the flame spray. The measurements in the flame spray downstream of this axial location were exceedingly difficult because of the high intensity flame radiation and extremely low data rate. For the PDPA system, the signal to noise ratio was too small to obtain any accurate information at low data rates. In general the mean axial velocity exhibits a peak at the centre and decays gradually in radial direction, similar to the cold spray results discussed above. However, for the spray flame the peak velocities are higher because of the increased temperature resulting from 121

144 heat release. Unlike, cold spray with swirl flow, the negative velocities for the droplets is not measured in flame spray, since they evaporate completely The mean axial velocity increases drastically in the flame spray compared to the cold spray. The difference of about 35-4 m/s is observed in the peak mean axial velocity for both the cases. The rapidly moving smaller drops in the mid spray region are attributed to the smaller drag drop forces and volumetric expansion of the hot gases. Slow moving larger drops in the outer edge of the spray result from high momentum of such drops. Similar trend is observed for the flame spray for VO at C as well. The peak mean axial velocities show a marginal difference for unheated and heated VO. Results show that fuel preheating increases mean axial velocity leading to improved atomization. Similar improvement is also observed in the combusting sprays compared to the non-burning cold sprays. (b) SMD Contours The SMD contour plot for open flame with unheated VO is presented in figure The larger droplets are seen close to the injector exit between Y = 5 and 1 mm, and also on the outer edge at farther downstream locations. The magnitude of the SMD ranges from about 6 µm to 37 µm. The maximum SMD of 37 µm in case of flame spray is significantly smaller compared to that of 6 µm for cold spray as seen in figure The decrease in the SMD is seen in the near field i.e. Y = 5 mm to 1 mm. The decrease in the droplet size can be attributed to the droplets passing through the flame zone. The high temperatures in the flame tend to reduce the droplet diameter throughout the spray including the core region near the jet centre where larger droplet sizes are observed in the cold spray. The SMD range in case of flame is narrower than the cold spray indicating high evaporation rate for the former case.farther downstream of the injector exit, SMD decreases rapidly indicating faster evaporation of droplets in the flame. Large droplets 122

145 do not evaporate completely and thus they are still able to migrate towards the outer edge of the spray. Figure 5.34 shows the SMD contour of an open flame for preheated VO at C. In this case the maximum SMD decreases to about 3 µm compared to 37 µm for unheated VO. The flame tends to reduce the radial jet spread and produces a narrower range of SMD indicating droplet evaporation associated with high flame and fuel temperatures. (c) Transverse Profiles of Mean and RMS Axial Velocity Figure 5.35 shows the mean axial velocity profiles in the cold spray and spray flame. The measurements were taken for axial locations of 5 mm to 4 mm in the flame, because of the low data rate at Y > 4 mm, indicating nearly complete fuel vaporization at these locations. In general, the mean axial velocity peaks at the centre and decreases towards the edge of the spray, similar to the trends in the cold spray discussed above. Figure 5.35 (a) compares the mean axial velocity for cold spray and flame spray at Y = 5 mm. The flame spray shows peak velocity of 7 m/s as compared to that of 45 m/s in the cold spray. Thermal expansion in the flame results in droplets with higher mean axial velocity. At Y = 3 mm, the peak axial velocity decreases to about 52 m/s as compared to peak axial velocity of 14 m/s for cold spray. The decrease in the axial velocity is attributed to the spray extending over a wider region at downstream locations. Figure 5.35 (h) shows that at Y = 4 mm, the peak axial velocity in the flame spray decreased to about 38 m/s compared to that of 1 m/s in the cold spray. Figure 5.36 shows the effect of flame on spray for VO at C. The observed trend for the axial velocities is similar to as discussed above. The mean axial velocity at Y = 5 mm, peaks to 75 m/s in the flame spray as compared to peak value of 48 m/s in the cold spray. At Y = 1 mm and Y = 15 mm, the profiles show a dip in the jet centre while peak velocity on either side of 123

146 centerline at X = 1 mm. Farther downstream in the spray at Y = 3 mm and Y = 4 mm, the flame spray velocity peaks at 58 m/s and 5 m/s compared to 2 m/s and 15 m/s in case of cold spray. Figure 5.35 shows that the RMS axial velocity is observed in the flame spray as well, as discussed above for cold spray. The flame spray profiles show a dip in the centre and peaks on either side of the centerline which represent the flame locations at X = 15 to 28 mm, unlike the cold spray having a decaying trend with peak at the centre. This is attributed due to the presence of higher turbulent fluctuations in case of the flame spray as compared to that of the cold spray. The magnitude of RMS axial velocity ranges from 1 m/s to 22 m/s which is consistent to the mean axial velocity discussed in the previous sections. The maximum RMS axial velocity values of about 22m/s, 18 m/s and 16 m/s are observed at Y = 5 mm, 1 mm and 15 mm, respectively. For the flame spray, the RMS axial velocity is higher than those of the cold spray on the outer edge of the spray while they are almost similar in magnitude in the center of the spray. Figure 5.36 shows the similar trend for flame spray for preheated VO at C. At the near field locations, the profiles peak nearly at same values for both flame spray and cold spray at Y = 5 mm, 1 mm and 15 mm. The flame sprays profiles show a slight asymmetry at the centre which can be attributed due to the asymmetry in the flame. The magnitude in the RMS axial velocity values varies from 5 m/s to 2 m/s for the axial locations of Y = 5 mm to 4 mm in the flame spray. As discussed above the RMS values are smaller in case of cold spray compared to that of flame. (d) Transverse Profiles of SMD Transverse profiles of SMD for flame spray and cold spray are presented in figure Figure 5.37 (a), shows the effect of flame on spray at axial location of Y = 5 mm. For the flame 124

147 spray the profiles of SMD show smaller droplets in the centre and slightly larger droplets on each side, and gradual decrease in droplet diameter in the transverse direction. This trend is same for near field axial locations of Y = 5 mm, 1 mm, 15 mm and 2 mm. The maximum SMD for the near field location in the flame spray is observed to be about35 µm at Y = 5 mm, 25 µm at 1 mm and 15 µm at 15 mm. Farther downstream, a noticeable change in the SMD trend is observed; smaller droplets are at the centre while the larger droplets on the outer edge. This result is attributed to the fact that the droplets at these locations are passing through the flame zone. The size difference farther away from the centerline is observed to be larger because at the outer edge of the spray the smaller droplets evaporate faster compared to the large droplets and the larger droplets with high inertia penetrate out of the flame. The escaped larger droplets are believed to burn in diffusion mode. Comparatively the SMD range for the burning spray is about 35 µm smaller than that of the cold spray for near field locations. The presence of flame increases the evaporation rate of the droplets hence producing smaller droplets. The radial spread for the flame spray is narrower compared to cold spray indicating that the droplets in case of cold spray possess higher initial momentum hence penetrating farther and wider. Figure 5.38 shows the radial profiles for SMD for fuel inlet temperature of C for flame and cold spray. Similar trend is observed for all the cases in figure 5.37 as discussed above. The effect of higher fuel inlet temperature is to decrease the fuel kinematic viscosity and surface tension, hence improving atomization characteristics by decreasing the mean droplet size. For cold spray the maximum SMD is 45 µm compared to 3 µm for the flame spray at Y = 5 mm. The SMD values decreases farther downstream in the spray compared to the near field locations. And this trend decreases at farther downstream locations in the spray. Evidently the higher fuel inlet temperature helped burning of the droplets much faster compared to unheated 125

