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This is to certify that the

thesis entitled

Detailed Flow Measurements within the Volute
of an Automotive HVAC Blower

presented by
Richard J. Prevost

has been accepted towards fulfillment
of the requirements for

MS degree in Mechanical Engineering

 
   
 

Major professor

Date August 21, 2003

0-7639 MS U is an Affirmative Action/Equal Opportunity Institution

PMCE IN RETURN BOX to remove this checkout from your record.
To AVOID FINES return on or before date due.
MAY BE RECALLED with earlier due date if requested.

 

DATE DUE

DATE DUE

DATE DUE

 

M

 

022105

 

 

 

 

 

 

 

 

 

 

 

 

6/01 c:/CIRC/DateDue.p65-p. 15

 

DETAILED FLOW MEASUREMENTS WITHIN THE VOLUTE
OF AN AUTOMOTIVE HVAC BLOWER

By

Richard J. Prevost

A THESIS

Submitted to
Michigan State University
in partial fulfillment of the requirements
for the degree of

MASTER OF SCIENCE

Department of Mechanical Engineering

2003

ABSTRACT

DETAILED FLOW MEASUREMENTS WITHIN THE VOLUTE
OF AN AUTOMOTIVE HVAC BLOWER

By

Richard J. Prevost

Experimental data that characterize the flow field within the volute of an automo-
tive HVAC (heating/ventilation and air conditioning) blower have been acquired and ana-
lyzed. The single inlet blower consisted of a forward curved (FC) blade impeller and
volute with a simple rectangular cross section. The performance characteristics of FC
blowers makes them especially desirable for automotive applications given the tight pack-
aging constraints of automotive systems. However, the F C blower is an inefficient flow
device and is the dominant contributor of noise in the vehicle cabin at high blower speeds.

Velocity and pressure measurements have shown the volute flow field to be highly
three-dimensional with significant turbulence levels. These measurements have also
revealed-the well known salient flow features associated with blowers, such as the "jet-
wake" pattern, the "double vortex" secondary flow, and the inefficient use of the impeller
width. These features are believed to be significant loss and noise generation mecha-
nisms. Insight on how the impeller/volute interaction may contribute to these features is
presented. The results of this study may be of particular interest to the designers of such

blowers and to those persons who wish to validate their CFD predictions.

© Copyright by
RICHARD J. PREVOST
2003

ACKNOWLEDGEMENTS

First, I would like to thank my advisor Prof. John Foss for his guidance and fman-
cial support during my graduate studies at MSU. My days at the Turbulent Shear Flows
Laboratory proved to be extremely beneficial in my development as an experimentalist.
My colleagues at the TSFL have also attributed greatly in this endeavour. I would like to
thank Douglas Neal, Al Lawrenz, Scott Treat, Matthew Maher, Kyle Bade, Mike Dusel
and all of the others that had passed through our laboratory doors during my studies. A
special thanks goes out to Prof. Scott Morris for first introducing me to the fine world of
fluid dynamics. Thanks also to my committee members, Professors Ahmed Naguib and
Charles Petty, for their input on my research.

I would like to thank Clemente Mesa of the DaimlerChrysler Corporation for
bringing this project to our attention and providing funding through the DaimlerChrysler
Challenge Fund Program. I would also like to thank his colleagues Stefan Wozniak and
Jenny Zheng for providing technical support.

Finally, I would like to thank my family and friends who have been extremely sup-

portive throughout my life.

iv

TABLE OF CONTENTS

 

 

 

 

 

LIST OF FIGURES- .................. - - - ..... - -- - - - - ...... viii
LIST OF TABLES ......... -- - - - - ...... . ......... xvi
NOMENCLATURE -- - - - - -- -- - ..... xvii
1.0 Introduction- - - ......... - 1
1.1 Automotive HVAC Systems and Centrifugal Blowers ................................ 1

1.2 Previous Investigations ................................................................................ 3

1.3 The Research Program ................................................................................. 6

1.3.1 Motivation and Background ............................................................ 6

1.3.2 Experimental Overview ................................................................... 8

1.3.3 Experimental Prelude ....................................................................... 9

2.0 Experimental Equipment ............ - -- .................... 18
2.1 Flow System .............................................................................................. 18

2.2 Data Acquisition and Processing Systems ................................................. 20

2.3 Sonic Nozzle Test Stand ............................................................................ 20

2.4 Pressure Measurements .............................................................................. 21

2.4.1 Pressure Transducers ...................................................................... 21

2.4.2 Microphones .................................................................................. 22

2.5 Velocity Measurements .............................................................................. 22

3.0

4.0

2.5.1 Hot-Wire Anemometry .................................................................. 22

 

 

2.5.1.1 Hot-Wires, Anemometers and Thennistor ......................... 22

2.5.1.2 Calibration Facility ............................................................ 23

2.5.2 Particle Image Velocimetry (PIV) .................................................. 24
Experimental Program ........................... - - 31
3.1 Instrument Panel Measurements ................................................................ 32
3.1.1 Procedures and Features ................................................................ 33

3.2 Experimental Flow System Performance Measurements (Uninstalled Unit)
................................................................................................................... 36

3.2.1 Procedures and Features ................................................................ 37

3.2.1.1 Inlet Study .......................................................................... 39

3.3 Volute Flow Field Measurement Techniques ............................................. 39
3.3.1 Surface-Streaking ........................................................................... 39

3.3.2 Hot-Wire Anemometry .................................................................. 40

3.3.2.1 Calibration and Processing ................................................ 41

3.3.2.2 Measurement Locations and Acquisition .......................... 42

3.3.3 Static Pressure Measurements ........................................................ 44

3.3.4 Microphone Measurements ............................................................ 45

3.3.5 Particle Image Velocimetry ............................................................ 47

3.3.5.1 Measurement Configurations ............................................. 49
Results---.---..-. - - ..... -- - ......... - -- ..----62
4.1 Data Analysis ............................................................................................. 62

vi

4.2 Performance ............................................................................................... 64

 

 

 

 

4.2.1 Inlet Study ...................................................................................... 67

4.3 Surface-Streaking ....................................................................................... 68

4.4 Hot-Wire Anemometry .............................................................................. 69

4.5 Scrolled Surface Static Pressure Distribution ............................................ 74

4.6 Microphone: F luctuating Pressure Measurements ..................................... 74

4.7 Particle Image Velocimetry ........................................................................ 76

5.0 Discussion ........ - ....... - -- - - 120
5.1 Primary Flow ........................................................................................... 120

5.2 Double Vortex (Secondary Flow) ............................................................ 122

5.3 Active Impeller Width ............................................................................. 128

5.4 Separated flow and Recirculation ............................................................ 132

5.5 Jet-Wake ................................................................................................... 134

5.6 Flow Field Fluctuation Intensity .............................................................. 138

6.0 Summary and Conclusions - -- - - 142
APPENDIX Aw: .... - - ........... _ _ 146
APPENDIX B - - - - ...... - - _ ..... - 149
REFERENCES- - -- - - -- -.167

 

vii

LIST OF FIGURES

FIGURE 1.1 Centrifugal Impeller Blade Types; F C (Forward Curved Blade), BC (Back-
ward Curved Blade), RB (Radial Blade). .................................................... 10

FIGURE 1.2 Schematic of a Typical Centrifugal Blower With F C Blades. .................... 10

FIGURE 1.3 "Jet-Wake" Velocity Pattern (Stationary Observer) at the Exit of a PC Impel-
ler [26]. ........................................................................................................ 11

FIGURE 1. 4 Schematic of the Separated Flow at the Impeller Inlet [27]. (See Figure 1.10
for definition of d) ...................................................................................... 11

FIGURE 1.5 Computed Velocity Fields of Arbitrary Radial Cross Sections From Two Dif-
ferent Blowers Showing (a) the "Double Vortex” [9] and (b) the "Single Vor-
tex" [18]. The Secondary Flow is Super-Imposed on to the Primary

Tangential Flow. Note: The Inlet flow direction is From Right to Lefi ..... 12
FIGURE 1.6 Instrument Panel and HVAC Unit. .............................................................. 12
FIGURE 1.7 HVAC Climate Controls: Mode Control; (a) Panel, (b) Bi-Level, (c) Floor,

((1) Defrost/Floor, (e) Defrost. ..................................................................... 13
FIGURE 1.8 HVAC Unit Inlet Attachment. ..................................................................... 13
FIGURE 1.9 HVAC Unit Features. .................................................................................. 14
FIGURE 1.10 Blower Schematic and Nomenclature, see Table 1.1 for Definitions. ..... 15

FIGURE 1.11 Volute Area as a Function of Azimuthal Location and Linear Curve Fit. 17

FIGURE 2.1 Experimental Prototype. .............................................................................. 26
FIGURE 2.2 Experimental Flow System. ........................................................................ 26
FIGURE 2.3 Reflective Optical Sensor Schematic. ......................................................... 27
FIGURE 2.4 Sonic Nozzle Test Stand. ............................................................................. 27
FIGURE 2.5 Typical Microphone Frequency Response Curve (Panasonic Specifications
Sheet). .......................................................................................................... 28
FIGURE 2.6 Microphone Op Amp Circuit Schematic. .................................................... 28

viii

FIGURE 2.7 Schematic of a Single Sensor Hot-Wire Probe. .......................................... 29

FIGURE 2.8 Hot-Wire Calibration Facility. .................................................................... 29
FIGURE 2.9 Hot-Wire Calibration F acility’s Pressure Measurement Correction. . ......... 30
FIGURE 3.1 Front View of the IP Flow Rate Measurement System. ............................. 52
FIGURE 3.2 Rear View of the IP Flow Rate Measurement System. .............................. 52

FIGURE 3.3 Schematic of the APE Measurement Configuration (Dimensions in mm). .53

FIGURE 3.4 Elbow Meter Calibration Curve and Fit. (Note: 1 in H20 = 248.84 Pa). ..... 53

FIGURE 3.5 Performance Measurement Experimental Configuration. ........................... 54
FIGURE 3.6 Performance Static Pressure Measurement Locations. ............................... 54
FIGURE 3.7 Sample Hot-Wire Calibration (a) Curve and (b) Curve Fit. ........................ 55
FIGURE 3.8 Hot-Wire Measurement Configuration. ...................................................... 56

FIGURE 3.9 Hot-Wire Probe and Traverse System for (a) Near-Impeller Measurements
and (b) Near-Surface Measurements. .......................................................... 56

FIGURE 3.10 Schematic of Hot-Wire Configurations for (a) Near-Impeller Measurements
and (b) Near-Surface Measurements. .......................................................... 57

FIGURE 3.11 Entry Port and Hot-Wire Measurement Locations. Entry Ports are Located
Along the Mid-Plane (z = 37.25 mm) at 0 = (a) 66° (b) 99° (0) 132° ((1) 165°
(e) 193° (0 220° (g) 248° and (h) 276°. Hot-Wire Measurement Locations
are Indicated by the Hash Marks. ................................................................ 58

FIGURE 3.12 Static Pressure and Microphone Measurement Configuration. ................. 58

FIGURE 3.13 Microphone Calibration Configuration (a) Before Insertion and (b) During
Calibration. .................................................................................................. 59

FIGURE 3.14 PIV Basic Principles. (Images obtained from Dantec’s website). ............ 59

FIGURE 3.15 PIV Measurement Planes. The Dashed Lines Represent the x-y Planes while
the Dotted Lines Represent the z-r Planes. The x-y Planes are located at z =
17.25 m, 37.25 mm, and 60.25 mm for the Rear, Mid, and Inlet Planes, re-
spectively. The z-r Planes are Located at 0 = 101°, 168°, and 252° for Re-
gions I, II, and 111, respectively. Entry Port Locations are shown for
Reference Purposes. .................................................................................... 6O

