Active Control of Hot, Supersonic Impinging Jets Using Microjets

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Electronic Theses, Treatises and Dissertations The Graduate School

2008

Active Control of Hot, SupersonicImpinging Jets Using MicrojetsSladana Lazic

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FLORIDA STATE UNIVERSITY

COLLEGE OF ENGINEERING

ACTIVE CONTROL OF HOT, SUPERSONIC IMPINGING

JETS USING MICROJETS

By

SLADANA LAZIC

A Thesis submitted to the

Department of Mechanical Engineering

in partial fulfillment of the

requirements for the degree of

Master of Science

Degree Awarded:

Summer Semester, 2008

ii

The members of the committee approve the thesis of Sladana Lazic defended on 3

rd of

June 2008.

________________________

Farrukh S. Alvi

Professor Directing Thesis

________________________

Juan C. Ordonez

Committee Member

________________________

Justin Schwartz

Committee Member

Approved:

__________________________________________________________________

Chiang Shih, Chairman, Department of Mechanical Engineering

__________________________________________________________________

C.J Chen, Dean, College of Engineering

The Office of Graduate Studies has verified and approved above named committee

members.

iii

ACKNOWLEDGEMENTS

I would like to thank my advisor Dr. Farrukh Alvi for providing me with the

opportunity to work at the Advance Aero Propulsion Laboratory. His guidance and

patience during the course of my project has been invaluable.

I also want to thank Dr. Rajan Kumar for answering all my questions patiently

and helping me understand the meaning behind the experiments done.

I am also thankfull to Dr. Juan Ordonez and Dr. Justin Schwartz for taking the

time to be on my thesis committee and also teaching me some of the fundamentals

necessary to reach my educational goals.

I am very grateful to Robert Avant in assisting me in machining anything

necessary for my experiments as well as for the invaluable advice when needed.

I appreciate all the support from my fellow graduate students at the Laboratory.

Last but not least I want to thank my family for the moral support and

encouragements, which were very much needed during the difficult times in my life as a

graduate student.

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TABLE OF CONTENTS

CHAPTER 1 ....................................................................................................................... 1

1.1 MOTIVATION......................................................................................................... 1

1.2 BACKGROUND ...................................................................................................... 3

1.2.1 SUPERSONIC FREE JET................................................................................. 3

1.2.2 IMPINGING JET FLOWFIELD ....................................................................... 7

1.2.3 CONTROL TECHNIQUES............................................................................. 14

1.2.4 OBJECTIVE OF THE CURRENT RESEARCH............................................ 16

CHAPTER 2 ..................................................................................................................... 17

2.1 FACILITY AND MODEL ..................................................................................... 18

2.1.1 GENERAL SETUP.......................................................................................... 18

2.1.2 CONVERGENT-DIVERGENT (C-D) NOZZLE ........................................... 22

2.1.3 LIFT PLATE.................................................................................................... 22

2.1.4 ACTIVE MICROJET CONTROL CONFIGURATION................................. 23

2.1.5 GROUND PLATE ........................................................................................... 25

2.2 MEASUREMENT AND INSTRUMENTATION ................................................. 26

2.2.1 MEAN PRESSURE MEASUREMENTS ....................................................... 27

2.2.2 UNSTEADY PRESSURE MEASUREMENTS.............................................. 28

2.2.3 TEMPERATURE MEASUREMENTS........................................................... 29

2.2.4 ACOUSTIC PRESSURE MEASUREMENTS ............................................... 30

2.2.5 DATA ACQUISITION.................................................................................... 31

2.2.6 EXPERIMENTAL UNCERTAINTY ............................................................. 31

2.2.7 TEST CONDITIONS....................................................................................... 32

CHAPTER 3 ..................................................................................................................... 33

3.1 MEAN PRESSURE MEASUREMENTS .............................................................. 33

3.1.1 EFFECT OF NPR ............................................................................................ 34

3.1.2 EFFECT OF TEMPERATURE....................................................................... 46

3.2 UNSTEADY PRESSURE LOADS ON THE GROUND PLATE AND LIFT

PLATE .......................................................................................................................... 50

3.2.2 UNSTEADY PRESSURE SPECTRA (SPL) .................................................. 58

3.2.3 OVERALL SOUND PRESSURE LEVELS (OASPL) ................................... 68

3.3 TEMPERATURE MEASUREMENTS.................................................................. 75

3.3.1 UNSTEADY TEMPERATURE LOADS........................................................ 75

3.3.2 MEAN TEMPERATURE DISTRIBUTION................................................... 82

3.4 LIFT LOSS ............................................................................................................. 85

CHAPTER 4 ..................................................................................................................... 90

APPENDICES...….……………………………………………………………………...94

REFERENCES...……………………………………………………………………….115

BIOGRAPHICAL SKETCH..………………………………………………………….121

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LIST OF FIGURES

Figure 1-1: F-35 Joint Strike Fighter (Courtesy of GEAviation.com)................................ 1

Figure 1-2: Schematic of a STOVL Aircraft in Hover Mode............................................. 3

Figure 1-3: Schematic of an Ideally Expanded Jet (Iyer 1999) .......................................... 4

Figure 1-4: Schematic of an Overexpanded Jet (Iyer 1999) ............................................... 5

Figure 1-5: Schematic of an Underexpanded Jet (Iyer 1999) ............................................. 6

Figure 1-6: Schematic of a Normal Impingement Flow on a Flat Plate (Iyer 1999) .......... 9

Figure 1-7: Schematic of a "Stagnation Bubble" Impinging Flow (Iyer 1999) ................ 12

Figure 1-8: Schematic of the Feedback Loop Mechanism (Alvi et al.)............................ 13

Figure 2-1: STOVL Supersonic Facility at the Advance Aero-Propulsion Laboratory ... 17

Figure 2-2: Schematic of the STOVL Facility Including Select Dimensions (Kumar et al.

2007) ......................................................................................................................... 18

Figure 2-3: Flow Control of the STOVL Supersonic Facility. (a) Manual Control Valve,

(b) Solenoid Shut Off Valve, (c) Dome Regulator, (d) Relief Valve, (e) Fisher

Controls Valve and (f) Heater................................................................................... 19

Figure 2-4: Control of High Pressure Air Supply using the Manual Ball Valve and the

Solenoid Control Valve............................................................................................. 20

Figure 2-5: Compressed Air Pressure is reduced using this TescomTM

Dome Regulator.21

Figure 2-6: Precise Regulation of the Air Pressure is achieved using a TescomTM

Low

Pressure Valve .......................................................................................................... 21

Figure 2-7: Mach 1.5, Convergent-Divergent Nozzle ...................................................... 22

Figure 2-8: Lift Plate with Steady Microjets .................................................................... 23

Figure 2-9: Lift Plate Flush Mounted to the Nozzle with Microjets Connected to the

Supply Stagnation Chamber ..................................................................................... 24

Figure 2-10: Ground Plate used for the Jet Impingement as well as for some of the

Measurement Techniques ......................................................................................... 25

Figure 2-11: Ground Plane Mounted on the Hydraulic Lift in Order to Simulate Hover 26

Figure 2-12: Schematic of t he Static Pressure Port Array on the Lift Plate .................... 27

Figure 2-13: Schematic of the Static Pressure Port Array on the Ground Plate ............... 27

Figure 2-14: Schematic of Unsteady Pressure Transducers mounted on the Ground Plate

................................................................................................................................... 28

Figure 2-15: Schematic of the Unsteady Pressure Transducer mounted on the Lift Plate 28

Figure 2-16: Schematic of the Thermocouples mounted in the Ground Plate.................. 29

Figure 2-17: Schematic of the Thermocouples mounted at the Lift Plate ........................ 29

Figure 2-18: Microphone Location with respect to the Lift Plate .................................... 30

Figure 3-1: Effect of Height on Mean Pressures of an Ideally Expanded Jet (NPR=3.7) at

TR=1.0 ...................................................................................................................... 36

Figure 3-2: Effect of Height on Mean Pressures of an Moderately Underexpanded Jet

(NPR=5) at TR=1.0................................................................................................... 37

Figure 3-3: Effect of NPR on a Cold Jet (TR=1.0) at h/d=2 ............................................ 38



Figure 3-6: Effect of Height on Mean Pressures of an Ideally Expanded Jet (NPR=3.7) at

TR=1.4 ...................................................................................................................... 41

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Figure 3-7: Effect of Height on Mean Pressures of an Moderately Underexpanded

(NPR=5.0) Jet at TR=1.4 .......................................................................................... 43

Figure 3-8: Effect of NPR on a Heated Jet (TR=1.4) at h/d=2......................................... 44

Figure 3-9: Effect of NPR on a Heated Jet (TR=1.4) at h/d=3......................................... 45

Figure 3-10: Effect of NPR on a Heated Jet (TR=1.4) at h/d=5....................................... 46

Figure 3-11: Effect of Temperature on an Ideally Expanded Jet (NPR=3.7) at (a) h/d=2,

(b) h/d=3, (c) h/d=5, and (d) h/d=8........................................................................... 47

Figure 3-12: Effect of Temperature on an Moderately Underexpanded Jet (NPR=5.0) at

(a) h/d=2, (b) h/d=3, (c) h/d=5, and (d) h/d= 8 ........................................... 49

Figure 3-13: Repeatability of Unsteady Pressure Data at TR=1.0.................................... 52

Figure 3-14: SPL at Different Locations in the Flow at h/d=3.5 and TR=1.0.................. 53

Figure 3-15: SPL at Different Locations in the Flow at h/d=3.5 and TR=1.6................. 54

Figure 3-16: Indication of Discrete Tone Frequency Prediction at h/d=3. 5 and TR=1.0 55

Figure 3-17: Impinging Tones as a Function of h/d and the Lines of Predicted

Frequencies for TR=1.0 ............................................................................................ 56

Figure 3-18: Indication of Discrete Tone Frequency Prediction at h/d=3.5 and TR=1.6. 57

Figure 3-19: Impinging Tones as a Function of h/d's and the Lines of Predicted

Frequencies at TR=1.6 .............................................................................................. 58

Figure 3-20: Influence of Temperature on the SPL Spectra in the Shear Layer of the Flow

(Lift Plate, x/d=2) ..................................................................................................... 59

Figure 3-21: Unsteady Pressure Spectra Obtained at the Lift Plate (x/d=2) at all TR's with

Frequency Scaled Using the Strouhal Number......................................................... 61

Figure 3-22: Influence of Temperature on the SPL Spectra of the Near-field Noise ....... 62

Figure 3-23: Near-field Narrowband Noise Spectra at Different TR's with Scaled

Frequency using the Strouhal Number ..................................................................... 63

Figure 3-24: Influence of Temperature on the SPL Spectra at the Stagnation Point of the

Flow, Ground Plane (x/d=0) ..................................................................................... 64

Figure 3-25: Effect of Microjet Control on Unsteady Pressure Spectra Obtained on the

Lift Plate (x/d=2) at h/d=2 for different jet temperatures (a) TR=1.0, (b) TR=1.2, (c)

TR=1.4 and TR=1.6 .................................................................................................. 65

Figure 3-26: Effect of Microjet Control on Near-field narrowband spectra at h/d=4 while

the temperature is varied (a) TR=1.0, (b) TR=1.2, (c) TR=1.4 and (d) TR=1.6 ...... 66

Figure 3-27: Effect of Microjet Control on the Stagnation Point of Jet Impingement at

h/d=4 for Different Temperatures Ratios (a) TR=1.0, (b) TR=1.2, (c) TR=1.4 and (d)

TR=1.6 ...................................................................................................................... 67

Figure 3-28: Effect of Temperature on the OASPL vs. h/d at different locations in the

flow ........................................................................................................................... 69

Figure 3-29: Influence of Temperature on the Average Fluctuating Pressure as h/d

changes at different locations in the flow (a) Ground plane (x/d=0), (b) Ground

Plane (x/d=1), (c) Lift Plate (x/d=2), and (d) Microphone (x/d=15) ........................ 70

Figure 3-30: Effect of Microjet Control on Pressure Fluctuation Intensities at the Ground

Plane (x/d=0) at (a) TR=1.0 and (b) TR=1.6 ............................................................ 72

Figure 3-31: Effect of Temperature on Delta OASPL as a Function of h/d at Different

Measurement Locations in the Flow (a) Ground Plane (x/d=0), (b) Ground Plane

(x/d=1), (c) Lift Plate (x/d=2), and (d) Microphone (x/d=15) .................................. 74

vii

Figure 3-32: Comparison of the unsteady thermal load distribution and the unsteady

pressure load distribution in the frequency domain at TR=1.0................................. 76

Figure 3-33: Comparison of the unsteady thermal load distribution and the unsteady

pressure load distribution in the frequency domain at TR=1.6................................. 77

Figure 3-34: Effect of jet temperature on the unsteady thermal loads at different locations

on the ground plane (a) x/d=0, (b) x/d=0.3, (c) x/d=1.2 and (d) x/d=1.8 ................. 79

Figure 3-35: Effect of Control on Unsteady Thermal Loads at the Stagnation Point (x/d=0)

on the Ground Plane at TR=1.0 and h/d=4 ............................................................... 80

Figure 3-36: Effect of Control on Unsteady Thermal Loads at the Stagnation Point on the

Ground Plane for a Hot Jet at TR=1.6 and h/d=4 ..................................................... 81

Figure 3-37: Effect of Temperature Ratio and Height (h/d) on the Stagnation Recovery

Factor ........................................................................................................................ 83

Figure 3-38: Effect of Control on the Vacuum Pressure Distribution on the Lift Plate at (a)

TR=1.0 and (b) TR=1.4 ............................................................................................ 86

Figure 3-39: Variation of Lift Loss with nozzle-to-plate distance at (a) TR=1.0 and (b)

TR=1.4 ...................................................................................................................... 87 Figure A- 1: Effect of NPR on a Cold Jet (TR=1.0) at h/d=4........................................... 94

Figure A- 2: Effect of NPR on a Cold Jet (TR=1.0) at h/d=6........................................... 95

Figure A- 3: Effect of NPR on a Cold Jet (TR=1.0) at h/d=8........................................... 96

Figure A- 4: Effect of NPR on a Cold Jet (TR=1.0) at h/d=12......................................... 97

Figure A- 5: Effect of NPR on a Heated Jet (TR=1.4) at h/d=4 ....................................... 98

Figure A- 6: Effect of NPR on a Heated Jet (TR=1.4) at h/d=6 ....................................... 99

Figure A- 7: Effect of NPR on a Heated Jet (TR=1.4) at h/d=8 ..................................... 100

Figure A- 8: Effect of NPR on a Heated Jet (TR=1.4) at h/d=12 ................................... 101

Figure A- 9: Effect of Temperature on an Ideally Expanded Jet (NPR=3.7) at (a) h/d=4,

(b) h/d=6, (c) h/d=12, and (d) Free Jet.................................................................... 102

Figure A- 10: Effect of Temperature on a Moderately Underexpanded Jet (NPR=5.0) at (a)

h/d=4, (b) h/d=6, (c) h/d=12, and (d) Free Jet ........................................................ 103

Figure B- 1: Influence of Temperature on the SPL Spectra at the Lift Plate (x/d=3)..... 104

Figure B- 2: Influece of Temperature on the SPL Spectra at the Ground Plate (x/d=1) 105

Figure B- 3: Unsteady Pressure Spectra Obtained at the Ground Plate (x/d=1) at all TR's

with Frequency scaled using the Strouhal Number ................................................ 106

Figure B- 4: Influence of Temperature on the SPL Spectra at the Ground Plate (x/d=3)

................................................................................................................................. 107

Figure B- 5: Unsteady Pressure Spectra obtained at the Ground Plate (x/d=2) at all TR's

with Frequency scaled using the Strouhal Number ................................................ 108

Figure B- 6: Effect of Temperature on Delta OASPL as a function of h/d at different

measurement locations in the Flow (a) Ground Plate (x/d=0), (b) Ground Plate

(x/d=1), (c) Lift Plate (x/d=2), and (d) Microphone (x/d=15) ................................ 109

Figure C- 1: Radial Distribution of the Recovery Factor at TR=1.0 .............................. 110

Figure C- 2: Radial Distribution of the Recovery Factor at TR=1.4 .............................. 111

Figure C- 3: Effect of Control on the Radial Recovery Factor Distribution at TR=1.0 . 112

Figure C- 4: Effect of Control on the Radial Recovery Factor Distribution at TR=1.4 . 113

Figure D- 1: Effect of Temperature on the Vacuum Pressure Distribution on the Lift Plate

for the (a) no control case and (b) controlled case............................................... 114

viii

Figure D- 2: Effect of Temperature on the Lift Loss variation versus nozzle-to-plate

distance for the (a) no control case and (b) controlled case.................................... 114

ix

ABSTRACT

Supersonic impinging jets, similar to the jets issued from a short takeoff and

vertical landing (STOVL) aircraft, generate a highly unsteady flow with high unsteady

pressure and thermal loads on the aircraft structure as well as the landing surface. These

high-pressure, high-temperature and acoustic loads are also accompanied by dramatic lift

loss, severe ground erosion and hot gas ingestion in the engine inlets. Previous studies

have concentrated on characterizing the impingement flow and its control for cold jets,

i.e. operating at ambient temperatures. They have shown that one of the major

characteristics of a supersonic impinging jet is the dominance of the feedback loop

mechanism. Previous work has also shown that active microjet control is successful at

attenuating the feedback loop and therefore the negative effects associated with it.

The current studies attempt to examine and investigate the flow properties of a

hot supersonic impinging jet issuing from a convergent-divergent, Mach 1.5 nozzle and

operating at more realistic, higher temperatures,. This ideally expanded jet was heated up

to a stagnation temperature of ~500K. The jet is impinging on a flat plate, called the

ground plate that is appropriately larger than the nozzle exit diameter. The ground plate

can be moved vertically in order to simulate different hover heights.

In order to compare the properties of a cold and a heated impinging jet, mean

pressure and unsteady pressure measurements, temperature measurements as well as

acoustic measurements were obtained. The mean and unsteady pressure measurements as

well as temperature measurements were performed on the lift plate (representing the

undersurface of a STOVL aircraft) as well as on the ground plate. Acoustic near-field

measurements were obtained using a microphone placed at 15-diameters away from the

nozzle exit.