148 case discussed in figure The minimum SMD measured for the flame spray is reduced to about 1 µm compared to 3 µm in the cold spray. The minimum for unheated flame spray case was about 2 µm. Also, the axial locations in the flame spray were measured to only about 4 mm, since most of the droplets were evaporated by the time they reached farther downstream. Hence the data rate was significantly low to measure any signal from the droplets above this axial locations in the spray. (e) Droplet Diameter Distribution Profiles Figure 5.39 shows the droplet diameter distribution profiles for flame spray and cold spray at near and far field locations of Y = 5 mm, 1 mm, 35 mm and 4 mm and X = mm.figure 5.39 (a) shows the distribution profile for near the injector exit in the centre of the spray at Y = 5 mm. Results show a significant difference in the droplet diameter distribution for flame and cold spray. The flame spray is observed to have most of the droplets of 5 µm while large droplets with > µm are found in the cold spray. The largest droplet in the flame spray is about µm where as for the cold spray it is about 2 µm. For the flame spray the droplets tend to start vaporizing due to the high flame temperature zones, impact of phase interactions and combustion producing smaller droplets and narrower distribution compared to the wider spread distribution as observed in the cold spray. The SMD for flame spray and cold spray is respectively 26 µm and 64 µm, indicating rapid evaporation of the droplets in the flame compared to the cold spray. Figure 5.39 (b, and c) show respectively the droplet diameter distribution profiles for Y = 1 mm and 35 mm and X = mm respectively. At Y = 1 mm, X = mm, the flame spray has most of the droplets of diameter <3 µm while the cold spray has majority of droplets 13 µm diameter. The largest diameter for flame spray is 75 µ m while that of cold spray is 175 µm. The 126

149 SMD is 19 µm and 5 µm, respectively, for flame spray and cold spray. The higher count of larger droplets dominates the higher SMD for cold spray. The larger numbers of smaller droplets in case of flame spray result in a decrease in SMD. At Y = 35 mm as shown in figure 5.39 (c), the largest droplet diameter in cold spray is 125 µm where as that in the flame spray is 6 µm. The respective SMD for these cases is 17 µm and 38 µm for flame spray and cold spray. Again there is a higher count of smaller droplets in case of flame spray resulting in the overall decrease in the SMD. As observed the distribution gets narrower with increase in axial distance, due to the droplets getting evaporated rapidly and combusted in the flame locations and leading to smaller number of larger droplets. The distribution profile is narrower in for far field locations in the spray for both the cases compared to the locations near the nozzle exit. Moreover the number density of the droplet further downstream decreases in the flame spray due to the droplet evaporation and burning. Figure 5.4 shows the droplet diameter distribution profiles for the flame spray for preheated VO at C. The largest droplet diameter for cold spray is observed to be 14 µm where as for flame spray it is 9 µm. The increase in fuel inlet temperature produces a narrow distribution profile with lower range of droplet diameter and hence with overall smaller SMD. The SMD s for both the cases is 28 µm and 37 µm for flame spray and cold spray respectively. At Y = 1 mm, X = mm the largest SMD for flame spray is of about 75 µm as well as for cold spray is of about 115 µm respectively. This decreases the global SMD to 22 µm and 33 µm for flame spray and cold spray. The increase in fuel inlet temperature aids to further reduce the SMD as compared to the unheated case. As we move axially farther downstream, the droplets distribution becomes narrower and the number density also reduces especially significantly for the flame spray attributed to the fact that the droplets get vaporized in the flame zone. Also, the 127

150 preheating of the oil produces the finer droplets to begin with since it improves the thermo physical properties of the fuel itself, and hence the finer droplets get vaporized and consumed faster. Farther downstream at Y = 35 mm, the largest droplet diameter is of about µm for cold spray and 4 µm for flame spray producing the SMD of 31 µm and 13 µm respectively. Similar to the previous discussions the far field region in the spray has narrower distribution with smaller droplet number density compared to that of the centre of the spray. This is consistent to that of the SMD profiles with dense centre liquid jet near the injector exit showing higher SMD values and decreasing trend as we move axially downstream with a wider and more dispersed spray. Also the fact that the droplets in case of flame spray get consumed faster and get evaporated rapidly, mostly the smaller ones compared to the larger drops and hence globally being responsible for smaller SMD values Effect of Enclosure on Spray Flame and Emissions Experiments were done to study the effect enclosure on flame spray characteristics and combustion emissions. The enclosure was insulated to provide thermal feedback to the VO flame and maintain a stable VO flame during all experimental conditions. It required about 2 hours to preheat the enclosure with a stable VO flame before the methane flow rate was reduced to about 3.8 SLPM which corresponds to φ =.89. This mode of operation resulted in no condensation of the droplets on the glass window of the enclosure. Practically in the continuous combustion operations of gas turbine applications the enclosed flames are used, and hence the present work also incorporates enclosure effect on the spray characteristics and emissions. At steady state insulation around the enclosure, retained sufficient heat within the system to raise the VO temperature at the injector exit to 9- C 128

151 (a) Transverse Profiles of Mean and RMS Axial Velocity Figure 5.41 shows the mean axial velocity profiles for enclosed flames of VO preheated to C and 15 C. As seen in figure 5.41 (a), at Y = 5 mm, mean axial velocity shows a peak at the centre and gradual decrease towards the edge of the spray. The peak mean axial velocity for both cases is nearly the same. For both cases, the axial velocity peaks at 7 m/s, and both the profiles almost overlap each other, with slight difference towards the edge of the spray. At Y = 1 mm and 15 mm the profiles are again observed to be overlapping, although at higher fuel inlet temperature the spread of the spray is observed to be narrower as compared to VO at C. The transverse spray for VO at C is extended widely from -4 mm to 3 mm while for VO at 15 C, the spread narrows down to a -15 mm 2 mm for Y = 5, 1,and 15 mm axially. Farther downstream, the peak axial velocity decreased down to 6 m/s for VO at 15 C and 5 m/s for VO at C. Unlike velocities, a significant difference is observed in the droplet sizes as shown in the SMD profiles for the enclosed flame. In enclosure, the co flow swirling air is the combustion air for the flame without any ambient entrainment. In open flame there is no heat feedback but there is a lot of ambient entrainment, while in enclosure heat feedback is there due to insulation. Figure 5.41 (b) shows the RMS axial velocity profiles for enclosed flame. For VO at C at Y = 5 mm, the peak RMS axial velocity is 22 m/s while for VO at 15 C the peak RMS axial velocity is 2 m/s. For VO at C, the profiles show a local minima at the jet centre line while double peaks at flame locations i.e. X = -3 mm to 15 mm, and Y = 1 mm to 35 mm. The magnitude of the RMS axial velocity range from about 2 m/s to 22 m/s for Y = 5 mm to Y = 35 mm. Slight asymmetry in the RMS axial velocity profiles is observed at far field locations in the spray i.e. Y = 25 mm to 35 mm. Comparatively for VO at 15 C, the RMS axial velocity profiles 129