FIGURE 3.16 Representative Contour Plot of Percent Vector Validation. Data are
Cropped at Volute and Impeller Boundaries. The "L" Shaped Region of Low
Validation In the Upper Right Comer of the Volute is Due To "Burned Out"

Pixels in the Camera’s CCD Chip. .............................................................. 61
FIGURE 4.1 Performance: Static Pressure Coefficient vs. Flow Coefficient. ................. 80
FIGURE 4.2 Performance: Static Efficiency vs. Flow Coefficient. ................................. 80
FIGURE 4.3 Performance: Reynolds Number vs. Flow Coefficient. .............................. 81

FIGURE 4.4 Performance: Normalized Average Net Flow Velocity vs. Flow Coefficient.
Collapse of Curves Show Confirmation of "Fan Laws". ............................ 81

FIGURE 4.5 Performance Inlet Study: Static Pressure vs. Flow Coefficient: Impeller speed

of 3000 rpm. ................................................................................................ 82
FIGURE 4.6 Performance Inlet Study: Static Efficiency vs. Flow Coefficient: Impeller

speed of 3000 rpm. ...................................................................................... 82
FIGURE 4.7 Surface-Streaking: Scrolled Surface. .......................................................... 83
FIGURE 4.8 Surface-Streaking: Inlet Surface (Images Cropped for Comparison). ........ 84
FIGURE 4.9 Surface-Streaking: Inlet Surface (Images Cropped for Comparison). ........ 84

FIGURE 4.10 Surface-Streaking: Rear Surface (Images Cropped for Comparison). ...... 85
FIGURE 4.11 Surface-Streaking: Rear Surface (Images Cropped for Comparison). ...... 85

FIGURE 4.12 Hot-Wire: Time Averaged Mean and rms Velocity Statistical Convergence

for Location g1. ........................................................................................... 86
FIGURE 4.13 Hot-wire: Normalized Velocity Histograms for Measurement Locations (a)

f1 (b) f5 and (c) f10. .................................................................................... 87
FIGURE 4.14 Hot-Wire: Normalized Velocity Distributions. ......................................... 88
FIGURE 4.15 Hot-Wire: Mesh Positions for Interpolated Values. .................................. 89
FIGURE 4.16 Hot-Wire: Normalized Mean Velocity Contour: 600 rpm. ...................... 90
FIGURE 4.17 Hot-Wire: Normalized Mean Velocity Contour: 3000 rpm. .................... 90
FIGURE 4.18 Hot-Wire: Normalized rms Velocity Contour: 600 rpm. ......................... 91

FIGURE 4.19 Hot-Wire: Normalized rms Velocity Contour: 3000 rpm. ....................... 91

FIGURE 4.20 Hot-Wire: Normalized Velocity Power Density Spectra for Measurement
Locations (a) f1 (b) f5 and (c) f10. Arrows in Figures a and b Indicate the

Spectral Peaks at 600 rpm. .......................................................................... 92
FIGURE 4.21 Hot-Wire: Normalized "Broadband" Energy Contour: 600 rpm. ............. 93
FIGURE 4.22 Hot-Wire: Normalized "Broadband" Energy Contour: 3000 rpm. ........... 93
FIGURE 4.23 Hot-Wire: Normalized "bpt" Energy Contour: 600 rpm. ......................... 94
FIGURE 4.24 Hot-Wire: Normalized "bpt" Energy Contour: 3000 rpm. ....................... 94

FIGURE 4.25 Hot-wire: Sample (a) Time Series, (b) TTL Signal, and (0) Phase Averaged
Quantities. Data are Taken From Measurement Location g1, 3000 rpm. .95

FIGURE 4.26 Hot-Wire: Phase Averaged Mean and ms Velocity Statistical Convergence
for Location g1 at or = 0° (TTL Signal Rise). ............................................. 96

FIGURE 4.27 Hot-Wire: Normalized Phase Averaged Mean Velocities for Entry Port f:
600 rpm. ...................................................................................................... 97

FIGURE 4.28 Hot-Wire: Normalized Phase Averaged Mean Velocities for Entry Port f:
3000 rpm. .................................................................................................... 97

FIGURE 4.29 Hot-Wire: Contour of Normalized Phase Averaged Mean Velocities for En-
try Port f: 600 rpm. (5f: r - r2 + 2) .......................................................... 98

FIGURE 4.30 Hot-Wire: Contour of Normalized Phase Averaged Mean Velocities for En-
try Port f: 3000 rpm. (5f: r - r2 + 2) ........................................................ 98

FIGURE 4.31 Hot-Wire: Contour of Normalized Phase Averaged rms Velocities for Entry
Port f: 600 rpm. (5f: r - r2 + 2) ................................................................ 99

FIGURE 4.32 Hot-Wire: Contour of Normalized Phase Averaged rrns Velocities for Entry
Port f: 3000 rpm. (5f: r - r2 + 2) .............................................................. 99

FIGURE 4.33 Hot-Wire: ms of Phase Averaged Velocity Normalized by Tip Velocity:
600 rpm. .................................................................................................... 100

FIGURE 4.34 Hot-Wire: rrns of Phase Averaged Velocity Normalized by Tip Velocity:
3000 rpm. .................................................................................................. 100

FIGURE 4.35 Static Pressure: Representative Static Pressure Statistical Convergence on
Mean for Entry Port c. ............................................................................... 101

xi

FIGURE 4.36 Static Pressure: Normalized Static Pressure Distributions, Lines are Plotted
to Show Connectivity. ............................................................................... 101

FIGURE 4.37 Microphone: Sample Fluctuating Pressure Statistical Convergence on rms
for Entry Port (1. ......................................................................................... 102

FIGURE 4.38 Microphone: Normalized Fluctuating Pressure Distributions, Lines are Plot-
ted to Show Connectivity. ......................................................................... 102

FIGURE 4.39 Microphone: Normalized Pressure Histograms for Entry Ports a, d, and h. .
.................................................................................................................... 103

FIGURE 4.40 Microphone: Normalized Pressure Power Spectrums, Entry Port a. ...... 104
FIGURE 4.41 Microphone: Normalized Pressure Power Spectrums, Entry Port (1. ...... 105
FIGURE 4.42 Microphone: Normalized Pressure Power Spectrums, Entry Port h. ...... 106

FIGURE 4.43 Microphone: Normalized Pressure Auto-Correlations for Entry Ports a, d,
and h (R*(or)=0.5 Given at Bottom Right of Each Plot). .......................... 107

FIGURE 4.44 Microphone: Normalized Pressure Cross-Correlations. Progression of Peak
Correlation is Shown (Left to Right) from Entry Port 3 to b, a to c and a to d.
.................................................................................................................... 108

FIGURE 4.45 Microphone: Normalized Pressure Cross-Correlations. Progression of Peak
Correlation is Shown (Left to Right) from Entry Port c to d, c to e and c to f.
.................................................................................................................... 109

FIGURE 4.46 Microphone: Normalized Pressure Cross-Correlations. Progression of Peak
Correlation is shown (Left to Right) from Entry Port e to f, e to g and e to h.
.................................................................................................................... 110

FIGURE 4.47 Microphone: Normalized Convection Velocity. Values are Based on the
Lag Time of the peak in the Cross-Correlations and the Separation Distance
Between Entry Ports a-b, b-c, c-d, d-e, e—f, f-g, and g-h. Lines are Plotted to
Show Connectivity. ................................................................................... 1 1 1

FIGURE 4.48 PIV: Time Averaged Mean and ms Velocity Statistical Convergence for
Region 11 (z = 43.3 mm, r = 103.9 mm) at Q = 600 rpm (97 Percent Valida-
fion) .......................................................................................................... 112

FIGURE 4.49 PIV: Normalized Velocity Magnitude Contour with Streamlines: Inlet Plane
(x-y) 600 rpm (See Figure 3.15 for Plane Definitions). ............................ 113

FIGURE 4.50 PIV: Normalized Velocity Magnitude Contour with Streamlines: Mid Plane

xii

(x-y) 600 rpm. ........................................................................................... 113

FIGURE 4.51 PIV: Normalized Velocity Magnitude Contour with Streamlines: Rear Plane
(x—y) 600 rpm. ........................................................................................... 114

FIGURE 4.52 PIV: Normalized Velocity Magnitude Contour with Streamlines: Region I
(z-r) 600 rpm (See Figure 3.15 for Region Definitions). .......................... 115

FIGURE 4.53 PIV: Normalized Velocity Magnitude Contour with Streamlines: Region I
(z-r) 1200 rpm. .......................................................................................... 115

FIGURE 4.54 PIV: Normalized Velocity Magnitude Contour with Streamlines: Region II
(z-r) 600 rpm. ............................................................................................ 116

FIGURE 4.55 PIV: Normalized Velocity Magnitude Contour with Streamlines: Region II
(z-r) 1200 rpm. .......................................................................................... 116

FIGURE 4.56 PIV: Normalized Velocity Magnitude Contour with Streamlines: Region III
(z-r) 600 rpm. ............................................................................................ 117

FIGURE 4.57 PIV: Normalized Velocity Magnitude Contour with Streamlines: Region III
(z-r) 1200 rpm. .......................................................................................... 117

FIGURE 4.58 PIV: Normalized Phase Averaged Velocity Magnitude Contour with
Streamlines: Inlet Plane (x-y) 600 rpm. .................................................... 118

FIGURE 4.59 PIV: Normalized Phase Averaged Velocity Magnitude Contour with
Streamlines: Mid Plane (x-y) 600 rpm. ..................................................... 118

FIGURE 4.60 PIV: Normalized Phase Averaged Velocity Magnitude Contour with
Streamlines: Rear Plane (x-y) 600 rpm. .................................................... 119

FIGURE 5.1 PIV: Sample Radial Velocity Contour with one isotach at 0 m/s Showing the
Distribution of Positive and Negative Values at the Impeller Outlet. ....... 129

FIGURE 5.2 PIV: Zoomed View of Region III z-r Plane. The Streamlines Reveal a Sepa-

ration of the Secondary Flow and a Region of Recirculation. .................. 133
FIGURE 5.3 Sample Phase Averaged Quantities from Figure 4.25c. ............................ 136
FIGURE 5.4 Hot Wire: Ratio of rms to Mean Velocities: 600rpm. ............................. 140
FIGURE 5.5 Hot Wire: Ratio of rms to Mean Velocities: 3000rpm. ........................... 141

FIGURE A.1 Installed Condition Evaporator Load Curves (Temperature Setting Lo-Lo).
.................................................................................................................... 147

xiii

FIGURE A.2 Installed Condition Evaporator Load Curves (Temperature Setting Lo-Hi). .
.................................................................................................................... 147

FIGURE A.3 Installed Condition Evaporator Load Curves (Temperature Setting Hi-Hi). .

.................................................................................................................... 148
FIGURE A.4 Installed Condition Evaporator Load Curves (Temperature Setting Hi-Lo). .