Active microjet control was implemented as a way to attenuate the adverse effects

of a jet impinging on a flat surface. It has already been shown that microjets are very

effective when introduced to an impinging flow of a cold supersonic jet. Another aim of

this study is to explore how effective microjet control is when the stagnation temperature

of the primary jet is heated.

The results clearly indicate that when the primary jet is heated, the pressure

fluctuations and the associated unsteady loads, are substantially higher then when the jet

x

is cold. These high unsteady loads also persist over larger nozzle to plate distances. The

hover lift loss at high temperatures increases dramatically as well, from ~50% of the

primary jet thrust at cold temperatures to an astounding ~75% of the primary jet thrust

when the jet is heated. The temperature recovery factor is strongly dependent on the

nozzle to plate distance and the temperature of the jet. There is an indication of an

increase in entrainment of ambient air when the jet is heated. Additionally the unsteady

thermal loads seem to increase in frequency as the stagnation temperature of the jet

increases.

These results show that the adverse side effects of an impinging supersonic jet are

even more dramatic when the jet is at higher temperatures – a trend that is expected to

continue as the temperatures are increased further to real jet exhaust conditions.

However, this study also demonstrates that the activation of microjets can provide an

effective way of reducing these negative effects even when the jet is heated. The pressure

fluctuations have been drastically reduced, where the discrete impinging tones have been

attenuated or even eliminated at both the cold and hot conditions. The overall pressure

levels on the ground plane have been reduced up to 20dB and on the lift plate up to 15dB

at small nozzle to ground plane distances while the jet was heated. Additionally up to

21% of the lift loss has been recovered at cold temperature jets and an astounding 35% of

the lift loss was recovered when the jet is heated. The temperature recovery factor

indicates a similar trend as the lift loss, which is that the entrainment of the ambient air is

decreased when the microjets are applied. The thermal unsteady loads have been

attenuated as well.

In summary, this study demonstrates that the adverse effects of impinging

supersonic jets are even more pronounced when the jet is heated, however microjet

control is very effective at both cold and hot conditions. These dramatic reductions due to

microjet control are achieved using microjets with a mass flow rate less than 0.5% of the

primary jet flux.

1

CHAPTER 1

INTRODUCTION

1.1 MOTIVATION

Impinging jets are defined as jets issuing from a nozzle, impinging on a solid

surface, in general significantly larger than the jet diameter. Few examples are the launch

of a rocket, the thrust vector control of solid rocket motor, cooling of electrical

components and the takeoff and landing of a STOVL aircraft. In order to efficiently

design and operate impinging jets, it is important to understand the flow characteristics of

these jets.

Figure 1-1: F-35 Joint Strike Fighter (Courtesy of GEAviation.com)

Although fundamental in nature, the research in this experiment is large part

geared towards exploring the negative effects of the STOVL aircraft in hover mode. A

good example of a STOVL aircraft is the F-35 Joint Strike Fighter (JSF) aircraft seen in

Figure 1-1. The STOVL aircraft is in hover mode during takeoff and landing. During

hover the hot supersonic jets, which are issued from the aircraft to produce thrust,

impinge on the ground. These high unsteady loads and oscillatory flows, Figure 1-2,

create a very harsh environment the ground surface causing ground erosion (Iyer 1999).

Other negative effects are sonic fatigue leading to failure of the nearby aircraft structure;

Impinging

Jet Flow

2

lift loss due to enhanced entrainment; hot gas ingestion into the engine inlets and noise

pollution due to the increased jet noise in the neighborhood of the impinging jets.

These adverse effects are created by the flow properties that occur in an

impinging jet as depicted in Figure 1-2. (A schematic of a STOVL aircraft with multiple

jets and various regions of problems is shown in Figure 1-2). These problematic effects

are briefly discussed below.

As the jets impinge on the ground, the flow moves away radially from the

impingement region, forming a radial wall jet. Wall jets are usually accompanied by high

unsteady pressure loads which together with the high unsteady thermal and pressure loads

created by the primary jet, can cause severe ground erosion. In this experiment, a single

jet arrangement is considered. However as seen in Figure 1-2, STOVL aircrafts are

usually equipped with multiple supersonic jets. In the case when two or more jets are in

close proximity, the radial wall jets created by each jet flow interact with each other and

form a “fountain”, which rises upward towards the aircraft. This fountain of high velocity

and high temperature fluid can cause high unsteady loads on the undersurface of the

aircraft as well as an unwanted increase in the surface temperature of the aircraft.

Additionally, if this hot air is ingested by the engine inlets it can lead to a deficiency in

aircraft engine performance due to the increase in inlet temperature. That is known as the

Hot Gas Ingestion (HGI) problem.

The high noise created by the impinging jet does not only lead to noise pollution

in the near-field of the jet but can cause acoustic loading on the structural elements of the

aircraft. This can ultimately lead to damage of the aircraft surfaces due to sonic fatigue.

Lift loss has been extensively studied since it can be very detrimental to the

performance of the aircraft (Margason et al. 1997). In case of STOVL aircraft, lift loss is

created due to the entrainment of the ambient air by the primary jets. This high velocity

entrainment of ambient air induces low surface pressures on the undersurface of the

aircraft which are lowered even further when the jet is at very small nozzle-to-ground

plate distances. These low surface pressures cause a force that acts in the opposite

direction of the jet thrust, and is commonly known as the suckdown force. This suckdown

force is responsible for the lift loss, which can be as high as 60% of the primary jet thrust

when the ground plane is very close to the jet exit (Krothapalli et al. 1999).

3

It is very obvious that it would be beneficial to control this unsteadiness in the

impinging jets and reduce all these adverse effects while maintaining the propulsive

force.

Jet Entrainment Flow

Fountain Upwash Flow

Ground Erosion Region

Impinging Jet Flow

Wall - Jet Flow

Unsteady Structural

Loads and Lift Loss




Impinging Jet Flow

Wall - Jet Flow

Unsteady Structural

Loads and Lift Loss




Impinging Jet Flow

Wall - Jet Flow

Unsteady Structural

Loads and Lift Loss

Figure 1-2: Schematic of a STOVL Aircraft in Hover Mode

In order to show how the control proposed in this experiment affects the

supersonic, impinging jets, whether it is isothermal or heated, the fundamental

characteristics of the flow must be discussed. In the following section the basic

characteristics of the supersonic free and impinging jet operating at design and off-design

conditions will be reviewed by considering some of the previous work done in this field.

Changes in those characteristics when the jet is heated will be briefly discussed as well.

1.2 BACKGROUND

1.2.1 SUPERSONIC FREE JET

In order to provide the propulsive force or thrust necessary for powered flight, a

nozzle is necessary. In case of STOVL aircraft axisymmetric convergent and convergent-

divergent (C-D) nozzles are used to produce this thrust. A convergent nozzle can at most

reach sonic conditions at the exit. In this experiment all the data was obtained at

4

supersonic conditions using a C-D nozzle. In a C-D nozzle the wall contour on the inlet

side converges to a minimum area also known as the throat. After that it diverges from

the throat to the nozzle exit. The convergent portion of the nozzle accelerates the

subsonic flow to sonic conditions at the throat, then it is further accelerated to supersonic

speeds in the divergent region until it reaches the nozzle exit. (Anderson 1990). The

design Mach number at the exit depends on the ratio of the nozzle exit area and the throat

area.

Each C-D nozzle when operated at ideal conditions is designed for one Mach

number. Whether the nozzle operates at the ideal condition depends on the nozzle

pressure ratio of the flow. The nozzle pressure ratio or NPR is defined as the stagnation

pressure of the jet divided by the ambient pressure. Design NPR is the pressure ratio at

which the nozzle operates at the ideal, design condition. The critical NPR is the minimum

NPR necessary in order to achieve supersonic flow at the exit of the nozzle.

IDEALLY EXPANDED JET

When the NPR is equal to the design pressure ratio for that particular nozzle, the

jet is said to be ideally expanded. A schematic of such a jet is shown in Figure 1-3. In this

case the pressure at the nozzle exit is equal to the ambient pressure. In the majority of the

experiments performed in the current setup the jet was nominally ideally expanded.

Figure 1-3: Schematic of an Ideally Expanded Jet (Iyer 1999)

5

OFF-DESIGN JETS

A C-D Nozzle can be operated at off-design conditions. When that is the case,

two different shock patterns in the core region are observed (Wishart 1995). First if the

nozzle pressure ratio is greater than the critical pressure but below the design pressure

ratio for that nozzle, the flow is considered overexpanded. Here, the exit pressure is lower

than the ambient pressure. In order to adjust to the ambient pressure the jet flow has to go

through a series of oblique shock waves. If the jet is highly overexpanded then a normal

shock also known as a Mach disc forms downstream of the nozzle (Garg 2001). A

schematic of a highly overexpanded jet is shown in Figure 1-4.

Figure 1-4: Schematic of an Overexpanded Jet (Iyer 1999)

If the exit pressure is higher than the ambient pressure, meaning that the nozzle is

operated at a NPR higher than the design NPR, the jet flow becomes underexpanded. In

order for the exit pressure to equal to the back pressure (ambient pressure) the jet flow is

expanded through a centered expansion fan, as shown in Figure 1-5. If the NPR is further

increased and the flow becomes highly underexpanded, the shock patterns in the flow

become even more complicated (Iyer 1999).

6

Figure 1-5: Schematic of an Underexpanded Jet (Iyer 1999)

HEATED JET

Much work has been done in order to investigate the features of a cold or

isothermal supersonic jet. Isothermal or cold jets are jets where the stagnation

temperature of the jet (To) is equal to the ambient temperature (Tamb). In those cases the

temperature ratio (TR), defined as stagnation temperature divided by ambient

temperature, is equal to one. However, most jets in aircrafts operate at very high

stagnation temperatures and therefore it is important to explore heated supersonic jets.

Some research has been done on the effect of temperature on supersonic free jets.

Lau (1981) reported that the spreading rate of the jet increases for supersonic jets when

the temperature ratio of the jet is increased above one. Additionally, he noticed a

shortening of the potential core of the jet when the jet was heated. Seiner et al. (1992)

also report a shortening of the potential core of the jet as well as an increase in mixing.

They also find that there is no significant change in the spreading rate of the shear layer

when the temperature of the jet is increased. There is however a significant decrease in jet

half-width when the temperature ratios are increased, up until about 10 jet diameters.

7

This suddenly changes at downstream locations when the jet half-width spreading starts

increasing.

Most recently Wishart and Krothapalli (1994) reported some similar results. They

have found that with increased temperature ratio there is no significant effect on the

development of the jet shear layer but there is a significant increase of the mixing of the

jet in the downstream region. For jets at off-design conditions the temperature ratio

seems to have no effect on the structure of the shock cell pattern in the initial region but

they show that the decay of the centerline Mach number is enhanced. At the end of the

potential core there is a slight difference due to the shear layers merging sooner for

higher temperature jets.

The above results were obtained using free jets at different temperature ratios. In

order to be beneficial to the STOVL aircraft design, more research is needed to determine

how the temperature of the jet influences the characteristics of an impinging jet. This is

the focus of the present study.

1.2.2 IMPINGING JET FLOWFIELD

Previous work has been done on isothermal, impinging, supersonic jets by

Donaldson and Snedeker (1971), Lamont and Hunt (1980), Powell (1988), Tam and

Ahuja (1990), Messersmith (1995) and Alvi and Iyer (1999), among others. According to

the study of Donaldson and Snedeker (1971) the flowfield produced by a turbulent,

axially symmetric jet, impinging perpendicular on a plate can be described by three

primary flow regimes (Figure 1-6):

1. The Free Jet Region or Primary Jet Flow which is upstream of any

strong local effects of impingement and therefore shows the same

characteristics as the before described free jet flow (Iyer 1999).

2. The Impingement Region/Zone which is around the stagnation point

of impingement where the strong, essentially inviscid, interaction of

the jet with the impingement plate causes a change in flow direction

(Iyer 1999).

8

3. The Wall Jet Region which is the radial flow along the impingement

surface beyond the point at which the strong interactions of

impingement produce local effects (Iyer 1999).

The flow of an impinging jet in the primary jet region behaves similar to a free jet

and therefore has already been describe in 1.2.1. The impingement region and wall jet

region will be explained further, in the following.

IMPINGEMENT REGION AND WALL JET REGION

The impingement region is very complex due to an intricate mix of shocks,

supersonic and subsonic regions. As the supersonic, underexpanded and axisymmetric jet

impinges on the ground plane, a plate shock is created. This plate shock interacts with the

shocks in the jet. The supersonic primary jet is reduced to subsonic velocity after it passes

through the shock. Once the jet reaches the ground plane the direction of the flow is

changed and a radial flow (also known as a wall jet) is created leading, away from the

stagnation point. The radial flow goes through a combination of expansion and

compression waves resulting in the secondary pressure peaks in the wall jet. These

complex interactions between the ground-plate shock and the jet shock, also known as the

normal impinging flow (Figure 1-6) create unsteadiness in the impingement region and

wall jet region. In order to understand this complex impinging flow and how to control its

unsteady nature many researchers have explored the impingement region. Donaldson and

Snedeker (1971), Gummer and Hunt (1971), Gubanova et al. (1973), Carling and Hunt

(1974), Kalghatgi and Hunt (1976) and Messersmith et al. (1995) have visualized this

complex flowfield using the shadowgraph technique and additionally they have

visualized the surface flow using lampblack and Day-Glo paints. They explored it even

further using surface pressure measurements along with temperature measurements.

Two categories of impinging flow have been defined, normal flow as described

above and the “stagnation bubble” flow which can be seen in Figure 1-7. In the normal

impingement flow the peak pressure occurs at the point where the centerline of the jet

impinges on the ground as can be seen in the graph in Figure 1-6.

9

Why a stagnation bubble, also known as a recirculation region is formed in the

impingement zone at certain conditions is a question that many studies have tried to

explain. According to Gummer and Hunt (1971), Gubanova et al. (1973), Ginzburg et al.

(1973) and Kalghatgi and Hunt (1976), the recirculating region is created due to the

primary jet flow being separated from the center of the ground plate surface by a thin

layer of fluid. Due to this, the primary jet impinges on the ground plane at a radial

distance from the stagnation point leading to maximum pressure at an annular location

away from the ground plate center, as see in the graph in Figure 1-7.

Figure 1-6: Schematic of a Normal Impingement Flow on a Flat Plate (Iyer 1999)

This recirculating region results in enhanced convection, leading to lower

temperatures as shown by Messersmith et al. (1995). Ginzburg et al (1973) and

10

Gubanova et al. (1973) performed experiments using C-D nozzles with different

divergence angles. Both used surface pressure measurements, schlieren, pressure probe

measurements to confirm the presence of reversible flow. They described the mechanism

of creation of the stagnation bubble as follows. When the jet shock and plate shock

intersect a tail shock is formed. This leads to a tangential discontinuity in velocity also

known as the slipline. This slipline divides the flow into two regions with the outer one

having a higher total pressure and higher velocity. While the slipline is not in contact

with the impinging plate, the mixing between the two regions occurs at the slipline. At

certain plate heights the slipline intersects with the ground plate and only some parts of

the fluid are drawn into the mixing zone and leave the central portion of the impingement

zone. The rest of the fluid, which is unable to overcome the pressure difference,

accumulates at the central point of the plate and forms a recirculating region also known

as the stagnation bubble. These studies have shown that the Mach number inside the

stagnation bubble can reach 0.4 and the bubble diameter can be as large as 80% of the jet

diameter. Ginzburg et al. (1973) noted the absence of the stagnation bubble at very small

y/d. The oscillating nature of the flow causes the appearance and dispersion of the

recirculating region between two and three diameters. According to Ginzburg et al.

(1973), once y/d=3.4 is reached, the flow becomes more stable and the appearance of the

stagnation bubble is well defined at this height. Gubanova et al. (1973) observed similar

results but in their case the stagnation bubble presence is well pronounced at y/d=3.

Kalghatgi and Hunt (1976) conducted similar test with nozzles of varying Mach

numbers. They proposed that the bubbles are caused by the intersection of the plate shock

and the jet shock waves either right before crossing the jet centerline or right after

crossing the jet centerline. They implied that the bubble can also be a product of some

surface imperfection, like a scratch in the nozzle surface and therefore can be easily

eliminated by publishing the surface.

Donaldson and Snedeker (1971) and Gummer and Hunt (1971) anticipated that

the stagnation bubble in an impinging jet is likely to be present when the impinging plate

is stationed around the position where the normal shock or a Mach disc are observed in a

free jet.

11

More recently, in the case where the jet is moderately underexpanded as at

NPR=5, Iyer (1999) found that there is evidence of a stagnation bubble at h/d=2, 3, and 5.

Garg (2001) confirmed Iyer’s results for the ideally expanded case at NPR=3.7, for a

converging-diverging nozzle. In the highly underexpanded case Garg also witnessed a

“flat” annular profile in the surface pressure distribution for h/d=3, 3.5, 4, and 5. A “flat”

annular profile in the surface pressure distribution is a good indication of the presence of

a stagnation bubble in the flow.

As shown in this review of previous studies, the existence of the stagnation

bubble in impinging flow is very sensitive to various experimental conditions, such as

plate distance, surface as well as divergence of the nozzle and slight variances in

unsteadiness and oscillation of the flow. The current objective is to explore what effect an

increase in the stagnation temperature of the jet will have on the presence and form of the

stagnation bubble.

12

Figure 1-7: Schematic of a "Stagnation Bubble" Impinging Flow (Iyer 1999)

Another mechanism that is associated with the impinging flow and is mainly

responsible for the high flow unsteadiness and noise in these flows, is the feedback loop.

In order to understand and control the impinging flow, understanding the feedback loop is

vital and therefore will be further described in the following section.

FEEDBACK LOOP

The highly unsteady nature of supersonic impinging flow field can be explained

by a feedback mechanism (Figure 1-8). The feedback loop is started with instability

waves that form in the shear layer of the jet near the nozzle exit. They grow as they travel

downstream and develop into large scale vertical structures. These structures can be

13

visible in flow visualization pictures like the shadowgraph technique (Alvi and Iyer

1999). As these structures travel downstream in the outer jet shear layer they entrain the

ambient air which leads to lift loss and mixing, as will be discussed further in the

following chapters. Once the large scale structures impinge on the ground plane they

generate acoustic waves which travel upstream in the ambient medium. These acoustic

waves in turn excite the shear layer near the nozzle exit resulting in the generation of

instability waves thus closing the feedback loop.

