EXPERIMENTAL VALIDATION OF A HOT GAS TURBINE PARTICLE DEPOSITION FACILITY. A Thesis. By: Christopher Stephen Smith, B.S. The Ohio State University

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1 EXPERIMENTAL VALIDATION OF A HOT GAS TURBINE PARTICLE DEPOSITION FACILITY A Thesis Presented in Partial Fulfillment of the Requirements for the Degree of Master of Science in the Graduate School of The Ohio State University By: Christopher Stephen Smith, B.S. Graduate Program in Aeronautical and Astronautical Engineering The Ohio State University 2010 Master s Examination Committee: Dr. Jeffrey P. Bons, Advisor Dr. James Gregory Dr. Ali Ameri

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3 ABSTRACT A new turbine research facility at The Ohio State University Aeronautical and Astronautical Research Lab has been constructed. The purpose of this facility is to recreate deposits on the surface of actual aero-engine Nozzle Guide Vane (NGV) hardware in an environment similar to what the hardware was designed for. This new facility is called the Turbine Reacting Flow Rig (TuRFR). The TuRFR provides air at temperatures up to 1200 C and at inlet Mach numbers comparable to those found in an actual turbine (~0.1). Several validation studies have been undertaken which prove the capabilities of the TuRFR. These studies show that the temperature entering the NGV cascade is uniform, and they demonstrate the capability to provide film cooling air to the NGV cascade at flow rates and density ratios comparable to the NGV design. Deposition patterns have also been created on the surface of actual NGV hardware. Deposition was created at different flow temperatures, and it was found that deposition levels decrease with decreasing gas temperature. Also, film cooling levels were varied from 0% film cooling to 4% film cooling. It was found that with increased rates of film cooling deposition decreased. With the TuRFR capabilities demonstrated, research on the effects of deposition on the aerodynamic performance of the NGV hardware was conducted. ii

4 Integrated non-dimensional total pressure loss values were calculated in an exit Re c range of 0.2x10 6 to 1.7x10 6 for a deposit roughened NGV cascade and a smooth cascade. The data suggests that deposition causes increased losses across the NGV cascade and possibly earlier transition. The data also suggests a possible region of separated flow in the NGV cascade which disappears at higher exit Reynolds numbers. These results are similar to those found in the literature. iii

5 ACKNOWLEDGMENTS I would like to take this opportunity to thank my advisor Dr. Jeffrey Bons for his patience, guidance, and continual support during this endeavor. It has truly been privilege learning from him. I would also like to acknowledge the help of Brett Barker, Carey Clum, and Josh Webb. Without their knowledge and willingness to solve problems the TuRFR would not be operational today. I would also like to thank the staff of AARL, especially Ken Copely, Ken Fout, Jeff Barton and Cathy Mitchell for their undying patience and help. I would also like to express my appreciation to Dr. James Gregory and Dr. Ali Ameri for serving on my graduate committee. Finally I would like to thank my friends and family for their support and encouragement along the way. iv

6 VITA July 15, Born in Texas City, Texas August 2007.B.S. Mechanical Engineering, The University of Texas at Austin 2007 Present Graduate Research Associate, Department of Aerospace Engineering The Ohio State University PUBLICATIONS 1. Smith, C., Barker, B., Clum, C., Bons, J., Deposition in a Turbine Cascade with Combusting Flow, to be presented at the 2010 Turbo Expo in Glasgow, Scotland. June 2010, paper # GT FIELDS OF STUDY Major Fields of Study: Aeronautical and Astronautical Engineering v

7 TABLE OF CONTENTS ABSTRACT... ii ACKNOWLEDGMENTS...iv VITA... v FIELDS OF STUDY... v TABLE OF CONTENTS...vi LIST OF FIGURES...ix LIST OF TABLES... xii LIST OF EQUATIONS... xiii NOMENCLATURE... xiv CHAPTER INTRODUCTION Background Literature Review Other Deposition Research...6 CHAPTER TuRFR DESCRIPTION AND VAILDATION STUDIES vi

8 2.1 Primary Flow Path Description Particulate Feed Sub-System TuRFR Validation Studies NGV Inlet Temperature Survey Film Cooling Validation Survey Exit Total Pressure Surveys CHAPTER ASH PARTICULATE Description of Particulate CHAPTER DEPOSITION TESTING RESULTS Deposition Testing Temperature Variation Film Cooling Variation Surface Roughness Measurements Deposit Structure and Chemical Composition CHAPTER AERODYNAMIC PERFORMANCE ASSESSMENT Aerodynamic Performance Assessment Background vii

9 5.1.1 Total Pressure and Exit Temperature Measurement Setup Inlet Total Pressure Measurement Aerodynamic Performance Results Rough Aerodynamic Performance Smooth Aerodynamic Performance Discussion of Results CHAPTER Conclusions Future Work REFERENCES APPENDIX A: UNCERTAINTY ANALYSIS OF THE γ CALCULATION viii

10 LIST OF FIGURES Figure 1 - Leading edge of turbine nozzle guide vanes exposed to high volcanic ash concentrations. Note the effects of increased surface heat transfer caused by clogged film cooling holes [2]...3 Figure 2 - Turbine Accelerated Deposition Facility (TADF) at Brigham Young University [11]...7 Figure 3 - TuRFR schematic showing main flow path Figure 4 - Flameholder inside combustion section of TuRFR Figure 5-3D cutaway view of upper section of TuRFR showing measurement locations Figure 6 - Top view of NGV cascade and film cooling reservoir showing film cooling temperature measurement locations Figure 7-3D CAD schematic of particulate feeder. Pressure equalization tube is not shown in the diagram Figure 8 - NGV cascade inlet temperature survey Figure 9 - Dimensionless temperature at vane exit with 4% film heating, red box shows approximate location of measurement plane ix

11 Figure 10 - Total pressure measurements at the exit of the NGV cascade with 0% film heating applied: a) Non-dimensional total pressure (P TOT /P amb ) and b) lines of total pressure loss at various span locations Figure 11 - Non-dimensional total pressure loss [(P TOTin P TOT )/0.5ρu 2 ] collected 13% true chord downstream of NGV exit plane indicating passage periodicity Figure 12 Total pressure measurements at the exit of the NGV cascade with 4% film heating applied: a) Non-dimensional total pressure (P TOT /P amb ) and b) lines of total pressure loss at various span locations Figure 13 - Particle size distribution results from Coulter counter for Bituminous coal fly ash Figure 14 - SEM image of Bituminous coal fly ash Figure 15 - Post test images of deposit formation for Low Temperature test, No Film Cooling test, and Film Cooling test Figure 16 - Representative surface roughness traces and roughness regions for deposition formed during NO Film Cooling test. CAD image of NGV doublet is not to scale Figure 17 - SEM photo highlighting deposit structure. The vane surface is not indicated in the image, but it is below the deposit flake Figure 18 - EDS element maps of deposit flake highlighting Cu, Fe, Al, Si, and O layers. SEM photo highlighting deposit structure. The vane surface is not indicated in the images, but it is below the deposit flake in all of the images Figure 19 - P totin vs. curve showing the experimental data, the curve fit, and the equation of the fit. m in the fit equation stands for inlet mass flow x

12 Figure 20 - Schematic of TuRFR used in the transit time estimates Figure 21 - Photo of exit plane indicating the approximate location and direction of the probe head travel Figure 22 - Photograph of NGV cascade surface with and without deposition Figure 23 - Non-Dimensional Total Pressure Loss vs. Position for Re c exit 0.5x Figure 24 γ vs. Re c exit for Blade 3 of the NGV cascade for the rough cases only Figure 25 - Total skin friction coefficient of sand roughened plate at zero incidence angle Figure 26 - γ vs. Re c exit for Blade 3 of the NGV cascade Figure 27 - γ vs. Re c exit for Blade 2 of the NGV cascade Figure 28 - Yarn tuft test showing that separation is not occurring in the Reynolds number range considered Figure 29 - Area averaged total pressure loss coefficient vs. Exit Reynolds number from Boyle and Senyitko [9] (filled symbols) with avearge γ between data presented in Figure 26 and Figure 27 (open symbols) xi

13 LIST OF TABLES Table 1 - Test conditions for TuRFR validation studies Table 2 - Particulate mass mean diameter and bituminous coal ash molecular composition Table 3 - Deposition testing conditions Table 4 - Operating conditions for deposition creation and surface characteristics of the NGV cascade before and after deposition Table 5 - Exit Reynolds numbers and gamma for Blade 2 and Blade 3 of the NGV cascade Table 6 - Typical values of the N and γ uncertainties using the error propagation analysis Table 7 - Values used in the γ variance determination xii