.................................................................................................................... 148
FIGURE 8.] Normalized ux rms Velocity Contour: Inlet Plane 600 rpm. .................... 149
FIGURE B.2 Normalized ux rms Velocity Contour: Mid Plane 600 rpm. .................... 149
FIGURE B.3 Normalized ux rms Velocity Contour: Rear Plane 600 rpm. .................... 150
FIGURE 3.4 Normalized uy rms Velocity Contour: Inlet Plane 600 rpm. .................... 150
FIGURE B.5 Normalized uy rms Velocity Contour: Mid Plane 600 rpm. .................... 151
FIGURE B.6 Normalized uy rrns Velocity Contour: Rear Plane 600 rpm. .................... 151
FIGURE B.7 Normalized Reynolds Shear Stress Contour: Inlet Plane 600 rpm. .......... 152
FIGURE 8.8 Normalized Reynolds Shear Stress Contour: Mid Plane 600 rpm. .......... 152
FIGURE B.9 Normalized Reynolds Shear Stress Contour: Rear Plane 600 rpm. .......... 153
FIGURE B.10 Normalized uz rms Velocity Contour: Region I 600 rpm. ..................... 153
FIGURE B.11 Normalized uz rms Velocity Contour: Region II 600 rpm. .................... 154
FIGURE B.12 Normalized uz rms Velocity Contour: Region III 600 rpm. ................... 154
FIGURE B. l 3 Normalized ur rms Velocity Contour: Region I 600 rpm. ...................... 155
FIGURE B. 14 Normalized ur rms Velocity Contour: Region II 600 rpm. .................... 155
FIGURE B.15 Normalized ur rms Velocity Contour: Region III 600 rpm. ................... 156
FIGURE B.16 Normalized Reynolds Shear Stress Contour: Region I 600 rpm. ........... 156
FIGURE B.17 Normalized Reynolds Shear Stress Contour: Region II 600 rpm. .......... 157
FIGURE B.18 Normalized Reynolds Shear Stress Contour: Region III 600 rpm. ........ 157
FIGURE B.19 Normalized uz rrns Velocity Contourzz Region I 1200 rpm. .................. 158

xiv

FIGURE B.20 Normalized uz rrns Velocity Contour: Region II 1200 rpm. .................. 158

FIGURE B.21 Normalized uz rrns Velocity Contour: Region III 1200 rpm. ................. 159
FIGURE B.22 Normalized ur rrns Velocity Contour: Region I 1200 rpm. .................... 159
FIGURE B.23 Normalized ur rms Velocity Contour: Region II 1200 rpm. .................. 160
FIGURE B.24 Normalized ur rrns Velocity Contour: Region III 1200 rpm. ................. 160
FIGURE B.25 Normalized Reynolds Shear Stress Contour: Region I 1200 rpm. ......... 161
FIGURE B.26 Normalized Reynolds Shear Stress Contour: Region II 1200 rpm. ........ 161
FIGURE B.27 Nomralized Reynolds Shear Stress Contour: Region III 1200 rpm. ...... 162

FIGURE B.28 Normalized Phase Averaged ux rrns Velocity Contour: Inlet Plane 600 rpm.
.................................................................................................................... 162

FIGURE B.29 Normalized Phase Averaged ux rms Velocity Contour: Mid Plane 600 rpm.
.................................................................................................................... 163

FIGURE B.30 Normalized Phase Averaged ux rrns Velocity Contour: Rear Plane 600 rpm.
.................................................................................................................... 163

FIGURE B.31 Normalized Phase Averaged uy rms Velocity Contour: Inlet Plane 600 rpm.
.................................................................................................................... 164

FIGURE B.32 Normalized Phase Averaged uy nns Velocity Contour: Mid Plane 600 rpm.
.................................................................................................................... 164

FIGURE B.33 Normalized Phase Averaged uy nns Velocity Contour: Rear Plane 600 rpm.
.................................................................................................................... 165

FIGURE B.34 Normalized Phase Averaged Reynolds Shear Stress Contour: Inlet Plane
600 rpm. .................................................................................................... 165

FIGURE B.35 Normalized Phase Averaged Reynolds Shear Stress Contour: Mid Plane 600
rpm. ........................................................................................................... 166

FIGURE B.36 Normalized Phase Averaged Reynolds Shear Stress Contour: Rear Plane
600 rpm. .................................................................................................... 166

XV

LIST OF TABLES

TABLE 1.1 Blower Nomenclature and Dimensions. ....................................................... 16
TABLE 4.1 Performance: Experimental Operating Point Values .................................... 66
TABLE 4.2 Performance: Normalized Experimental Operating Point Values ................ 67

TABLE 5.1 PIV: Impeller Active Width, i.e. Percentage of Impeller Outlet Width (b2) With
a Positive Radial Velocity Profile (ur(H < z < H + b2, r2) > O). ................. 129

TABLE A.1 Operating Conditions of the Instrument Panel (Installed Condition) at the
Highest Blower Control Setting. ................................................................. 146

xvi

emu

A,B,n

game

2 Z

PThrottle Box

NOMENCLATURE

Hot-wire calibration coefficients

Scroll Area

Blade Chord Length

Hot-wire anemometer output voltage

Frequency

Tip clearance

Hub to impeller distance

Mach number

Number of blades, Number of samples in a measured time record

Number of complete impeller revolutions for a given time series

Atmospheric Pressure

Microphone calibration pressure
Nozzle exit pressure of the hot-wire calibration facility

Supply plenum pressure

Microphone calibration reference pressure

Mean gauge pressure on the scrolled surface of the volute

Time resolved fluctuating pressure on the scrolled surface of the volute

rrns of the time resolved fluctuating pressure

Gage Pressure of the throttle box

xvii

RCC

R(9)

UR
Uconvection
Ur

U

’0’

UZI’
Ui
U1

U2

Total Pressure

Reynolds number based on blade cord length

Scroll Shape

Cutoff Clearance, Scale factor (image to measurement) PIV

Volumetric flow rate

Isentropic stagnation temperature

Magnitude of the x-y velocity vector measured by the hot-wire probe

Resampled hot-wire time series
Convection velocity

Impeller tip speed

Magnitude of the x—y velocity vector measured by PIV
Magnitude of the z-r velocity vector measured by PIV
Spatially averaged flow velocity at the volute inlet
Spatially averaged flow velocity at the impeller inlet
Spatially averaged flow velocity at the impeller outlet

Hot-wire calibration velocity

Microphone circuit output voltage rms level

Impeller width

Volute width
Impeller Inlet Width

Impeller Outlet Width

xviii

po

1'2

ux,u uz,ur

y,
x,y,z
z,r, 0

GREEK

(1

BI

92

Particle displacement as measured by PIV
Cutoff diameter

Volute inlet diameter

Impeller inlet diameter

Impeller outlet diameter

Blower motor voltage
Blower motor current

Microphone sensitivity

Hot-wire calibration facility calibration constant
Mass flow rate

Isentropic stagnation pressure

Input power to the blower motor

Impeller outlet radius

Time, blade thickness

Components of the velocity vector (subscript specifies coordinate)
Cartesian coordinates

Cylindrical coordinates

Degree of impeller rotation

Blade inlet angle

Blade outlet angle

xix

ophase ms of the phase averaged mean hot-wire distributions
AP Pressure-rise

APB Pressure-rise across blower

APH Pressure-rise across HVAC unit

APE Evaporator pressure differential

APem Pressure Differential across elbow meter radii

At Laser pulse separation time (image pair separation time) PIV
Q Impeller rotational speed [rpm]

p Density of air

v Kinematic viscosity of air

11 Efficiency percent

(1) Flow coefficient

w Pressure coefficient

(r) Vorticity

W

For a quantity g:

g Mean Value

g nns value

g2 Variance (Reynolds stress tensor nomral component)

XX

 

g xg y Reynolds stress tensor shear component
g* Normalized value
g' Instantaneous deviation from mean value
<g> Phase averaged value
Eg(F) One-sided power spectral density function
8182 Cross-correlation function (auto-correlation gl=g3)
JED BB Broadband contribution to the normalized velocity fluctuation intensity
i130 bpf Narrowband contribution to the normalized velocity fluctuation intensity
w
B Blower
H HVAC
inlet Volute inlet
5 Scrolled surface
1 Impeller inlet
2 Impeller outlet
AQBQEXME
A/D Analog to digital
bpf Blade passing frequency
DC Direct current
HVAC Heating/Ventilation and Air Conditioning
IP Instrument Panel

xxi

LED

PIV

SPL

Light emitting diode
Particle Image Velocimetry
Sound pressure level

root mean square
revolutions per minute

Turbulent Shear Flows Laboratory, Michigan State University

xxii

1.0 Introduction

1.1 Automotive HVAC Systems and Centrifugal Blowers

The major components of an automotive HVAC (heating/ventilation and air condi-
tioning) system are the centrifugal blower, evaporator, heater core, and the associated
ductwork. The blower, evaporator, and heater core are typically integrated into a single
unit which is designed to meet packaging constraints and a number of performance speci-
fications. These performance specifications include flow rate, pressure-rise, temperature,
and humidity control and sound emission. Next to the air quality, the sound quality within
the automobile cabin plays a very important role in passenger comfort. Progress in the
overall sound quality of the vehicle has emphasized the need to lower the sound emission
from the HVAC system. The noise emitted from the HVAC system, as pointed out by J ef-
frey et a1. [14], is the single most dominating contributor of interior noise for several oper-
ating conditions; the worst condition being low engine speed and high HVAC impeller
speed. It is well known that the major noise source within the HVAC system is the cen-
trifugal blower [2].

Centrifiigal blowers are typically categorized according to blade type. These blade
types inelude forward curved (FC), backward curved (BC), radial blade (RB), and others
that lie between these examples, see Figure 1.1. The type of blower used depends on the
performance requirements for a particular application. FC centrifugal blowers are typi-
cally used in all types of HVAC systems since FC blowers deliver considerably more vol-
ume flow and also produce higher static pressures than other fans of the same size and

speed. These benefits are, however, at the expense of lower efficiencies [4]. Specifically,

F C blowers can run at about half the speed to produce comparable ranges of volume flow
and static pressure [4]. Generally, there are also higher turbulence and noise levels associ-
ated with PC blowers for a given impeller speed and diameter [28]. However, Eck [7]
pointed out that there is no other fan or blower which operates as quietly at comparable
pressures. These performance characteristics make the F C centrifugal blowers especially
desirable for automotive application given their tight packaging constraints and their rela—
tively large system pressure losses.

The large volume flow and pressure rise associated with a FC blower stems from

the curvature of the blades in relation the rotation of the impeller. The forward curvature

of the blades provides an additional centrifugal forceI to the fluid residing in the blade
gaps. As shown by Eck [7], an idealized vector diagram at the tip of an FC blade can
result in exit velocities on the order of two times the tip velocity.

A schematic of a simple, single inlet, rectangularly shaped FC centrifugal blower
is shown in Figure 1.2. These type of blowers are also commonly known as "Sirocco run-
ners" or '"squirrel cage blowers". The blower has two components; the impeller and the
volute. The basic (i.e. simplified) operation of the blower is as follows. Flow enters the
impeller through the volute inlet in an axial direction and then turns radially toward the
impeller’blades. The impeller then imparts kinetic energy to the flow and is discharged
into the volute at all points around the impeller. Once in the volute, the flow primarily

takes on a tangential direction. The majority of the flow is then directed to the blower out-

let. However, at the cutoffz, a portion of the flow is recirculated through the volute.

 

1. An "effective force" as seen by a non-inertial reference frame observer.

2. The portion of the volute closest to the impeller.

The volute’s scrolled shape is based on the streamlines of an inviscid isolated
source vortex [7]. Specifically, a streamline is replaced with a solid surface that confines
the impeller after one full revolution. The result is a housing with an increasing area to
account for the additional mass being passed through the volute as the point of observation
moves counterclockwise around the blower, see Figure 1.2. A properly designed volute
will have a uniform pressure distribution at the perimeter of the impeller which in turn will
give a uniform velocity distribution at the parimeter of the impeller. An important func-
tion of the volute is the conversion of kinetic energy (fluid Velocity) to flow work (pres-
sure/density). This is especially true for the FC blower given the high velocity magnitudes
at the impeller exit. In fact, F C blowers only function properly in a volute housing, which
is unlike BC blade impellers which can be readily used for radial discharge, as in plug fans

or roof ventilators [4].