Figure 1-8: Schematic of the Feedback Loop Mechanism (Alvi et al.)

In the 1950s Powell was the first one to describe the physics of the formation of

the feedback loop in free supersonic jets and additionally he provided a formula for

predicting the frequency of discrete tones (screech tones) generated by this mechanism.

Powell’s feedback formula is provided in Equation 3.

14

These discrete tones are called impinging tones when the jet is impinging on a flat

plate. It was shown by Neuweth (1974), Ho and Noisseir (1981), Powell (1988), Tam and

Ahuja (1990), Henderson and Powell (1993), and Krothapalli et al. (1999) that these

impinging tones are also generated by the feedback loop and therefore their frequency

can be predicted using Powell’s formula (Equation 3).

Due to the difficulty of measuring the convective velocity Ci of the downstream

traveling large scale structures most researchers assume the convective velocity to be 60-

70% of the jet flow velocity when using Powell’s formula. Krothapalli et al. (1999)

obtained more accurate values for Ci using the PIV technique. They reported that the

convection velocity of the large scale structures for the free jet to be about 60% of the

mean jet exit velocity. For impinging jets, the convection velocity was about 50% of the

primary jet velocity and it increases with plate height. They also suggested that for an

impinging jet the convection velocity increases when the jet is underexpanded or

overexpanded due to the presence of a more pronounced shock cell structures.

1.2.3 CONTROL TECHNIQUES

The unsteadiness of the jet flow field leads to many adverse effects on the aircraft

as discussed earlier. Increase in lift loss and noise pollution added to ground erosion and

structural fatigue can cause serious issues to the aircraft as well as to the surroundings.

An adequate technique that can eliminate or contain these effects remains a problem that

is to be addressed when designing an efficient supersonic STOVL aircraft. The feedback

loop is the mechanism behind the flow unsteadiness that is the cause for these negative

side affects. Therefore to control or attenuate these effects, the feedback loop has to be

weakened or interrupted.

Garg (2001) has summarized the possibilities used by researches for attenuating

the feedback loop, which are:

- Intercept the upstream/downstream propagating acoustic waves so that

they can not complete the feedback loop.

- Manipulate the shear layer near the nozzle exit so that they are less

receptive to the acoustic wave excitation.

15

- Disrupt the interaction between the large structures and the acoustic

field by pulsating high energy source that excites the nozzle shear layer.

One of the examples of methods based on these ideas is two plates placed normal to the

jet centerline by Karamcheti et al. (1969) to suppress the edge tones in low speed flows.

Samimy et al. (1993) placed small tabs protruding normally into the jet flow at the nozzle

exit of an axisymmetric jet ranging in Mach numbers of 0.3-1.81. They achieved

reasonable reduction in OASPL (~6.4dB) for a 4-tab configuration. They suggested that

the jet spread rate is enhanced by mixing and leads to destruction of the flow symmetry at

the nozzle exit which is necessary for the feedback loop to be established (Tanna 1977).

However use of tabs leads to thrust loss which can be a detrimental to the aircraft design.

Elvarasan et al (2000) managed to recover up to 16 % of Lift Loss by introducing a

control plate just outside the nozzle exit of a supersonic impinging jet. Additionally the

near-field OASPL was reduced by 6-7 dB, which suggests that the feedback loop has

been weakened. Kweon et al. (2006) attenuated the screech tones and the broadband

level noise over a range of nozzle pressure ratios by placing two thin wires orthogonally

to the jet axis. All of these techniques have proven to be reasonably effective but only

over a limited range of geometry and require major modifications to the aircraft design.

More recently, a new way to interrupt and weaken the feedback mechanism has

been developed by Alvi et al. (2003). They have shown that an array of high momentum

microjets placed near the nozzle exit can effectively attenuate the high energy impinging

tones associated with the feedback loop in an impinging jet. Lou et al. (2003), with the

help of PIV measurements, have shown that microjets introduce streamwise vorticity at

the expense of azimuthal vorticity of the main jet. This leads to an attenuation of the

feedback loop. These experiments have demonstrated, how effective high momentum

microjets can be. They are very effective at reducing or even eliminating the negative

effects associated with the feedback loop in an impinging, supersonic flow field.

However, these experiments were performed with the jet at ambient temperatures, which

are not a realistic simulation of jets used in STOVL aircraft. This leads us to the objective

of the current research.

16

1.2.4 OBJECTIVE OF THE CURRENT RESEARCH

The present objective is to explore a more realistic case of supersonic, impinging

jet flow, in order to provide a control solution for the adverse effects created by the

feedback loop associated with this kind of flow. How an increase in stagnation

temperature of the jet affects the unsteadiness of the flow field or the feedback loop can

provide some insight in the effective use of microjet control at heated conditions.

In order to characterize and understand the unsteady behavior of supersonic high

temperature impinging flow field, the jet was heated up to a total temperature of ~500K.

Mean pressure, unsteady pressure, temperature and acoustic measurements were

performed at several stagnation temperatures, varying plate-to-nozzle distances as well as

in some cases several nozzle pressure ratios.

Effectiveness of microjet control was explored at different temperature ratios as

well. As shown before for cold jets, microjet control is found to be very effective in

reducing the unsteadiness in the flowfield by interrupting the feedback loop in a cold,

supersonic, impinging jet.

17

CHAPTER 2

EXPERIMENTAL SETUP

The experiments described in this study were performed at the Short Take-Off

Vertical Landing (STOVL) supersonic facility of Advanced Aero-Propulsion Laboratory

(AAPL), Department of Mechanical Engineering, at the Florida State University. A

picture of the STOVL facility can be seen in Figure 2-1. More details about this facility

can also be found in Iyer (1999), Garg (2001) and Lou (2005).

Figure 2-1: STOVL Supersonic Facility at the Advance Aero-Propulsion Laboratory

18

2.1 FACILITY AND MODEL

The STOVL facility is used to simulate the STOVL aircraft during hover and it is

capable of running single and multiple jets at design and off-design conditions up to

M=2.2. A schematic of the STOVL facility is presented in Figure 2-2. Details of the

experiment setup as well as the various measurement methods are to follow

Figure 2-2: Schematic of the STOVL Facility Including Select Dimensions (Kumar et al. 2007)

2.1.1 GENERAL SETUP

The supersonic STOVL facility is supplied with dry, high pressure air from large

storage tanks with a total volume of 10m3 and a maximum pressure of 2000psi. The

compressed air is provided by a MakoTM

compressor (Model 5436-60E3). Before the

ambient air enters the compressor it is dried and filtered in order to remove moisture and

19

particulates. This high pressure air is supplied to the STOVL facility via steel piping. A

picture of the overall flow control inside the STOVL facility is shown in Figure 2-3.

Figure 2-3: Flow Control of the STOVL Supersonic Facility. (a) Manual Control Valve, (b) Solenoid

Shut Off Valve, (c) Dome Regulator, (d) Relief Valve, (e) Fisher Controls Valve and (f) Heater

The supply from the high pressure tanks can be shut off using a manual ball valve

and a solenoid valve installed right after the manual shut off (Figure 2-4). The solenoid

valve can be controlled from the control room for additional safety. A TescomTM

dome

regulator (Figure 2-5) installed in series after the solenoid valve is used to reduce the high

pressure of the compressed air. The regulator is designed so that the operator can provide

an input pressure from a nitrogen bottle, to the dome side of the regulator. The regulator

is set to balance the output pressure so that it is equal to the input pressure. The output

pressure in this experiment was varied from 150-300 psi according to the conditions of

each individual experiment. Once the pressure was selected it is kept constant at one

(a)

(b)

(c)

(d)

(e)

(f)

20

pressure throughout the experiment. The regulator is supplying a FisherTM

low pressure

valve (Figure 2-6) with this air at a constant pressure. The FisherTM

valve is used to more

precisely control the pressure that is being supplied to the facility. The opening of this

valve and therefore the pressure of the air can be adjusted remotely to ensure that the jet

is constantly operated at the appropriate nozzle pressure ratio, NPR. In order to assure

that the conditions of the experiment can be repeated and that the temperature of the air is

kept constant throughout the experiment duration the air is routed through an electrical

heater. The voltage input to the heater can be controlled from the control computer

allowing for the temperature in the heater to be adjusted during the experiment. The

heater not only assures that the jet can be run at isothermal conditions but also allows us

to heat the jet up to ~500K in order to achieve more realistic jet temperature conditions.

Figure 2-4: Control of High Pressure Air Supply using the Manual Ball Valve and the Solenoid

Control Valve

21

Figure 2-5: Compressed Air Pressure is reduced using this TescomTM

Dome Regulator.

Figure 2-6: Precise Regulation of the Air Pressure is achieved using a TescomTM

Low Pressure Valve

Once the air has been conditioned to the appropriate stagnation temperature and

pressure it is fed into a stagnation chamber. The role of the stagnation chamber is to

allow the air to become laminar before it enters the nozzle. A thermocouple and an

OmegaTM

pressure transducer allows for constant monitoring of stagnation temperature

and stagnation pressure of the jet, respectively. Air at a specific stagnation temperature

22

and pressure enters a C-D nozzle. More details about the nozzle will be given in the

following section.

2.1.2 CONVERGENT-DIVERGENT (C-D) NOZZLE

The air leaving the stagnation chamber is issued into a converging-diverging,

axisymmetric nozzle. The throat and exit diameters (d, de) of the nozzle are 2.54cm and

2.75cm, respectively. In Figure 2-7 a C-D nozzle is shown that when run at ideal

conditions can reach a design Mach number of 1.5. The diverging section of the nozzle is

straight conic with a 3˚ divergence angle from the throat to the nozzle exit. The flow

coming out of this nozzle is considered ideally expanded when the nozzle pressure ratio

(NPR, where NPR=stagnation pressure/ambient pressure) is 3.70.

Figure 2-7: Mach 1.5, Convergent-Divergent Nozzle

The nozzle is connected to a circular plate called the ‘lift plate’. The purpose and

design of the lift plate will be explained as follows.

2.1.3 LIFT PLATE

A circular plate, shown in Figure 2-8, was flush mounted with the nozzle exit.

The diameter of the lift plate is 25.4cm, approximately 10 times the nozzle throat

diameter, d. The lift plate is used to represent the underside of an aircraft as well as for

23

mounting of different measurement transducers, which will be described in detail in later

sections of this chapter. A central hole, equal to the nozzle exit diameter, allows for the

jet to be issued from the nozzle exit. Additionally, microjets are mounted around the

nozzle exit on the lift plate. More details about the microjets are provided in the

following section.

Figure 2-8: Lift Plate with Steady Microjets

2.1.4 ACTIVE MICROJET CONTROL CONFIGURATION

Nitrogen from compressed nitrogen tanks is supplied first to a large stagnation

chamber where the flow is settled and the stagnation pressure of the microjets is

measured using an OmegaTM

pressure transducer. Nitrogen is issued from the stagnation

Microjets

KuliteTM

Slot

24

chamber to four small plenum chambers in order for the flow to become as stable as

possible. Each one of these plenum chambers is connected to four stainless steel tubes

(diameter of 400µm), which are used as microjets in this setup. These stainless steel tubes

are mounted circumferentially around the nozzle exit on the lift plate at a 60˚ inclination

angle with respect to the main jet axis. The pressure at which these sixteen microjets are

issuing the nitrogen into the jet flow can be controlled throughout the experiment by

adjusting the regulator on the nitrogen tank supply. The microjet set up can be observed

in Figure 2-9.

Figure 2-9: Lift Plate Flush Mounted to the Nozzle with Microjets Connected to the Supply

Stagnation Chamber

So far we have described how the supersonic, heated flow is achieved and the

microjet control setup has been explained. In order to simulate the STOVL aircraft

impinging flow, an appropriately large, flat plate is needed.

Small Plenum

Chambers

Microjet Stagnation

Chamber

25

2.1.5 GROUND PLATE

The ground at which the supersonic jet is impinging is simulated using a 1m x 1m

x 25mm aluminum plate shown in Figure 2-10. This ground plate is mounted on a

hydraulic lift as can be see in Figure 2-11 and is centered directly under the nozzle exit.

This hydraulic lift enables the ground-to-nozzle exit distance (h) to be varied from 2d to

35d. A Rexroth (DLC-100) hydraulic controller was used to control the vertical motion of

the ground plane. This controller is connected to the control computer and can be used to

adjust the height during the running of the jet. This is used to replicate hover conditions

of a STOVL aircraft.

Figure 2-10: Ground Plate used for the Jet Impingement as well as for some of the Measurement

Techniques

KuliteTM

Slot

26

Figure 2-11: Ground Plane Mounted on the Hydraulic Lift in Order to Simulate Hover

An overview of how a hot, supersonic, impinging jet is simulated at different

hover heights has been provided in this section. As described, most of the variable

parameters can be controlled and monitored during the experiment form the control room

using a control LabviewTM

program and several displays. This is very important in order

to perform stable and repeatable measurements. A more detailed description of

measurement techniques used throughout this study will be provided in the next few

sections.

2.2 MEASUREMENT AND INSTRUMENTATION

In order to characterize the performance of the impinging jet flow at room

temperature as well as at higher temperatures, several carefully designed measurement

setups were used. All the measurements were obtained with the microjets turned off and

with microjets turned on as well. All sensors were carefully calibrated prior to data

acquisition.

Hydraulic Lift

Ground Plate

27

2.2.1 MEAN PRESSURE MEASUREMENTS

Mean pressure measurements were obtained on the lift plate as well as on the

ground plane. In order to measure the static pressure distribution on the lift plate, 19

pressure ports along a radial line (Figure 2-12) were scanned using a ScanivalveTM

scanner. The ScanivalveTM

was connected to a ±1psid range ValidyneTM

strain gauge

transducer.

Lift Plate

de

Microjet array

Static

Pressure

Ports

Lift Plate

de

Microjet array

Static

Pressure

Ports

Lift Plate

de

Microjet array

Static

Pressure

Ports

Figure 2-12: Schematic of t he Static Pressure Port Array on the Lift Plate

A similar set up was used to measure the mean pressure distribution along the

centerline on ground plate. In this case only seventeen pressure ports were scanned using

a ScanivalveTM scanner, which was connected to 100psia OmegaTM pressure

transducer.

x

Static Pressure Ports

x

Static Pressure Ports

Figure 2-13: Schematic of the Static Pressure Port Array on the Ground Plate

28

2.2.2 UNSTEADY PRESSURE MEASUREMENTS

Unsteady pressure measurements on the ground plane and lift plate were made

using high frequency response, miniature KuliteTM

pressure transducers. In order to

mount the Kulites two specific slots were manufactured for both the ground plane and lift

plate. On the ground plate high temperature (up to 535ºF) KuliteTM

(Model XTEH-10L-

190-100A) pressure transducers were used with an operating range of 100 psia. The

location of the Kulites on the ground plane (Figure 2-14) was as follows, x/d=0, 1 and 2

with x/d=0 being the stagnation point of the impinging flow on the ground plate.

x

Unsteady Pressure

Transducers

x/d

=2

x/d

=1

x

Unsteady Pressure

Transducers

x/d

=2

x/d

=1

Figure 2-14: Schematic of Unsteady Pressure Transducers mounted on the Ground Plate

On the lift plate the operating range of the Kulites (Model XCS-062-5D) is 5psid.

Unsteady pressure field on the lift plate was measured using three pressure transducers

mounted at x/d=2, 3 and 4, from the nozzle centerline (Figure 2-15).

de

Microjet array

x/d

=4

x/d

=2

x/d

=3

de

Microjet array

x/d

=4

x/d

=2

x/d

=3

Figure 2-15: Schematic of the Unsteady Pressure Transducer mounted on the Lift Plate

Unsteady Pressure

Transducers

29

2.2.3 TEMPERATURE MEASUREMENTS

The temperature distribution on the ground plane and lift plate was measured

using K-type thermocouples. The thermocouples were mounted in a similar fashion as the

unsteady pressure transducers, using a slot that can be interchanged.

Eight thermocouples are mounted along the centerline of the ground plate. The

exact location can be seen in Figure 2-16, with x=0 being the point where the centerline

of the jet intersects with the ground plate (stagnation point).

xThermocouples

x = -1.5d, -0.9d, -0.6d, 0, 0.3d, 0.6d, 1.2d, 1.8d

xThermocouples

x = -1.5d, -0.9d, -0.6d, 0, 0.3d, 0.6d, 1.2d, 1.8d

xThermocouples

xThermocouples

x = -1.5d, -0.9d, -0.6d, 0, 0.3d, 0.6d, 1.2d, 1.8d

Figure 2-16: Schematic of the Thermocouples mounted in the Ground Plate

The lift plate was fitted with six thermocouples as shown in Figure 2-17. The

exact locations of the thermocouples were x/d=0.83, 1.16, 1.49, 2.16, 2.83 and 3.5 away

from the nozzle centerline. The first thermocouple was chosen at x/d=0.83 since that was

the closest possible location to the nozzle exit.

de

Microjet array

Static

Pressure

Ports

Thermocouples

de

Microjet array

Static

Pressure

Ports

Thermocouples

de

Microjet array

Static

Pressure

Ports

Thermocouples

Figure 2-17: Schematic of the Thermocouples mounted at the Lift Plate

30

2.2.4 ACOUSTIC PRESSURE MEASUREMENTS

In addition to temperature and pressure measurements, near field acoustic

measurements were made using a 0.635cm diameter B&K microphone (Figure 2-18)

which was connected to a preamplifier (Model 2633). The microphone was placed at

x/d=15 away from the nozzle centerline, in a direction 90° with respect to the jet axis

(Figure 2-2).

Figure 2-18: Microphone Location with respect to the Lift Plate

In order to reflect, mean pressure, unsteady pressure, temperatures as well as

acoustic measurements were conducted in this study in order to characterize the flow of

cold and hot, supersonic, impinging jets. A summary of all the measurement techniques

used throughout this study is shown in Table 1.