14 LIST OF EQUATIONS Equation Equation Equation Equation Equation Equation Equation Equation Equation Equation Equation xiii

15 NOMENCLATURE A * Reference cross sectional area required to reach sonic velocity A in NGV inlet cross sectional area A vs C u2 Cross sectional area of view section of TuRFR Absolute tangential flow velocity in turbine rotor passage C D Drag Coefficient Cu CAD EDS L LDV Chemical Symbol for Copper Computer Aided Drawing Energy dispersive spectroscopy Total width of wake used in the γ calculation Laser Doppler Velocimetry M guess Mach number at NGV inlet guess M vs Mach number in view section of TuRFR NGV P amb P inlet Nozzle guide vane Ambient pressure Average total pressure at the cascade inlet P op P pipe Pressure in main air line just downstream of the orifice plate flow meters Pressure in the main air line xiv

16 P TOT meas P TOT Measured total pressure at the exit Total pressure at NGV exit P TOTin PSP R Ra Re Re c Average total pressure at NGV inlet Pressure Sensitive Paint Specific gas constant Average relative roughness Reynolds number Reynolds number based on the true chord of the NGV Re cx Reynolds number based on the axial chord of the NGV Re c exit Exit Reynolds number based on the true chord of the NGV Re p SEM Reynolds number of flow around a spherical particle Scanning electron microscope T ex Temperature of air exiting NGV cascade Average inlet temperature T FC T meas Film cooling temperature Measured temperature TADF TuRFR Turbine Accelerated Deposition Facility Turbine Reacting Flow Rig U 2 Turbine rotor tangential velocity Ideal shaft work done by turbine xv

17 X XRD Y a c Pitchwise grid dimension X-ray diffraction Spanwise grid dimension Speed of sound in TuRFR during operation NGV true chord (6.8cm est.) c x d p NGV axial chord (3.6cm est.) Diameter of coal ash particle dx k Spacing between measurement locations of P TOT meas Ratio of specific heats Mass flow of fluid ppmw-hr t Parts per million by weight hour Total transit time for air in TuRFR to go from the orifice plates to the NGV exit plane u exit Flow velocity at the NGV exit u vs Flow velocity in view section of TuRFR w ax Axial flow velocity in a turbine cascade x Δg τ γ ζ t Location along main air line being considered for the liner drop in P pipe Total head loss Time constant of particle reaction to abrupt change in gas flow Integrated Non-Dimensional Total Pressure Loss Total pressure loss coefficient xvi

18 µ Dynamic Viscosity ν Kinematic Viscosity ρ g ρ p Density of gas flowing around particle Density of particle ρ vs Fluid static density in view section of TuRFR xvii

19 CHAPTER 1 INTRODUCTION 1.1 Background With the decreasing availability of traditional oil based jet engine or power generation fuels, modern engineers are turning their attention to alternative fuels. Some of these alternative fuels are based on biomasses, such as grass or organic waste. Other fuels are based on a more abundant source: coal. Improvements in the direct liquefaction and gasification of coal are bringing its use as an alternative fuel for jet propulsion and power generation applications to the forefront of possibility. Political pressures are also forcing oil dependent nations to rely on domestic sources of energy, and among the most abundant sources of energy in some countries is coal. The many advocates of using coal as an alternative fuel cite its relatively low cost and availability as the primary reasons for converting to coal energy. However, there are opponents who have justifiable reasons not to use coal as an alternative energy source. Among the arguments against coal usage are the negative impacts on the environment from coal combustion and coal acquisition. Other reasons coal is having trouble gaining momentum in its fight to become the fuel of choice is the negative effect of coal combustion products on power generation hardware. 1

20 When coal is burned it can sometimes produce airborne particulate known as fly ash. Often coal is used in steam power generation plants as the fuel source for steam generation. The fly ash produced by the combustion of coal, deposits onto the walls of the boiler system and degrades the system over time. Periodically the entire plant must be shut down and the boiler over-hauled in order to lengthen the life of the power generation system and to protect the workers from an unexpected catastrophic failure. If coal is used as an alternative fuel in a gas turbine the combustion products could deposit onto the surface of the turbine hardware. Removing deposition on turbine hardware requires weeks of intense labor to disassemble the turbine, clean the components, and reassemble the turbine. Not to mention the risks to personnel and equipment safety. Understanding how deposition forms on the surface of turbine hardware can help to reduce it, therefore increasing the time between overhauls and reducing costs. In much the same way liquid and gaseous coal based fuels can create deposits when burned in the combustion chamber of a jet engine or power generation system [1]. Turbine machinery is much more complicated and sensitive to minor changes in geometry which could result from the deposition of fly ash onto its surfaces. Some of the side effects of deposition are increased heat transfer due to increases in surface roughness, and the blockage of essential film cooling holes. These side effects, combined with constantly increasing turbine inlet gas temperatures, make the problem of deposition more and more prevalent. 2

21 1.2 Literature Review As previously mentioned, particulate entrained into a turbine can erode or corrode the turbine material, or it could deposit onto the turbine blade surface, causing severe damage and potential failure. The effects of deposition on turbine hardware are well documented in the literature. Kim et. al. [2] looked at the effects of large amounts of volcanic ash being ingested into the hot sections of two aero-engines. It was found that deposition can cause clogging of critical film cooling holes, which can result in the deterioration of the turbine hardware. As the hardware degrades it could potentially release in large portions and cause large amounts of damage to downstream components. Figure 1 shows a possible outcome of blocked film cooling holes. A significant portion of the tip region is eroded away causing losses in aerodynamic performance and part integrity. Figure 1 - Leading edge of turbine nozzle guide vanes exposed to high volcanic ash concentrations. Note the effects of increased surface heat transfer caused by clogged film cooling holes [2]. 3

22 Another effect of deposition inside a turbine is made evident in a model proposed by Wenglarz [3]. He states that decreasing the throat area of nozzle guide vanes with deposition reduces the overall power generated by the turbine because of the reduction in mass flow through the turbine. This fact is made apparent by examining Equation 1. Overall power output is directly proportional to total mass flow passing through the engine, therefore a decrease in total mass flow results in a decrease in overall power. Equation 1 Also, Wenglarz states that deposit buildup in the trailing edge region can reduce the aerodynamic performance of turbine airfoils. Other studies conducted by Bogard et. al. [4] have shown that surface roughness can increase heat transfer as much as 50%. Sundaram and Thole [5] have shown that deposition near an endwall film cooling hole can sometimes increase the film cooling effectiveness, but film cooling effectiveness decreases with greater deposit heights. They have also shown that cooling effectiveness is decreased when cooling holes are either partially or completely blocked. Schlichting [6] presents the fundamental principles for boundary layer growth on rough surfaces. He shows that boundary layer thickness increases substantially as the parameter known as equivalent sand roughness increases. Schlichting developed the parameter equivalent sand roughness in an attempt to unify roughness measurements. Bammert and Sandstede [7] quote rough surface momentum thicknesses at the trailing edge of a fifty percent reaction turbine blade three times as great as those on a smooth turbine blade. The equivalent sand roughness used to create such a marked increase in momentum thickness was 3.3x10-3 µm. The momentum thickness is directly related to the skin friction drag 4

23 along a surface, and with an increase in drag comes a loss in performance of turbomachinery. Zhang, et al. [8] show that the wake behind a symmetric airfoil, simulating the pressure surface of a turbine vane, increases in width with increasing surface roughness, and that non-dimensional total pressure losses also increase with increasing surface roughness. They attribute the decreasing performance to thicker boundary layers at the airfoil trailing edge, and increased turbulent diffusion in the transverse direction within the wake as it travels downstream. Boyle and Senyitko [9] show that as Reynolds number increases, at constant Mach number, a uniformly rough turbine vane will produce up to a 60% loss increase when compared to a smooth vane. They also show a slight loss benefit at low Reynolds numbers due to the surface roughness, which was attributed to laminar separation being prevented by earlier boundary layer transition. Zhang et. al. [10] demonstrates a 60% increase in total pressure loss with increased surface roughness, and a marked increase in the width of the vane wake for a scaled up version of a turbine vane. They performed tests on uniformly rough vanes and vanes with variable roughness. Zhang et. al. claim that the roughness elements cause earlier boundary layer transition from laminar to turbulent, thicker boundary layers, and increased turbulent diffusion, which combine to increase the losses for the rough vanes. On top of performance degradation caused by deposition, cleaning and repairing turbine hardware is expensive and time consuming. It is clear from current literature that particulate deposition poses significant problems for the operation of turbines using dirty fuels. 5