1.2 Previous Investigations

Many researchers have studied the flow and noise characteristics of centrifugal
blowers. Most of the studies in this area have concentrated on the near impeller and the
cutoff regions. Other work has described the gross flow behaviors at the impeller inlet and
within the volute. However, there does not appear to be "exhaustive" investigations of the
flow field in the domain between the impeller and volute.

The "jet-wake" flow pattern, in the discharge of the impeller, has received consid-
erable attention; see, e.g. [7,24,26,31]. Gradients in the velocity profile with respect to
circumferential direction and for radii close to the impeller, exhibit sharp minima and
maxima due to blade wakes. This jet-wake profile is shown schematically in Figure 1.3.

The figure demonstrates the high velocity "jet" flow near the blade pressure surface, while

the low velocity "wake" flow is near the blade suction surface. In an extensive hot-wire
study at the exit of the impeller, Raj [26] found that the jet-wake profile is a function of
both axial and tangential location. His results also conclude that the profile is improved
(i.e. larger jet zone, see Figure 1.3) if the flow rate is increased, however, the "active
impeller width" (discussed below) decreases with flow rate.

Researchers, e. g. [7,26, 27], have also pointed out the inefficient use of the impel-
ler width. Specifically, the presence of separated flow at the inlet of the impeller results in
an "active impeller width" which is less than the full width of the impeller, see Figure 1.4.
Raj [26] found that the active width is a function of tangential location and flow rate. A
widely accepted solution to this flow behavior is to reduce the ratio of the impeller width
to the impeller diameter, see Bleier [4] and Eck [7]. However, Vadari et a1. [27] found that
significant reductions of the emitted noise and the size of the separation region could be
obtained by adding an annular inlet guide.

A secondary flow occurs within the volute of blowers which is known to be a
source of inefficiencies. This secondary flow can take the form of either a "double vortex"
or a "single vortex", see Figure 1.5. The secondary flow is superimposed on the primary
flow direction which is in the "desired" (i.e. tangential) flow direction. The kinetic energy
of the secondary motion does not, therefore, contribute to the mass flux and its subsequent
dissipation contributes to the inefficiencies of the system. Many authors (e.g.
[7,9,10,1.8,26]) have reported on these secondary flow structures, however, no quantifica-
tion of the strength of the secondary flow relative the primary flow have been identified in

the author’s search of the literature.

The aerodynamic noise of blowers is formally divided into two components; tonal
and broadband. The tonal noise has been extensively researched and the source is well
understood, which makes it generally easy to remedy without a significant loss in perfor-
mance. The source of the broadband noise, however, is not well understood. Attempts to
reduce the broadband noise have either reduced the performance to unacceptable levels or
are not cost effective for the reduction in noise obtained.

The tonal noise derives from the impulse given to the air each time a blade passes
a given point. Hence, the tonal noise is formed as a series of discrete tones at the funda-
mental blade passing frequency (bpf) and its harmonics [28]. In the case of centrifugal
blowers, the impulse given to the air stems from the "j et" (discussed above) impinging on
the cutoff at a repetition rate of the bpf thus producing tonal noise at the bpf and its har-
monics. Many papers on the subject of blower noise reduction were reviewed in the liter-
ature survey by Niese [24]. His survey concluded that the shape and position of the cutoff
are the two main design features with respect to blower tonal noise. Two of the sugges-
tions provided were: i) distancing the cutoff farther from the impeller and ii) either slop-
ing the cutoff or the impeller blades. By distancing the cutoff further from the impeller,
the extent to which the cutoff experiences the maxima and minima of the jet-wake flow is
reduced. By sloping either the cutoff or the impeller blades, the impulse of air from the
impeller blades is no longer evenly distributed across the cutoff and the tonal noise is sig-
nificantly reduced. Sloping of the cutoff is much more cost effective than sloping of the
impeller. By sloping the cutoff, Humbad et al. [13] successfully reduced the bpf noise by

roughly 9 dBA.

The broadband noise is most commonly believed to be largely due to vortices
shedding from the impeller blades and from blade surface pressure fluctuations caused by
flow separation on the suction surface. Neise [24] reported the most successful attempt in
reducing the broadband noise was accomplished by mounting meshes along both the inner
and outer circumference of the impeller. The meshes were meant to shifi the point of sep-
aration toward the trailing edge of the suction surface and to smooth the outlet velocity
profile. However the decrease in efficiency was not cost effective for the reduction in
noise. Neise concluded by stating:

"The most important task of future research work is to reduce the random compo-
nent of centrifugal noise without losing fan efficiency. The first step in this direction
would be to locate the main regions of broadband noise. There are some hints that the

rotating impeller is not the dominant source, but that a significant contribution comes from
the fan casing: some reduction in broadband noise was gained by modifying the casing

while flow through the impeller remained unchanged; see references [17] and [19]3."

1.3 The Research Program

1.3.1 Motivation and Background

The current research subject was brought (in 1999) to the attention of the author by

Mr. C. Mesa of the DaimlerChrysler Corporation. The TSFL4 was asked to study the
HVAC system that was in service in the 1999 Jeep Grand Cherokee (WJ model). The
problem proposed was one for which the HVAC system produced maximal noise with lit-
tle increase in airflow at the highest blower speed.

An Instrument panel (IP) from the 1999 Jeep Grand Cherokee, see Figure 1.6, was

supplied by DaimlerChrysler in order to establish the appropriate flow variables of the

 

3. References [17] and [19] of the present document were not used by the present author. They are provided
as a reference source for Neise’s [24] conclusion.

4. Turbulent Shear flows Laboratory, Michigan State University.

HVAC system. The main component of the HVAC system is the fullyiintegrated DENSO
HVAC unit assembly, also shown in Figure 1.6. Specifically the HVAC unit integrates the
blower, evaporator, and heater core within a single housing. The HVAC unit installed
within the IP will be termed "installed" while the uninstalled HVAC unit will be termed
"uninstalled".

The climate controls for the HVAC system consist of a series of rotary knobs, see
Figure 1.7. These controls can be set to obtain desired interior conditions for both the
driver and passenger side. That is, the HVAC system has a dual-zone climate control in
which the driver and passenger can maintain separate climate conditions. In Figure 1.7
the controls, from left to right, are the blower control, the temperature controls, and the
mode control. The blower control has ten impeller speed settings from "L0" to "H1". The
temperature controls are also set from "L0" to "Hi"; the left temperature control is used by
the driver and the right control is used by the passenger. The mode control has five set-
tings which set the outlet paths of the air flow. Referring to Figure 1.6, these mode control
settings are:

0 Panel: Air flows through the Panel Registers.

- Bi-Level: Air flows through the Panel and Floor Registers.

' Floor: Air flows through the Floor Registers.

° Defrost/Floor: Air flows through the Defrost and Floor Registers.

° Defrost: Air flows through the Defrost Registers.

The HVAC system also has the option to draw air from outside of the vehicle,
"fresh-air mode", or draw air from the vehicle cabin, "recirculation mode". The recircula-
tion mode is typically used to rapidly cool, or heat, the inside of the vehicle. This option is

achieved through an HVAC unit inlet attachment, see Figure 1.8. The attachment has two

inlets, one of which is blocked by the "air switch door" while the other is in use.

The internal components of the HVAC unit are shown in Figure 1.9. Air is drawn
in through the inlet by the impeller and collected in the volute. From the volute the air
passes through the evaporator which cools the air when the air conditioning is in opera-
tion. The flow is split into two paths: one to the driver and one to the passenger, down-
stream of the evaporator. Air switch doors immediately downstream of the evaporator,
can be proportionally opened such that air can pass through the heater core if a higher tem-
perature is desired. The exhaust end of the unit then sets the path for the given mode con-
trol (i.e. panel, bi-level, etc.).

A power supply was used to apply 13.8 volts at the IP fuse buss in order to operate
the installed HVAC system. This arrangement allowed for the blower motor and the air
switch door actuators to be fully operational. The system was operated in isothermal con-
ditions, i.e. operational temperature refrigerant and coolant was not pumped through the
heat exchangers. Given the open laboratory environment in which the IP was operating,
the configuration is one that would closest match an idle (stationary) vehicle with the win-

dows fully open.

1.3.2 Experimental Overview

The focus of this study is on the blower section of the uninstalled HVAC unit.
Specifically, the flow field domain between the impeller and volute surface was examined
in detail to identify sources of inefficiencies and noise generation. Three factors played a
role in deciding on this focus area of the HVAC system; these are described as:

1. Packaging Constraints: The HVAC system‘s flow paths rely heavily on pack-

aging constraints which change from vehicle to vehicle. However, these con-
straints have a significantly lower effect on the blower since the basic

components (impeller and volute) remain relatively unchanged from vehicle to
vehicle.

2. Noise Generation: The blower was the major source of noise in the system.
The observed noise did not have the tonal characteristic of the blade passing
frequency, but it did have a very loud "rumble" that is characteristic of low fre-
quency broadband noise.
3. Flow Source: The blower is the source of flow generation in the HVAC sys-
tem. Hence, the overall performance of the system relies heavily on the blower.
A detailed schematic and dimensions of the blower under study are given in
Figure 1.10 and Table 1.1. Centrifugal fans are often described by three important geo-

metric parameters, these are given as; number of blades (N=41 FC blades), impeller out-

side diameter (d2=150 mm), and cutoff clearance (S=10.5 mm). The scroll shape is also
an important parameter that is often described by the area (A) enclosed by the impeller

and the volute surfaces as a function of azimuthal location (0) from the cutoff. This area

is given as:

Av(0) = bv - (R(0)—d2), (1.1)

where bV is the volute width and R(0) is the volute radius as a function of azimuthal loca-

tion (0) from the cutoff, see Figure 1.10 and Table 1.1. This area function is plotted along

with a linear curve fit in Figure 1.11.

1.3.3 Experimental Prelude

The following chapters present the experimental equipment and program for the
numerous measurement techniques used in this study. These chapters are followed by the
experimental results which are presented on a technique by technique basis for organiza-
tional purposes only. The Discussion will then integrate the results and discuss the salient
flow features. The conclusions, presented in the final chapter, will summarize the major

results of this study.

FC BC RB

FIGURE 1.1 Centrifugal Impeller Blade Types; FC (Forward Curved Blade), BC
(Backward Curved Blade), RB (Radial Blade).

 
   

  

 

 
  
  

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 
  

  

 

 

 

ll Volute\
Tangential A l 1 Radial Flow
Flow
Axial Flow
#_ lmpelle
Outlet 3
'—>

FIGURE 1.2 Schematic of a Typical Centrifugal Blower With FC Blades.

10

Jet

Wake

\
Wake Suction
Zone ' Surface
{Ct . Pressure —""'
Lone Surface
K

FIGURE 1.3 "Jet-Wake" Velocity Pattern (Stationary Observer) at the Exit of a FC
Impeller [26].