152 looks more symmetrical with local minima dip in the centerline as well as double peaks on the flame locations. This result is attributed to the fact that the jet centre is subjected to the small turbulent fluctuations with lower RMS values compared to that at the edge of the spray. Near the injector exit i.e. at Y = 5 mm to 15 mm, the peak RMS axial velocity for both the cases overlaps with no significant difference in the velocity values. At locations farther downstream, the peak RMS axial velocities for 15ºC are higher at X = 2 mm, indicating higher velocity fluctuations compared to VO at C. (b) Transverse Profiles of SMD Figure 5.42 shows the transverse profiles of SMD for enclosed VO flame at c and 15 C. Figure 5.42 (a) shows the SMD profile at Y = 5 mm. The profiles show peak SMD value at the centre of the spray. The maximum SMD for VO at C is 34µm, while for VO at 15 C is 32 µm. At Y = 15 mm and 2 mm, the range of SMD values decreases. The profiles no longer show a well-defined central peak as compared to near the injector exit. SMD values range from 1 µm to 25 µm for VO at C while for VO at 15 C the SMD ranges from 18 µm to 28 µm. Farther downstream, at Y = 25 mm to 35 mm, the larger droplets are observed to have moved towards the periphery of the spray causing higher SMDs at the outer edge of the spray as seen in figurer 5.44 (e and f).the profiles show a central depression as compared to central peak observed in the near injector locations (Y = 5 and 1 mm). For VO at C, the SMD ranges from 16 µm to 25 µm, which is larger compared to 18 µm to 22 µm for VO at 15 C. For VO at 15 C, the transverse distribution of SMD is narrower compared to VO at C. This result is attributed to the higher fuel inlet temperature and improved vaporization of the droplets thus reducing the SMD at downstream locations in the flame. At Y = 35 mm, the SMD profile for VO at C by 5 µm compared to that for VO at 15 C. For VO at C, the SMD values show a 13

153 decrease of about 12 µm with axial distance while that for VO at 15 C, the SMD decreases by about 15 µm. (c) Droplet Diameter Distribution Profiles Figure 5.43 shows the droplet size distribution profiles for enclosed VO flame at C and 15 C respectively. At Y = 5 mm, X = mm, the largest diameter for VO C is 15 µm while that for VO at 15 C is 6 µm. For VO at 15 C, a narrow distribution profile is observed with greater number of smaller droplets. 9 % of the droplet sizes are in the range of < 5 µm. The presence of few large drops highly influences the SMD in case of lower fuel inlet temperature, i.e. VO at C. The SMD for VO at 15 C is 28 µm while for VO at C is 32 µm. At Y = 1 mm, X = mm, the distribution profile is almost similar for the drops < 3 µm.a minor difference in droplet diameter distribution is observed in the droplets > 5 µm. The lower fuel inlet temperature i.e. VO at C, has more large droplets compared to that for VO at 15 C. The SMD for VO at C is 27 µm and for VO at 15 C is 25 µm. Largest diameter for VO at 15 C is of about 6 µm and for VO at C is of about 8 µm respectively. At Y = 35 mm, X = mm, VO at 15 C shows higher count of smaller droplets for drops < 11 µm compared to that of VO T C. The largest droplet for VO 15 C is 7 µm and for VO C 58 µm respectively. For drops > 4 µm, the distribution in case of VO 15 C is higher than that of VO at C. The respective SMDs for both the cases are 2 µm and 18 µm which are influenced by the presence of few large droplets in the case of VO at 15 C. At Y = 35 mm, X = mm, the largest diameter for VO at 15 o C is 6 µm where as for VO o C is 55 µm. The respective SMDs for both the cases are 16 µm and 2 µm, which is attributed to the fact of larger percentage of smaller droplets for higher fuel inlet temperature i.e. VO at 15 o C compared to that of lower fuel inlet temperatures. As the axial distance increases the droplet 131

154 distribution is observed to be narrower and also more number of smaller droplets is observed for both the cases. Comparatively the higher fuel inlet temperature shows improved atomization with higher percentage of smaller drops compared to that of lower fuel inlet temperature. (d) Effect of Enclosed Spray on Combustion Emissions The emissions of CO and NOx were measured for three different axial locations in the flame. The radial profiles are plotted for axial locations at Y = 43.5 cm, 41. cm and 38. cm. The combustor enclosure is of a regular pentagonal shape of about 15 cm measured diagonally. Figure 5.44 shows the CO profile for VO at C and 15 C at Y = 43.5 cm, 41. cm and 38. cm at the combustor exit plane. At C, the CO emissions are seen to be higher on negative transverse location of the flame as compared to positive transverse location as seen in the plot. This is likely due to the emissions probe extending across the combustor exit plane, which results in heating of the probe due to exposure to the flame at the exit plane. This causes further dissociation of the CO2, resulting in higher CO emissions as compared to the positive transverse locations where the reactions get cooled down rapidly resulting in lower CO emissions. The CO emissions for VO at15 C range from 2 to 5 ppm while those of VO at C range from 4 to 12 ppm. The CO emissions at the combustor exit plane indicate that sufficient flow mixing has taken place, forming a homogenous product gas mixture. The CO emissions show a minor decreasing trend in axial direction but mostly they are in similar range. The decrease in emissions for VO at 15 C is about 8 ppm compared to VO at C, which is attributed to the smaller droplet diameter for flame with higher fuel inlet temperature. Figure 5.45 shows the NOx emissions profile at the aforementioned axial locations in the flame. A significant reduction in NOx emissions is observed for VO at 15 C compared to VO at C. The NOx emissions for VO at 15 C are constant at about 25 ppm, where as for VO at 132

155 C, the NOx emissions are significantly higher, in the range of 12 ppm to 16 ppm. The NOx emissions for different fuel temperatures are shown in chapter 3, which show a similar trend of emissions for a given ALR i.e. for VO at 99 C, the NOx emissions range from 13 ppm to 155 ppm. An increase in fuel inlet temperature decreases the kinematic viscosity of the fuel hence improving the atomization process, having significant effect on emissions. Spray droplet characteristics cannot be directly compared to the emissions, since the emissions data were taken at the axial locations farther downstream near the combustor exit plane. Most of the droplets were consumed and vaporized in the flame and the PDPA system could not detect the droplets, because of the low data rate after Y = 4 mm. near the injector exit, the CO emissions could not be measured because it was not possible to quench the reactions consuming CO. The reactions near the injector exit are supposed to be incomplete, due to continuous combustion and the emissions probing within the flame should give detectable CO emissions. However in our case the probe was exposed to a high temperature environment, which allowed the reactions to proceed further within the sampling probe. Since we did not use a water-cooled sampling probe, the measurement of CO emissions was not possible within the flame. (e) Enclosure Exterior Surface Temperature Distributions Figure 5.46 shows the infra red camera image of the enclosure. The exterior surface temperature data is inferred from the thermal radiation captured by the infra red image as shown in figure Image reveals the transverse profiles of temperature across the enclosure surface at 5 different axial locations. The surface temperature profile shows lower temperature for points 1 and 2 and higher for the points 3 and 4 and again lower for point 5. This decrease can be observed as point 5 is exposed to exit of the combustor close to the exhaust ambient air while for points 1 and 2 the heat loss is seen to occur due to the uninsulated glass area. Point 4 has the 133