24 1.2.1 Other Deposition Research The studies conducted which show the negative effects of deposition on turbine hardware used either massive amounts of particulate in a fully operational jet engine or prefabricated, large scale, deposition patterns taken from in-service turbine hardware. They do not allow the detailed study of the mechanisms behind particulate deposition or the physical characteristics of the deposits. However, several other facilities are equipped to examine the physical characteristics of deposits and the mechanisms behind deposit formation. Jensen et. al. [11] describes a facility, called the Turbine Accelerated Deposition Facility (TADF), used to study the mechanisms of deposition on one-inch diameter cooled and un-cooled turbine material samples. The TADF, shown in Figure 2, brings air into its base which is heated using natural gas to temperatures comparable to turbine inlet temperatures. Particulate is injected into the flow prior to the flow being accelerated to turbine inlet Mach numbers. The hot particulate laden flow is exhausted to atmosphere where the cooled and un-cooled turbine material samples are fixed. The flow impinges on the samples and deposits are formed on the surface. 6

25 Figure 2 - Turbine Accelerated Deposition Facility (TADF) at Brigham Young University [11] Crosby et. al. [12] used the TADF to study how gas temperature, particle size, and metal temperature affect the physical characteristics of the deposit. They found that a deposition threshold exists at 960 C, and that deposition increases with particle size, and deposition decreases with metal temperatures. Ai et al. [13] evaluated both impingement 7

26 and film cooling effects using the same facility. They found that deposition could be reduced substantially with proper cooling. Due to the obvious complexities of high temperature deposition experiments, several other research groups have recently designed low temperature deposition simulators. For example, Lawson and Thole [14] describe a low speed and low temperature facility used to simulate deposition on a flat plate with a row of film cooling holes. This facility uses low melting temperature wax to form deposition patterns on a flat plate with film cooling holes. Adiabatic effectiveness is evaluated near the cooling holes using infrared thermography. It was found that cooling effectiveness at low momentum flux ratios was decreased as deposition increased. Another facility constructed by Vandsburger and Tafti [15] uses Teflon or PVC particles in an elevated temperature gas flow to simulate syngas ash particulate entrained in a power turbine. The Teflon or PVC particles were chosen such that their momentum and thermal Stokes numbers match that of coal ash; this is to ensure they behave in a way similar to coal ash in turbine flow paths. These facilities begin to inspect the mechanisms behind the deposition process and the effects of particulate deposition on turbine hardware and film cooling effectiveness. However, there has not been a facility that is capable of simulating deposition on actual turbine hardware in a controlled flow environment similar to the turbine environment. Testing actual turbine hardware is extremely useful in that the test results are direct indicators of what could happen in service. A new facility is needed in order to conduct these tests in a timely manner. Kramer [16] has detailed the construction of such a facility called the Turbine Reacting Flow Rig (TuRFR). This thesis presents an overview of this facility, 8

27 some validation studies of the facility, and the initial deposit test results found by using the facility. This thesis also presents the effects of deposits on the aerodynamic performance of the turbine hardware. 9

28 CHAPTER 2 TuRFR DESCRIPTION AND VAILDATION STUDIES 2.1 Primary Flow Path Description To study the effects of particulate deposition on turbine Nozzle Guide Vanes (NGV) a facility was constructed that is capable of attaining flow conditions representative of those found exiting the combustion section of a power generation turbine. The facility, named the Turbine Reacting Flow Rig (TuRFR), utilizes current NGV technology for deposition tests and is capable of supplying film cooling air at flow rates and temperatures comparable to those of the specific NGV design. One goal of the TuRFR is to reproduce deposition patterns found on turbine NGVs, which have been in service for thousands of hours, in a 1 to 2 hour accelerated test. A brief description of the TuRFR and its sub-systems is provided below. Figure 3 shows a schematic of the TuRFR and the primary flow path. The main air supply comes from two 16 MPa, 21 m 3, storage tanks that are maintained at constant pressure by three Ingersol-Rand compressors. Air from these tanks passes through a regulator and two parallel choked orifice flowmeters. Pressure and temperature measurements at the orifice plate yield an uncertainty in mass flow rate of 3-5% over a 10

29 range of massflows from 0.5 to 1 kg/s. Following the orifice meters, the airflow is split into four branches: main airflow, film cooling flow, fuel premix flow, and particulate flow. The main airflow is brought into the base of the combustor at four evenly-spaced locations. The air passes through a pebble bed and honeycomb system to evenly disperse and straighten the flow. At this point, the air is seeded with particulate (a description of the particulate feed system is provided later on in Sec. 2.2 Particulate Feed Sub- System). Next the air is heated, via the combustion of natural gas, to similar temperatures found in power generation turbines. Natural gas is taken from a local low pressure supply line. Natural gas predominantly consists of methane, but it can contain significant amounts of several other hydrocarbons. The detailed chemical makeup and molecular weight of the natural gas was supplied by Columbia Gas. Natural Gas used in this particular study has specific gravity of 0.592, a heating value of 38.6 MJ/m 3, and consists of 95% methane, 2.35% ethane, 1.17% CO 2, and 1.48% other hydrocarbons. Fuel gas is compressed to 20 MPa, using two Fuelmaker reciprocating compressors, and stored in an array of 32 high pressure gas cylinders. Fuel gas is passed through a choked orifice meter similar to the one used for the main air supply. Temperature and pressure measurements at the orifice plate yield an uncertainty in fuel flow rate of 3% at flow rates up to 0.03 kg/s. Fuel is brought into the combustion section using eight fuel lines evenly dispersed around the circumference of the TuRFR. At the end of each fuel line is a set of four flame-holders (32 flame-holders in all). The flame holders, shown in Figure 4, are necessary to insure flame stability and mitigate sooting. The flameholders are constructed of stainless steel to increase oxidation resistance. 11

30 Figure 3 - TuRFR schematic showing main flow path. After heating, the air is accelerated through a 60 axisymmetric cone with an inlet-to-exit area ratio of 68:1. The cone accelerates the heated air to Mach numbers in the range of 0.1 to 0.3, depending on the flow rate and temperature of the air. Attached to the cone is the equilibration tube. This tube allows the particulate to come to kinetic and 12

31 thermal equilibrium prior to entering the NGV cascade. The length of the tube, 0.79 m, was prescribed so that a 40 micron diameter particle would be at thermal and kinetic equilibrium with the airflow upon exiting the tube. Figure 4 - Flameholder inside combustion section of TuRFR. The heated air leaves the equilibration tube and passes through a round to rectangular transition prior to entering the viewing section. At its circular lower end, the round to rectangular transition incorporates a circumferential sliding seal which accommodates the thermal growth of the TuRFR. The sliding seal allows for approximately 5cm of relative thermal growth between the upper TuRFR components 13

32 (round to rectangular transition, view section, and vane holding section) and the lower TuRFR components (the combustion section, cone, and equilibration tube). There are two viewing ports on the side of the TuRFR which allow access for various optical measurement techniques, such as infrared thermography and Laser Doppler Velocimetry (LDV), and for video recording of deposition on the vane leading edges as it is occurring (Figure 5). One of the optical cavities is equipped with a static pressure port and thermocouple to measure the gas properties just upstream of the final inlet contraction. The uncertainties of these measurements are 3% for the pressure measurement and 2% for the temperature measurement. This same optical view section has two isolation cavities located on either side of the main gas path. The cavities accept lower temperature air from a separate source and are separated from the main gas path with removable plates. These plates can be perforated to allow cooler air to mix with the primary gas path. This capability replicates pattern factors (spanwise temperature variations) created by dilution jets found in modern combustor liners. For the present study, the removable plates were solid, allowing no dilution flow. Zero dilution flow was selected for this study due to lack of information on pattern factors seen by the NGV cascade in use. Finally, the air enters a contraction which transitions from the rectangular viewing/dilution section to the annular inlet area of the NGV cascade. The contraction ratio for this final transition is roughly 2:1 with the NGV inlet area being approximately m 2. The NGV hardware used in this study is from a CFM56-5B production engine. The two nozzle guide vane doublets are mounted at the exit of this rectangular to 14

33 annular transition piece. Figure 6 shows a top view of the NGV housing with the cover removed. Figure 5-3D cutaway view of upper section of TuRFR showing measurement locations. Film cooling air is supplied to the NGV cascade from an auxiliary flow path. The total film cooling flowrate is measured using an inline pneumatic volumetric flow meter with an uncertainty of 7% of reading at flow rates equal to 12% of the cascade inlet flow rate. The outer casing and inner hub film cooling cavities are fed by separate lines. The casing film cooling cavity is metered using an in-line pneumatic flow meter with an uncertainty of 7% at flow rates up to 8% of the cascade inlet flow. A backflow valve is used to balance the temperatures for the two film cooling reservoirs. The secondary flow path for the film cooling air is kept separate from the main flow path to prevent premature heating of the film cooling air, and to maintain adequate density ratios. The 15