Active :~———-—— b2
Impeller \
Width w\

\ \fifl‘éfi

E I ___- Inlet

fl?“ Area=£§i

 

 

 

 

 

 

I'll] \9
l][/ b \9

Impeller
Blades

 

 

Separated Flow

FIGURE 1.4 Schematic of the Separated Flow at the Impeller Inlet [27]. (See
Figure 1.10 for definition of di)

11

 

 

 

 

 

 

 

.1: ”7’23““... pr???“
"\1 7'7:-\§ a; ‘i‘x+$,
1‘? \\\\*’«‘:‘\\ ‘ ‘ln’m'
‘ l \h") K *. ‘1 ‘
[ Volute :I‘FCSK'K" + x
k‘l—j \4£\ ’<'-’\:K{ ‘3' $ '
Ir \ =4 ._ '0‘,‘

 

W-

(a) Impeller Outlet (b . ‘

FIGURE 1.5 Computed Velocity Fields of Arbitrary Radial Cross Sections From
'IWo Different Blowers Showing (a) the "Double Vortex” [9] and (b) the "Single
Vortex" [18]. The Secondary Flow is Super-Imposed on to the Primary Tangential
Flow. Note: The Inlet flow direction is From Right to Left

Defrost Registers

3"“
1“-

Climate
Controls

). HVAC Unit
Exhaust

"" Uninstalled

 

FIGURE 1.6 Instrument Panel and HVAC Unit.

Temperature “
Controls

 

FIGURE 1.7 HVAC Climate Controls: Mode Control; (a) Panel, (b) Bi-Level,
(c) Floor, (d) Defrost/Floor, (e) Defrost.

" Air Switch. 1 v
Door ‘ Flow
Direction

l

 

FIGURE 1.8 HVAC Unit Inlet Attachment.

 

 

 

 

FIGURE 1.9 HVAC Unit Features.

     

:ECEEm

59:20 wEEum

 

   

74ch SE:
5:: 5:995

 
     

r”””>>>”\r”>”’)’)’

 

‘1‘“{4V‘4V‘4V4V‘4V‘AV‘A‘AV‘AV‘AV‘

EDESC wczzom

Cut
.2 ow: Q

 

 

 

U

 

 

 

 

Batsm
5:;

 

    

 

oumtzm
vuzobm

 
 

FIGURE 1.10 Blower Schematic and Nomenclature, see Table 1.1 for Definitions.

15

TABLE 1.1 Blower Nomenclature and Dimensions.

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Impeller Width 75.0 mm

Volute Width V 74.5 mm

Impeller Inlet Width b] 60.5 mm

Impeller Outlet Width b2 65 mm

Blade Chord Length 18.0 mm

Cutoff Diameter dCO 18.5 mm

Volute Inlet Diameter dl- 129.5 mm

Impeller Inlet Diameter (1] 119.0 mm

Impeller Outlet Diameter d2 150.0 mm

Tip Clearance G 3.5 mm

Hub to Blade Distance H 4.0 mm

Number of Blades N 41 blades

Sm“ Shapea Rm) 4x10‘7e3 — 3><10-592 + 0.15949 + 85.202
(0° 5 0 3 290°) (units: mm)

Cutoff Clearance S 10.5 mm

Blade Thickness t 1.25 mm

Blade Inlet Angle l3: 80°

Blade Outlet Angle [32 150°

Impeller Rotational Speed Q varies (units: rpm)

Cartesian Coordinates x, y, 2 mm, mm, mm

Cylindrical Coordinates r, 9, 2 mm, deg, mm

 

 

 

a. Microsofi Excell Trendline Fit (Coefficient of Determination (R2 = 0.997) )

l6

 

 

 

 

50 ...____._..._......fl _ . . - W 1
Volute Area Distributionl 0'

 

 
  
 

40 Av (0)=0.1370+6.185
R2 = 0.995

 

 

 

 

Av (0) [cm2]
8

 

 

0 50 100 ISO 200 250

0 (degrees)

FIGURE 1.11 Volute Area as a Function of Azimuthal Location and Linear Curve
Fit.

l7

2.0 Experimental Equipment

This aim of this chapter is to describe the experimental apparatus and equipment
used in this body of work. The laboratory configuration for the uninstalled HVAC unit
will be given first followed by the main experimental equipment. A detailed description

of how the equipment was employed will be given in Chapter 3.0.

2.1 Flow System

An experimental prototype of the HVAC unit described in Section 1.3.1 was pro-
vided by DaimlerChrysler in order to gain optical accessibility to the flow field within the
volute. The experimental prototype was a production HVAC unit with the volute housing
replaced by optically clear surfaces, see Figure 2.1.

The inlet and hub surfaces of the volute were made of flat 6 mm thick Plexiglas
and permanently attached to the body of the HVAC unit. The hub surface was painted
flat-black to provide a contrasting background for observations made within the volute
and to reduce glare when PIV measurements were employed. The inlet surface was also
equipped with a bellmouth cut from a production unit.

The scrolled surface was made of 3 mm thick Plexiglas formed to the shape of the
production scroll. This scrolled surface began at and included the cutoff and extended to
the exit of the volute. The flexibility of the scrolled surface allowed for its accurate posi-
tioning in relation to the impeller. This flexibility could also be used to explore other
scroll shapes (which was outside the scope of this work). A 1 mm thick rubber seal was

attached to the edges of the scrolled surface to form an airtight seal between the scrolled

18

surface and the inlet and hub surfaces. The scrolled surface was held in place by five 4" 1/
4-20 bolts which pulled the inlet and hub faces together.

A second scrolled surface was fabricated from RenShape #440. RenShape is a
syntactic polyurethane material which is ideal for prototyping due to its ease of machin-
ing. This second scrolled surface was needed during the pressure and hot-wire measure-
ments which required entry ports through the surface of the scroll. The thickness of the
RenShape scroll was 8 mm, installation and all other dimensions were identical to that of
the Plexiglas scroll.

To simulate the pressure drop associated with the duct-work and registers the
experimental unit was mounted on a throttle box, see Figure 2.2. The box had one outlet
but was otherwise completely sealed and it encased all flow outlets of the experimental
unit. For a given rpm the back-pressure could then be adjusted by positioning the throttle
plate on the side of the box to achieve the value desired for operation. The static pressure
rise across the HVAC unit was monitored by a static pressure tap on the top surface of the
box.

The impeller was driven by a DENSO 12.0 volt DC motor. This is the standard
motor for the production unit. The motor was powered by a Hewlett-Packard 20 volt
adjustable voltage DC power supply (model #6264B) in the present experiment.

The impeller rpm was monitored by an OMRON EE-SFS reflective optical sensor.
The basic components of this sensor are emitter and detector elements, see Figure 2.3.
The emitter element emits infrared light from an LED. When a reflective object passes
over the sensor the light is reflected back to the detector element (a phototransistor). The

output of this device is a 0-5 volt TTL signal that was acquired by the A/D board and mon-

19

itored by a Fluke 87 digital multimeter. The reflective object was a thin piece of sheet-
metal placed on the impeller hub between two consecutive impeller blades. The sensor
was mounted to the hub surface of the volute housing at the same radius as that of the
reflective object. This arrangement provided one 5 volt pulse per revolution. The time

between each pulse was then used to calculate the rpm and keep track of blade position.

2.2 Data Acquisition and Processing Systems

With the exception of PIV and sonic nozzle measurements all data were acquired
by an Iotech WaveBook/516 A/D board. This is a 16 bit differential A/D board. The
board was equipped with a sample and hold card able to record up to eight channels simul-

taneously at a total sampling rate of 1 MHz. The board recorded in bipolar mode with a
maximum range of i10 volts and could be set for gain factors of x1, x2, x5, x10, x20,

x50, x100. The typical noise level of the board was :2 bits. All data were recorded with

a Dell Latitude CPi Pentium 11 laptop computer.
All primary data were transferred to a Dell OptiPlex Gle Pentium III computer
for processing. Processing of these data was executed using "in-house" MatLab programs.

Post-processing programs include MatLab, Tecplot and Excel].
2.3 Sonic Nozzle Test Stand

Flow rate measurements were accomplished using a Testec model 11359 sonic

nozzle test stand, see Figure 2.4. Known mass flow rates (m) between approximately 3.6

kg/hr and 750 kg/hr could be produced by means of opening and closing a set of nine noz-

zles located in the sonic nozzle chamber. When a sufficient pressure differential exists

20

across these nozzles a sonic (M=1) or choked flow condition occurs at the nozzle throat.

This condition produces precisely known mass flow rates for known upstream p0, To val-

ues. Any combination of the nine nozzles could be used for a possible 29 independent
flow rate measurements in increments of approximately 3.6 kg/hr. This results in a maxi-
mum flow rate of 1843 kg/hr [20]. The facility however was not able to produce the low
pressures needed downstream of the nozzles to achieve flow rates above about 750 kg/hr.
The only acquisition means available for the flow rate readings were to take them by eye
from the display screen. This method proved to be adequate since the flow rate readings

fluctuated in the third decimal place.
2.4 Pressure Measurements

2.4.1 Pressure Transducers

Static pressure measurements were employed to monitor operating conditions and
obtain pressure distributions of the HVAC unit and for the calibration of hot-wires. This
required the use of several differential pressure transducers. Validyne transducers were
used for the monitoring of operating conditions. The transducers, models DP15-20 and
DP15-22 equipped with CD15 carrier demodulators, produce a ilO volts DC analog out-
put proportional to the full scale pressure range. Validyne reports a full scale accuracy of
i025 % for the DPlS series. MKS Baratron transducers were used to obtain pressure dis-

tributions and for the calibration of hot-wires. The large range of pressures encountered
required the use of 1, 1.0 and 100 torr MKS transducers. The model series 398HD-

00001SP05 with a type 270 signal conditioner were used in this study. Baratron reports an

accuracy of i005 % of reading for these units.

21

2.4.2 Microphones

Eight flush mounted microphones were used to measure the fluctuating pressure
on the scrolled surface of the volute. Specifically, Panasonic omnidirectional electret con-
denser microphone cartridges, model WM-62CC were used in this investigation. These
microphones are 6 mm in diameter and 2.7 mm thick with a sensing diameter of 2 mm.
Panasonic reports a sensitivity of -45 324 dB (0 dB = l volt/Pa) and a nominally flat fre-
quency response over the range of 20-16,000 Hz, see Figure 2.5. A custom built circuit
was used to operate the microphones, a schematic of this circuit is shown in Figure 2.6. A
Tripp Lite line conditioner, model LC-1200, was used to suppress line noise to the circuit

power supply.
The microphones were calibrated with a Class 1 Larson Davis CAL200 sound

level calibrator. The CAL200 provides a choice of calibration sound pressure levels

(SPL) of94.0 and 114.0 J_r0.2 dB (re: 20 uPa) at a frequency ofl kHz i1 %.

2.5 Velocity Measurements
2.5.1 Hot-Wire Anemometry

2.5.1.1 Hot-Wires, Anemometers and Therrnistor

Single sensor hot-wire probes were employed to make point wise flow velocity
measurements within the volute housing. These probes were fabricated in the TSFL at
Michigan State University. A schematic representation of a single sensor hot wire probe
is shown in Figure 2.7. A probe consists of a 3 mm length of 5 um diameter tungsten

wire soldered at each end to stainless steel broaches. The ends of the wire are copper

22

plated to 30 um leaving an active sensing region of 1 mm at the center of the wire. The
broaches are mounted to the probe body and act as the electrical connection to the BNC
cables.

DISA type 55M10 constant-temperature anemometers were used to control the
hot-wires. The electrical noise of these anemometers were typically 2 mV RMS [22]. The
anemometers were tuned to give frequency responses above 35 kHz at a flow speed of 80
m/s.

A thermistor was used to compensate for any temperature fluctuations occurring

from pre to post-calibrations of the hot-wires. The thermistor had a sensitivity of 2.03 K/
Kohm, an accuracy of i0.2 K, and a frequency response of 10 Hz at 293 K [5]. The labo-

ratory and experimental facility provided nearly isothermal conditions and any tempera-

ture fluctuations are on the order of hours so the response time was adequate.