156 higher average temperature of about 132 C which indicated that most of the heat is retained in the flame zone at that axial location. The average temperature as seen from the figure varies from a minimum of 113 C to 127 C. The minimum temperature observed on each axial location would in fact be higher than the indicated because of variation in view factor, since the enclosure is not a flat surface facing normal to the viewing direction. The principle on which IR camera works is based on the radiosity readings, and hence the view factor, emissivity of the object in the field of view would affect the temperature readings. The IR image provided with an idea on how the exterior surface temperature varied. The red contour region in the figure shows the highest temperature which is for the steel bars used to support the glass in the enclosure. Again a significant heat loss is there due to the surface exposed to the ambient air and also the small window kept uninsulated for optical access. However, the insulation helped retaining heat with minimum heat loss through the surface and thus allowing minimum supply of methane to achieve a stable VO flame throughout the experimental duration. The temperature profile at five different axial locations is shown in figure Conclusions The present study investigates the spray characteristics of heated and unheated vegetable oil. The measurements of mean axial velocity, RMS axial velocity, SMD, and droplet diameter distributions are presented to explain the effect of swirling air on a non-evaporating spray, effect of flame on spray, and effect of enclosure on spray. The mean axial velocity for the swirling air flow is observed to be higher than the non swirling air flow. The axial velocity further decrease in the farther downstream locations in the spray. The swirling spray shows higher axial velocity in the near field region of the spray; whereas for far field regions the non swirling spray shows the axial velocity 134

157 profiles with higher magnitudes. Similar trend is observed for RMS axial velocity profiles; higher RMS axial velocity near the injector exit compared to farther downstream locations in the spray. And higher RMS axial velocity in the centre of the spray while lower RMS axial velocity at the edge of the spray. Swirling air flow in non evaporating spray lowers the SMD. The reduction is significantly in the far field locations of the spray i.e. Y = 25 mm and higher. In the near field location of the spray i.e. Y = 5 mm to 2 mm, the swirling air flow tends to increase the radial extent of the spray hence further disintegrating the droplets. The effect of swirling air flow is more prominent near the injector exit while the spray flattens out farther downstream, overall improving the atomization compared to the non swirling air flow. Higher flame temperatures tend to accelerate the droplets leading into increase in mean axial velocity at all the measured axial locations compared to the cold spray. The mean axial velocity peaked at the centre while they showed a decreasing trend at the outer edges of the spray. RMS axial velocity increases for flame spray compared to cold spray. The droplet diameter is observed to be smaller in case of flame spray compared to the cold spray. SMD significantly reduces with axial distance. A decrease of about 4-5 µm is observed in the SMD for the flame spray compared to cold spray with swirl air flow. This result is attributed to the improved vaporization because of high flame temperatures versus cold spray. The effect of fuel inlet temperature improves the fuel kinematic viscosity and surface tension hence improving the atomization leading to reduction in SMD. This result is observed in both cold spray and also for the flame spray with heated and unheated VO. 135

158 Increase in mean and RMS axial velocity is observed for higher VO temperature compared to unheated VO. The greater number of larger droplets in case of unheated VO dominates the higher droplet diameter compared to VO at C. Higher fuel inlet temperatures aids to improve droplet vaporization and hence combustion reactions thus improving the overall spray characteristics.this further improves chemical reactions thus reducing the emissions due to the thermal feedback from the flame delineating the effect of heat release rate on the fuel vaporization. The effect of enclosed flame was seen on two different fuel inlet temperatures with higher fuel inlet temperature producing finer droplets lowering the SMD, increasing the mean and RMS axial velocity. The insulated enclosure helped to provide heat feedback which enabled the PDPA measurements avoiding any fuel condensation in the measurement area. The confinement of the flame due to enclosure reduced the radial spread compared to open flame conditions. There was no significant difference in the mean axial velocities and SMD values. Based on this study we conclude that the smaller drop size distribution results in lower emissions of CO and NOx. Larger droplets tend to burn in diffusion mode compared to finer drops that lead to premix mode of combustion. Larger droplets that are burning in diffusion mode results into higher local temperatures and higher flame temperatures result in increase in NOx and the more likely chances of fuel pyrolysis, fuel coking problems and fuel decomposition resulting in higher CO emissions. It can be seen from the present investigation, that the flame environment enhances the reduction of droplet size and smaller droplet size distribution, significantly altering the 136

159 mixing and entrainment due to high flame temperature, as well as faster fuel vaporization as a secondary effect of high temperatures consuming larger droplets. A significant change is not observed in the SMD and axial velocity for open and enclosed flame because of the addition of large amount of methane to sustain a stable VO flame in open conditions which results, in significant heat feedback in the near field region of the flame spray, thus neutralizing the influence of the insulated enclosure, effect that is observed in other cases. 137

160 Figure 5.1 Flame structure of an air blast atomizer 138

161 Flame Fuel Vapor Spray Boundary Air Droplet Boundary Layer Air Supply Air Supply Fuel Atomizer Figure 5.2 Vaporization of a typical droplet in an idealized spray flame 139

162 Figure 5.3 Schematic of the combustor experimental set-up 14

163 (a) (b) Pentagonal Enclosure Insulation Stable VO Flame 8 cm Φ=15 cm 46 cm Stainless Steel key stock bars to support the glass plates Figure 5.4 Photographic view of the enclosure (a) without and (b) with insulation and schematic of the top view and vertical view of the pentagonal enclosure 141

164 Figure 5.5 Airblast Injector Details 142

165 Fuel Heater Thermocouple Fuel Pump Liquid Fuel Combustor Pulse Dampener Fuel Reservoir Thermal Logger Figure 5.6 Schematic of Liquid fuel supply system. Pressure Regulator LFE Control Valve Air Water Traps / Filters MFC Methane Needle Valve MFM Check Valve Figure 5.7 Schematic of gaseous fuel air supply system. 143

166 Diffusion flame Exhaust Duct Cold Spray Diffusion flame torches Laser Beams passing through the spray Non-evaporating VO Spray Figure 5.8 Photographic representation of the Flare system 144

167 Figure 5.9 (a) Emissions Analyzer; (b) Emissions Measurement Traversing system 145

168 Figure 5.1 Schematic of the PDPA system 146

169 Axial velocity Radial velocity Figure 5.11 Experimental set-up of a PDPA system mounted on a 3-way traversing system 147

170 Traversing rail Direction of motion of the probe volume Y Z Origin of the X coordinate system at the injector exit Pentagonal Enclosure Probe volume Combustion chamber Figure 5.12 Plan view of the PDPA traversing mechanism in radial and axial coordinates 148

171 Receive r Probe Traversing system Transmitte r Probe Non evaporating VO Spray Combust or Figure 5.13 Photographic view of the PDPA system integrated with the combustor assembly 149

172 Figure 5.14 Photographic Image of the MikroScan 72v Infrared Camera 15

173 Axial Location (mm) Mean axial velocity (m/s) No swirl flow Figure 5.15 Mean axial velocity contour for cold spray without swirl 151

174 Axial Locaiton (mm) Mean axial velocity (m/s) Swirl flow Figure 5.16 Mean axial velocity contour for cold spray with swirl 152

175 Axial location (mm) Mean axial velocity (m/s) Swirl flow T f = o C Figure 5.17 Mean axial velocity contour for cold spray with swirl at T f = C 153

176 Axial location (mm) RMS axial Velocity (m/s) No swirl flow Figure 5.18 RMS axial velocity contour for cold spray without swirl 154