34 density ratio was calculated as the NGV cascade inlet to film cooling reservoir temperature ratio. This assumes the pressure of the film cooling air is equal to the free stream pressure, which is not exactly the case. Thus density ratios calculated represent conservative estimates of the actual density ratio. Film cooling air from the hub and casing cavities enters the hollow vanes through cooling cavities that are equipped with impingement inserts, similar to industrial practice. Figure 6 - Top view of NGV cascade and film cooling reservoir showing film cooling temperature measurement locations. The NGV cascade exhausts to atmosphere from the test cell through a retractable ceiling and large bay doors. Due to the high temperatures required of the TuRFR all of the components downstream of, and including, the combustion section are made from Inconel 600 series alloys. These exotic metals are capable of retaining strength at extremely high temperatures which makes them aptly suited for this type of application. 16

35 The strength of the alloys used in the TuRFR design allow it to operate at temperatures up to 1200 C for three to four hour tests. Based on deposition threshold temperatures cited by previous researchers [1,2,12] the TuRFR is capable of producing gas temperatures and particulate impact velocities within the limits required to produce deposition. 2.2 Particulate Feed Sub-System One of the main goals of the TuRFR is the duplication of particulate deposition seen on in-service hardware, inside a laboratory setting. Without the ability to supply particulate to the main air flow passing through the turbine hardware, the TuRFR would not be able to accomplish its goal. The particulate feed sub-system, shown in Figure 7, consists of a particulate hopper, an auger (with agitator), a driving motor, and a gearbox. The particulate hopper is a 6 foot section of clear two inch diameter PVC pipe. The PVC pipe is clear to allow for gauging of the particulate level during a test. The hopper was designed to hold more than enough particulate for a 2.5 hour test at particulate concentration of 80 ppmw in a 1.18 kg/s main air flow. If the particulate flows out of the hopper through the auger the pressure in the hopper will decrease due to the increase in volume. Therefore the space above the particulate inside the hopper must be kept at the same static pressure as the air being bled from the main flow. Pressure equalization is accomplished via a small piece of tubing connecting the hopper to the main air bleed line (not indicated in Figure 7). 17

36 Figure 7-3D CAD schematic of particulate feeder. Pressure equalization tube is not shown in the diagram. Directly below the hopper is the auger. The auger is powered by a small DC motor connected to a speed reducing gearbox. A speed reducing gearbox is used to maintain the low particulate flow rates required to match the net particle loading for a one to two hour test. The small DC motor is powered by a variable power supply so that the particulate feed rate can be controlled. The auger forces particulate into the air stream that has been 18

37 bled from the main air flow. The particulate is picked up by the secondary air stream passing over the auger and is injected into the TuRFR in the direction of the flow directly beneath the flameholders. To aid in particulate pickup a small contraction in this secondary flow path reduces the static pressure in the vicinity of the auger exit. This static pressure decrease helps to pull the particulate from the auger and reduce clogging in the feeder. A common problem with any type of particulate injection system is a phenomenon known as bridging. This is when the particulate in the hopper, in essence, clogs by forming a bridge strong enough to support the weight of the particulate above. When this occurs the particulate below the bridge flows normally while the particulate above the bridge is prevented from ever reaching the auger. To mitigate this effect, the auger was fitted with agitators, welded to its end, which stir the particulate as it approaches. 2.3 TuRFR Validation Studies Prior to conducting deposition studies using the TuRFR, several validation studies were required. The validation studies included a survey of the temperature entering the inlet of the NGV cascade, a temperature survey at the exit plane of the cascade for a nonheated main flow with heated film cooling air, and a total pressure survey at the exit plane of the cascade for a non-heated flow. The purpose of these tests are to characterize and validate the inlet conditions entering the NGV cascade, to ensure the absence of flow anomalies, and to prove that film cooling air is in fact exiting the film cooling holes and is affecting the free stream flow in an expected manner. For the remainder of this thesis 19

38 the term film heating will refer to heated film cooling air being supplied through the film cooling holes. A traverse mechanism was constructed at the exit of the NGV cascade. LabView software was configured to control the traverse system, acquire the pressure and temperature data, and monitor other important parameters of the TuRFR during operation. The actual grid used for each survey is presented in the respective sections. All of the grids were rectangular. None of the measurement grids were able to traverse the entire inlet or exit flow field due to spatial constraints of the measurement probes. Test conditions are indicated in Table 1, again with non-combusting flow and film heating. Inlet Velocity (m/s) Inlet Mach Number Inlet Temperature ( C) Inlet Temperature Map Temperature Map 4% Film Heating Pressure Maps 0% Film Heating 4% Film Heating Exit Re c x Film Heating Temperature N/A 244 N/A 245 ( C) Film Heating Density Ratio N/A N/A Table 1 - Test conditions for TuRFR validation studies. 20

39 2.3.1 NGV Inlet Temperature Survey For the studies presented herein a uniform inlet temperature distribution was required. If the inlet temperature is not uniform it must be characterized so that any unexpected non-uniformities in deposition can potentially be explained. For this validation study the NGV cascade was removed from the TuRFR and the main air flow exhausted to atmospheric conditions through the annular hole where the NGV cascade is fixed. The traverse system was mounted away from the annular exit of the TuRFR in order to protect it from the hot main air flow. A metal arm was mounted to the traverse which held a 1.58 mm diameter, Type K, thermocouple inside the main air flow. A grid of 9 cm in the pitch dimension, and 5.4 cm in the span direction with 3 mm spacing between each point was used. An outline of the grid used in this test is shown by the rectangular box in Figure 8. It is important to note that the thermocouple tip started approximately 5 mm from the hub radius inside the flow path. When the test was complete the thermocouple tip was about 3 mm from the hub radius outside the flow path. This offset is represented in the test data by a very cold region, which has a similar shape to the hub radius of the annular hole in the TuRFR. The thermocouple probe was traversed in a plane that was approximately 3 mm above the cascade inlet plane. Figure 8 shows the result of the inlet temperature survey non-dimensionalized by the average temperature of the inlet area. The average temperature was calculated using the region indicated by the line and arrows in Figure

40 Figure 8 - NGV cascade inlet temperature survey. From these results it is evident that there is a spanwise temperature variation in the inlet temperature. This variation is at most 10% of the average inlet temperature. An inlet temperature variation may cause differences in the formation of the deposits along the vane span because deposit formation is highly dependent upon temperature Film Cooling Validation Survey The film cooling validation survey was conducted by supplying heated film cooling air to an unheated main flow. Conducting the validation study in this way allows the air exiting the film cooling passages to stand out from the main flow without the need for combustion. Also, the danger to the traverse mechanism incurred by a heated flow required that the flow be unheated. Film heating was supplied to the NGV cascade at a rate of 4% of the inlet mass flow. The inlet mass flow for this test was set to 1.27 kg/s, which required a film cooling 22

41 flow rate of 0.05 kg/s. The traverse mechanism was set with a grid of 2.7 cm in the spanwise direction and 6.9 cm in the pitchwise direction. The grid spacing was 3 mm in both the spanwise and pitchwise dimensions in order to resolve the effects of the heated film cooling air on the temperature distribution. The main flow temperature was kept low (-15 C) and an auxiliary heat source heated the coolant to approximately 244 C. This provided a temperature difference between the coolant and free stream without exceeding the temperature limitations of standard measurement probes. The probe was traversed through the wakes of the NGV cascade as indicated by the red rectangle in Figure 9. Figure 9 shows the exit temperature contour map, normalized by the difference between the film cooling air and the exit gas, taken 20% axial chord downstream of the vane trailing edge plane. Temperature measurements were taken with a.203 mm diameter type K thermocouple. Dark blue regions in Figure 9 represent regions in which the temperature measured is further from the exit temperature, T ex (exit temperature is averaged across a line through the passage center). Non-dimensional temperatures less than unity indicate a measured temperature closer to the film cooling temperature. Figure 9 shows that coolant is confined to the vane wake and endwall regions in the cascade, as expected. The region of higher temperature near the end wall of the cascade is from the additional film heating air being applied at the hub endwall. 23

42 Y coordinate 20 FC meas FC gas T T FC FC T T meas ex T ex = -16 C X = 1.25*Pitch Y = 0.6*Span T ex = F T fc = F X coordinate Figure 9 - Dimensionless temperature at vane exit with 4% film heating, red box shows approximate location of measurement plane Exit Total Pressure Surveys Finally, an assessment of the total pressure losses incurred across the NGV cascade, with and without film heating, was made in order to make a cursory comparison of results with current literature (a more detailed comparison will be made of aerodynamic performance in following chapters). Also, total pressure losses across the cascade help to determine aerodynamic performance of the NGV cascade. The total pressure loss data can be taken prior to deposition then re-taken after deposition, in unheated flows, to estimate the effects of deposition on aerodynamic performance. The measurement grid used for this test was 5.4 cm in the spanwise direction and 9 cm in the pitchwise direction. The grid spacing was set to 3 mm as with previous tests. Total pressure data were taken using a Pitot tube oriented at approximately the vane surface 24