Microphone

31

Table 1: Summary of Test Cases and Measurement Techniques used

Diagnostic Technique Location TR=1.0 TR=1.2 TR=1.4 TR=1.6

Ground Plate (8 TC's) √ √ √ √ Thermocouples

Lift Plate (6 TC's) √ √ √ √ Ground Plate (x/d=0) √ √ √ √ Ground Plate (x/d=1) √ √ √ √ Ground Plate (x/d=2) √ √ √ √

Lift Plate (x/d=2) √ √ √ √ Lift Plate (x/d=3) √ √ √ √

Kulites

Lift Plate (x/d=4) − − − √

Microphone h/d=0, x/d=15 √ √ √ √ Ground plane (17) √ − √ − Mean Pressure

Measurement Lift plate (17) √ − √ −

2.2.5 DATA ACQUISITION

The acoustic, unsteady pressure as well as temperature data for the ground plate

and lift plate were acquired though a National Instruments digital data acquisition board

using LabviewTM

software and were processed offline using MatlabTM

software. The

transducer outputs were filtered at 30kHz using Stanford Research System low-pass

filters (Model SR640). The signal was sampled at 70 kHz thus satisfying the Nyquist

criteria, and maintaining the data free from aliasing errors. Standard Fast Fourier

Transform (FFT) analysis was used to obtain spectra and overall sound pressure levels

(OASPL) as well as temperature power spectra density (PSD). A total of 100 FFT’s of

4096 samples each were obtained, giving us a resolution of 17.2 kHz, in order to obtain

statistically reliable narrow-band spectra.

2.2.6 EXPERIMENTAL UNCERTAINTY

All the transducer used in this experiment where carefully calibrated prior to data

acquisition, nevertheless an error associated with each transducer has to be considered

when analyzing the corresponding data. Additionally, the uncertainty associated with the

data acquisition card needs to be added to the total measurement error as well. In case of

unsteady measurements, where the FFT analysis mentioned above is used, a statistical

32

error related to the number of data sets (in our case 100 data sets were averaged, see

above 2.2.5) has to be considered.

The input resolution of the 12-bit data acquisition card used in this experiment is

2.44mV for an input range of ±5V. The resulting error in pressure and temperature

associated with the data acquisition card needs to be added to the uncertainty in the

corresponding transducer measurements, in order to obtain the total error. The size of the

symbols in the graphs shown here, are chosen appropriately to represent the error in the

corresponding measurement.

2.2.7 TEST CONDITIONS

The experiments presented in the report were conducted mostly at nearly ideally

expanded jet flow condition, where the NPR is equal to 3.7 (stagnation pressure of the jet

being around 55psi) as appropriate for a Mach 1.5 nozzle. Some of the mean pressure

measurements were also conducted at NPR=5 (moderately underexpanded case,

stagnation pressure = 73psi). The jet stagnation temperature was varied from 300K to

480K. This corresponds to temperature ratios (TR, stagnation temperature/ambient

temperature) of 1.0 to 1.6. The test Reynolds number based on exit velocity and nozzle

diameter of the jet is estimated at 7 x 105. The nozzle-to-ground plane distance (h/d) was

varied form 2 to 12 for hover conditions and h/d=35 corresponding to the free jet.

A total of 16 microjets were implemented as described before in 2.1.4 around the

main jet to implement active flow control. The microjets were operated at a stagnation

pressure of 100 psia, leading to a combined mass flux from all the microjets of less than

0.5% of the primary jet mass flux.

This summarizes the experimental setup including measurements techniques used

and test conditions applied. This leads us into the discussion of the results obtained using

this setup and the before described measurements.

33

CHAPTER 3

RESULTS AND DISCUSSION

This chapter presents a discussion of the relevant results of this study. The results

were obtained at a wide range of temperatures and hover heights.

The occurrence of the stagnation bubble on the ground plane is explored at cold

and hot temperature first for an ideally expanded jet as well as for a moderately

underexpanded jet. After the mean pressure distribution, the effect of temperature and

control on the unsteady pressure loads are considered on the ground plane as well as on

the lift plate. The thermal unsteady loads on the ground plane are discussed for cold and

hot jets, with and without microjet control. The effect of temperature, height and control

on the temperature recovery factor is shown next. And at the end the changes in lift loss

when the jet is heated are presented and how lift loss can be recovered using microjet

control.

3.1 MEAN PRESSURE MEASUREMENTS

The mean properties of a single jet impinging on the ground plane are presented in

this section. The experiment was conducted for cold and hot temperatures and the mean

pressure measurements were recorded using the setup described in the previous chapter.

The cold tests were carried out at room temperature (TR=1.0, ~290K) and TR=1.4

(~420K) was used as a representative of the hot stagnation temperature conditions.

Temperature ratio, TR was defined before as the stagnation temperature divided by

ambient temperature. The jet was operated ideally expanded at NPR=3.7 as well as at

NPR=5.0. The moderately underexpanded condition, NPR=5.0 was considered only for

the surface pressure measurements since it was shown by Iyer (1999) and Garg (2001)

that the stagnation bubble is clearly visible at the underexpanded case and only at certain

heights (h/d’s). Therefore only certain heights were considered in this discussion. The

objective of this portion of the study was to determine what effect a higher jet

34

temperature has on the mean pressure field of an ideally expanded and highly

underexpanded impinging jet. All results are presented in a non-dimensional form using

the surface pressure coefficient for supersonic flows, Cp where

2

5.0

)(

j

oj

j

MP

PPP

P

Cp∗∗∗

−∗

=∞

∞

γ Equation 1: Surface pressure coefficient for supersonic flows

P : measured averaged surface pressure

Poj : measured averaged stagnation pressure of the jet

Pj : the desired stagnation pressure for the jet,

for NPR=3.7, Pj=55psi

for NPR=5.0, Pj =73psi

Note: Pj/Poj=1, for an ideal jet

P∞: ambient pressure which was measured before each run

Mj : Jet Mach number

for NPR=3.7, Mj =1.5

for NPR=5.0, Mj =1.7

γ : ratio of specific heats, for air γ=1.4

3.1.1 EFFECT OF NPR

As previously stated it was reported by Iyer and Garg that the stagnation bubble is

clearly visible at NPR=5.0, while for NPR=3.7 there was no stagnation bubble reported.

Iyer found that in order for a stagnation bubble to form a certain level of underexpansion

is required. At NPR=3.7, the jet is almost ideally expanded and therefore no stagnation

bubble should be observed. In the case where the jet is moderately underexpanded as at

NPR=5, Iyer found that there is evidence of a stagnation bubble at h/d=2, 3, and 5. Garg

confirmed Iyer’s results for the ideally expanded case at NPR=3.7, for a converging-

diverging nozzle. In the highly underexpanded case Garg also witnessed a “flat” annular

profile for h/d=3, 3.5, 4, and 5. Garg and Iyer performed their experiments at room

temperature, i.e. TR=1.0 (“cold” jet) consequently, the goal of this part of the study is to

examine the influence of a higher stagnation temperature on the surface pressure

distribution and occurrence of a recirculating region. In all the plots in this section the

vertical axis is representative of the non-dimensional mean pressure values (Cp) and the

35

horizontal axis indicates the radial location (x/d) of the port where the pressure was

measured (Fig in Chapter 2), with x/d=0 being at the centerline of the impinging jet.

TR=1.0

Figure 3-1 is a plot of the non-dimensionlized mean pressures on the ground plane

for the ideally expanded case, NPR=3.7 for the cold jet, TR=1.0. Several heights are

plotted in order to show that there is a well defined high pressure peak at the point where

the centerline of the jet impinges. This elevated pressure at the centerline is a good

indication of the absence of a recirculation region (Alvi and Iyer, 1999). It is interesting

to note that this peak pressure at the stagnation point stays constant as the height is

increased up to h/d=8. Then it drops significantly until it reaches the free jet condition,

where it recovers to ambient pressure. This is to be expected since the average potential

core of a supersonic jet is 8-10d long; after that there are no shocks to interact with the

ground plane. In the radial direction, away from the stagnation point, Cp values reduce as

the jet spreads. Higher pressures are observed close to the stagnation point and adjust to

the ambient pressure thereafter, as the Cp value goes to zero.

When the jet is highly underexpanded at NPR=5.0, as depicted in Figure 3-2, we

can see an annular pressure distribution for h/d=3 and 5. Additionally the Cp value at the

stagnation point of heights h/d=3 and 5 is 1.1 and 1.2, respectively whereas at h/d=2, 4,

and 6 the Cp value at the stagnation point is significantly higher and stays constant at

about 1.6. The existence of a flat profile and low Cp values at h/d=3 and 5 points to the

presence of a strong plate shock (see sketch in Figure 1-7) and therefore the existence of

a recirculation region. It is also evident that at h/d=3 the profile is flatter and the Cp value

is slightly lower than at h/d=5. This might be due to the weakening of the plate shock and

therefore the stagnation bubble as the impinging plane is moved further away from the

nozzle exit.

36

x x x x x x x x x x x x x x x x x

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8 h/d=2h/d=3

h/d=4h/d=5h/d=6

h/d=8h/d=12Free Jet, h/d=35x

TR=1.0 and NPR=3.7

Figure 3-1: Effect of Height on Mean Pressures of an Ideally Expanded Jet (NPR=3.7) at TR=1.0

Again, after h/d=8 the mean pressures, at the heights with “normal” impingement,

are reduced considerably until they recover back to the ambient pressure at h/d=35 which

corresponds to the free jet condition. As we move away from the centerline in the radial

direction the distribution of the mean pressures is comparable to the ideally expanded

case except that the wall jet region in this case shows stronger mean pressure

fluctuations, which persist further away from the centerline. These radial peak and

valleys are due to the stronger combinations of compression and expansion waves, as

discussed in Iyer (1999).

37


x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8 h/d=2h/d=3

h/d=4h/d=5h/d=6


TR=1.0 and NPR=5

Figure 3-2: Effect of Height on Mean Pressures of an Moderately Underexpanded Jet (NPR=5) at

TR=1.0

38

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8 NPR=3.7NPR=5.0

TR=1.0, h/d=2

Figure 3-3: Effect of NPR on a Cold Jet (TR=1.0) at h/d=2

With the aim of exploring the difference between ideally expanded and

moderately underexpanded case further, h/d=2, 3, and 5 are plotted separately. In Figure

3-3 both NPR cases are shown for h/d=2. It is observed that the Cp values for the

moderately underexpanded case are slightly higher then for the ideally expanded case due

to the higher stagnation pressure at NPR=5. It is interesting to note that both the

underexpanded case as well as the ideally expanded case, recover back to ambient

pressure at the same x/d distance (around 1). It is again obvious that there is no indication

of a stagnation bubble in either case at h/d=2, since the profile has a central peak

39

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8 NPR=3.7NPR=5.0

TR=1.0, h/d=3


In order to show the difference between ideally expanded and moderately

underexpanded jet behavior at the heights where the stagnation bubble occurs, the heights

h/d=3 and 5 are plotted in Figure 3-4 and Figure 3-5, respectively. It is apparent that the

highly underexpanded cases have much lower Cp values and a flatter profile close to the

stagnation point when compared to their ideally expanded counterparts. The obvious

presence of a low, flat profile in the underexpanded case strongly suggests that there is a

stagnation bubble formed at these two heights. In contrast, there is no hint of a stagnation

bubble for the ideally expanded profile at h/d=3 as well as h/d=5. Both profiles have a

well defined peak at the center point of the jet. As previously mentioned, the wall jet

Pressure perturbation due to

the shock cells in the radial

wall jet

40

region has more features in terms of shock cells in the case of moderate underexpansion

as is clearly apparent in Figure 3-4.

When comparing the underexpanded jet at h/d=3 and h/d=5 there is evidence of a

weakening of the stagnation bubble at h/d=5. As seen here, the maximum Cp value at

h/d=5 is slightly higher than at h/d=3 and the profile of the mean pressure distribution is

more “rounded” close to the stagnation point. This hints towards a weakened or smaller

stagnation bubble. This could be due to the impinging jet approaching the free jet

condition as the ground plane is moved further away from the nozzle exit, and the jet is

spreading and entraining more of the ambient air.

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8 NPR=3.7NPR=5.0

TR=1.0, h/d=5


41

TR=1.4

The behavior of the jet when it is heated to a temperature of 420 K, which

corresponds to a temperature ratio of 1.4 is considered next. Figure 3-6 is a plot of non-

dimensionlized mean pressure, Cp versus the radial location, starting from the jet

centerline. Different height-to-nozzle distances are plotted to indicate the change in mean

pressures as the impingement distance is varied from very small (h/d=2) up to what is

considered a free jet at h/d=35.


x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

h/d=2h/d=3h/d=4h/d=5h/d=6h/d=8

h/d=12Free Jet, h/d=35x

TR=1.4 and NPR=3.7

Figure 3-6: Effect of Height on Mean Pressures of an Ideally Expanded Jet (NPR=3.7) at TR=1.4

There is no suggestion of a stagnation bubble at the higher temperature for the

ideally expanded case. The central peak is clearly visible in Figure 3-6 and the mean

pressures are decreasing in the axial direction until they revert to ambient condition

42

around x/d=1.5-2. The central peak pressures are greatly elevated up to height-to-nozzle

distance of eight, at which point they start to decrease until they reach ambient when the

nozzle is far from the ground plane and can be considered a free jet.

In Figure 3-7 the mean pressures for a highly underexpanded and heated jet are

shown. As before several heights are shown to determine the heights at which there is a

stagnation bubble. It seems that even when the jet is heated there is only sign of a

stagnation bubble at h/d=3 and 5. The remaining heights show a well defined peak at the

centerline of the jet. In this heated case the Cp values for the so called “normal”

impingement cases decrease h/d=6, which is sooner than the cold jet, where the peak Cp

decay at h/d=8. Furthermore, for the higher temperature jet the Cp values at h/d=8 are

reduced even more drastically than the TR=1.0 case. This could be attributable to the

potential core of the jet being shorter as the entrainment of the ambient air is higher due

to a stronger feedback loop; this will be further explored when we look at the unsteady

pressures. Higher entrainment of ambient air could lead to a more rapid decay to the

ambient pressure when increasing the height-to-nozzle distance.

For the cases where there is no stagnation bubble as at h/d=2, depicted in Figure

3-8, one can see that there is a very high Cp maxima at the stagnation point of the

impingement. As the pressure in the radial direction away from the stagnation point is

considered, the Cp value is reduced due to the entrainment of ambient air and the radial

wall jet decays to ambient pressure at about 1d away from the stagnation point. However,

it is noticeable that the maximum value for the underexpanded jet at Cp=1.6 is much

higher when compared to the maximum value of Cp=1.4 that the ideally expanded jet

reaches. Furthermore, the wall jet created at NPR=5 has again more pressure

perturbations than the wall jet that is being formed due to the impingement of the ideally

expanded jet; this behavior is analogous to the cold jet experiment results.

43


x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

h/d=2h/d=3h/d=4h/d=5h/d=6


TR=1.4 and NPR=5

Figure 3-7: Effect of Height on Mean Pressures of an Moderately Underexpanded (NPR=5.0) Jet at

TR=1.4

44

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

NPR=3.7NPR=5.0

TR=1.4, h/d=2

Figure 3-8: Effect of NPR on a Heated Jet (TR=1.4) at h/d=2

The stagnation bubble is visible again when the jet is heated at heights h/d=3 and

5, as plotted in Figure 3-9 and Figure 3-10, respectively. Low Cp values at the stagnation

point of the underexpanded jet are a good indication of the existence of a recirculation

region. The wall jet created by the underexpanded jet impinging on the ground plane

shows once more a higher fluctuation of mean pressures when compared to the ideally

expanded case at h/d=3 in Figure 3-9. At h/d=5 in Figure 3-10 the wall jet displays less

defined fluctuations in mean pressures in the underexpanded case when compared to its

counterpart h/d=3.

45

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

NPR=3.7NPR=5.0

TR=1.4, h/d=3


46

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

NPR=3.7NPR=5.0

TR=1.4, h/d=5


3.1.2 EFFECT OF TEMPERATURE

Some differences were noted between the heated and cold impinging jet in the

previous section. In this section the changes in mean characteristics of the impinging jet

flow will be examined more closely while only the temperature is varied and the height

and NPR are kept constant.

NPR=3.7

In Figure 3-11 there are four different plots. Each plot represents a mean pressure

distribution at one height, starting with (a) h/d=2, (b) h/d=3, (c) h/d=5, and (d) h/d=8. All

47

data was obtained at the ideally expanded condition of NPR=3.7 and at two temperature

ratios, TR=1 and 1.4. The goal is to highlight the effect of heating the stagnation

temperature of an ideally expanded jet as we increase the ground-plate-to nozzle distance.

Figure 3-11: Effect of Temperature on an Ideally Expanded Jet (NPR=3.7) at (a) h/d=2, (b) h/d=3, (c)

h/d=5, and (d) h/d=8

As see here, at the smallest distance (h/d=2) the peak Cp values of the heated jet

are slightly higher than the peak Cp values of the cold jet. As we move further away from

the nozzle exit to h/d=3 and 5, there is a decrease in Cp values of the heated jet. Once

h/d=8 is reached both the heated and the cold jet Cp values are reduced immensely at the

stagnation point. However the value at the heated jet was reduced from over 1.4 at h/d=2

down to about 1 at h/d=8 whereas in the cold case the Cp value was reduced from about

1.4 at h/d=2 to about 1.1. This faster decay in stagnation pressure as h/d is increased

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

TR=1.0

TR=1.4

NPR=3.7 and h/d=5

(c)

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

TR=1.0

TR=1.4

NPR=3.7 and h/d=2

(a)

x/dC

p

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

TR=1.0

TR=1.4

NPR=3.7 and h/d=3

(b)

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

TR=1.0

TR=1.4

NPR=3.7 and h/d=8

(d)

48

could be due to a higher entrainment rate of the ambient air in the heated jet, which could

have been caused by a stronger feedback loop. This is to be explored more by looking at

the unsteady pressure data in the further sections of this chapter.

NPR=5.0

Again, we look at four different plots in Figure 3-12, in this case however the jet

is moderately underexpanded. Each plot features one nozzle-to-ground-plate distance

starting with the smallest (a) h/d=2 up until (d) h/d=8.