177 Axial location (mm) RMS axial velocity (m/s) Swirl flow Figure 5.19 RMS axial velocity contour for cold spray with swirl 155

178 Axial location (mm) RMS axial velocity (m/s) Swirl flow T f = o C Figure 5.2 RMS axial velocity contour for cold spray with swirl at T f = C 156

179 Axial location (mm) SMD (microns) No swirl flow Figure 5.21 SMD contour for cold spray without swirling air 157

180 Axial locations (mm) SMD (microns) Swirl flow Figure 5.22 SMD contour for cold spray with swirling air 158

181 Axial location (mm) SMD (microns) Swirl lfow T f = o C Figure 5.23 SMD contour for cold spray with swirling air at T f = C 159

182 Swirling air No swirling air 5 5 Mean axial velocity (m/s) Swirling air No swirling air RMS axial velocity (m/s) Figure 5.24 (a and b) Transverse profiles of mean axial velocity and RMS axial velocity for cold spray with and without swirl at Y = 5 mm 16

183 Swirling air No swirling air 5 5 Mean axial velocity (m/s) Swirling air No swirling air RMS axial velocity (m/s) Figure 5.24 (c and d) Transverse profiles of mean axial velocity and RMS axial velocity for cold spray with and without swirl at Y = 1 mm 161

184 Swirling air No swirling air 5 5 Mean axial velocity (m/s) Swirling air No swirling air RMS axial velocity (m/s) Figure 5.24 (e and f) Transverse profiles of mean axial velocity and RMS axial velocity for cold spray with and without swirl at Y = 15 mm 162

185 Swirling air No swirling air 5 5 Mean axial velocity (m/s) Swirling air No swirling air RMS axial velocity (m/s) Figure 5.24 (g and h) Transverse profiles of mean axial velocity and RMS axial velocity for cold spray with and without swirl at Y = 2 mm 163

186 Swirling air No swirling air 5 5 Mean axial velocity (m/s) Swirling air No swirling air RMS axial velocity (m/s) Figure 5.24 (i and j) Transverse profiles of mean axial velocity and RMS axial velocity for cold spray with and without swirl at Y = 25 mm 164

187 Swirling air No swirling air 5 5 Mean axial velocity (m/s) Swirling air No swirling air RMS axial velocity (m/s) Figure 5.24 (k and l) Transverse profiles of mean axial velocity and RMS axial velocity for cold spray with and without swirl at Y = 3 mm 165

188 Swirling air No swirling air 5 5 Mean axial velocity (m/s) Swirling air No swirling air RMS axial velocity (m/s) Figure 5.24 (m and n) Transverse profiles of mean axial velocity and RMS axial velocity for cold spray with and without swirl at Y = 35 mm 166

189 Swirling air No swirling air 5 5 Mean axial velocity (m/s) Swirling air No swirling air RMS axial velocity (m/s) Figure 5.24 (o and p) Transverse profiles of mean axial velocity and RMS axial velocity for cold spray with and without swirl at Y = 4 mm 167

190 Swirling air No swirling air 5 5 Mean axial velocity (m/s) Swirling air No swirling air RMS axial velocity (m/s) Figure 5.24 (q and r) Transverse profiles of mean axial velocity and RMS axial velocity for cold spray with and without swirl at Y = 45 mm 168

191 Swirling air No swirling air 5 5 Mean axial velocity (m/s) Swirling air No swirling air RMS axial velocity (m/s) Figure 5.24 (s and t) Transverse profiles of mean axial velocity and RMS axial velocity for cold spray with and without swirl at Y = 5 mm 169

192 Swirling air No swirling air 5 5 Mean axial velocity (m/s) Swirling air No swirling air RMS axial velocity (m/s) Figure 5.24 (u and v) Transverse profiles of mean axial velocity and RMS axial velocity for cold spray with and without swirl at Y = 6 mm 17

193 Swirling air No swirling air 5 5 Mean axial velocity (m/s) Swirling air No swirling air RMS axial velocity (m/s) Figure 5.24 (w and x) Transverse profiles of mean axial velocity and RMS axial velocity for cold spray with and without swirl at Y = 7 mm. 171

194 Swirling air No swirling air 5 5 Mean axial velocity (m/s) Swirling air No swirling air RMS axial velocity (m/s) Figure 5.24 (y and z) Transverse profiles of mean axial velocity and RMS axial velocity for cold spray with and without swirl at Y = 75 mm. 172

195 Swirling air; unheated VO 5 5 Swirling air, T f = o C Mean axial velocity (m/s) Swirling air, unheated VO 25 Swirling air, T f = o C 25 RMS axial velocity (m/s) Figure 5.25 (a and b) Transverse profiles of mean axial velocity and RMS axial velocity for swirling air cold spray; unheated and VO at T f = C Y = 5 mm 173

196 Swirling air; unheated VO 5 5 Swirling air, T f = o C Mean axial velocity (m/s) Swirling air, unheated VO 25 Swirling air, T f = o C 25 RMS axial velocity (m/s) Figure 5.25 (c and d) Transverse profiles of mean axial velocity and RMS axial velocity for swirling air cold spray; unheated and VO at T f = C Y = 1 mm 174

197 Swirling air; unheated VO 5 5 Swirling air, T f = o C Mean axial velocity (m/s) Swirling air, unheated VO 25 Swirling air, T f = o C 25 RMS axial velocity (m/s) Figure 5.25 (e and f) Transverse profiles of mean axial velocity and RMS axial velocity for swirling air cold spray; unheated and VO at T f = C Y = 15 mm 175

198 Swirling air; unheated VO 5 5 Swirling air, T f = o C Mean axial velocity (m/s) Swirling air, unheated VO 25 Swirling air, T f = o C 25 RMS axial velocity (m/s) Figure 5.25 (g and h) Transverse profiles of mean axial velocity and RMS axial velocity for swirling air cold spray; unheated and VO at T f = C Y = 2 mm 176

199 Swirling air; unheated VO 5 5 Swirling air, T f = o C Mean axial velocity (m/s) Swirling air, unheated VO 25 Swirling air, T f = o C 25 RMS axial velocity (m/s) Figure 5.25 (i and j) Transverse profiles of mean axial velocity and RMS axial velocity for swirling air cold spray; unheated and VO at T f = C Y = 25 mm 177

200 Swirling air; unheated VO 5 5 Swirling air, T f = o C Mean axial velocity (m/s) Swirling air, unheated VO 25 Swirling air, T f = o C 25 RMS axial velocity (m/s) Figure 5.25 (k and l) Transverse profiles of mean axial velocity and RMS axial velocity for swirling air cold spray; unheated and VO at T f = C Y = 3 mm 178

201 Swirling air; unheated VO 5 5 Swirling air, T f = o C Mean axial velocity (m/s) Swirling air, unheated VO 25 Swirling air, T f = o C 25 RMS axial velocity (m/s) Figure 5.25 (m and n) Transverse profiles of mean axial velocity and RMS axial velocity for swirling air cold spray; unheated and VO at T f = C Y = 35 mm 179

202 Swirling air; unheated VO 5 5 Swirling air, T f = o C Mean axial velocity (m/s) Swirling air, unheated VO 25 Swirling air, T f = o C 25 RMS axial velocity (m/s) Figure 5.25 (o and p) Transverse profiles of mean axial velocity and RMS axial velocity for swirling air cold spray; unheated and VO at T f = C Y = 4 mm 18