43 exit angle. The probe was traversed in a plane that was 30% axial chord downstream of the vane trailing edge. Figure 10a shows that this unheated flow is approaching sonic velocities because the non-dimensional pressures, P TOT /P amb, are greater than the critical pressure ratio to achieve choked flow, Also, Figure 10a shows the wake region and parts of two passages; suggesting passage periodicity. Passage periodicity is confirmed by examining the trace of dimensionless total pressure loss taken at approximately midspan over two and a half vane wakes (Figure 11). Note that the values of non-dimensional total pressure loss in the two passages (between the wakes) are both approximately equal to zero. The vane wake regions show different non-dimensional total pressure loss values between blade 2 and blade 3. This difference is most likely due to the limitations of the traverse mechanism. The probe was traversed along a straight line in the exit plane of the NGV cascade, however, the cascade is annular. This results in the probe head measuring exit total pressure in different span locations for each blade. Also this means that the probe head, which begins perpendicular to the trailing edge of vane 2, is no longer perpendicular to the trailing edge of vane 3. Despite these differences passage periodicity is still preserved. The vane wake region can be seen as the areas of positive total pressure loss in Figure 10b. Total pressure loss was calculated for the data in Figure 10b by using Equation 2. P TOTin was taken to be the average total pressure across the passage span in the region unaffected by the wake, and was 270 kpa. The sum of total pressure losses in the vane wake region, between the lines in Figure 10a, for 0% film heating applied was %. 25

44 Figure 10 - Total pressure measurements at the exit of the NGV cascade with 0% film heating applied: a) Non-dimensional total pressure (P TOT /P amb ) and b) lines of total pressure loss at various span locations. 26

45 Figure 11 - Non-dimensional total pressure loss [(P TOTin P TOT )/0.5ρu 2 ] collected 13% true chord downstream of NGV exit plane indicating passage periodicity. Equation 2 After taking total pressure data with zero film heating, total pressure data with 4% film heating was taken using the same measurement grid and spacing as the test without film heating. Figure 12 shows the results of this test. Comparing Figure 12b with Figure 10b shows that there are fewer total pressure losses incurred across the cascade when film cooling is not applied. The sum of the total pressure losses in the wake region, indicated 27

46 by the lines in Figure 12a, for 4% film heating is %. Also, the wake region in Figure 10b is wider than the wake region of Figure 12b, but the losses are more intense in the wake shown in Figure 12b. The wide wake of the 0% film heating test (Figure 10b) is most likely caused by the main flow passing over hundreds of film cooling holes which are not exhausting film heating air. Total pressure losses being higher when film heating is applied compared to when film heating is not applied qualitatively matches with studies conducted by Day et. al. [19]. They used a film cooled transonic turbine NGV cascade, at an exit Re cx of 2 x 10 6, to compare measured aerodynamic losses between a case without coolant supplied to the cascade and a case with coolant supplied to the cascade. The data showed a 4% increase in loss between the 0% film cooling case and the case with 5% of the inlet mass flow being ejected from the cylindrically shaped film cooling holes. In the film heating cases presented above a 22% increase in loss was observed between the 0% film heating case and the 4% film heating case. The difference in aerodynamic loss increase could be due to differences in how the losses were estimated above and how Day et. al. estimates losses. The difference could also be due to the fact that different vane geometries were used in the two studies. 28

47 Figure 12 Total pressure measurements at the exit of the NGV cascade with 4% film heating applied: a) Non-dimensional total pressure (P TOT /P amb ) and b) lines of total pressure loss at various span locations. 29

48 CHAPTER 3 ASH PARTICULATE 3.1 Description of Particulate One of the primary goals of this study is to duplicate deposition found on in service turbine NGVs that have been subjected to particle laden air. The particulate can come from a number of sources including, but not limited to, degradation of components upstream of the turbine, incomplete combustion, combustion of synthetic fuels, and airborne particulate in the environment. This study focuses on the particulate generated by the combustion of synthetic, coal based, fuels. It has been shown in other research that bituminous coals produce some of the highest yields of liquid fuel from a direct liquefaction process. Lignite coals are easier to liquefy, but they do not give as high of a liquid fuel yield [17]. This suggests that bituminous coal is likely to be used, when available, for the direct liquefaction of coal. Bituminous fly ash was selected as the particulate of choice for this study due to its likelihood of being used in synthetic fuels derived from coals and because of its availability. The bituminous coal fly ash used in these experiments was donated from a local power plant and was mechanically ground before use. The molecular composition of the 30

49 particulate was determined using X-Ray Diffraction (XRD), the elemental composition was identified using Energy Dispersive Spectroscopy (EDS), and the mean diameter was measured using a laser based Coulter Counter. The chemical composition of the bituminous fly ash is presented in Table 2 along with the mass mean diameter of the fly ash particles. Figure 13 shows the Coulter counter particle sizing distribution. Advanced filtration systems in modern power turbines typically remove all particulate above 10µm, thus nearly half of the ground bituminous ash used in this study would not reach the turbine if proper filtration procedures were followed. The primary molecules present in the fly ash are SiO 2, Fe 2 O 3, and Al 2 O 3. An image of the particulate, taken with a Scanning Electron Microscope (SEM), is shown in Figure 14. Mass Mean Diameter (µm) Bituminous 14 Bituminous Coal Ash Composition (Wt %) SiO CaO 2.93 Fe 2 O Al 2 O SO K 2 O 2.48 Table 2 - Particulate mass mean diameter and bituminous coal ash molecular composition. 31

50 Mass % 6.00% 5.00% 4.00% 3.00% Mean = µm Median = µm Size (µm) Mass % % % % % % % 2.00% 1.00% 0.00% Diameter (µm) Figure 13 - Particle size distribution results from Coulter counter for Bituminous coal fly ash. Particle loading inside a gas turbine engine is sometimes measured as a particulate concentration. Normal concentrations of particulate for power turbines are very low, less than 0.1 parts per million by weight (ppmw) [18]. The TuRFR cannot be operated at these low particulate concentrations for long enough periods to duplicate deposition patterns found on in service hardware. Fortunately, Jensen et. al. [11] have shown that net particulate throughput, not test time or particulate concentration independently, is the dominant factor in the development of deposits similar to those found in service. Therefore, the TuRFR is operated with a much higher particle concentration for a shorter duration. This concept gives rise to a unit of measure known as net particle loading, given in terms of ppmw-hr. Net particle loading is a measure of the particulate concentration seen by the NGV cascade throughout the duration of exposure. In other words, a one hour TuRFR test at a particulate concentration of 250 ppmw would yield the 32

51 same net particle loading for approximately 4000 hours of service at a particulate concentration of ppmw. This unit of measurement allows the TuRFR to operate short duration, high particle concentration, tests that accurately simulate the deposition found on actual in service turbine hardware. Figure 14 - SEM image of Bituminous coal fly ash. 33

52 CHAPTER 4 DEPOSITION TESTING RESULTS 4.1 Deposition Testing One goal of this work is to show that the TuRFR can produce deposition patters seen in service and those created by other researchers. Tests conducted for the TuRFR deposition study examined the differences between deposit formations created with varying levels of inlet temperature and film cooling percentages. During each test the TuRFR is slowly brought to the steady state test conditions summarized in Table 3. A test begins once steady state conditions have been met and particulate is added to the flow. In all three cases, the full particulate was dispensed over 45 minutes at an approximately constant rate. The TuRFR was then shut down over a period of several hours to prevent thermal shock. The vanes were photographed prior to removing them from the facility for analysis. After all post-test deposit analysis had been completed the NGV cascade was cleaned using a sandblaster, similar to industry practices. 34

53 Low Temperature No Film Cooling Film Cooling Inlet Velocity (m/s) Inlet Mach Number Inlet Temperature ( C) Exit Re c x Film Coolant Temperature ( C) N/A N/A 677 Film Cooling Density Ratio N/A N/A Centerline Average Roughness of Pressure surface (Ra, microns) Net Particle Loading (ppmw-hr) Table 3 - Deposition testing conditions. In order to estimate the inlet velocity, the static pressure and recovery temperature of the particle laden air stream was measured at the locations shown in Figure 5. The recovery temperature of the inlet flow was measured using a 1.58 mm diameter, type K, thermocouple with an uncertainty of 10 C at test temperatures. The static pressure was measured with 3% uncertainty. The flow reference temperature given in Table 3 was taken from the vane inlet thermocouple. The vane inlet velocity was estimated given the air and fuel massflow measurements and assuming isentropic flow between the view section conditions and vane inlet. The calculation is an iterative process. First the properties of the air and 35