2.5.1.2 Calibration Facility

A special hot-wire calibration facility, shown in Figure 2.8, was constructed in
order to reproduce the relatively high flow velocities encountered within the volute hous-
ing. The facility is capable of producing calibration flow velocities from approximately
0.5 to 90 m/s. This was accomplished by means of a Spencer Turbo-Compressor, Lot NO.
37215, and a throttle plate. The facility is also capable of holding a probe at angles rang-
ing from i36° in increments of 6° and at i90°. The design of the facility permitted all
connections of the probes to remain in contact throughout the experiments.

A flow conditioner equipped with filter material was placed upstream of the cali-

bration nozzle to provide a low disturbance calibration flow. The use of this flow condi-

23

tioner introduced friction to the incoming flow, therefore use of the Bernoulli equation

directly between Palm and PCI: (calibration facility’s static pressure taken at a tap located
near the exit of the nozzle) was not valid. To remedy this the total pressure (PT) at the exit
of the nozzle, referenced to PCF, was used to find a calibration constant (m) for the desired

measurement of (Patm - PCF). A "modified" Bernoulli equation:

v = [gmdmm— FUJI "2 (2.1)

could then be used to calculate the nozzle exit velocity (V). The calibration information is

given in Figure 2.9.

2.5.2 Particle Image Velocimetry (PIV)

A Dantec two component planar PIV system was used to further investigate the

flow field within the volute housing. This was a fully integrated Dantec PIV system‘.
The major components of this system are the camera. lasers, processor, PC, and seed gen-
erator.

The camera used was a Kodak Megaplus ES 1.0. This is a greyscale 8-bit digital
camera capable of 15 double-frames/s. The camera contains a 1K x 1K pixel Class II
CCD (charge couple device) chip and was equipped with a 532 nm band-pass camera fil-
ter.

A New Wave MiniLase III, 15 Hz 40-50 mJ double cavity Nd:YAG laser equipped

with light-sheet optics was used in this investigation. The output was a 532 nm pulsed

light-sheet with a pulse width of 5-7 ns and a variable pulse time delay (At).

 

l. A subcomponent of the Multi-Phase Flow Facility (MF F) ofthe College of Engineering at MSU.

24

The Dantec F lowMap 2000 processor was the main "work-horse" of the system.
The processor housed the synchronization controls and a correlator unit which provided
correlated image maps in near real-time.

The entire system was controlled by F lowManager software interfaced with a Dell
Dimension XPS Pentium II PC. The software was used to set up experimental conditions
which include, component timing, image scaling and image correlation options. Post-pro-
cessing options for the resulting vector maps include validation schemes, filtering
schemes and statistical analysis. This work however only utilized FlowManager’s post-
processing for validation purposes, all other post-processing was accomplished using Mat-
Lab programs.

A Rosco 1600 fog machine seeded the flow field of interest. The fog machine had
a variable output and uses Rosco fog fluid which is a nontoxic glycol, de—ionized water
mixture. This made it safe to work in the vicinity of the experiments as they were con-

ducted. The output particle size is reported to be 0.25 - 60 microns.

25

Inlet,

' 1Hub

   
 

f; '-_ I} Scrolled
1'. f i ‘. Surface

 

FIGURE 2.1 Experimental Prototype.

 

FIGURE 2.2 Experimental Flow System.

26

Vl/l/l/I/I/I/A Reflective

/\
\F

 

 

 

 

 

LED ' i ’ ,Phototransistor

FIGURE 2.3 Reflective Optical Sensor Schematic.

l l. I ‘
Display -"
Screen

Chamber

 

FIGURE 2.4 Sonic Nozzle Test Stand.

27

Relative Response (dB)

 

+20

 

+10

 

 

O
t

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

20 .50 100 200 500 100'.) 2000 5000 10000 20000
Freque nc y oi Hz 2'

FIGURE 2.5 Typical Microphone Frequency Response Curve (Panasonic
Specifications Sheet).

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

2 +2.6V
+1 V
o—‘LM317 / IOOKQ
—/\/VV‘
$820!) $31“) +12V
0.1uF
gr 2- 7 BNC
7f LF411 6
IKQ 1K9 3Vl4/
<g> ?
Microphone 1K9

FIGURE 2.6 Microphone Op Amp Circuit Schematic.

28

   
  

Q30 um
Copper Plating 35 um
Tungsten Wire

\
JlmL

‘—-3mm

 
 
   
 

 

 

  

 

Broaches

 

 

 

 

/
Probe Body

 

 

O\/

FIGURE 2.7 Schematic of a Single Sensor Hot-Wire Probe.

 

 

 

FIGURE 2.8 Hot-Wire Calibration Facility.

29

120 “’ *’

 

  

(PT — PCP) (Pa)

 
 

 

 

100 '
80 ' ’
60 i ” ]
(PT ' PCP) : 1-0024 (Palm ' PCF)
40 2 T‘
R ==l
20 ’ ” A’x‘ ' 4*
L l
0 , , -- z z _ a
O 20 4O 60 80 100 120

(Patm ‘ PCP) (Pa)

FIGURE 2.9 Hot-Wire Calibration Facility’s Pressure Measurement Correction.

30

3.0 Experimental Program

This chapter describes the experimental configurations and procedures used to
obtain the various HVAC flow field measurements. The first section describes the meth-
ods used to obtain the appropriate operating characteristics (i.e. rpm, pressure and flow
rate) of the installed condition. The section that follows describes the methods used to
obtain characteristic curves of the uninstalled unit while the remaining section describes
the methods used to obtain detailed flow measurements within the volute of the unin-
stalled unit. The uninstalled unit was configured to mimic the downstream flow restric-
tions of the installed condition. The detailed experimental methods include surface-
streaking, hot-wire anemometry, static pressure, microphone, and particle image velocim-
etry (PIV). Details on major components of the experimental equipment used in this study
were given in Chapter 2.0.

Difficulties were encountered in obtaining PIV measurements (discussed in
Section 3.3.5.1) at impeller speeds higher than Q = 600 rpm. However, measurements at
multiple impeller speeds using other techniques were readily acquired. In predicting the
effect of changing the operating conditions of fans (or blowers) it is important to include
the independent parameters of the flow field given by the "fan laws". For a given geome-
try these parameters include fluid density and viscosity, fan diameter and fan rotational
speed. Given that rotational speed is the only variable in this study the governing scaling

parameter is the impeller tip speed given by:

ndZQ
Ut = 60 . (3.1)

 

31

Hence, the flow field’s Reynolds number dependence could be inferred from the full set of

data by sealing the results on Ut. That is, if the flow field were to show a weak depen-
dence on Reynolds number, then the normalized PIV results could be applied to the higher
impeller speeds with reasonable reliability. The Reynolds number for this study is defined

as:

ReC = C82, (3.2)

 

where C is blade cord length and U2 is the spatially averaged flow velocity at the impeller

outlet. U2 is given by:

 

Cl
:‘l

.90
NO“

2 = , (3.3)

where Q is volumetric flow rate and d2 and b2 are the impeller outlet diameter and width,

respectively. These observations will be used in chapters 4.0 and 5.0 to aid in the display

and discussion of the results.

3.1 Instrument Panel Measurements

The instrument panel (IP) described in Section 1.3.1 was used to gain an under-
standing of the HVAC unit as it operated in the installed condition. Specifically, it was of
interest to obtain rpm, pressure and flow rate measurements for each of the climate control
settings. By knowing these operating characteristics, the uninstalled unit could be

arranged to closely reproduce them. For these measurements the HVAC system was oper-

32

ated in the "fresh-air mode" only. This was due to the lack of access to the inlet used for

the "recirculation mode" which was tightly positioned behind the glove box.

3.1 .1 Procedures and Features

The experimental configuration of the IP flow rate measurement system is shown
in Figures 3.1 and 3.2. The measurement system consisted of four major components in
conjunction with the IP; an auxiliary blower, an air delivery system, an "elbow meter", and
a supply plenum. The basic operation of the system is as follows. The air register louvers
on the IP were first fully opened and the HVAC system was turned on and set to a given
climate control setting (see Section 1.3.1 for setting configurations). The auxiliary blower
then delivered a metered flow rate to the supply plenum. By adjusting the delivered flow

to a condition of Pplcnum=Patm, the flow rate is that which would exist in the absence of the

supply plenum. Finally, along with the flow rate, pressure and rpm data of the HVAC unit
were recorded to establish the appropriate flow variables that describe the system’s operat-
ing characteristics.

A Cincinnati blower, model #PB-14A, provided the supply air. The blower was
equipped with a throttle plate which was used to adjust the amount of air entering the sys-
tem. From the blower the flow was ducted through the air delivery system made from 3"
PVC piping. The air delivery system entered the supply plenum from the side and was
directed downward. This arrangement allowed the exhaust to impinge on the bottom sur-
face of the plenum to facilitate removal of excess kinetic energy from the entering air

stream.

33

The supply plenum encased the passenger-side portion of the cowl box and was
completely sealed. The cowl box is the first element of a typical HVAC system which is
intended to separate and drain rainwater from the incoming air stream [1]. Upstream of
the cowl box the supply plenum was equipped with flow conditioners to ensure a uniform
flow distribution. The flow conditioners included filter material, honeycomb, and screens.

The supply plenum pressure, as previously stated, was pressurized to maintain Patm. This

pressure was initially monitored at three locations (see Figure 3.2) to confirm an equal
pressure distribution and subsequently at one location (the rear tap) when this condition
was confirmed. A Validyne pressure transducer and a Fluke digital multimeter were used

to monitor these measurements.

The pressure differential across the evaporator (APE ), located just downstream of

the volute, was recorded at each of the climate control settings. The reason for these mea-
surements was to have a quantity that could be used to reproduce flow restrictions in the
uninstalled configuration. The techniques used to acquire the data and the locations of the
pressuremeasurements were selected in an effort to keep the IP and HVAC unit intact
(that is, without having it disassembled). A static pressure tap was used in the upstream
location since it was possible to assure a flush mount fit with the inner surface of the
HVAC unit which could be reproduced in the uninstalled unit. However, on the down-
stream side of the evaporator there was no access to the inside surface to assure a flush
mount of a static tap. To accommodate this situation a special "pressure probe" was fabri-
cated which could be inserted into the side of the HVAC unit. Figure 3.3 shows the con-
figuration for these pressure measurements. The pressure probe essentially became a

cylinder‘in a cross-flow. The advantage of having the pressure tap located on the down-

34

stream side of the probe is that it lies in the aft region where the pressure remains nearly
constant [30]. This characteristic is advantageous for purposes of reproducing the condi-
tions in the uninstalled unit. That is, with this configuration, small deviations in probe
positioning would not affect the results.

Volumetric flow rate (Q) measurements were accomplished with an "elbow meter"
located in the air delivery system just upstream of the supply plenum, see Figure 3.2.

Static pressure taps located at the inner and outer radii provided the means to obtain a

pressure differential (APcm) from which the flow rate could be inferred. Calibration of

the elbow meter was executed using the sonic nozzle test stand described in Section 2.3.
The exhaust end of the entire PVC air delivery system was mounted to the test stand inlet.
This configuration was used to ensure identical flow conditions at the elbow meter to that
of the measurement configuration. Additional care was taken by adding flow conditioners
(honeycomb and screens) to the straight sections upstream of the elbow meter. Calibra-
tion pressure measurements were acquired with a Validyne pressure transducer sampled at
300 Hz for 10 seconds while flow rates were taken by eye from the display screen as
described in Section 2.3. The calibration data, shown in Figure 3.4, was fit with a fourth

order polynomial:

Q = — 0.523AP:m + 3.071 Mg”, — 6.221 AP2 + 7.876AP

C [Tl C in

+ 0.538, (3.4)

which was used to infer the flow rates of the HVAC system.
The impeller speed was recorded at the highest blower control setting only, with

the reflective optical sensor described in Section 2.1. The acquisition rate was based on

acquiring two samples per blade gap (41 blade gaps) at a typical impeller speed (Q) of

35

about 3200 rpm, resulting in a sampling rate of 4400 Hz. The rpm and pressure data were
acquired simultaneously for 10 seconds.