At small h/d distances the mean pressure distribution seems to be quite similar at

TR=1.0 and TR=1.4. Once the distance is increased to h/d=3, the influence of the higher

stagnation temperature of the jet on the stagnation bubble is visible. The peak Cp values

of the high temperature jet are higher at than the maximum Cp value of the cold jet.

These peak values seem to be decaying more rapidly in the radial direction, close to the

stagnation point, in the heated case which leads to a less flat profile. These two

observations lead to the conclusion that the stagnation bubble at h/d=3 is present in both

cases but seems to be weakened when the jet is heated. When the distance increased to

h/d=5, it is apparent that the stagnation bubble in the cold jet is weakened now as the

peak value is increased (but still much lower than the normal impingement values).

There are less fluctuations in the mean pressures in the wall jet region in both the

cold and the hot jet at h/d=5 when compared to h/d=3. It seems that the wall jet region is

not influenced by temperature, but only by NPR and the occurrence of a stagnation

bubble. Since there is no evidence of a stagnation bubble at h/d=8 we are comparing this

height to the “normal” impingement heights (similar to h/d=2). Therefore, when

compared to h/d=2, both the cold and hot case maximum Cv values seem to be reduced

drastically. In the heated case the rate of decay seems to be increased as in this case the

decay of the peak Cp values starts already at h/d=6 (Figure 3-7). This could be due to a

higher entrainment rate of the ambient air which could cause a weaker normal shock and

therefore less interaction between the normal shock and the ground plate leading to a

weaker stagnation bubble, as seen in Figure 3-12 (b), as well more rapid adjustment to

ambient pressure as seen in Figure 3-12 (d).

49

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

TR=1.0

TR=1.4

NPR=5 and h/d=2

(a)

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

TR=1.0

TR=1.4

NPR=5 and h/d=5

(c)

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

TR=1.0

TR=1.4

NPR=5 and h/d=3

(b)

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

TR=1.0

TR=1.4

NPR=5 and h/d=8

(d)

Figure 3-12: Effect of Temperature on an Moderately Underexpanded Jet (NPR=5.0) at (a) h/d=2, (b)

h/d=3, (c) h/d=5, and (d) h/d= 8

In conclusion, the stagnation bubble seems to be weakened when the jet is heated

due to the higher entrainment rate of the ambient air. The heights at which the stagnation

bubble occurs are however not impacted by the higher stagnation temperatures since

there is evidence of a recirculation region only at h/d=3 and 5 for both cases, hot and

cold. The wall jet unsteadiness is mainly influenced by the underexpansion of the main

jet as well as the occurrence of a stagnation bubble. Also, the higher entrainment rates

for hot impinging jets lead to a faster decay in peak Cp values at higher temperature jets.

50

3.2 UNSTEADY PRESSURE LOADS ON THE GROUND PLATE

AND LIFT PLATE

In order to analyze the unsteadiness in an impinging, supersonic jet flow the

unsteady pressures of an almost ideally expanded supersonic jet were measured, while

varying the height-to-nozzle distance (h/d) from 2 to 12 and the temperature ratio (TR)

from 1.0 to 1.6. The average unsteady pressure or RMS pressure, Prms quantifies the

unsteadiness of the flow. In order to present the RMS pressures in a non-dimensional

form we convert them to the decibel (dB) scale using Equation 2. In this case the

reference pressure, Pref is 20μPa and the number 6895 is the conversion factor for

converting pressure values in psi to Pascal.

)6895*

log(*20ref

rms

P

PdB = Equation 2: Converting RMS pressures into the dB scale

In this section results are presented concerning the effect of temperature and

height on the unsteadiness of the jet and the effectiveness of microjet control at the same

conditions. Two different quantities are used to show the unsteadiness of the jet flow.

First the unsteady pressure spectra or Sound Pressure Levels (SPL) to show the energy

distribution in the frequency domain, this will also be used to illustrate the presence of

the feedback loop and the influence of the temperature and control on it. Second, the

Overall Sound Pressure Levels (OASPL) will give us an overview of the influence of

height, temperature, and control efficiency on the overall magnitude of unsteadiness of

the flow. OASPL is the integral of the area below an unsteady pressure spectra or the

total magnitude of energy of a fluctuating signal at a given condition.

Unsteady pressure measurements were obtained using KuliteTM

pressure

transducers at several locations along a radial line on the ground plane as well as on the

lift plate (the locations can be seen in Figure 2-14 and Figure 2-15). In Table 2, the

pressure transducer locations are summarized for a better overview. Acoustic

measurements are acquired using a microphone that in the horizontal direction was

mounted 15 inches away from the nozzle centerline and in the vertical direction it was at

51

the same plane as the nozzle exit (see Figure 2-2). A detailed description of the hardware

used can be found in 2.2.

Table 2: Spatial Distribution of the Transducers (also see Figure 2-14 and Figure 2-15)

KuliteTM

Transducer Location Distance from the Nozzle Centerline

GP 1 Ground Plane 0 inches



LP 1 Lift Plate 2 inches



REPEATABILITY

The experimental data that is compared throughout this section was obtained over

a year. Hence, it is important to show the repeatability of the pressure measurements at

the ground plane and lift plate as well as the repeatability of the acoustic measurements

by the microphone. For clarity reasons only data at one ground plane location (GP 1) and

one lift plate (LP 1) location is presented together with the microphone data.

In Figure 3-13, unsteady pressure spectra is shown for h/d=3.5 and TR=1.0. As

seen, two sets of GP 1 data, acquired 6 days apart, as well as two sets of each, LP1 and

microphone data acquired 3 days apart is shown. It is clear that for all three measurement

locations the graphs have identical characteristics for the different dates. The high

amplitude tones occur at the same frequency with matching magnitudes. Broadband

levels are very comparable as well. Therefore, it is with confidence that can be stated the

data acquired throughout this experiment is repeatable, allowing for comparisons

between data obtained from different tests.

52

Frequency (kHz)

SP

L(d

B)

5 10 15 20 25 30 3580

100

120

140

160

180 Ground Plate, 03/11/08

Ground Plate, 03/17/08

Lift Plate, 03/23/08Lift Plate, 03/26/07Microphone, 03/23/08

Microphone, 03/26/07

TR=1.0, h/d=3.5

Figure 3-13: Repeatability of Unsteady Pressure Data at TR=1.0

GLOBAL TREND

As mentioned before, data was recorded at different locations in the flow for all

the diverse conditions. To demonstrate that the flow unsteadiness is global when the jet is

cold as well as when it is heated, measurements from different locations and at different

TR’s are compared. In order to keep this section brief data for temperature ratio 1.0 and

1.6 is shown only, for the other TR’s one can refer to the Appendix.

In Figure 3-14, the narrowband pressure spectra is graphed for the ground plane,

lift plate, and near-field microphone at h/d=3.5 and TR=1.0, representing a cold jet. As

seen here, the high amplitude tones which correspond to the impinging tones caused by

the feedback loop have the same frequency at all transducer locations. The magnitude of

53

these discrete tones does change with location, as expected, since the intensity is changed

which is attributable to how far the location is from the jet. As the microphone is the

farthest away from the jet centerline at x/d=15, the microphone unsteady spectra is the

lowest, followed by the lift plate data at x/d=2. The ground plane KuliteTM GP 1 is

mounted directly where the jet centerline is impinging and therefore measures the highest

unsteady loads and hence increased levels in the SPL.

Frequency (kHz)

SP

L(d

B)

5 10 15 20 25 303580

100

120

140

160

180Ground Plate, x/d=0Lift Plate, x/d=2

Microphone, x/d=15 and y/d=0

TR=1.0, h/d=3.5

Figure 3-14: SPL at Different Locations in the Flow at h/d=3.5 and TR=1.0

In Figure 3-15 an analogous graph is offered, where the jet is heated to stagnation

temperature of 480K, corresponding to temperature ratio is 1.6. Again, the high

amplitude tones have identical frequencies pointing towards a common source and

54

therefore indicating that they are global characteristic of the flow in case of hot as well as

cold jets.

Frequency (kHz)

SP

L(d

B)

5 10 15 20 25 30 3580

100

120

140

160

180Ground Plane, x/d=0Lift Plate, x/d=2Microphone, x/d=15

TR=1.6, h/d=3.5

Figure 3-15: SPL at Different Locations in the Flow at h/d=3.5 and TR=1.6

As described in the Introduction it is believed that the feedback loop mechanism

as described by Powell (1953a) is the phenomenon governing the unsteady flow

properties of impinging jets. Powell has proposed that the discrete tone frequency can be

predicted using Equation 3.

∫ +=± h

ain C

h

C

dh

f

pn

0 Equation 3: Impingement Tone Frequency Prediction Formula

h: nozzle exit-to-ground plate distance (h=2in-12in)

Ci: convection velocity of the downstream traveling large structures or Ci=0.5*Uj

(Krothapalli, 1999)

Ca: is the speed of the upstream traveling acoustic waves

55

n: arbitrary integer, n=1,2,3,… (when n=1 corresponds to mode1, n=2 corresponds to

mode 2, etc.)

P: phase lag between the acoustic wave and convected disturbance and is chosen

according to jet conditions, p= -0.4

To confirm that the high amplitude discrete tones seen in the unsteady pressure

spectra are due to the impinging tones created by the feedback loop, the tonal frequencies

are compared with those predicted by Equation 3 in Figure 3-16 and Figure 3-17.

In Figure 3-16, the vertical dashed lines indicate the frequencies predicted by

Equation 3 for n-values or mode numbers. In Figure 3-17, the impinging tone results are

summarized at all heights and the predicted frequency lines created by Equation 3 are

shown for comparison.

Frequency (kHz)

SP

L(d

B)

5 10 15 20 25 30 3580

100

120

140

160

180

Ground Plate, x/d=0

Lift Plate, x/d=2

Microphone, x/d=15 and y/d=0

TR=1.0, h/d=3.5

n=2 n=3 n=4

n=5

n=6 n=9n=10

n=15 n=19

Figure 3-16: Indication of Discrete Tone Frequency Prediction at h/d=3. 5 and TR=1.0

As seen in Figure 3-17, at all h/d measured in this study, the discrete tonal

frequencies are predicted fairly well by the formula proposed by Powell, indicating that at

56

TR=1.0 the discrete tones that are seen in the unsteady pressure spectra are the impinging

tones created by the feedback loop and therefore changes in those tones can be related to

changes in the feedback loop.

h/d

Fre

qu

en

cy

(kH

z)

2 3 4 5 6 7 8 9 10 11 12

5

10

15

20

25

3035

TR=1.0, Liftplate x/d=2p =- 0.4, n=2-20

n=2

n=20

Figure 3-17: Impinging Tones as a Function of h/d and the Lines of Predicted Frequencies for

TR=1.0

In order to determine whether the same is for jets at higher temperatures similar

graphs are presented in Figure 3-18 and Figure 3-19 but at TR=1.6. Again, in Figure 3-18

one can see that the predicted frequencies at h/d=3.5 for different integers (n) are very

indicative of the frequency of the discrete tones at all three different points of

measurement. This points again to a common noise source for all the transducers and

therefore these tones are indicative of a global nature in the flow at TR=1.0 from Figure

3-16 as well as a TR=1.6 from Figure 3-18.

57

Frequency (kHz)

SP

L(d

B)

5 10 15 20 25 30 3580

100

120

140

160

180

Ground Plane, x/d=0Lift Plate, x/d=2Microphone, x/d=15

TR=1.6, h/d=3.5

n=2 n=3 n=4 n=5

n=8

n=11 n=15

Figure 3-18: Indication of Discrete Tone Frequency Prediction at h/d=3.5 and TR=1.6

In Figure 3-19 one can see that the discrete tones frequency can be used in order

to predict impinging tones even at higher temperature jets at all the heights. In both

Figure 3-17 and Figure 3-19 the filled symbols indicate the peak impinging frequencies at

the different heights. As seen here, they vary from 5 kHz to 8 kHz at both temperature

ratios. Attenuation of those tones can lead to a reduction of the adverse effects caused by

the feedback loop.

In summary, it has been shown that the discrete tones seen in the unsteady

pressure spectra of this impinging jet have a common source for all the transducer and are

representative of the feedback loop. Therefore it is reasonable to consider changes in

these tones to be an indication of the changes in the feedback loop. Next effects of

58

temperature, height and control on the feedback loop will be considered by examining the

changes in the high amplitude tones in the unsteady pressure spectra.

h/d

Fre

qu

en

cy

(kH

z)

2 3 4 5 6 7 8 9 10 11 12

5

10

15

20

25

3035

TR=1.6, Lift Plate x/d=2p=-0.4, n=2-20

n=2

n=20

Figure 3-19: Impinging Tones as a Function of h/d's and the Lines of Predicted Frequencies at

TR=1.6

3.2.2 UNSTEADY PRESSURE SPECTRA (SPL)

This section will focus on the unsteady pressure spectra. It is important to

understand that the impinging tones that are created by the feedback loop are manifested

in the unsteady pressures spectra as high amplitude tones and harmonics. In addition,

broadband levels are a good indication of the intensity of the overall unsteadiness in the

flow. Therefore, as the following graphs are explored, the objective will be to determine

how increased temperature and microjet control changes the peak tones as well as the

broadband levels. In the first part of this section, the temperature effect on the unsteady

59

pressures at different locations in the flow is examined. In the second part of this section,

the focus is on the effectiveness of control for cold and hot jets.

EFFECT OF TEMPERATURE

Previously in 3.1 it was noted that the entrainment rate of the ambient air seems to

increase when the stagnation temperature of the jet is increased. This is the reason for the

increase in lift loss as we will see later ( 3.4). The effects of increased stagnation

temperature on the unsteady pressures, which are directly influenced by the change in the

feedback loop mechanism, are considered in Figure 3-20.

Frequency (kHz)

SP

L(d

B)

5 10 15 20 25 30 35100

120

140

160

180TR=1.0TR=1.2

TR=1.4

TR=1.6

Lift Plate, x/d=2No Control, h/d=4

Impinging Tones

Figure 3-20: Influence of Temperature on the SPL Spectra in the Shear Layer of the Flow

(Lift Plate, x/d=2)

60

Here, the unsteady pressure spectra is shown for four different temperature ratios

at h/d=4 and is measured at the lift plate location closest to the nozzle exit (x/d=2 from

the jet centerline). As seen in this plot, the broadband levels are increased as the jet is

heated. There is a very obvious trend for the peak tones of each pressure spectra as well.

As the temperature increases, these tones increase in frequency. These high amplitude

tones are due to the impinging tones that are created by the feedback loop; this leads us to

believe that the feedback loop may be completed faster at higher stagnation temperatures

of the jet. A faster feedback loop created by faster large-scale structures in the flow leads

to higher entrainment rates of the ambient air. It is interesting to note that there is not a

significant change in frequency and amplitude of the peak tones when we increase the

temperature from TR=1.4 to 1.6. This might be an indication that the increase in

frequency due to the increase of temperature either is a step function that levels out at

certain temperatures or that there is a maximum peak frequency for each height. This

might only be explored further by increasing the temperature of the jet beyond the range

examined in this research.

The shift in frequency due to the change in temperature from TR=1.0 to 1.6 can

be accounted for by converting the frequency to the Strouhal Number using Equation

4 :

U

dfSt

•= Equation 4: Strouhal Number

In Equation 4 f stands for frequency in Hz, d is a characteristic length, in this case the

throat diameter of the nozzle, and U is the exit velocity of the jet, calculated using

Equation 5.

RTMU γ*=

Equation 5: Exit Velocity

M: Mach Number, 1.5

γ: Ratio of Specific Heats, 1.4 for air

R: Universal Gas Constant, 287 J/kg*K

T: Static Temperature in the jet at the nozzle exit

For TR=1.0, T=206.9K

For TR=1.2, T=248.3K

61

For TR=1.4, T=289.7K

For TR=1.6, T=331.1K

The exit temperature of the jet is directly related to the stagnation temperature, therefore

the unsteady pressure spectra versus the Strouhal number is shown. Strouhal number is a

extract the effect of temperature on frequency.

In Figure 3-21, unsteady pressure spectra versus the scaled the frequency using

the Strouhal number is shown. All the high amplitude peaks overlap fairly well for

different temperature ratios, indicating that the increase in the frequency of the impinging

tones are related to the increase in stagnation temperature of the jet. In order to \assess

whether the impingement tone frequency increases as the temperature is increased, data

obtained with the microphone mounted 15 inches away from the nozzle centerline will be

considered next. One example of the near-field of the jet flow is shown in Figure 3-22.

ST=(f*d)/U

SP

L(d

B)

0.5 1 1.5 2100

120

140

160

180 TR=1.0TR=1.2TR=1.4TR=1.6

Lift Plate, x/d=2No Control, h/d=4

Figure 3-21: Unsteady Pressure Spectra Obtained at the Lift Plate (x/d=2) at all TR's with Frequency

Scaled Using the Strouhal Number

62

The effect of temperature on the near-field narrowband noise spectra can be seen

in Figure 3-22. The effect of a faster feedback loop is clearly visible. The discrete high

amplitude tones are well defined at all the temperature ratios. These impinging tones are

increasing in frequency and amplitude as the temperature of the jet is increased. Note that

even in the near-field there is little change in frequency between tones at TR=1.4 and 1.6,

but the broadband level has increased significantly.

Frequency (kHz)

SP

L(d

B)

5 10 15 20 25 30 3580

90

100

110

120

130

140

150TR=1.0TR=1.2

TR=1.4TR=1.6

Microphone, x/d=15No Control, h/d=4

Impinging Tones

Figure 3-22: Influence of Temperature on the SPL Spectra of the Near-field Noise

(Microphone, x/d=15)

Again, to confirm that the increase in frequency of the peak tones is due to the

change in temperature, frequency from Figure 3-22 has been scaled using the Strouhal

number (Figure 3-23). As seen here again, the peak tones and their harmonics

63

approximately overlap for all the different temperature ratios, once the frequency is

scaled with respect to temperature.