54 combustion products mixture are determined. Then the Mach number in the view section is found using Equation 3. Equation 3 With the Mach number in the view section determined the reference area, A*, is calculated using the isentropic relation given in Equation 4. Using Equation 5, the value of A*, and guesses for M in, A in is calculated and compared with the inlet area of the NGV cascade. Once the calculated A is within 0.1% the iterations are stopped and the inlet velocity is determined using M in and the speed of sound in the mixture. Equation 4 Equation 5 The exit Reynolds numbers presented throughout this thesis are calculated based on the true chord of the NGV cascade, and are given by Equation 6; where dynamic viscosity, µ, is a function of temperature according to Sutherland s law. Equation 6 36

55 4.1.1 Temperature Variation The first deposition test conducted in this study was at a low temperature (908 C) as shown in Table 3. No deposits were observed with this test. Crosby et. al. [7] observed that there was a significant decrease in deposits below the critical temperature, found in their studies to be about 960 C. The two subsequent tests were at sufficient temperature to generate deposition on the NGV cascade. Figure 15 shows images of the NGV cascade after the three tests given in Table 3. These results confirm that the critical temperature for deposit formation with bituminous coal ash in the TuRFR is in the same range as that found by Crosby et. al. [12] in their TADF for subbituminous coal ash. Figure 15 - Post test images of deposit formation for Low Temperature test, No Film Cooling test, and Film Cooling test. 37

56 4.1.2 Film Cooling Variation For the two high temperature deposition tests the level of film cooling supplied to the NGV cascade was varied from 4% of the cascade inlet mass flow down to approximately zero. Deposition occurred for both of these tests. The No Film Cooling test produced large enough deposits that could easily be removed to perform chemical analysis. Roughness results presented in Table 3 combined with surface images from Figure 15 show that increasing the film cooling rate provides a marked decrease in deposition. Research performed by Crosby et. al. [12] and Ai et. al. [13] confirms these results. The average thickness of deposits removed after the No Film Cooling test was measured to be 240 m using SEM images. At first glance, another interesting result can be gleaned by comparing deposition images from Figure 15 to those found by Kim et. al. [1] (Figure 1). Kim et. al. show that the most degradation occurred near the tip region of the vane leading edge because the larger dust particles were centrifuged toward the casing side of the NGV due to the combustor geometry. Examining the deposition created by the No Film Cooling test, much of the deposition also occurs on the leading edge near the casing side of the NGV cascade. This could be explained by centrifuging of the particulate due to the more aggressive combustor exit angle on the hub wall compared to the casing, a feature replicated from the actual engine design. However, this is only true for particles which are slow to react to abrupt changes in the flow. Kim et. al. provides a method for determining the time constant for a particle in a flow based on a drag analysis. In this 38

57 analysis it is assumed that so that the time constant can be given by Equation 7. The Re p range for this assumption is between 0 and 100. Re p for 14 μm diameter particulates is approximately 8.5, well within the range of the assumed C D. Equation 7 The time constant for a 14 µm diameter particle was determined to be approximately 0.6 ms, while the residence time of a particle in the final contraction of the TuRFR leading up to the NGV cascade inlet is approximately 2.25 ms. Therefore a 14 µm diameter particle would actually not be moved toward the casing region of the NGV cascade, but it would follow the flow path of the TuRFR. In order for a particle to be forced to the casing region, due to the geometry of the TuRFR, it would have to be larger than 30 µm. The mean diameter of the bituminous coal ash was 14 μm, however significant levels of particulate larger than 30 μm were found to be present in the ash (Figure 13). These larger particles would not react as quickly to the abrupt changes in the flow path and, as a result, could move toward the casing region of the NGV cascade. These larger particles could be depositing in the casing region of the NGV cascade, causing a local increase in the levels of deposition. 4.2 Surface Roughness Measurements The primary measure of surface deposition in this study is the centerline average roughness (Ra). A Mitutoyo Surftest surface profilometer was used to measure the vane 39

58 surface roughness at multiple locations after each test. Roughness measurements were acquired in the same vane locations between tests, and Ra values were calculated at each location. Ra values represent the average height of surface non-uniformities relative to a spline fit on the profilometer data. The spline fit characterizes large scale surface geometries (such as vane curvature) so that small scale surface roughness can be emphasized. The height data used in the Ra calculation is relative to the spline fit of the raw height data [20]. Figure 16 shows regions of surface roughness obtained from the No Film Cooling test. Table 3 shows the average relative roughness values on the pressure surface for each test. The roughness values demonstrate that inlet temperature and film cooling percentage play significant roles in the formation of deposits. 40

59 Profile height ( m) Profile height ( m) Profile height ( m) Profile height ( m) PS M1 LE Suction Surface Trace Ra = microns Ra = 23.2 micron microns PS M2 Ra = microns PS M1 Ra=17.01 m PS M1 Ra=17.01 PS M2 Ra=17.98 m mss PS M2 Ra=17.98 PS M1 LE Ra=23.2 PS m M1 Ra=17.01 m m LE PS M2 Ra=17.98 Ra=23.2 m LE Ra=23.2 SS Ra=2.32 PS M2 Ra=17.98 m m SS Ra=2.32 m LE SS Ra=2.32 m Ra=23.2 m SS Ra=2.32 m Ra = 2.31 microns R Surface tangential distance in mm Surface 3 tangential distance 4 in mm Surface tangential distance in mm Surface tangential distance in mm Figure 16 - Representative surface roughness traces and roughness regions for deposition formed during NO Film Cooling test. CAD image of NGV doublet is not to scale. Surface roughness values for the four locations shown in Figure 16 compare well with maximum roughness values found in an exhaustive roughness survey of serviced turbine hardware conducted by Bons et. al. [20]. As shown in Figure 16, the leading edge and near trailing edge regions of the vane pressure surface exhibit large values of Ra, while the suction surface is nearly free of deposits. Bons et. al. report that these two regions of the pressure surface usually exhibit the largest values of surface roughness. They also observed little if any suction surface deposits beyond the leading edge region. 41

60 This indicates that the deposition created using the TuRFR is a good model of the deposition that forms on in service hardware. 4.3 Deposit Structure and Chemical Composition Deposit flakes from the No Film Cooling test were examined using Energy Dispersive Spectroscopy (EDS), XRD, and SEM microscopy. The EDS results show the elemental composition of the deposit. However, EDS only indicates which elements are present; not the relative amount of each element present. XRD helps to determine what molecules make up the deposit and how much of that molecule is present. So, EDS combined with XRD can provide the complete chemical makeup of the deposit, which can be compared to that of the ash used in the test. Figure 17 - SEM photo highlighting deposit structure. The vane surface is not indicated in the image, but it is below the deposit flake. 42

61 The elements found in the deposit flakes (Fe, Al, Si, and O) match those found in the bituminous ash used in the test at approximately the same concentrations. Figure 17 shows an SEM photo taken of the deposit flake on edge. The chemical analysis of the Low Film Cooling test deposits also shows a large Copper (Cu) presence. This was unexpected because the bituminous fly ash used in the test did not contain Cu. The Cu originated from the failure of previous versions of the TuRFR flameholders, made from Cu, throughout the test. The current flameholders are the ones described in Section 2.1. The Cu flameholders oxidized heavily and the oxidized Cu produced a heavy layer of Cu deposition on the NGV surface. The Cu layer likely aided in the deposition of the bituminous ash. The deposition created on the vanes during these tests was completely removed after each deposition test. In this way, each time deposition was created on the surface of the NGV cascade the initial surface was smooth and free of deposition from previous tests. O Cu Fe Al Si SEM Image Figure 18 - EDS element maps of deposit flake highlighting Cu, Fe, Al, Si, and O layers. SEM photo highlighting deposit structure. The vane surface is not indicated in the images, but it is below the deposit flake in all of the images. 43