Although the data acquired in this section were an integral part in obtaining the
appropriate flow variables necessary to mimic the installed conditions in the uninstalled
configuration, only one operating configuration was used in the remaining study. For this
reason the results for the other operating conditions are given in APPENDIX A.

The operating configuration selected for the remaining study was determined
based on audible noise and flow rate at the highest blower speed. It was found that one
climate control setting produced the highest levels of both conditions. The configuration

of this setting was: Bi-level mode with both temperature controls set to L0 (i.e. Lo—Lo con-
figuration)‘. This resulted in: Q = 3074 rpm, APE = 90.84 Pa, and Q = 6.274 m3/min.

The operating pressure and blower speed found above were only used to set the
downstream flow restriction for the detailed studies (described in Section 3.3). This was

accomplished by adjusting the throttle plate of the experimental flow system (see

Figure 2.2) until APE = 90.84 Pa was attained at 3074 rpm. Once the downstream flow

restriction was set, measurements could be carried out at any blower speed desired.

3.2 Experimental Flow System Performance Measurements (Uninstalled Unit)

In an effort to gain a better understanding of how the uninstalled HVAC unit per-

forms in the Bi-level / Lo-Lo configuration, characteristic curves were obtained at four

impellerspeeds: Q = 600, 1200, 2400 and 3000 rpm. More specifically, the performance

 

I. The current study was performed in isothermal conditions. i.e. the temperature controls were used to set
flow paths not temperature levels.

36

characteristics of pressure-rise (AP), blower motor input power ( [to ) and volumetric flow

rate (Q) were obtained and plotted as (AP vs. Q) and ( go vs. Q). These characteristic
curves provided the means to obtain the operating points at which the HVAC unit oper-

ated when the exhaust end of the unit was placed in the mocked downstream environ-

mentz. This was advantageous since detailed measurements were carried out at three of

the four impeller speeds specified above.

3.2.1 Procedures and Features
The experimental configuration for the performance measurements is shown in
Figure 3.5. The experimental flow system was mounted to the sonic nozzle test stand with

the exhaust3 of the throttle box ducted to the inlet of the test stand by a 6" PVC elbow.
This arrangement allowed for HVAC unit to operate against a back pressure at a known

flow rate for a given impeller speed.

The steps used to set the conditions for each measurement are as follows. First the
maximum mass flow rate (mmax) for a given impeller speed was determined by increas—
ing the flow rate until PThrome BonPatm- This is the condition that would exist if the
HVAC unit were exhausting to the atmosphere and thus the lowest restriction to which it

would be exposed. Once nan... was found, measurements began starting at shut-off

(m = 0) and continued in increments of nominally 516mm,, until mm, was attained. At

 

2. The mocked downstream environment refers to the simulation of the installed downstream flow restric-
tion in the uninstalled configuration.

3. The throttle plate of Figure 2.2 was not used for these measurements since the operating pressures were
predetermined by flow rate, set by the sonic condition of the test stand. and impeller speed.

37

each increment the blower motor input voltage was adjusted to maintain the given impel-
ler speed. This procedure was then repeated for each Q condition.

Two pressure measurements relative to Patm were acquired for this experiment.

These were the pressure-rise across the blower (APB) and the pressure-rise across the
HVAC unit (APH ). The static tap for the APB measurement was located along the volute

mid-plane at an angle of 0 = 315° from center of the volute cutoff, while the static tap
for the APH measurement was located at the top of the throttle box, see Figure 3.6. Vali-

dyne pressure transducers were utilized for these measurements.
Input power ( g) ) to the blower motor was obtained from measurements of motor

voltage (e) and current (i) and then calculated by the following summation:

N
1 .
X9 = E}: e, .11., (3.5)
121 ‘

where N is the number of samples. The input voltage was measured directly across the
blower motor while the current was measured in series with the power circuit using a
shunt DC current meter. The current meter produced an output voltage signal directly pro-
portional to current and had a calibration constant of 5 amps / volt. The high voltage
requirement (>10 volts) of the 3000 rpm impeller speed condition required the use of a 2:1
voltage divider.

Data were acquired at 360 samples/revolution. Acquisition rates were therefore

determined by impeller speed. These rates are 3600, 7200, 14400 and 18000 Hz for the

600, 1200, 2400 and 3000 rpm impeller speeds respectively. All measurements (APB ,

38

APH , e, i, and rpm) were acquired simultaneously for 10 seconds. Details on the flow rate

and impeller speed measurements were given in Chapter 2.0.

3.2.1.1 Inlet Study

Although all detailed measurements (described in the next section) in this study
were conducted with a "free inlet" condition, second and third performance tests were exe—
cuted at Q = 3000 rpm with the HVAC unit inlet attachment. These measurements were
performed in an effort to gain insight on the effects of varying inlet conditions. Tests were
conducted with the inlet attachment set to both fresh-air mode and recirculation mode, see

Figure 1.8.

3.3 Volute Flow Field Measurement Techniques

This section describes the methods used to obtain detailed flow measurements
within the volute of the uninstalled HVAC unit. Depending on the measurement tech-
nique used, data were acquired at impeller speeds of; Q = 600, 1200, 2400 and 3000 rpm.
All data 'were taken with the unit set to the Bi-level / Lo-Lo configuration determined in

Section 3.1 and performed using the experimental flow system shown in Figure 2.2.

3.3.1 Surface-Streaking

Surface-streaking is a technique used to resolve the time mean, near-surface
streamline pattern. This technique, although qualitative, can reveal (in a relatively short
period of time) important salient features of the flow that can otherwise go unnoticed

using other "more sophisticated" quantitative techniques.

39

Surface-streaking results were obtained using a mixture of kerosene and a fluores-

cent pigment (AX-13-5 Rocket Red)4. The impeller was removed and a thin coat of this
mixture was applied to the inside surface of the volute. The scrolled surface of the volute
was first covered with a mylar sheet before application of the mixture. The purpose of the
mylar sheet was to have a removable surface that could be laid flat for ease of viewing.
Once the mixture was applied, the impeller was installed and set to the desired rpm. As
the mixture dried, the shear of the fluid acting upon the surface interface conformed the
mixture to the near-surface, or limiting, streamline patterns.

Once the mixture had dried, the HVAC unit was removed from the throttle box and
disassembled to reveal the streamline patterns on the hub and inlet surfaces. The mylar
sheet was then removed from the scrolled surface and coated with a clear protective spray.

Results were fluoresced under a "black" (UV) lamp and recorded with a digital camera.

The surface-streaking experiments were performed at two operating speeds: Q = 1200

and 3000 rpm.

3.3.2 Hot-Wire Anemometry

Point-wise time resolved flow velocity data were acquired within the volute using
single sensor hot-wire anemometry. The hot-wire experiments were performed at impel-
ler speeds of Q = 600 and 3000 rpm. This section provides details on calibration, mea-
surement locations, and data acquisition. Details on the sensors, anemometers, thermistor
and calibration facility was given in Section 2.5.1. The reader is referred to reference [6]

for further reading on hot-wire technique and theory.

 

4. A Product of Better World Manufacturing, Inc., Fresno. California.

40

3.3.2.1 Calibration and Processing

Preliminary hot-wire experiments revealed that a large range of flow velocities
(from about 1 to 60 m/s) occur within the volute over the impeller speeds of Q = 600 and
3000 rpm. In order to calibrate the hot-wires over this velocity range a three step calibra-
tion had to be implemented. Calibration velocities, calculated with Equation 2.1, required
the direct measurement of pressure. It is these pressure measurements that dictated the
three step approach. The three steps of this calibration are termed low, medium, and high
range. Pressure measurements of the low, medium, and high ranges implemented the use
of l, 10, and 100 torr pressure transducers, respectively. Although the 100 torr transducer
was capable of measuring the full range encountered, the lower range transducers were
employed for increased accuracy at the lower pressures.

Calibrations of the hot-wires were performed in the calibration facility shown in
Figure 2.8. The calibration employed a "quasi-steady" method used by Morris [23] and
was accomplished by slowly opening the throttle plate over the desired pressure range for
a period of 60 seconds. This produced a continuously variable flow speed that was calcu-
lated by the "modified" Bernoulli equation given in Equation 2.1. A basic assumption of
this equation is steady-state flow. Experiments by Morris [23] have shown that steady-
state calibrations show "no measurable differences" to that of the quasi-steady calibra-
tions. Hot-wire, thermistor and pressure calibration data were acquired at a sampling rate
of 300 Hz.

Calibrations were performed for the low (1-15 m/s), medium (10-45 m/s), and high
(35-85 m/s) ranges. Sample calibration data of anemometer output voltage (E) vs. veloc-

ity (V) can be seen in Figure 3.7a. The overlap of these ranges were cropped at 95% of

41

the anemometer output voltage range. For example, if the low range anemometer output
had a range of 1.5-5.5 volts, the signal would be cropped at 5.3 volts (given by 1.5+(5.5-
1.5)*0.95) and the medium range would begin just above 5.3 volts. Once cropped the data

from each calibration range were fit to the power law

132 = A + 13vn (3.6)

using a least-squares linear fit. The exponent (n) was cycled through 0.1 to 0.6 in incre-
ments of 0.0125 while fitting the calibration data to Equation 3.6. The values of A, B, and

n which produced the smallest standard deviation of (Vealculated-Vmeamed) provided the

calibration constants used for velocity field measurements. Typical standard deviation
values were 0.05, 0.17, and 0.32 m/s for the low, medium and high ranges, respectively.
These values provide an indication of the uncertainty in the measured velocity values.
Specifically, the maximum uncertainty in the mean values is typically within 0.5% and
5% at the upper and lower end of each calibration range, respectively. Sample curve fits
for the three calibration ranges are given in Figure 3.7b.

The sensors were calibrated before (pre-calibration) and after (post-calibration)
each experiment. Velocity data calculated from the pre-calibration and post-calibration
always agreed to better than 1%. The calibration constants used for the measurements

were obtained by combining the pre and post calibrations.

3.3.2.2 Measurement Locations and Acquisition

The enclosed volume of the volute housing required the hot-wire probes to
traverse through the scrolled surface as seen in Figure 3.8. A traverse and special probe

bodies and guiding sleeves were fabricated to accomplish this. The traverse, shown in

42

Figure 3.9, was equipped with a dial that was capable of positioning the probe in incre-
ments of 0.01 mm. The probe and traverse employed a key and slot system to assure
proper alignment of the probe as it traversed. The key was a small diameter cylinder that
ran the full length of the probe which fit snugly into slots machined into the traverse and
guiding sleeve.

The basic operation of the traversing system is shown schematically in
Figure 3.10. The traverse was mounted to the scrolled surface and the guiding sleeve was
positioned to be flush with the inside surface of the volute. Measurements were divided
into two regions; "near-impeller" and "near-surface". For the near-impeller measurements
(see Figure 3.10a) the guiding sleeve allowed the probe body to traverse fully through the
scrolled surface to reach within 2 mm of the impeller. However, the guiding sleeve for the
near-surface measurements (see Figure 3.10b) only allowed for the hot-wire probe
broaches to traverse through the scrolled surface. By allowing only the broaches to pass
through, near—surface measurements could take place within 2 mm of the scroll without
flow disturbances from cavities (void surfaces), as would be the case with a fully retracted
probe using the near-impeller guiding sleeve of Figures 3.9 and 3.10.