ST=(f*d)/U

SP

L(d

B)

0.5 1 1.5 280

90

100

110

120

130

140

150TR=1.0

TR=1.2TR=1.4

TR=1.6

Microphone, x/d=15No Control, h/d=4

Figure 3-23: Near-field Narrowband Noise Spectra at Different TR's with Scaled Frequency using

the Strouhal Number

In the graph below (Figure 3-24) data obtained at the stagnation point of the flow

on the ground plane, has been shown in form of unsteady pressure spectra. There are well

defined peak tones and their harmonics visible in the spectra showing influence of the

feedback loop at the stagnation point of the flow. However, there is a small shift in

frequencies and amplitudes of the impinging tones at this location, as the jet is heated.

This suggests that the unsteadiness at the stagnation point of the flow might be influenced

by other characteristics of the jet, such as the oscillation of the jet core and the stagnation

bubble, as well.

64

Frequency (kHz)

SP

L(d

B)

5 10 15 20 25 30 35120

130

140

150

160

170

180

190TR=1.0TR=1.2TR=1.4TR=1.6

Ground Plane, x/d=0No Control, h/d=2

Impinging tones

Figure 3-24: Influence of Temperature on the SPL Spectra at the Stagnation Point of the Flow,

Ground Plane (x/d=0)

Microjet control is believed to interrupt the feedback loop and therefore reduce or

eliminate these high amplitude tones in the spectra. In the next segment the efficiency of

this method of control when the jet is heated, is addressed.

EFFECT OF CONTROL

How microjet control effects the feedback loop mechanism at all the different

temperature ratios, TR’s at the same height, h/d=4.0 is graphically presented in Figure

3-25 (a) to (d). Each plot is representative of one temperature ratio, starting with (a)

TR=1.0 to (d) TR=1.6. For each of these temperature ratios two data sets are shown, one

is the baseline case and the other is the controlled case (depicted by the green line in all

plots). Microjet control is applied at a stagnation pressure of 100psia and at an angle of

65

60deg with respect to the jet centerline. A more detailed description of the setup is

provided in Chapter 2.

Frequency (kHz)

SP

L(d

B)

5 10 15 20 25 303580

100

120

140

160

No ControlWith Control

TR=1.0, h/d=4Lift Plate, x/d=2

(a)

Frequency (kHz)

SP

L(d

B)

5 10 15 20 25 303580

100

120

140

160



(b)

Frequency (kHz)

SP

L(d

B)

5 10 15 20 25 303580

100

120

140

160 No ControlWith ControlTR=1.4, h/d=4

Lift Plate, x/d=2

(c)

Frequency (kHz)

SP

L(d

B)

5 10 15 20 25 303580

100

120

140

160



(d)

Figure 3-25: Effect of Microjet Control on Unsteady Pressure Spectra Obtained on the Lift Plate

(x/d=2) at h/d=2 for different jet temperatures (a) TR=1.0, (b) TR=1.2, (c) TR=1.4 and TR=1.6

In Figure 3-25 data obtained at the lift plate is presented. These results

substantiate the fact that for all the TR’s this control method is very effective. One can

see that the high amplitude tones are attenuated or even eliminated in some cases.

Microjet control interrupts/weakens the feedback loop which is causing these impinging

tones to disappear. This dramatic effect is even stronger for higher temperature jets,

where the tonal reduction due to control is greater. Distinct tones are the cause for sonic

fatigue and ground erosion in STOVL aircrafts and their operational environment. Hence,

these results indicate that microjet control is should be effective at reducing or

66

eliminating those adverse effects even at more realistic conditions. There is also

reduction in the broadband levels due to microjet control. This is noteworthy since most

control methods reduce tonal noise with no reduction in the broadband levels.

Frequency (kHz)

SP

L(d

B)

5 10 15 20 25 303580

90

100

110

120

130

140

150


TR=1.0, h/d=4Microphone, x/d=15

(a)

Frequency (kHz)S

PL

(dB

)

5 10 15 20 25 303580

90

100

110

120

130

140

150

No controlWith Control


(b)

Frequency (kHz)

SP

L(d

B)

5 10 15 20 25 303580

90

100

110

120

130

140

150

No Control

With Control


(c)

Frequency (kHz)

SP

L(d

B)

5 10 15 20 25 303580

90

100

110

120

130

140

150



(d)

Figure 3-26: Effect of Microjet Control on Near-field narrowband spectra at h/d=4 while the

temperature is varied (a) TR=1.0, (b) TR=1.2, (c) TR=1.4 and (d) TR=1.6

A comparable graph is presented in Figure 3-26 showing the effect of microjets in

the near-field of the jet. One can see that in the near-field of the jet the impingement

tones are completely eliminated when microjets are turned on. The fact that the peak

tones are eliminated in the near-field sound noise indicates that the microjet control is

very effective in reducing the noise pollution that can be created by supersonic,

impinging jets whether they are cold, Figure 3-26 (a) or heated, Figure 3-26 (d). Again,

there is also a significant reduction in the broadband levels as well.

67

Earlier in Figure 3-24, the unsteadiness of the jet at the stagnation point of the

flow on the ground plane was discussed. As a reminder, it was concluded that even

though this point in the center of jet impingement is influenced by the feedback loop

effects in the shear layer, it may be also affected by the changes in the impingement zone.

Frequency (kHz)

SP

L(d

B)

5 10 15 20 25 3035110

130

150

170

190


TR=1.0, h/d=4Ground Plane, x/d=0

(a)

Frequency (kHz)

SP

L(d

B)

5 10 15 20 25 3035110

130

150

170

190

No Control

With Control


(b)

Frequency (kHz)

SP

L(d

B)

5 10 15 20 25 3035110

130

150

170

190



(c)

Frequency (kHz)

SP

L(d

B)

5 10 15 20 25 3035110

130

150

170

190



(d)

Figure 3-27: Effect of Microjet Control on the Stagnation Point of Jet Impingement at h/d=4 for

Different Temperatures Ratios (a) TR=1.0, (b) TR=1.2, (c) TR=1.4 and (d) TR=1.6

In Figure 3-27 the efficiency of microjet control on the unsteadiness of the jet at

the stagnation point is shown. At all the temperatures the high amplitude tones are

reduced, pointing again towards the effect of microjet control on the feedback loop.

There is a large decrease in broadband levels when microjet control is applied. This could

be due to the decrease in unsteadiness of the core jet oscillation or other changes in the

impingement region, but would need more thorough characterization of the impingement

68

region. Therefore, it may be noted again that the unsteadiness of the jet at the stagnation

point is not only affected by the feedback loop but also by the structure of the

impingement zone.

In order to explore the effect of temperature and control on the overall

unsteadiness of the flow at all locations and height-to-nozzle distances, the Overall Sound

Pressure Levels are considered in a similar manner as the spectra.

3.2.3 OVERALL SOUND PRESSURE LEVELS (OASPL)

To measure the magnitude of the overall unsteadiness of a supersonic impinging

jet at different stagnation temperatures (TRs) and height-to-nozzle distances (h/d’s) one

can calculate the total energy of the fluctuating pressure spectra at each condition (TR

and h/d). When this unsteady pressure Prms is converted to dB, this is often referred to as

Overall Sound Pressure Levels, OASPL. In the following section the influence of

temperature and control on OASPL is considered, while the h/d is increased. First, the

effect of increasing stagnation temperature on the overall magnitude of unsteadiness in

the jet is examined. Second, the effectiveness of microjet control on OASPL at different

locations in the jet flow will be shown.


In Figure 3-28, OASPL versus the height-to-nozzle distance (h/d) is shown at

different temperature ratios (a) TR=1.0, (b) TR=1.2, (c) TR=1.4, and (d) TR=1.6. At all

temperatures the highest fluctuating pressures are measured at the stagnation point of the

jet (Ground Plane, x/d=0), it is followed by the two ground plane locations (x/d=1,2),

then the lift plate and near-field microphone at x/d=15. The feedback loop associated

with the impinging jets at small h/d’s has been observed to be very strong, therefore the

OASPL is expected to be highest at small h/d’s and decrease as the height is increased.

69

h/d

OA

SP

L(d

B)

0 2 4 6 8 10 12100

120

140

160

180

200

Ground Plane, x/d=0

Ground Plane, x/d=1Ground Plane, x/d=2Lift Plate, x/d=2

Lift Plate, x/d=3Microphone, x/d=15TR=1.0

(a)

h/d

OA

SP

L(d

B)

0 2 4 6 8 10 12100

120

140

160

180

200

Ground Plane, x/d=0Ground Plane, x/d=1

Ground Plane, x/d=2Lift Plate, x/d=2Lift Plate, x/d=3

Microphone, x/d=15TR=1.2

(b)

h/d

OA

SP

L(d

B)

0 2 4 6 8 10 12100

120

140

160

180

200

Ground Plane, x/d=0Ground Plane, x/d=1Ground Plane, x/d=2

Lift Plate, x/d=2

Lift Plate, x/d=3Microphone, x/d=15

TR=1.6

(d)

h/d

OA

SP

L(d

B)

0 2 4 6 8 10 12100

120

140

160

180

200

Ground Plane, x/d=0Ground Plane, x/d=1Ground Plane, x/d=2

Lift Plate, x/d=2

Lift Plate, x/d=3Microphone, x/d=15

TR=1.4

(c)

Figure 3-28: Effect of Temperature on the OASPL vs. h/d at different locations in the flow

There are non-monotonic variations in OASPL in form of “hills and valleys” due

to the change in resonance in the feedback loop. This resonance is more pronounced at

higher temperatures and is decreased as h/d is increased in both cold and hot jets.

In Figure 3-28 it is especially obvious that the OASPL levels are higher at small

h/d’s and have better defined hills and valleys at the lift plate. The same trend is not

apparent at the ground plane therefore a closer look is taken at the overall sound pressure

levels at all the locations separately in the following figures.

70

h/d

OA

SP

L(d

B)

0 2 4 6 8 10 12150

155

160

165

170

175TR=1.0TR=1.2TR=1.4

TR=1.6

Lift Plate, x/d=2

(c)

h/d

OA

SP

L(d

B)

0 2 4 6 8 10 12130

135

140

145

150

155TR=1.0

TR=1.2TR=1.4

TR=1.6

Microphone, x/d=15

(d)

h/d

OA

SP

L(d

B)

0 2 4 6 8 10 12170

175

180

185

190

195TR=1.0

TR=1.2TR=1.4

TR=1.6

Ground Plane, x/d=1

(b)

h/d

OA

SP

L(d

B)

0 2 4 6 8 10 12170

175

180

185

190

195

TR=1.0

TR=1.2

TR=1.4

TR=1.6Ground Plane , x/d=0

(a)

Figure 3-29: Influence of Temperature on the Average Fluctuating Pressure as h/d changes at

different locations in the flow (a) Ground plane (x/d=0), (b) Ground Plane (x/d=1), (c) Lift Plate

(x/d=2), and (d) Microphone (x/d=15)

In Figure 3-29 (a) the OASPL is plotted versus the increasing height-to-nozzle

distance at different temperature ratios where the data shown is only at the ground plane

location, x/d=0. It would be expected that the unsteadiness of the jet increases when the

stagnation temperature is increased. However, the data in Figure 3-29 (a) shows us a

different picture; the higher temperature jets have a minima at small h/d’s and the

fluctuating pressures are lower in the hot jets than in the colder jets. This might be due to

the fact that this point of measurement is the stagnation point and therefore might be

more influenced by the structure of the impingement region (see Figure 1-6 and Figure

1-7) than the shear layer. The shear layer is the primary region where the effect of the

71

feedback loop is evident. As shows in Section 3.4 the heated jet generates an increased

entrainment rate and higher vacuum pressure near nozzle exit. This is due to the

strengthening of the feedback loop when the jet is heated, which we showed previously in

the unsteady pressure spectra. These increased differential pressures lead to an increase in

the underexpansion ratio, as defined by Iyer (1999), in the jet and therefore a change in

the characteristic shock cell structure of the jet when it is heated. Following this

reasoning, it could be considered that the point of interaction in the shock cell with the

impinging plate shock changes, thus modifying the structure in the stagnation region

(impingement region) when the jet is heated. This can lead to a change in unsteadiness

that could result in a decrease of fluctuating pressure at the stagnation point.

Additionally, the underexpansion of the jet causes a change in the structure of the jet core

oscillations and has an effect on the unsteadiness we see at the stagnation point. In a

future study, it would be interesting to consider a way of visualizing the oscillations of

the jet when it is cold and heated and see whether an increase in stagnation temperature

reduces the oscillations of the core of the jet.

In Figure 3-29 (b) the data was measured one inch away from the stagnation

point, one can see that the unsteady pressures are now influenced by the shear layer and

therefore the plot shows an increase in unsteadiness as the stagnation temperature of the

jet is increased. A similar effect is observed in Figure 3-29 (d), which shows data

measured using a microphone placed 15 inches away from the nozzle centerline. Overall,

the high unsteady pressures start decreasing as the nozzle exit is moved further away

from the ground plane, which is overlapping with the feedback loop weakening when h/d

is increased.

The graph in Figure 3-29 (c) represents the data measured at the lift plate location

two inches away from the nozzle centerline. In this plot it is clearly apparent that the

OASPL has a maximum at small h/d’s and reduces monotonically thereafter.

Additionally, one can observe that the levels at higher temperatures are increased by

more than 8dB when compared to the cold temperature jets. These peak levels stay

constant up to h/d=6 at TR=1.2, 1.4 and 1.6, which shows more persistence, than the high

levels that are maintained only up to h/d=3 at TR=1.0. Collectively, this is evidence of

the feedback loop strengthening as well as the feedback loop being persistent over a

72

larger range of h/d when the jet is heated. Both of these effects are of practical

significance since they are the main cause of ground erosion, sonic fatigue, noise

pollution and lift loss in a STOVL aircraft.

Next it will be explored how microjet control efficiency is affected by this

strengthened and more persistent feedback loop.

EFFECT OF MICROJET CONTROL

High fluctuating loads are the source of many unwanted effects such as ground

erosion, sonic fatigue and they are also a major source of noise pollution. In the above

section, the evidence was shown that when the jet is at a more realistic or higher

stagnation temperature, the feedback loop becomes stronger leading to higher

unsteadiness in the jet. In prior studies, good results were achieved using microjet control

on ambient temperature jets (Iyer 1999, Garg 2001, Lou 2005). Therefore, it could be

expected that microjet control will also be effective in reducing the unsteadiness of a hot,

impinging jets and the adverse effects mentioned before. In this section, the effectiveness

of microjet control for higher temperature jets will be examined and compared to the

efficiency of control for cold jets.

h/d

OA

SP

L(d

B)

0 2 4 6 8 10 12160

170

180

190

200No Control

With ControlGround Plane, x/d=0TR=1.6

(b)

h/d

OA

SP

L(d

B)

0 2 4 6 8 10 12160

170

180

190

200No Control

With ControlGround Plane, x/d=0TR=1.0

(a)

delta OASPL

Figure 3-30: Effect of Microjet Control on Pressure Fluctuation Intensities at the Ground Plane

(x/d=0) at (a) TR=1.0 and (b) TR=1.6

73

The unsteady pressures were measured at the same locations and at the same

conditions as described before, with and without applying microjet control. In order to

see the effect of the microjets on the intensity of the fluctuating pressures, OASPL as a

function of h/d is shown, with and without control, at TR=1 and TR=1.6, in Figure 3-30

(a) and (b) respectively. A reduction in unsteadiness for both cold and hot jets at small

h/d's is observed. This reduction in unsteadiness is decreasing as the nozzle-to-ground

plane distance is increasing.

In order to determine how the change in stagnation temperature affects the

efficiency of control, the difference in OASPL or “delta OASPL” will be examined next.

In Figure 3-31, “delta OASPL” as a function of h/d is shown. Delta OASPL is measured

between the microjet control and no control cases, while the test conditions (h/d, TR)

remain the same. Please refer to the Figure 3-30 for a graphical presentation of “delta

OASPL”.

In Figure 3-31(a), pressure fluctuation intensities at different TRs were measured

at the stagnation point of the jet on the ground plane. There is a significant reduction in

unsteadiness at all the temperature ratios at small h/d. However, at this point of the flow

the reduction is highest at the ambient temperature jet. As discussed before (see 3.2.2),

the unsteadiness of the jet at the jet centerline may in part be affected by the

underexpansion of the jet.

The underexpansion is reduced when microjet control is applied due to the

reduced entrainment rate as the azimuthal vorticity is eliminated at the cost of streamwise

vorticity, as described in 3.4, leading to a reduction in broadband levels (see Figure 3-27)

and therefore a reduction in OASPL. As the stagnation point is influenced only partially

by the feedback loop as shown in the spectra graphs (see Figure 3-24), the reduction in

the overall unsteadiness of the jet when microjet control is applied is higher for cold jet

than hot jets. This may be due to the effects of the impingement zone on the unsteadiness

outweighing the effects of the feedback loop at the stagnation point of the impingement

when the jet is heated.

74

h/d

de

lta

OA

SP

L(d

B)

0 2 4 6 8 10 120

2

4

6

8

10

12

14

16

18

20

22TR=1.0TR=1.6

Ground Plane, x/d=1

(b)

h/d

de

lta

OA

SP

L(d

B)

0 2 4 6 8 10 120

2

4

6

8

10

12

14

16

18

20

22 TR=1.0

TR=1.6Lift Plate, x/d=2

(c)

h/d

de

lta

OA

SP

L(d

B)

0 2 4 6 8 10 120

2

4

6

8

10

12

14

16

18

20

22TR=1.0

TR=1.6Microphone, x/d=15

(d)

h/d

de

lta

OA

SP

L(d

B)

0 2 4 6 8 10 120

2

4

6

8

10

12

14

16

18

20

22TR=1.0TR=1.6

Ground Plane, x/d=0

(a)

Figure 3-31: Effect of Temperature on Delta OASPL as a Function of h/d at Different Measurement

Locations in the Flow (a) Ground Plane (x/d=0), (b) Ground Plane (x/d=1), (c) Lift Plate (x/d=2), and

(d) Microphone (x/d=15)

In Figure 3-31 (b), the effect of control on OASPL measured at 1 inch away from

the stagnation point is shown. The control in this case is more effective for cold jet only

at certain nozzle-to-ground plane distances. There seems to be a maximum reduction at

h/d=3. This shows that the effect of the feedback loop and main jet shear layer is more

prominent as we move away from the center of the jet. Therefore, the microjet control is

becoming slightly enhanced at points where the unsteadiness is primarily due to the

feedback loop being strengthened.