203 Swirling air; unheated VO 5 5 Swirling air, T f = o C Mean axial velocity (m/s) Swirling air, unheated VO 25 Swirling air, T f = o C 25 RMS axial velocity (m/s) Figure 5.25 (q and r) Transverse profiles of mean axial velocity and RMS axial velocity for swirling air cold spray; unheated and VO at T f = C Y = 5 mm 181

204 Swirling air; unheated VO 5 5 Swirling air, T f = o C Mean axial velocity (m/s) Swirling air, unheated VO 25 Swirling air, T f = o C 25 RMS axial velocity (m/s) Figure 5.25 (s and t) Transverse profiles of mean axial velocity and RMS axial velocity for swirling air cold spray; unheated and VO at T f = C Y = 6 mm 182

205 Swirling air; unheated VO 5 5 Swirling air, T f = o C Mean axial velocity (m/s) Swirling air, unheated VO 25 Swirling air, T f = o C 25 RMS axial velocity (m/s) Figure 5.25 (u and v) Transverse profiles of mean axial velocity and RMS axial velocity for swirling air cold spray; unheated and VO at T f = C Y = 7 mm 183

206 Swirling air; unheated VO 5 5 Swirling air, T f = o C Mean axial velocity (m/s) Swirling air, unheated VO 25 Swirling air, T f = o C 25 RMS axial velocity (m/s) Figure 5.25 (w and x) Transverse profiles of mean axial velocity and RMS axial velocity for swirling air cold spray; unheated and VO at T f = C Y = 75 mm 184

207 Swirling air 9 No swirling air 8 8 Sauter mean diameter (μm) Swirling air 9 No swirling air 8 8 Sauter mean diameter (μm) Figure 5.26 (a and b) Transverse profiles of SMD for cold spray with and without swirl at Y = 5 mm and 1 mm 185

208 Swirling air 9 No swirling air 8 8 Sauter mean diameter (μm) Swirling air 9 No swirling air 8 8 Sauter mean diameter (μm) Figure 5.26 (c and d) Transverse profiles of SMD for cold spray with and without swirl at Y = 15 mm and 2 mm 186

209 Swirling air 9 No swirling air 8 8 Sauter mean diameter (μm) Swirling air 9 No swirling air 8 8 Sauter mean diameter (μm) Figure 5.26 (e and f) Transverse profiles of SMD for cold spray with and without swirl at Y = 25 mm and 3 mm 187

210 Swirling air 9 No swirling air 8 8 Sauter mean diameter (μm) Swirling air 9 No swirling air 8 8 Sauter mean diameter (μm) Figure 5.26 (g and h) Transverse profiles of SMD for cold spray with and without swirl at Y = 35 mm and 4 mm 188

211 Swirling air 9 No swirling air 8 8 Sauter mean diameter (μm) Swirling air 9 No swirling air 8 8 Sauter mean diameter (μm) Figure 5.26 (i and j) Transverse profiles of SMD for cold spray with and without swirl at Y = 45 mm and 5 mm 189

212 Swirling air 9 No swirling air 8 8 Sauter mean diameter (μm) Swirling air 9 No swirling air 8 8 Sauter mean diameter (μm) Figure 5.26 (k and l) Transverse profiles of SMD for cold spray with and without swirl at Y = 6 mm and 7 mm 19

213 Swirling air 9 No swirling air 8 8 Sauter mean diameter (μm) Figure 5.26 (m) Transverse profiles of SMD for cold spray with and without swirl at Y = 75mm 191

214 Swirling air, unheated VO 9 Swirling air, T f = o C 8 8 Sauter mean diameter (μm) Swirling air, unheated VO 9 Swirling air, T f = o C 8 8 Sauter mean diameter (μm) Figure 5.27 (a and b) Transverse profiles of SMD for swirling air cold spray; unheated and VO at T f = C Y = 5 mm and Y = 1 mm 192

215 Swirling air, unheated VO 9 Swirling air, T f = o C 8 8 Sauter mean diameter (μm) Swirling air, unheated VO 9 Swirling air, T f = o C 8 8 Sauter mean diameter (μm) Figure 5.27 (c and d) Transverse profiles of SMD for swirling air cold spray; unheated and VO at T f = C Y = 15 mm and Y = 2 mm 193

216 Swirling air, unheated VO 9 Swirling air, T f = o C 8 8 Sauter mean diameter (μm) Swirling air, unheated VO 9 Swirling air, T f = o C 8 8 Sauter mean diameter (μm) Figure 5.27 (e and f) Transverse profiles of SMD for swirling air cold spray; unheated and VO at T f = C Y = 25 mm and Y = 3 mm 194

217 Swirling air, unheated VO 9 Swirling air, T f = o C 8 8 Sauter mean diameter (μm) Swirling air, unheated VO 9 Swirling air, T f = o C 8 8 Sauter mean diameter (μm) Figure 5.27 (g and h) Transverse profiles of SMD for swirling air cold spray; unheated and VO at T f = C Y = 35 mm and Y = 4 mm 195

218 Swirling air, unheated VO 9 Swirling air, T f = o C 8 8 Sauter mean diameter (μm) Swirling air, unheated VO 9 Swirling air, T f = o C 8 8 Sauter mean diameter (μm) Figure 5.27 (i and j) Transverse profiles of SMD for swirling air cold spray; unheated and VO at T f = C Y = 5 mm and Y = 6 mm 196

219 Swirling air, unheated VO 9 Swirling air, T f = o C 8 8 Sauter mean diameter (μm) Swirling air, unheated VO 9 Swirling air, T f = o C 8 8 Sauter mean diameter (μm) Figure 5.27 (k and l) Transverse profiles of SMD for swirling air cold spray; unheated and VO at T f = C Y = 7 mm and Y = 75 mm 197

220 Swirling air, SMD = 64 μm 1 3 No swirling air, SMD = 76 μm Diameter Count Diameter (microns) Swirling air, SMD = 5 μm 1 3 No swirling air, SMD = 52 μm Diameter Count Diameter (microns) Figure 5.28 (a and b) Droplet Distribution Profile for cold spray with and without swirl at Y = 5 mm, X = mm and Y = 1 mm, X = mm 198

221 Swirling air, SMD = 37 μm 1 3 No swirling air, SMD = 46 μm DiameterCount Diameter (microns) Swirling air, SMD = 36 μm 1 3 No swirling air, SMD = 44 μm D iameter Count Diameter (microns) Figure 5.28 (c and d) Droplet Distribution Profile for cold spray with and without swirl at Y = 4 mm, X = mm and Y = 6 mm, X = mm 199

222 Axial location (mm) Mean axial velocity (m/s) Flame Figure 5.29 Mean axial velocity contour for flame spray 2

223 Axial location (mm) Mean axial velocity(m/s) Flame T f = o C Figure 5.3 Mean axial velocity contour for flame spray for VO at C 21

224 Axial location (mm) RMS axial location (m/s) Flame Figure 5.31 RMS axial velocity contour for flame spray 22

225 Axial location (mm) RMS axial velocity (m/s) Flame T f = o C Figure 5.32 RMS axial velocity contour for flame spray for VO at C 23