62 Figure 18 shows a series of element maps obtained for a Low Film Cooling test deposit sample. These maps show two layers of Cu, indicated in the image labeled Cu in Figure 18, surrounding a layer containing Fe, Al, Si, and O. The elements found in the Fe, Al, Si, and O layer match those found in bituminous ash used in the test indicating that this layer is comprised of coal ash. Figure 18 also shows an SEM photo taken of the same region as the EDS element maps. The SEM photo demonstrates a difference in the structure of the Cu deposit layer versus the ash deposit layer. The lower porosity of the Cu layers could be a result of a higher fraction of molten Cu particles, while the higher porosity of the ash layer potentially indicates a lower fraction of molten ash particles. These tests have shown that the TuRFR can in fact supply an approximately uniform temperature to the inlet of real engine hardware, the TuRFR can supply film cooling air to the engine hardware at comparable density ratios and mass flows, and that the TuRFR can supply particulate to the heated flow in order to make deposition on the surface of the engine hardware. In short, these tests show that the TuRFR is capable of accomplishing its goal of creating real deposition on real engine hardware. 44

63 CHAPTER 5 AERODYNAMIC PERFORMANCE ASSESSMENT 5.1 Aerodynamic Performance Assessment Background With the TuRFR validation tests complete, determination of the effects of real deposition on the aerodynamic performance of real engine hardware could begin to be explored. Aerodynamic performance of a turbine NGV can be assessed in many different ways, from measurements of total pressure losses in the wake of the NGV, in much the same way as Boyle and Senyitko [9] or Hummel et. al. [22], to aerodynamic efficiency similar to Abuaf et. al. [23]. In this thesis aerodynamic performance is assessed using an integrated (area-weighted) non-dimensional total pressure loss given by Equation 8. Equation 8 A positive value of this parameter indicates that available total pressure is lost across the NGV. The details of how the integrated non-dimensional total pressure loss is measured are described in the following sections. It has been shown by many authors [7, 9, 22, 23] that roughness along the surface of a turbine blade increases the losses incurred 45

64 across the blade. Boyle and Senyitko [9], among others, also show that total pressure loss values change with Reynolds number. In the interest of comparing results to published data, γ was calculated over a range of Reynolds numbers. From the work previously mentioned it is expected that values of γ increase with increasing roughness, when compared with a smooth surface. It is impossible to know, though, exactly what to expect for this particular NGV cascade simply because its geometry differs from all of those found in the literature. Another significant difference between the NGV cascade used in this study to those found in literature is that the NGV cascade used here is actual turbine hardware, not a scaled up version made into a linear cascade. Also, the deposition patterns used in this study are comparable to those found on in service turbines, whereas the deposition patterns used in the referenced literature use a uniform array of discrete roughness elements [4] or a uniform distribution of random roughness elements (i.e.: particulate blown onto a layer of adhesive [7, 9,10, 22]) Total Pressure and Exit Temperature Measurement Setup In order to calculate γ, total pressure, inlet mass flow, and exit temperature must be measured. The inlet mass flow is measured using choked flow orifice plates as described earlier in Section 2.1. Total pressure measurements are made using the same pressure transducer used to measure static pressure in one of the optical cavities. A total pressure probe, with a head diameter of approximately mm, is attached to the same two axis traverse mechanism used previously, and situated outside of the exit flow path. The plane in which all of the total pressure measurements were taken for these aerodynamic performance assessments was 13% true chord downstream of the NGV exit plane. 46

65 In order to assess the effects of deposition on the aerodynamic performance of the NGVs in the cascade, γ must be calculated before and after deposition is created on the surface of the NGV. The probe, however, will not survive the extreme temperatures required to create deposition, which means that it must be removed while deposition is being created. Also, the close proximity of the traverse mechanism dictates that it too must be removed in order to keep it from being damaged during a heated test. As a result a system of probe relocation was developed that ensured the probe head was placed in the same location once the traverse mechanism and probe were reinstalled after a heated test. The design of the NGV cascade being used for these measurements has a 3D profile. locations. This type of geometry has different blade cross sections at different span These differences could result in different vane wakes at different span locations. Hence, the relocation of the probe head and traverse mechanism after creating deposition is critical. In order to be assured that the probe head is relocated to the same position every time the traverse mechanism and/or probe is removed, several scribe marks were made onto the exit hole of the TuRFR locating the non-wetted surfaces of the NGV cascade adjacent onto the TuRFR s annular exit hole. The probe head was then aligned with the same scribe mark prior to and after creating deposition. Finally the probe head was traversed the same distance in the spanwise and pitchwise directions to reach the location where measurements could begin. Total pressure data was collected in the same span location each time so that any γ variation could be attributed to roughness and not to different span measurement locations. 47

66 Exit temperature was measured using a K-type thermocouple of the same size as used in the optical viewing cavity, with the same uncertainty. The thermocouple was placed in the exit flow of the NGV cascade in a position that would not affect the exit total pressure measurement Inlet Total Pressure Measurement The inlet total pressure is determined from total pressure measured in the center of a passage at the exit plane. This is made possible through the assumption that the flow remains essentially unaffected in the center of the passage, and there is negligible total pressure loss. Inlet total pressure was initially determined by averaging the exit total pressure across a passage, but if the inlet mass flow fluctuated while the total pressure probe was traversing the passage, the inlet total pressure would be affected. For instance, the exit total pressure can be related to the square of the inlet mass flow, assuming constant density along a streamline, following Equation 9. Equation 9 Hence the inlet total pressure, and as a result γ, is heavily dependent on the inlet mass flow. The inlet total pressure increases with increasing inlet mass flow. As a result, it was extremely important to keep the mass flow constant as exit total pressure data wass being collected. It is difficult, though, to keep the inlet mass flow constant enough to eliminate fluctuations. The difficulty arises in the fact that the air supply is of finite volume and the air inside the tanks is not being replenished at exactly the same rate as it 48

67 is being exhausted through the TuRFR. This means that the inlet mass flow rate must be adjusted during a test to keep it constant. Another way of determining inlet total pressure is to calculate it from a curve fit of an inlet total pressure versus inlet mass flow plot. The data collected for the inlet total pressure versus inlet mass flow curve, the curve fit, and the equation of the fit are given in Figure 19. Figure 19 - P totin vs. curve showing the experimental data, the curve fit, and the equation of the fit. m in the fit equation stands for inlet mass flow. 49

68 With inlet total pressure determined using a curve fit it can be calculated directly from an inlet mass flow measurement. This means that no matter what the inlet mass flow is at a given exit total pressure measurement location, the inlet total pressure can accurately be determined for that specific location. This process of determining the inlet total pressure removes some of the importance of constant inlet mass flow throughout data collection. Another question yet remains, however. The inlet mass flow measurement is taken far upstream of the NGV cascade, assurance must be made that the inlet mass flow measurement corresponds to the exit total pressure measured. The transit time of the air between the mass flow meters and the NGV cascade now becomes important to ensure that the exit total pressure data collected corresponds to the correct inlet mass flow. Figure 20 shows a schematic of the piping leading up to the TuRFR and the area changes in the TuRFR used in the transit time estimate. The total distance traveled between the orifice plate flow meters and the exit plane of the NGV cascade is approximately 16 m. While the total change in height is approximately 3m. It is important to note that the following transit time calculations are only estimates due to the assumptions required to perform the calculations simply. The assumptions required to perform the transit time estimate are as follows: 1) the flow is incompressible, 2) orifice plate temperature throughout the system remains constant, 3) mass flow is constant, 4) the pipe area is constant, 5) all of the area changes happen abruptly, 6) pressure drops linearly through the pipe, and 7) gravitational effects are neglected. 50

69 Figure 20 - Schematic of TuRFR used in the transit time estimates. Transit time is estimated by using the definition of velocity, combined with the continuity equation at the different areas indicated in Figure 20 and the length of the sections shown. All of the transit times are then summed to achieve a total time required for the air to travel from the flow meters to the NGV cascade. Equation 10 is used for the linear pressure drop through the pipe, and Equation 11 is used to estimate the air transit time in the pipe; where P op is the pressure just downstream of the orifice plate flow meters, P can is the pressure in the base of the TuRFR, L pipe is the total length of the pipe, A pipe is the area of the pipe, R is the specific gas constant for air, T is the air temperature at the orifice plate, x is the distance along the pipe, and is the mass flow through the pipe. Equation 10 51

70 Equation 11 Using Equations 10 and 11 yields transit times of approximately 0.9 s for = kg/s, 1.0 s for = kg/s, and 1.3 s for = kg/s. In order to ensure that the inlet mass flow measured corresponds to the exit total pressure measured at a given measurement location, the exit total pressure data should be collected a certain amount of time the mass flow measurement is collected. Due to the nature of the transit time estimates the amount of delay cannot be accurately known. Still, this method of data collection does not solve the problem. The exit total pressure will always lag the mass flow measurement. Hence, exit total pressure data and inlet mass flow data were collected simultaneously for a total of 4 seconds with the understanding that the majority of the exit total pressure data corresponds to the measured inlet mass flow. 5.2 Aerodynamic Performance Results Exit total pressure was collected over a range of Reynolds numbers before and after deposition had been created on the surface of the NGV cascade. Data labeled smooth indicates that it was taken prior to deposition, while data labeled rough indicates that it was taken after deposition. For all of the data presented below, exit total pressure was measured every 1.5 mm along the same line in the exit flow of the NGV, approximated by the red line in Figure 21. Exit Reynolds number is based on the true chord of the NGV cascade, and is varied by changing the inlet mass flow. Exit Mach number and exit Reynolds number can not be independently variedin this facility. As the 52