When considering locations close to the shear layer, which is almost exclusively

affected by the feedback loop (Lift plate, x/d=2), as depicted in Figure 3-31 (c), one can

see that the microjet control is very effective at decreasing unsteadiness due to the

feedback loop. In Figure 3-31 (c), OASPL measured at the lift plate two diameters away

Larger and

more

persistent

reduction

75

from the jet centerline, is plotted as a function of h/d. It has been described earlier that

there is evidence of the feedback loop being completed faster at higher temperatures and

being persistent over a larger range of h/d’s. It is observed in Figure 3-31 (c) that the

microjet control attenuates this stronger feedback loop which leads to a higher reduction

in OASPL at higher stagnation temperature. There is as much as 14dB reduction in

OASPL at h/d=4.0 for TR=1.6, which is very large when compared to the 6dB reduction

at TR=1.0, at the same height. The reduction, due to microjet control, also persists for

longer distances (up to h/d=5) when the jet is heated, which can be explained by the

feedback loop persisting for a larger range of h/d’s when the jet is heated as mentioned

previously. This is clear evidence that the microjet control is efficient at interrupting the

feedback loop for hot and cold jets, and therefore decreasing the adverse effects caused

by it. In Appendix X the reduction at TR=1.2 and 1.4 is shown.

3.3 TEMPERATURE MEASUREMENTS

In addition to the unsteady and steady pressure measurements, the variations of

temperature on the ground plane and lift plate are measured using thermocouples. The

results obtained on the ground plane using an array of seven thermocouples will be

discussed in the following sections. The exact location of the thermocouples can be found

in the 2.2.3. Similar to most of the measurement techniques used throughout this

experiment, the thermocouples were used to attain data at all four temperature ratios

(TR=1, 1.2, 1.4 and 1.6). For conciseness, only data at select TRs, representative of the

cold and hot impinging supersonic jet, will be considered. Results of all other conditions

are shown in Appendix xx. The consequence of increasing the stagnation temperature in a

supersonic impinging jet on the unsteady thermal loads on the ground plane will be

shown as well as how microjet control affects the thermal loads.

3.3.1 UNSTEADY TEMPERATURE LOADS

The unsteadiness in the temperature on the ground plane in the form of power

spectra density (PSD) is compared to the unsteadiness in the pressure data. Power spectra

76

density in this case represents the unsteady thermal load distribution in the frequency

domain. The peak tone frequencies seen before in 3.2.2 should be obvious in this spectra

in form of peak amplitude thermal loads, thus confirming that the feedback loop effects

are felt in the unsteady thermal load distribution.

Frequency (kHz)

TP

SD

(K^2

/Hz)

SP

L(d

B)

5 10 15 20 25 30 3510

-6

10-5

10-4

10-3

10-2

10-1

120

130

140

150

160

170

180

Thermocouple, x/d=0Kulite, x/d=0

TR=1.0, h/d=4No Control

Figure 3-32: Comparison of the unsteady thermal load distribution and the unsteady pressure load

distribution in the frequency domain at TR=1.0

In Figure 3-32 the unsteady pressure spectra and the unsteady temperature

distribution are presented in the frequency domain. Both data sets have been acquired at

the same temperature ratio (TR=1.0), the same height (h/d=4) and at the same location in

the ground plane, namely the stagnation point of the impinging jet (x/d=0). Here, the

spectra are similar, with the high amplitude pressure tones and peak thermal loads

occurring at the same frequencies. This points toward the conclusion that at TR=1.0, the

unsteadiness felt in the thermal loads as well as in the pressure are influenced by the same

77

phenomenon. That phenomenon is the feedback loop created by the jet impinging on the

ground.

A similar graph is presented in Figure 3-33 for a heated jet at TR=1.6. Even when

the jet is heated, the frequency of the high amplitude thermal loads and the high

amplitude pressure tones remains the same. These high amplitude tones are due to the

impinging tones created by the feedback loop. Following that, it is safe to conclude that

the high amplitude thermal loads seen in the power density spectra are also caused by the

impinging tones created by the feedback loop. Previously, in Figure 3-24 we have seen

that the frequency of these high amplitude unsteady pressure loads increases when the

stagnation temperature is increased. The same can be expected from the unsteady thermal

loads.

Frequency (kHz)

TP

SD

(K^2

/Hz)

SP

L(d

B)

5 10 15 20 25 30 3510

-6

10-5

10-4

10-3

10-2

10-1

120

130

140

150

160

170

180

Thermocuple, x/d=0

Kulite, x/d=0TR=1.6, h/d=4.0No Control

Figure 3-33: Comparison of the unsteady thermal load distribution and the unsteady pressure load

distribution in the frequency domain at TR=1.6

78

EFFECT OF TEMPERATRUE

In Figure 3-34 (a) to (d) unsteady thermal spectra is presented at different

locations on the ground plane. Figure 3-34 (a) being the stagnation point (x/d=0),

followed by spectra at several points away from the jet centerline, with the furthest one in

Figure 3-34 (d) at x/d=1.8. Each plot depicts data obtained at all the different temperature

ratios at a specific location on the ground plane. This allows for a comparison of the

changes in the unsteady thermal loads as the temperature of the jet is increased.

There is no apparent change in the broad band levels in the temperature spectra.

However, there is an obvious trend where at all the different locations on the ground

plane, the frequency of the peak unsteady thermal loads is increasing as the jet is heated.

This is the same trend that was observed in unsteady pressure spectra mentioned before

(see Figure 3-24). This again is an indication that the feedback loop, which is the cause of

these high amplitude thermal loads, is completed faster when the jet is heated. As the

distance to the stagnation point is increased there does not seem to be much of an

influence on the spectra except for the changes in the amplitude of the peak thermal

loads. However, there is no discernable trend to be observed.

79

Frequency (kHz)

TP

SD

(K2̂

/Hz)

5 10 15 20 25 30 3510

-6

10-5

10-4

10-3

10-2

10-1

TR=1.0

TR=1.2

TR=1.4TR=1.6

h/d=4, No ControlThermocouple at x/d=0

(a)

Frequency (kHz)

TP

SD

(K2̂

/Hz)

5 10 15 20 25 30 3510

-6

10-5

10-4

10-3

10-2

10-1

TR=1.0

TR=1.2

TR=1.4TR=1.6

h/d=4, No ControlThermocouple at x/d=0.3

(b)

Frequency (kHz)

TP

SD

(K2̂

/Hz)

5 10 15 20 25 30 3510

-6

10-5

10-4

10-3

10-2

10-1

TR=1.0TR=1.2

TR=1.4

TR=1.6


(c)

Frequency (kHz)

TP

SD

(K^2

/Hz)

5 10 15 20 25 30 3510

-6

10-5

10-4

10-3

10-2

10-1

TR=1.0

TR=1.2

TR=1.4

TR=1.6


(d)

Figure 3-34: Effect of jet temperature on the unsteady thermal loads at different locations on the

ground plane (a) x/d=0, (b) x/d=0.3, (c) x/d=1.2 and (d) x/d=1.8

The next step is to consider the effectiveness of microjet control in reducing

unsteady thermal loads at different jet stagnation temperatures.

EFFECT OF CONTROL

High unsteady thermal loads are the cause of ground erosion. It was shown in the

previous section these unsteady thermal loads increase in frequency as the temperature of

the jet is increased. Reducing these high amplitude loads would result in a reduction of

ground erosion caused by the jets impinging on the ground, making the aircrafts that use

vertically impinging jets usable on different terrains and extending the lifetime of the

80

landing surface. In addition, the fact that the impinging jet generate high unsteady loads

also suggests that impinging jet can be used to enhance heat transfer.

In Figure 3-35 the unsteady temperature is plotted at TR=1.0 and h/d=4, where

the data was obtained by the thermocouple mounted at the stagnation point in the ground

plane.

Frequency (kHz)

TP

SD

(K^2

/Hz)

5 10 15 20 25 303510

-7

10-6

10-5

10-4

10-3

10-2

10-1

No Control

With Control

TR=1.0, h/d=4.0Thermocouple at x/d=0

Figure 3-35: Effect of Control on Unsteady Thermal Loads at the Stagnation Point (x/d=0) on the

Ground Plane at TR=1.0 and h/d=4

The case where the microjets are not turned on is presented as the “no control”

case and the data obtained when the microjets are in full effect is presented as the “with

control” case. Unsteady thermal loads have been reduced overall with the most

significant reduction at frequencies corresponding to the impinging tones. This implies

that microjet control is very effective at colder temperatures. Most real life jets however

Reduction in Peak

Amplitude at TR=1.0

81

are hot, therefore the heated jet at TR=1.6 will be considered next in Figure 3-36, in order

to see whether this form of control is still effective.

Frequency (kHz)

TP

SD

(K^2

/Hz)

5 10 15 20 25 303510

-7

10-6

10-5

10-4

10-3

10-2

10-1

No Control

With Control

TR=1.6, h/d=4.0Thermocouple at x/d=0

Figure 3-36: Effect of Control on Unsteady Thermal Loads at the Stagnation Point on the Ground

Plane for a Hot Jet at TR=1.6 and h/d=4

In Figure 3-36 the effect of control is shown for a heated jet at TR=1.6. The same

nozzle-to-ground plane distance of four is presented as an example of all the different

heights. There is a dramatic reduction in the unsteady thermal loads even when the jet is

heated and the most noticeable reduction occurs at frequencies corresponding to the

impinging tones.

It is very promising to show that microjet control can be effective in reducing

thermal loads and therefore ground erosion not only for cold jets but also for more

Reduction in Peak

Amplitude at

TR=1.6

82

realistic cases, such as the heated jet at TR=1.6. In order to further explore this, the mean

temperature distribution on the ground plane is considered next.

3.3.2 MEAN TEMPERATURE DISTRIBUTION

High temperature impinging jets can be very damaging to the ground and

therefore it is important to determine what the temperature distribution is on the ground

plane is and how it is affected by the different nozzle-to-ground distances, different

temperatures of the jet, and how effective microjet control can be in reducing these

temperatures.

The measured temperatures are expressed in a dimensionless form as the recovery

factor, r. The recovery factor was calculated using Equation 6. This is a common way

to study the wall temperature distribution (Goldstein, 1986).

dT

TTr

)(1 0−

+= Equation 6: Recovery Factor

T: wall temperature or the temperature measured on the ground plane

T0: stagnation temperature of the jet

Td: dynamic temperature of the jet, calculated using Equation 7.

0

2

22

2

11

2

1

2T

M

M

C

UT

P

j

d

⎟⎠⎞

⎜⎝⎛ −

+

⎟⎠⎞

⎜⎝⎛ −

==γ

γ

Equation 7: Dynamic Temperature

The effect of the temperature on the stagnation recovery factor with and without

microjet control will be considered in the following section.


The stagnation recovery factor, r0 will be considered and how it changes when h/d

is increased from 2 to 12 as well as TR is increased from 1.0 to 1.6. The Stagnation

recovery factor is calculated using Equation 6 and temperatures measured at x/d=0.

83

h/d

r 0

0 1 2 3 4 5 6 7 8 9 10 11 120.6

0.8

1

1.2

1.4No Control

With Control(d) TR=1.6

h/d

r 0

0 1 2 3 4 5 6 7 8 9 10 11 120.6

0.8

1

1.2

1.4No Control

With Control(c) TR=1.4

h/d

r 0

0 1 2 3 4 5 6 7 8 9 10 11 120.6

0.8

1

1.2

1.4 No Control

With Control(b) TR=1.2

h/d

r 0

0 1 2 3 4 5 6 7 8 9 10 11 120.6

0.8

1

1.2

1.4 No Control

With Control

Goldstein et al.

(a) TR=1.0

Figure 3-37: Effect of Temperature Ratio and Height (h/d) on the Stagnation Recovery Factor

In Figure 3-37 (a)-(d) each separate plot represents data obtained as the

temperature ratio is increased from (a) TR=1.0 to (d) TR=1.6. The stagnation recovery

factor is plotted versus the nozzle to impingement plate distance. For each of the four

temperature ratios, the base line case is plotted (filled symbols) as well as the controlled

case (open symbols). Additionally, the recovery factor for a subsonic (M=0.47, Reynolds

number=1.24 x 105) impinging jet from Goldstein et al. (1986) has also been plotted for

comparison in Figure 3-37 (a).

At temperature ratio TR=1.0, the stagnation recovery factor stays close to unity

only for h/d=2 and 2.5. It increases after that as h/d is increased. This increase in the

stagnation recovery factor is due to the increased entrainment of warmer ambient air, as

84

the jet static temperature is reduced below the ambient temperature in the case of a cold

jet. Please note that the Goldstein et al. (1986) data shows traits very similar to the data

obtained in this experiment. Additionally, similar to the observations made by Goldstein

et al. (1986) for a circular impinging jet, the data in this case also shows that as the jet

impingement distance is increased, the amount of warmer air that is entrained by mixing

is increased. This leads to a boost in the static temperatures of the jet as the h/d is

increased, which can be seen in the increase of the stagnation recovery factor in Figure

3-37 (a) as the h/d becomes larger.

When the microjets are turned on at TR=1.0, the stagnation recovery factor stays

close to unity for a larger extent of h/d’s, up to h/d=6. After that it increases but stays

slightly below the uncontrolled case. It has been shown by Lou et al. (2003) with the help

of PIV measurements that microjet control reduces the azimuthal vorticity and redirects it

into the streamwise direction. This leads to a decrease in the entrainment velocities

responsible for the mixing. Decreased entrainment of warmer ambient air leads to a

stagnation factor closer to unity.

In the case of higher temperatures as seen in Figure 3-37 (b) through (d) the

stagnation recovery stays close to unity for small nozzle to ground distances, up to h/d=5.

In these heated jet cases without control, the temperature recovery factor decreases for

larger h/d. This decrease in the recovery factor is again due to the increase in mixing

downstream of the jet. In this case however the static temperature of the jet is higher than

the ambient air temperature. Therefore as the colder ambient air is entrained the static

temperature of the jet is decreased, which leads to a decrease in the stagnation recovery

factor for larger h/d. As the control is applied in the case of hot jets there is now an

increase in the stagnation recovery factor. This is again due to the decrease in

entrainment velocities when microjet control is applied. A decrease in the entrainment of

the colder ambient air leads to a stagnation recovery factor closer to unity.

In summary, for cold jets there is a net heat flux coming in from the warmer

ambient air leading to an increase in the recovery factor as the mixing is increased at

higher h/d. For hot jets there is a net heat flux coming out due to the ambient air being

colder than the static temperature of the jet. This leads to a decrease in the recovery factor

85

as the mixing is enhanced for higher h/d. For both cases, microjet control reduced the

mixing or entrainment of ambient air, which affects the recovery factor accordingly.

Another side effect of the feedback loop that illustrates the change in entrainment

velocities, as the jet is heated and microjet control is activated, is the lift loss.

3.4 LIFT LOSS

As described in Chapter 1, the presence of large structures in the shear layer

enhances the entrainment of the ambient air especially close to the nozzle exit. This

increased entrainment rate causes vacuum pressures in the region around the jet. These

vacuum pressures are the cause of the suck down force that is acting opposite to the lift

force. High lift loss can be a significant problem when designing an aircraft that

implements Short-Take-Off-Vertical-Landing.

To simulate an aircraft structure around the nozzle in this experimental setup, a

circular plate was attached to the nozzle and flush mounted with the nozzle exit plane. As

described in Chapter 2, 19 ports along a radial line were used to measure the static

pressure. The exact pressure port distribution is described in Chapter 2 as well as in

Fig. ?. The data acquisition program records the differential pressure at each port.

Differential pressure is the difference between the surface pressure at the port and the

ambient pressure. In Figure 3-38, one can see a typical differential pressure distribution at

a particular ground-plate-to-nozzle-exit distance. This vacuum pressure which is obtained

at one height is integrated by dividing the lift plate into 19 circular strips. Each strip

begins in the middle of two pressure ports and ends in the middle of the following two

pressure ports. The area of each of these strips is multiplied by the vacuum pressure

obtained at the common pressure port, therefore giving the force on that particular strip of

the lift plate. By adding all those individual forces the total force can be calculated. In

this case this total force is also called the suck-down force, since it acts in a direction

opposite to the jet thrust. In order to show how this suck-down force influences the thrust

the lift loss is normalized by the total thrust of this jet. Assuming this is an ideally

expanded jet with a Mach 1.5 velocity, the jet thrust was calculated to be 189 N, using the

nozzle exit conditions and isentropic relations.

86

radial location, mmV

acu

um

Pre

ssu

re,P

v0 20 40 60 80 100 120

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7


h/d=1.5NPR=3.7, TR=1.4

(b)

radial location, mm

Va

cu

um

Pre

ssu

re,P

v

0 20 40 60 80 100 1200

0.1

0.2

0.3

0.4

0.5

0.6

0.7


h/d=1.5NPR=3.7, TR=1.0

(a)

Figure 3-38: Effect of Control on the Vacuum Pressure Distribution on the Lift Plate at (a) TR=1.0

and (b) TR=1.4

Two graphs are shown in Figure 3-38, both of them represent a typical vacuum

pressure distribution for a cold jet (a) and a hot jet (b) along a radial with respect to the

centerline of the jet. Each plot has two curves where the open symbols represent the

pressure distribution without any control and the filled symbols, the pressure distribution

once the microjet control is applied.

First thing one notices, for the no control cases is that when the jet stagnation

temperature is increased, the vacuum pressures that are acting downwards on the lift plate

are increased as well. In case of the hot jet where these pressures close to the nozzle exit

are elevated it hints towards the jet being slightly more underexpanded. These higher

vacuum pressures might be due to the strengthening of the feedback loop at higher

temperatures which leads to higher entrainment rates, as seen in the previous chapter, and

therefore increased vacuum pressures.