226 Axial location (mm) SMD (microns) Flame Figure 5.33 SMD contour for flame spray 24

227 Axial location (mm) SMD (microns) Flame T f = o C Figure 5.34 SMD contour for flame spray for VO at C 25

228 Flame Swirl spray Mean axial velocity (m/s) Flame Swirl spray RMS axial velocity (m/s) Figure 5.35 (a and b) Transverse profiles of axial mean velocity and RMS axial velocity for flame spray and cold spray at Y = 5 mm 26

229 Flame Swirl spray Mean Axial Velocity (m/s) Flame Swirl spray RMS axial velocity (m/s) Figure 5.35 (c and d) Transverse profiles of axial mean velocity and RMS axial velocity for flame spray and cold spray at Y = 1 mm 27

230 Flame Swirl spray Mean axial velocity (m/s) Flame Swirl spray RMS axial velocity (m/s) Figure 5.35 (e and f) Transverse profiles of axial mean velocity and RMS axial velocity for flame spray and cold spray at Y = 15 mm 28

231 Flame Swirl spray Mean axial velocity (m/s) Flame Swirl spray RMS axial velocity (m/s) Figure 5.35 (g and h) Transverse profiles of axial mean velocity and RMS axial velocity for flame spray and cold spray at Y = 2 mm 29

232 Flame Swirl spray Mean axial velocity (m/s) Flame Swirl spray RMS axial velocity (m/s) Figure 5.35 (i and j) Transverse profiles of axial mean velocity and RMS axial velocity for flame spray and cold spray at Y = 25 mm 21

233 Flame Swirl spray Mean axial velocity (m/s) Flame Swirl spray RMS axial velocity (m/s) Figure 5.35 (k and l) Transverse profiles of axial mean velocity and RMS axial velocity for flame spray and cold spray at Y = 3 mm 211

234 Flame Swirl spray Mean axial velocity (m/s) Flame Swirl spray RMS axial velocity (m/s) Figure 5.35 (m and n) Transverse profiles of axial mean velocity and RMS axial velocity for flame spray and cold spray at Y = 35 mm 212

235 Flame Swirl spray Mean axial velocity (m/s) Flame Swirl spray RMS axial velocity (m/s) Figure 5.35 (o and p) Transverse profiles of axial mean velocity and RMS axial velocity for flame spray and cold spray at Y = 4 mm 213

236 Flame, T f = o C Swirl spray, T f = o C Mean axial velocity (m/s) Flame, T f = o C Swirl spray, T f = o C RMS axial velcoity (m/s) Figure 5.36 (a and b) Transverse profiles of mean axial velocity and RMS axial velocity for flame spray and cold spray for T f = C at Y = 5 mm. 214

237 Flame, T f = o C Swirl spray, T f = o Mean axial velocity (m/s) Flame, T f = o C Swirl spray, T f = o RMS axial velcoity (m/s) Figure 5.36 (c and d) Transverse profiles of mean axial velocity and RMS axial velocity for flame spray and cold spray for T f = C at Y = 1 mm. 215

238 Flame, T f = o C Swirl spray, T f = o Mean axial velocity (m/s) Flame, T f = o C Swirl spray, T f = o RMS axial velcoity (m/s) Figure 5.36 (e and f) Transverse profiles of mean axial velocity and RMS axial velocity for flame spray and cold spray for T f = C at Y = 15 mm. 216

239 Flame, T f = o C Swirl spray, T f = o Mean axial velocity (m/s) Flame, T f = o C Swirl spray, T f = o RMS axial velcoity (m/s) Figure 5.36 (g and h) Transverse profiles of mean axial velocity and RMS axial velocity for flame spray and cold spray for T f = C at Y = 2 mm. 217

240 Flame, T f = o C Swirl spray, T f = o Mean axial velocity (m/s) Flame, T f = o C Swirl spray, T f = o RMS axial velcoity (m/s) Figure 5.36(i and j) Transverse profiles of mean axial velocity and RMS axial velocity for flame spray and cold spray for T f = C at Y = 25 mm. 218

241 Flame, T f = o C Swirl spray, T f = o Mean axial velocity (m/s) Flame, T f = o C Swirl spray, T f = o RMS axial velcoity (m/s) Figure 5.36 (k and l) Transverse profiles of mean axial velocity and RMS axial velocity for flame spray and cold spray for T f = C C at Y = 3 mm. 219

242 Flame, T f = o C Cold spray, T f = o C Mean axial velocity (m/s) Flame, T f = o C Cold spray, T f = o C RMS axial velcoity (m/s) Figure 5.36 (m and n) Transverse profiles of mean axial velocity and RMS axial velocity for flame spray and cold spray for T f = C at Y = 35 mm. 22

243 Flame, T f = o C Cold spray, T f = o C Mean axial velocity (m/s) Flame, T f = o C Cold spray, T f = o C RMS axial velcoity (m/s) Figure 5.36 (o and p) Transverse profiles of mean axial velocity and RMS axial velocity for flame spray and cold spray for T f = C C at Y = 4 mm. 221

244 Flame Swirl spray Sauter mean diameter (μm) Flame Swirl spray Sauter mean diameter (μm) Figure 5.37 (a and b) Transverse profiles of SMD for flame spray and cold spray at Y = 5 mm and 1 mm. 222

245 Flame Swirl spray Sauter mean diameter (μm) Flame Swirl spray Sauter mean diameter (μm) Figure 5.37 (c and d) Transverse profiles of SMD for flame spray and cold spray at Y = 15 mm and 2 mm. 223

246 Flame Swirl spray Sauter mean diameter (μm) Flame Swirl spray Sauter mean diameter (μm) Figure 5.37 (e and f) Transverse profiles of SMD for flame spray and cold spray at Y = 25 mm and 3 mm 224

247 Flame Swirl spray Sauter mean diameter (μm) Flame Swirl spray Sauter mean diameter (μm) Figure 5.37 (g and h) Transverse profiles of SMD for flame spray and cold spray at Y = 35 mm and 4 mm 225

248 Flame, T f = o C Swirl spray, T f = o Sauter mean diameter (μm) Flame, T f = o C Swirl spray, T f = o Sauter mean diameter (μm) Figure 5.38(a and b) Transverse profiles of SMD for flame spray and cold spray at Y = 5 mm and 1 mm for T f = C. 226

249 Flame, T f = o C Swirl spray, T f = o C Sauter mean diameter (μm) Flame, T f = o C Swirl spray, T f = o C Sauter mean diameter (μm) Figure 5.38 (c and d) Transverse profiles of SMD for flame spray and cold spray at Y = 15 mm and 2 mm for T f = C. 227

250 Flame, T f = o C Swirl spray, T f = o C Sauter mean diameter (μm) Flame, T f = o C Swirl spray, T f = o C Sauter mean diameter (μm) Figure 5.38 (e and f) Transverse profiles of SMD for flame spray and cold spray at Y = 25 mm and 3 mm for T f = C. 228

251 Flame, T f = o C Swirl spray, T f = o C Sauter mean diameter (μm) Flame, T f = o C 8 Swirl spray, T f = o C 8 Sauter mean diameter (μm) Figure 5.38 (g and h) Transverse profiles of SMD for flame spray and cold spray at Y = 35 mm and 4 mm for T f = C. 229

252 Flame, SMD = 26μm 1 3 Cold spray,smd = 64μm Diameter Count Diameter (microns) Flame, SMD = 19 μm 1 3 Cold spray, SMD = 51 μm Diameter Count Diameter (microns) Figure 5.39 (a and b) Droplet Distribution Profile for flame spray and cold spray at Y = 5 mm, X = mm and Y = 1 mm, X = mm 23