71 inlet massflow is increased, there is even a slight drop in air temperature, thus the exit Mach number increases with increasing exit Reynolds number. Figure 21 - Photo of exit plane indicating the approximate location and direction of the probe head travel Rough Aerodynamic Performance Deposition was first created on the surface of the NGV cascade. This was done by operating the TuRFR with the conditions given in Table 4. Also presented in Table 4 is the average Ra value for the pressure surface. Ra values for the suction surface do not seem to be significantly altered by the particle laden flows, therefore they are not presented. 53

72 Operating Conditions Inlet Velocity (m/s) 62.9 Inlet Mach Number Inlet Temperature ( C) 990 Exit Re c (x10 6 ) 0.36 Film Coolant Temperature ( C) N/A Film Coolant Density Ratio N/A Net Particle Loading (ppmw-hr) 283 Surface Characteristics Ra on Pressure Surface w/deposition (μm) 4.55 Ra/c on Pressure Surface w/deposition 1.27 x 10-4 Ra on Pressure Surface w/o Deposition (μm) 1.38 Ra/c on Pressure Surface w/o Deposition 3.85 x 10-5 Table 4 - Operating conditions for deposition creation and surface characteristics of the NGV cascade before and after deposition. It is important to note that the Re c exit of the flow used to create the deposition is on the lower end of those found in a gas turbine engine. Figure 22 is a photograph of the NGV surface with deposition and without deposition. Deposition is evident in the photograph, especially on the leading edge of the airfoil. 54

73 Figure 22 - Photograph of NGV cascade surface with and without deposition. In comparison to other tests using the TuRFR, the level of roughness generated during this test is relatively small. This could be due to the way that particulate was added to the system during operation. There have been significant problems with bridging inside the particulate feeder. A bridge is essentially a clog in the particle feeder that must be removed by either repeated striking of the feeder or by adjusting the air valves upstream of the feeder so that backpressure behind the bridge breaks the bridge. Occasionally the addition of the back pressure forces a large amount of particulate through the feeder very quickly. This is what occurred while deposition was being created in this particular instance. For the remainder of the test however, particulate was added to the flow continuously at the same rate. It is unclear what kind of effect the particulate addition step change had on the actual deposition pattern, but it is important to 55

74 note what could have caused the different surface roughness level in the event that future tests produce significantly different levels of deposition. Figure 23 - Non-Dimensional Total Pressure Loss vs. Position for Re c exit 0.5x10 6. Figure 23 shows a typical non-dimensional total pressure loss plot for Re c exit 0.5 x The wake of the NGV airfoil is indicated in this plot by a positive value of the non-dimensional total pressure loss, while the passage flow is shown by a zero value of the non-dimensional total pressure loss. The values for the integrated non-dimensional total pressure loss for each of the exit total pressure surveys are also shown in Figure 23. The development of the uncertainty of the integrated non-dimensional total pressure is 56

75 given in APPENDIX A:. A plot of γ versus Re c exit (Figure 24) was generated for three exit Re c, and the average values of γ between two runs are tabulated in Table 5. Figure 24 γ vs. Re c exit for Blade 3 of the NGV cascade for the rough cases only. 57

76 Rough Exit Re c (x10 6 ) γ - Blade 2 γ - Blade ± ± ± ± ± ± Smooth Exit Re c (x10 6 ) γ - Blade 2 γ - Blade ± ± ± ± ± ± ± ± ± ± ± ± Table 5 - Exit Reynolds numbers and gamma for Blade 2 and Blade 3 of the NGV cascade. The observed trend is that as Re c exit increases the integrated non-dimensional total pressure loss decreases. This trend does not seem intuitive at first, but when compared to data from literature one can see that this trend is to be expected. As shown in Schlichting [6] (Figure 25), total skin friction decreases with increasing Re for flow through an unseparated turbine cascade. Schlichting states that for an unseparated flow in a turbine cascade the effect of Reynolds number on the loss coefficient is similar to that of a single airfoil because most of the losses are caused by the boundary layer. In this thesis the loss coefficient is defined by the integrand of Equation 8, which is similar in form to that which is defined in Schlichting. 58

77 Figure 25 - Total skin friction coefficient of sand roughened plate at zero incidence angle. Schlichting later goes on to say that the losses are then affected by the Reynolds number in about the same way as the skin friction coefficient of a flat plate at zero incidence and are proportional to [Re] -1/2 for laminar flow, becoming [Re] -1/5 in turbulent flow [6] Smooth Aerodynamic Performance After exit total pressure data was collected for the three Re c exit shown earlier the NGV cascade was sand blasted to remove the deposition. The average Ra and the Ra/c values for the pressure surface after sand blasting are also presented in Table 4. As mentioned earlier exit total pressure data was collected for three Re c exit and the integrated non-dimensional total pressure for each of these runs. However, at this point an unexpected phenomenon began to appear. It seemed that at Re c exit 1x10 6 the losses for 59

78 the smooth cases were higher than that of the rough (Figure 26 & Figure 27). This was unexpected, so exit total pressure data was collected for several more Re c exit. Figure 26 and Figure 27 show the integrated non-dimensional total pressure loss, for both the rough and the smooth cases, for both blades in the NGV cascade, and for all of the Reynolds numbers considered. Figure 26 - γ vs. Re c exit for Blade 3 of the NGV cascade. 60

79 Figure 27 - γ vs. Re c exit for Blade 2 of the NGV cascade. For the reasons mentioned in the previous section, the trends in Figure 25 should be analogous to the trends expected in the NGV cascade used in these studies. As roughness levels increase the transition to the fully rough regime moves lower on the Re scale. Comparing this trend to the data shown in Figure 26 and Figure 27, a potential reason for the sudden increase in γ becomes apparent. The roughness on the NGV cascade, created by the deposition, could force transition to happen earlier along the airfoil surface; hence the transition to the fully rough regime would happen closer to the leading edge in the rough cases. The sudden spike in γ could be due to transition to the 61

80 fully rough regime, from the transition regime, for the roughness levels present on the smooth NGV cascade. While transition may be a cause of the sudden increase in γ, boundary layer separation could also play a role. The sudden spike in γ values seen in the smooth vane data, at exit Re c of about 1 million, could be caused by a region of separated flow. As the Reynolds number is increased the separation bubble could reattach to the vane, causing a decrease in γ. The sudden increase in γ is not replicated in the rough vane data possibly because roughness on the vanes could be creating a boundary layer that is not as likely to separate. As a result further tests must be conducted in order to characterize the boundary layer, to conclusively rule out or support the possibility of separation occurring along the airfoil. As a preliminary attempt to determine if massive separation was occurring for this NGV cascade tufts of yarn were attached to the suction surface of the NGV airfoils. The tufts of yarn were not attached to the pressure surface because of the low probability of the boundary layer separating. The boundary layer on the pressure side is less likely to separate because of the accelerating flow along the pressure side. Figure 28 shows the tufts of yarn at several Re c exit conditions. All of the tufts are following the flow path very well, and none of them are indicating massively separated flow at the Re c exit shown. These simple tests show that the flow is not massively separated. The yarn tuft separation test does have its limitations, however. A separation bubble could still be present which might be on the order of the yarn tuft size. If such a 62

81 separation bubble existed on the suction surface of the airfoil the tuft or the adhesive strip holding it in place, could affect how the separation bubble forms (or prevent it from forming all together). In short, the yarn tufts are not conclusive evidence that separation is not a contributing factor to the sudden spike in γ, seen in Figure 26 and Figure 27. Another method of boundary layer characterization is needed to rule out separation as a cause for the smooth vane γ spike at Rec = 1 million. Oil flow visualization techniques could be used along the suction surface of the airfoil to determine if a separation bubble forms. With the ability to view both the pressure and suction surfaces of the NGV cascade, Pressure Sensitive Paint (PSP) could be used to obtain pressure measurements over the entire surface of the airfoil. These pressure measurements would conclusively show where a separation bubble forms, if it forms at all. Figure 28 - Yarn tuft test for Reynolds number range from 0.5 to 1.5 million. Finally it is important to acknowledge a difference in the integrated nondimensional total pressure values from blade 2 to blade 3. Earlier, passage periodicity validation studies have shown that this type of variation should not exist, so variation in γ 63