Next is to look at what the effect of control is at the higher stagnation

temperature. At the cold condition (TR=1.0), the amazing reduction of these vacuum

pressures can be observed in Figure 3-38(a). Additionally in Figure 3-38(b), even though

the no control case leads to higher vacuum pressures in the hot case, microjet control

87

reduces these vacuum pressures almost to the same levels as in the cold jet. The

importance of the reduction in these vacuum pressures is that these pressures are

responsible for the lift loss or loss of a part of the thrust force. When these pressures,

which act opposite to the thrust force are reduced the efficiency of the STOVL aircraft

design is improved.

One interesting trend in both of the above graphs is that as the radial distance

increases the vacuum pressures stay fairly constant, until there is a sudden increase at the

last few pressure ports, known as the end effect. This vacuum pressure is created due to

the entrainment of the ambient air. As the ports move away from the jet centerline the

entrainment should be reduced and the pressure should adjust to ambient pressure,

meaning the gauge pressure should go to zero. Hence, in an actual aircraft, the

undersurface is very large compared to the lift plate, in this study. It is safe to presume

that the surface pressure would eventually go to ambient pressure levels if the surface of

the lift plate was to be increased by a sufficient amount. It was shown by Garg (2001)

that for a larger lift plate with the same set up as used in this trial, the differential

pressures do indeed relax back to ambient pressure at about 200 mm away from the

centerline.

H/D

-Lift

Lo

ss/T

hru

st

0 2 4 6 8 100

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

No Control

With Control

TR=1.0, NPR=3.7

(a)

H/D

-Lift

Lo

ss/T

hru

st

0 2 4 6 8 100

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

No Control

With Control

TR=1.4, NPR=3.7

(b)

Figure 3-39: Variation of Lift Loss with nozzle-to-plate distance at (a) TR=1.0 and (b) TR=1.4

88

In Figure 3-39 normalized lift loss changes as a function of h/d are plotted. There

are two graphs, where each one represents one stagnation temperature condition. The

cold jet [Figure 3-39(a)] is represented as TR=1.0 and the hot jet [Figure 3-39(b)] is

represented by TR=1.4. Concentrating on the cold jet with no control, it is apparent that

as the distance between the nozzle exit and the impinging plate is increased the lift loss is

decreased. There is a maximum of lift loss at h/d=1.5 at about 53% of the isentropic

thrust of the jet. This is to be expected as the entrainment velocities are higher at the

smaller h/d. The activation of microjets for this cold case leads to a lift loss recovery of

about 40% at h/d=1.5. It has been shown before that microjet control redirects azimuthal

vorticity into streamwise direction which leads to reduced entrainment velocities, and

therefore a reduction in the vacuum pressures responsible for the lift loss.

The next important question is what influence the increased stagnation

temperature of the jet has on the lift loss as well as on the lift loss recovery. To answer

this, the graph in Figure 3-39 (b) is considered next. The no control line in the heated

flow shows the same trend as the cold case. However, the levels of the lift loss are

significantly higher when the jet is heated. The maximum lift loss for the jet at TR=1.4 is

again at h/d=1.5, but it is almost 76% of the jet thrust. As in the cold case, the lift loss

decreases dramatically as the jet is moved away from the impinging plate, therefore

indicating that the problematic lift loss is only an issue at small h/d (corresponding to

hover), as previously discussed. At this point it has been shown that the lift loss is very

large at small h/d and it increases with increased stagnation temperature of the jet. Thus it

can be concluded that the feedback loop is stronger when the jet is heated as well as at

small h/d. This is similar to the observations made in the unsteady pressure

measurements and temperature measurements discussed before.

How well does the microjet control work at this increased temperature? Microjet

control interrupts the feedback loop and eliminates/reduces the azimuthal vorticity and

redirects it into streamwise direction, therefore it comes as no surprise that when the

control is applied in the heated case the recovery at small h/d is even better at 46% (at

h/d=1.5) than in the cold flow.

In summary, one can state that the lift loss caused by the vacuum pressures is

maximum at small h/d regardless of the temperature of the jet. It is however higher when

89

the jet is heated. On a positive note, half of the lift loss is recovered when the microjet

control is activated at cold and hot jet flow conditions.

90

CHAPTER 4

CONCLUSION

The main objective of this study was to continue the work done by Alvi et

al. (2000, 2003) and Lou et al. (2003) by exploring the effect of temperature on

supersonic, impinging jets and their control. The highly oscillatory flow field and the

associated feedback loop are responsible for many of the adverse effects when employed

on a STOVL aircraft. In the following summary of the results obtained during this study,

the change in those adverse effects of the feedback loop with increasing temperature will

be discussed. As well as, how effective implementing microjet control at these conditions

may be.

In the present studies, the effect of a number of parameters was considered. Mean

pressure measurements were obtained on the ground plate for an ideally expanded jet as

well as for a moderately underexpanded jet. Unsteady pressure measurements were

conducted at the lift plate and at the ground plate, in order to determine how the unsteady

pressure loads change with temperature, hover height and control. Temperature

measurements were also obtained at the lift plate and ground plate, and are indicative of

the amount of thermal loading that is imposed on the landing surface as well as on the

undersurface of the aircraft. Mean pressure measurements on the lift plate provided the

necessary data to calculate the lift loss endured when the jet is heated but also the lift loss

recovered when microjet control is applied.

The mean pressure distribution on the ground plane shows that there is evidence

of a stagnation bubble at hover heights h/d=3 and 5 for both, cold and hot jets. This is

very similar to the results obtained by Iyer (1999) and Garg (2001) for cold jets.

However, the stagnation bubble seems to be weakened when the jet is heated due to the

higher entrainment rate of the ambient air. The mean pressure fluctuations in the wall jet

are created by the shock cells in the primary jet and therefore depend largely on the

underexpansion of the primary jet as well as on the occurrence of the stagnation bubble.

Additionally, higher entrainments rates in hot jets are responsible for the faster decay in

peak Cp values.

91

Unsteady pressure loads cause sonic fatigue, ground erosion, hot gas ingestion as

well as noise pollution. The results of this study show that the frequency of the discrete

tones, which can be found in unsteady pressure spectra as well as in narrowband noise

spectra is increased as the temperature of the jet is increased. The increase in frequency is

more pronounced at the lift plate. However, data at the stagnation point of the flow shows

some influence by the structure of the impingement zone, in addition to the influence of

the feedback loop. When microjet control is applied, the distinct impinging tones are

attenuated at every temperature as well as at all the measurement locations. There is also

a significant reduction in broadband levels which is pointing towards an overall decrease

in jet unsteadiness.

Overall sound pressure levels were considered in order to assess the effect of

temperature and control on the total unsteady pressure loads at different impingement

distances. At the lift plate, overall pressure levels increase significantly (up to 9dB at

h/d=4) with increasing temperature. Additionally for hot jets, these high OASPL are

maintained for a larger range of hover heights, up to h/d=6. This suggests that the

feedback loop is becoming stronger and more persistent when the temperature of the jet is

elevated. Similar results are seen in the near-field region.

However, at the stagnation point of the flow, the higher temperature jets seem to

have a minimum at small h/d’s and the fluctuation pressures are lower in the hot jet than

in the colder jets. This might be due to the stagnation point being not only affected by the

feedback loop but also by the changes in structure in the impingement zone.

Again, microjet control is effective at reducing the overall unsteady pressure

levels at every point of measurement and at every temperature of the jet. Up to 20dB

reduction is reached at the stagnation point and up to 14dB of reduction is achieved on

the lift plate at small nozzle to ground distances. Noise levels, as measured with the

microphone in the near-field, were also reduced by up to 8dB for certain conditions. The

reduction seems to persist for a larger range of h/d at the lift plate for hot jets.

The temperature spectral density indicates similar unsteady thermal load behavior

to the ones seen in the pressure spectra. These peak thermal loads increase in frequency

as the temperature of the jet is increased. This can result in very high unsteady thermal

loads, which can negatively affect the landing surface as well as damage the undersurface

92

of the aircraft. However, this study is able to show that employing microjet control leads

to a significant reduction in these high unsteady thermal loads for cold and hot jets.

The temperature recovery factor that was calculated shows that the height as well

as the temperature has a large impact on the entrainment of the ambient air. This increase

in entrainment with increase in temperature, points once again towards the strengthening

of the feedback loop. . When control is applied we notice a decrease in entrainment of the

ambient air. The decrease in entrainment when control is applied indicates again the

effect that microjets have on the feedback loop.

Finally, one of the most adverse effects that are created by the feedback loop is

the lift loss. The large scale structures that drive the feedback loop in the shear layer are

responsible for the high entrainment velocities. These high velocities lead to high vacuum

pressures close to the nozzle exit, which produce the negative force that induces lift loss

in STOVL aircrafts. The results in this study show that the lift loss can be as much as

~50% of the primary jet thrust for cold jets and is increased to 75% of the jet thrust when

the stagnation temperature of the jet high. This amount of lift loss can cause severe

inefficiency for an aircraft. However, it has been shown before by Lou (2005) that

microjets generate streamwise vorticity at the expense of the azimuthal vorticity, which is

responsible for the increased entrainment of the ambient air in the cold impinging jet.

Therefore, microjet control is able to recover most of this lift loss. At cold temperature

(TR=1.0), 40% of the lift loss was recovered with the microjets. For the hot jet, when

microjet control was applied an amazing 46% of the lift loss was recovered.

The above results show that the heated, impinging jet is overall more unsteady

and causes more negative effects when compared to the cold jet. This overall

unsteadiness is attributed to the strengthening of the feedback loop. The strengthening of

the feedback loop might be due to the large scale structures in the shear layer growing

faster. Along with the more rapid growth, they travel downstream at higher velocities due

to higher jet temperature. This appears to result in a strengthening of the feedback loop

that can persist for a larger range nozzle-to-ground distances. However, it was shown that

most of the effects of this strengthened and more persistent feedback loop can be

attenuated using microjets with a mass flow rate less than 0.5% or the primary jet flux.

93

In summary, even though the adverse effects are more pronounced for hot jets,

this study has shown that microjet control is comparably more effective in reducing them.

94

APPENDIX A :

MEAN PRESSURE GRAPHS

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8 h4, NPR 3.7h4, NPR 5

TR=1.0

Figure A- 1: Effect of NPR on a Cold Jet (TR=1.0) at h/d=4

95

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8 h6, NPR 3.7h6, NPR 5

TR=1.0


96

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8 h8, NPR 3.7h8, NPR 5

TR=1.0


97

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8 h12, NPR 3.7h12, NPR 5

TR=1.0


98

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

h4, NPR 3.7h4, NPR 5

TR=1.4

Figure A- 5: Effect of NPR on a Heated Jet (TR=1.4) at h/d=4

99

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8


TR=1.4


100

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8


TR=1.4


101

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

h12, NPR 3.7h12, NPR 5

TR=1.4


102

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

TR=1.0TR=1.4

NPR=3.7 and h/d=4

(a)

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

TR=1.0TR=1.4

NPR=3.7 and h/d=6

(b)

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

TR=1.0TR=1.4

NPR=3.7 and h/d=12

(c)

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

TR=1.0TR=1.4

NPR=3.7 and Free Jet

(d)

Figure A- 9: Effect of Temperature on an Ideally Expanded Jet (NPR=3.7) at (a) h/d=4, (b) h/d=6, (c)

h/d=12, and (d) Free Jet

103

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

TR=1.0TR=1.4

NPR=5 and h/d=6

(b)

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

TR=1.0TR=1.4

NPR=5 and h/d=12

(c)

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

TR=1.0TR=1.4

NPR=5 and h/d=4

(a)

x/d

Cp

0 1 2 3 4-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

TR=1.0TR=1.4

NPR=5 and Free Jet

(d)

Figure A- 10: Effect of Temperature on a Moderately Underexpanded Jet (NPR=5.0) at (a) h/d=4, (b)

h/d=6, (c) h/d=12, and (d) Free Jet

104

APPENDIX B :

UNSTEADY PRESSURE GRAPHS

Frequency (kHz)

SP

L(d

B)

5 10 15 20 25 303580

100

120

140

160 No ControlWith ControlTR=1.4, h/d=4

Lift Plate, x/d=2

(c)

Frequency (kHz)

SP

L(d

B)

5 10 15 20 25 303580

100

120

140

160



(d)


(b)

Frequency (kHz)

SP

L(d

B)

5 10 15 20 25 303580

100

120

140

160

TR=1.0TR=1.2

TR=1.4

TR=1.6

h/d=4Lift Plate, x/d=3

(a)

Figure B- 1: Influence of Temperature on the SPL Spectra at the Lift Plate (x/d=3)

105

Frequency (kHz)

SP

L(d

B)

5 10 15 20 25 30 35120

130

140

150

160

170

180

190TR=1.0TR=1.2TR=1.4TR=1.6


Figure B- 2: Influece of Temperature on the SPL Spectra at the Ground Plate (x/d=1)

106

ST=(f*d)/U

SP

L(d

B)

0.5 1 1.5 2100

110

120

130

140

150

160

170

180

190TR=1.0

TR=1.2TR=1.4TR=1.6


Figure B- 3: Unsteady Pressure Spectra Obtained at the Ground Plate (x/d=1) at all TR's with

Frequency scaled using the Strouhal Number

107

Frequency (kHz)

SP

L(d

B)

5 10 15 20 25 30 35120

130

140

150

160

170

180

190TR=1.0TR=1.2TR=1.4TR=1.6


Figure B- 4: Influence of Temperature on the SPL Spectra at the Ground Plate (x/d=3)

108

ST=(f*d)/U

SP

L(d

B)

0.5 1 1.5 2100

110

120

130

140

150

160

170

180

190TR=1.0

TR=1.2TR=1.4TR=1.6


Figure B- 5: Unsteady Pressure Spectra obtained at the Ground Plate (x/d=2) at all TR's with

Frequency scaled using the Strouhal Number

109

h/d

de

lta

OA

SP

L(d

B)

0 2 4 6 8 10 120

2

4

6

8

10

12

14

16

18

20

22TR=1.0

TR=1.2

TR=1.4

TR=1.6

Ground Plane, x/d=1

h/d

de

lta

OA

SP

L(d

B)

0 2 4 6 8 10 120

2

4

6

8

10

12

14

16

18

20

22TR=1.0

TR=1.2TR=1.4

TR=1.6

Microphone, x/d=15

h/d

de

lta

OA

SP

L(d

B)

0 2 4 6 8 10 120

2

4

6

8

10

12

14

16

18

20

22 TR=1.0

TR=1.2

TR=1.4TR=1.6

Lift Plate, x/d=2

h/d

de

lta

OA

SP

L(d

B)

0 2 4 6 8 10 120

2

4

6

8

10

12

14

16

18

20

22TR=1.0TR=1.2

TR=1.4

TR=1.6

Ground Plane, x/d=0

Figure B- 6: Effect of Temperature on Delta OASPL as a function of h/d at different measurement

locations in the Flow (a) Ground Plate (x/d=0), (b) Ground Plate (x/d=1), (c) Lift Plate (x/d=2), and (d)

Microphone (x/d=15)

110

APPENDIX C :

RADIAL TEMPERATURE RECOVERY FACTOR

x/d

Re

co

ve

ryF

acto

r,r

0 0.5 1 1.5 20.8

0.9

1

1.1

1.2h/d=2

h/d=3h/d=4h/d=6

h/d=8h/d=10h/d=12

TR=1.0, No Control

Figure C- 1: Radial Distribution of the Recovery Factor at TR=1.0

111

x/d

Re

co

ve

ryF

acto

r,r

0 0.5 1 1.5 20.6

0.7

0.8

0.9

1

1.1

1.2 h/d=2h/d=3h/d=4h/d=6h/d=8h/d=10h/d=12

TR=1.4, No Control

Figure C- 2: Radial Distribution of the Recovery Factor at TR=1.4

112

x/d

Re

co

ve

ryF

acto

r,r

0 0.5 1 1.5 20.8

0.9

1

1.1

1.2 h/d=2, No Control

h/d=2, With Controlh/d=6, No Controlh/d=6, With Control

TR=1.0

Figure C- 3: Effect of Control on the Radial Recovery Factor Distribution at TR=1.0

113

x/d

Re

co

ve

ryF

acto

r,r

0 0.5 1 1.5 20.6

0.7

0.8

0.9

1

1.1

1.2h/d=2, No Controlh/d=2, With Controlh/d=4, No Controlh/d=4, With Controlh/d=12, No Controlh/d=12, With Control

TR=1.4

Figure C- 4: Effect of Control on the Radial Recovery Factor Distribution at TR=1.4

114

APPENDIX D :

LIFT LOSS GRAPHS

radial location, mm

Va

cu

um

Pre

ssu

re,P

v

0 20 40 60 80 100 1200

0.1

0.2

0.3

0.4

0.5

0.6

0.7

TR=1.4

TR=1.0

h/d=1.5, No Control

(a)

radial location, mm

Va

cu

um

Pre

ssu

re,P

v

0 20 40 60 80 100 1200

0.1

0.2

0.3

0.4

0.5

0.6

0.7

TR=1.4

TR=1.0

h/d=1.5, With Control

(b)

Figure D- 1: Effect of Temperature on the Vacuum Pressure Distribution on the Lift Plate for the

(a) no control case and (b) controlled case

h/d

-Lift

Lo

ss/T

hru

st

0 2 4 6 8 100

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8TR=1.0

TR=1.4

No Control

h/d

-Lift

Lo

ss/T

hru

st

0 2 4 6 8 100

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8TR=1.0

TR=1.4

With Control

Figure D- 2: Effect of Temperature on the Lift Loss variation versus nozzle-to-plate distance for the

(a) no control case and (b) controlled case

115

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BIOGRAPHICAL SKETCH

The author was born in Sarajevo, Yugoslavia on July 24th

, 1981. In 1993 she

relocated to Munich, Germany were she completed the majority of her High School

education.

In 2000 she moved to Tallahassee, Florida and completed her senior year at Leon

High School. She started her freshman year at Florida State University in 2001, majoring

in Mechanical Engineering. She received her undergraduate degree in Mechanical

Engineering in 2006.

In 2006 she was admitted into graduate standing and started her work as a

graduate research assistant at the Advanced Aero Propulsion Laboratory. Dr. Alvi was

kind enough to serve as her graduate advisor throughout the next two years.

Active Control of Hot, Supersonic Impinging Jets Using Microjets

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