253 Flame, SMD = 17 μm 1 3 Cold spray, SMD = 38 μm Diameter Count Diameter (microns) Flame, SMD = 19μm 1 3 Cold spray, SMD = 37μm Diameter Count Diameter (microns) Figure 5.39 (c and d) Droplet Distribution Profile for flame spray and cold spray at Y = 35 mm, X = mm and Y = 4 mm, X = mm 231

254 Flame, T f =o C, SMD = 28μm 1 3 Cold Spray,T f = o C, SMD = 37μm Diameter Count Diameter (microns) Flame,T f =o C, SMD = 22μm 1 3 Cold Spray,T f = o C, SMD = 33μm Diameter Count Diameter (microns) Figure 5.4 (a and b) Droplet Distribution Profile for flame spray and cold spray for T f = C at Y = 5 mm, X = mm and Y = 1 mm, X = mm 232

255 Flame,T = f o C, SMD = 13μm 1 3 Cold spray,t f = o C, SMD = 3μm Diameter Count Diameter(microns) Flame,T f =o C, SMD = 14μm 1 3 Cold spray,t f = o C, SMD =29μm Diameter Count Diameter (microns) Figure 5.4 (c and d) Droplet Distribution Profile for flame spray and cold spray for T f = C at Y = 35 mm, X = mm and Y = 4 mm, X = mm. 233

256 Enclosed flame, T f = o C Enclosed flame, T f =15 o C Mean axial velocity (m/s) Enclosed flame, T f = o C Enclosed flame, T f =15 o C RMS axial velocity(m/s) Figure 5.41 (a and b) Transverse profiles of mean axial velocity and RMS axial velocity for enclosed flame for T f = C and T f = 15 C at Y = 5 mm 234

257 Enclosed flame, T f = o C Enclosed flame, T f =15 o C Mean axial velocity (m/s) Enclosed flame, T f = o C Enclosed flame, T f = 15 o C RMS axial velocity(m/s) Figure 5.41 (c and d) Transverse profiles of mean axial velocity and RMS axial velocity for enclosed flame for T f = C and T f = 15 C at Y = 1 mm 235

258 Enclosed flame, T f = o C Enclosed flame, T f =15 o C Mean axial velocity (m/s) Enclosed flame, T f = o C Enclosed flame, T f = 15 o C RMS axial velocity(m/s) Figure 5.41 (e and f) Transverse profiles of mean axial velocity and RMS axial velocity for enclosed flame for T f = C and T f = 15 C at Y = 15 mm 236

259 Enclosed flame, T f = o C Enclosed flame, T f = 15 o C Mean axial velocity (m/s) Enclosed flame, T f = o C Enclosed flame, T f = 15 o C RMS axial velocity(m/s) Figure 5.41 (g and h) Transverse profiles of mean axial velocity and RMS axial velocity for enclosed flame for T f = C and T f = 15 C at Y = 2 mm 237

260 Enclosed flame, T f = o C Enclosed flame, T f =15 o C Mean axial velocity (m/s) Enclosed flame, T f = o C Enclosed flame, T f =15 o C RMS axial velocity(m/s) Figure 5.41 (i and j) Transverse profiles of mean axial velocity and RMS axial velocity for enclosed flame for T f = C and T f = 15 C at Y = 25 mm 238

261 Enclosed flame, T f = o C Enclosed flame, T f = 15 o C Mean axial velocity (m/s) Enclosed flame, T f = o C Enclosed flame, T f = 15 o C RMS axial velocity(m/s) Figure 5.41 (k and l) Transverse profiles of mean axial velocity and RMS axial velocity for enclosed flame for T f = C and T f = 15 C at Y = 3 mm 239

262 Enclosed flame, T f = o C Enclosed flame, T f =15 o C Mean axial velocity (m/s) Enclosed flame, T f = o C Enclosed flame, T f = 15 o C RMS axial velocity(m/s) Figure 5.41 (m and n) Transverse profiles of mean axial velocity and RMS axial velocity for enclosed flame for T f = C and T f = 15 C at Y = 35 mm 24

263 Sauter mean diameter (microns) Enclosed flame, T f = o C Enclosed Flame, T f = 15 o C Sauter mean diameter (microns) Enclosed flame, T f = o C Enclosed Flame, T f = 15 o C Figure 5.42 (a and b) Transverse profiles of SMD for enclosed flame for T f = C and T f = 15 C at Y = 5 mm and Y = 1 mm. 241

264 Sauter mean diameter (microns) Enclosed flame, T f = o C Enclosed Flame, T f =15 o C Sauter mean diameter (microns) Enclosed flame, T f = o C Enclosed Flame, T f =15 o C Figure 5.42(c and d) Transverse profiles of SMD for enclosed flame for T f = C and T f = 15 C at Y = 15 mm and Y = 2 mm. 242

265 Sauter mean diameter (microns) Enclosed flame, T f = o C Enclosed Flame, T f = 15 o C Sauter mean diameter (microns) Enclosed flame, T f = o C Enclosed Flame, T f =15 o C Figure 5.42 (e and f) Transverse profiles of SMD for enclosed flame for T f = C and T f = 15 C at Y = 25 mm and Y = 3 mm. 243

266 Sauter mean diameter (microns) Enclosed flame, T f = o C Enclosed Flame, T f = 15 o C Figure 5.42 (g) Transverse profiles of SMD for enclosed flame for T f = C and T f = 15 C at Y = 35mm. 244

267 T f = o C, SMD = 32μm 1 3 T f =15 o C, SMD = 28μm DiameterCount Diameter (microns) T f = o C, SMD = 27 μm 1 3 T f = 15 o C, SMD = 26μm D iameter Count Diameter (microns) Figure 5.43(a and b) Droplet Distribution Profile for Enclose Flame for T f = C and T f = 15 C at Y = 5 mm, X = mm and Y = 1 mm, X = mm 245

268 T f = o C, SMD = 18 μm 1 3 T f =15 o C, SMD = 19μm DiameterCount Diameter(microns) T f = o C, SMD = 2 μm 1 3 T f = 15 o C, SMD = 17 μm D iameter Count Diameter (microns) Figure 5.43 (c and d) Droplet Distribution Profile for Enclosed Flame for T f = C and T f = 15 C at Y = 3 mm, X = mm and Y = 35 mm, X = mm 246

269 Y=43.5cm,T f = o C Y=41.cm,T f = o C Y=38.cm,T f = o C Y=43.5cm,T f =15 o C Y=41.cm,T f =15 o C Y=38.cm,T f =15 o C CO (ppm) Transverse location (cm) Figure 44. Transverse profiles of CO for enclosed VO flame at C and 15 C 247

270 NOx (ppm) Y=43.5cm,T f = o C Y=41.cm,T f = o C Y=38.cm,T f = o C 75 Y=43.5cm,T f =15 o C Y=41.cm,T f =15 o C 75 Y=38.cm,T f =15 o C Transverse distance (cm) 4 6 Figure 45. Transverse profiles of NOx for enclosed VO flame at C and 15 C 248

271 . C Axial location (mm) Min C Max C Avg C Range C Line Line Line Line Line Figure 5.46.IR Image of the Enclosure surface temperature 249