Probe access was gained by eight azimuthally located entry ports (a—h) along the
mid-plane of the scrolled surface, see Figure 3.11. Measurements were acquired at ten
"semi-radial" equally spaced points for each entry port for a total of 80 measurement loca-
tions. These measurements spanned 2 mm from the impeller to 2 mm from the scrolled
surface. The semi-radial positioning is a direct result of the traverse direction being nor-
mal to the scrolled surface. That is, the normal direction does not pass through the center

of the impeller due to the spiral shape of the scroll. Measurement locations will be identi-

43

fied by the entry port letter (a-h) followed by a numeral (1-10). For instance the measure-
ment location nearest the impeller at entry port (f) will be termed fl and the measurement
location farthest from the impeller will be termed fl 0.

The hot-Wire’s single sensing element was oriented parallel to the impeller rota-
tional axis, as seen schematically in Figure 3.10. This orientation provided a means of
obtaining x-y component velocity magnitudes. That is, this technique could only provide

velocity magnitudes given by:

U = ./ui + uj, (3.7)

where the u)( and u), represent the x and y components of velocity, respectively. Given an

interest in obtaining the spectral content of the velocity fluctuations, data were sampled at

8 samples per blade gap for 1,000 revolutions of the impeller. This results in a sampling
rate of 16,400 Hz for 20 seconds and 3,280 Hz for 100 seconds at Q = 600 and 3000

rpm, respectively. Hot-wire, thermistor, and rpm data were recorded simultaneously for

these experiments.

3.3.3 Static Pressure Measurements

Static pressure measurements were acquired in order to gain the mean gage surface
pressure (PS) distribution along the scrolled surface of the volute. Each static tap was

machined into a "plug" which also housed a microphone (discussed in the next section).
The plugs were inserted into entry ports (a) through (h) for a total of 8 measurement loca-

tions. Care was taken to ensure no flow disturbances by mounting each plug flush

44

(iOl mm) to the scrolled surface with the static tap upstream of the microphone, see
Figure 3.12.

Measurements were carried out at impeller speeds of Q = 600, 1200, 2400 and

3000 rpm. The measurements were acquired using the 1 torr MKS transducer for

Q = 600 and 1200 rpm and the 10 torr MKS transducer for Q = 2400 and 3000 rpm.

Data were sampled at 4 samples/revolution for 1000 revolutions of the impeller. These

sampling rates and times are 40, 80, 160, and 200 Hz, for 100, 50, 25, and 20 seconds at

impeller speeds of Q = 600 , 1200, 2400 and 3000 rpm, respectively.

3.3.4 Microphone Measurements

Time resolved fluctuating pressure (P's) measurements were acquired along the
scrolled surface of the volute using Panasonic omnidirectional electret condenser micro-
phones. These measurements were used to obtain both rrns (P's) and spectral pressure

quantities at impeller speeds of Q = 600 , 1200, 2400 and 3000 rpm.
Microphones were calibrated at both 94.0 and 114.0 dB SPL (sound pressure
level) at a frequency of 1000 Hz, see Figure 3.13 for calibration configuration. Each
microphone output voltage signal was sampled at 20 kHz for 10 seconds while being

exposed to the output SPL of the calibrator. The calibrator SPL is given as:

 

~

SPL = 2010g£ PC ], (3.8)
Pref

45

 

where SPL is the given calibrator sound pressure level, PC is the calibrator pressure signal

(given in Pa ms), and Pref is the reference pressure (20 pPa ). Solving for the calibrator
pressure signal gives:

S l’ L 20

13. = fircr- 10 , (3.9)

which was used to obtain the microphone sensitivity (km). Specifically, the microphone

sensitivity is given as:

k = _, (3.10)

where Vm is the microphone circuit output voltage rms level. Sensitivities measured

from the two calibrator SPLs, which agreed to better than 2% for each microphone, were
averaged together to obtain the individual microphone sensitivities.

The microphones were reported to have a nominally flat frequency response over

the range of 20- 1 6,000 Hz, see Figure 2.5. However, a direct measure of uncertainty over

this range could not be obtained given that the microphones were calibrated at only one

frequency. A measure of this uncertainty was given by Hudy [11] using similar micro-

phones and measurement techniques. Hudy reported an uncertainty of 7% of the mean

sensitivity. This uncertainty was computed by the 2*nns value associated with the scatter

'11 sensitivity over six calibration frequencies. Also reported was a time delay of 10 to
3 us over the 80 microphones used in the study. It is important for this delay to be sub-
.antially smaller than any convective time scale in the flow if cross-correlation or cross-

)ectra estimates are desired. Given that the lowest convective time scale of the current

46

 

study is 1,155ps, the microphone time delay is two orders of magnitude (z 100x)

smaller than the lowest convective time scale. The convective time of 1,155 us was cal-

culated by the average distance between two consecutive microphones (63.5 mm) divided

into the maximum velocity measured (hot-wire results at Q = 3000 rpm) in the volute

(55 m/s). It is also important to note the microphone drop-off in frequency response

below 20 Hz. The effect of the drop off can be inferred from the resolved spectral quan-

tities, which will be shown in Section 4.6 to have an insignificant effect on the measured
values.

After calibration, the microphones were inserted into the same plugs used for the

static pressure measurements and flush mounted at the 8 entry port locations, see

Figure 3.12. The microphone data were acquired simultaneously at 8 samples per blade

gap for 1,000 revolutions of the impeller. These sampling rates and times are 3,280,

6,580, 13,120, and 16,400 Hz, for 100, 50, 25, and 20 seconds at impeller speeds of

Q = 600 , 1200, 2400 and 3000 rpm, respectively.

3.3.5 Particle Image Velocimetry

Whole field velocity data were acquired within the volute using planar particle
image velocimetry (PIV)5. The majority of these measurements were carried out at an
impeller speed of Q = 600 rpm while some results were also obtained at an impeller speed

f Q = 12 00 rpm. The following section provides details on measurement configurations.

 

The resolved vectors from this technique are given as U : tix + fry.

47

 

 

Details of the PIV system used for this study was given in Section 2.5.2. The following
paragraphs give the basic principles of this measurement technique.

The basic idea behind PIV is to gain a set of instantaneous whole field velocity
measurements for a desired region. This is accomplished by first seeding the flow field
with tracer particles that will track the motion of the fluid, see Figure 3.14. The nominal
requirements of the seeding are that the particles are to be essentially neutrally buoyant

and homogeneously distributed within the flow. A thin slice of the flow field is then illu-
minated by a laser light-sheet defining the measurement plane. The illuminated particles
scatter the light and provide a signal that is detected by a camera placed at a right angle to

the measurement plane. This process happens twice (by quickly pulsing the light-sheet) at

a known time interval (At) resulting in two images of the particle positions at times (t)

and (t + At).
The two camera images are then processed to resolve the velocity vector map of

the flow field. The images are first subdivided into interrogation regions. The displace-
ments (3 ) of the particle groups from image 1 to image 2 of each interrogation region are

then measured using Fast Fourier Transform (FFT) cross-correlation techniques.6 The

velocity vector (U ) of each region can then be calculated by:
U = S— _, (3.11)

where S is the scale factor of the measurement to image area.

 

._ For the current study subpixel interpolation was used in this step to resolve particle displacement accu-
rate to 1/64 of the pixel pitch.

48

 

 

3.3.5.1 Measurement Configurations

Difficulties with seeding proved to be the limiting factor for acquiring data at
impeller speeds higher than Q = 1200 rpm. The fog particles used in this study would

collect on the Plexiglas surfaces of the volute soon after introduction to the flow. A maxi-

mum of about 100 image pairs at Q = 600 rpm could be acquired before these collected
particles impeded the optical clarity of the volute. This would not only obstruct the view
of the camera but also block portions of the light-sheet causing shadows in the measure-
ment plane. Therefore, the impeller had to be removed and the volute wiped clean before
further measurements could be made. This process was repeated until a total of 500 image

pairs were acquired. At higher impeller speeds the particles collected on the surfaces at
much higher rates. Although some effort was put towards acquiring data at Q = 1200

rpm, the majority of the data were acquired at Q = 600 rpm.

The PIV measurements were acquired in three regions of the volute. These
regions are termed Region I, II, and III as seen in Figure 3.15. Each region contained four

measurement planes, three x-y planes (Rear, Mid, and Inlet) and one z-r plane.
Two measurements were carried out at Q = 600 rpm for each x-y plane. The
first of these measurements imaged an area which extended from the impeller to the
scrolled surface and were used to gain time averaged quantities. The second measurement
imaged a smaller area at the impeller outlet and was phase locked to blade position to gain
ibase averaged quantities. Time averages were acquired to reveal the net flow behavior,
rhile phased quantities were acquired to reveal the blade-to—blade flow dynamics at the

npeller outlet. Measurements of the z-r planes were carried out at impeller speeds of

49

 

Q = 600 and 1200 rpm. Each of these planes were acquired to gain time averaged quan-
tities. A total of 500 vector maps were acquired for each measurement location. Sam-
pling rates were 2 Hz for the time averaged data and 10 Hz for the phase referenced data.

The phase referenced sampling rate was dictated by impeller speed (optical sensor trigger
rate) and maximum camera sampling rate. Image pair time separation (At) varied from

35 to 150 us depending on impeller speed, image magnification, and measurement plane.
The acquired images were processed using 32x32 pixel interrogation regions with
a 25% overlap resulting in a 41x42 grid vector map (or 1722 calculated vectors per image
pair). The size of the interrogation region (32x32 pixels), together with the measurement
to image area scaling factor, defines the spatial resolution for each vector calculation. The
overlap is used to, in a sense, "over-sample" the acquired images for increased vector map
resolution.
A number of factors can lead to the calculation of spurious vectors. These are
incorrect vectors resulting from, among other things, noise peaks in the correlation func-
tion, poor seeding, and regions of high velocity gradients. The "raw" vector maps must
therefore be subsequently validated. That is, the spurious vectors must be tagged as to not
be included in post-processing calculations.
A two step validation scheme was used in this study to tag the spurious vectors.
The first of these was a "peak-height validation". This validation method compares the
highest peak with the second highest peak of the cross-correlation function. A detectabil-
ty criterion of 1 .2 was used for this method as recommend by Kean et al. [16].
The second method used was a "moving-average validation". This validation

method uses an iterative approach which validates or rejects vectors based on a compari-

50

 

son between neighboring vectors. An implicit assumption of this method is continuity of
the flow field. The idea behind this method is that the calculated vectors have a correla-
tion with the surrounding vectors, i.e. the magnitude and direction of the surrounding vec-
tors do not vary significantly with the vector under scrutiny. Details on the validation
algorithms can be found in reference [8].

Figure 3.16 shows a representative "percent validated" contour map calculated
from a set of 500 validated vector maps. For a given interrogation region the percent vali-
dated refers to the percentage of vectors that passed the validation scheme relative to the
total number of "raw" vectors. A low validation percentage near the flow boundaries was
common for this study. This was due to glare of the incident light—sheet causing low sig-
nal to noise ratios. In areas of low validation care should be used when assessing the
results since there is a much lower probability of statistical convergence. The low valida-
tion of vectors near the surface of the volute, as seen in Figure 3.16, was a common fea-

ture for both the x-y and z-r measurements planes.

51

 

 

'slz'

    
   

Air Delivery]

System \

 

 

FIGURE 3.2 Rear View of the IP Flow Rate Measurement System.

52

 

 

Pressure Tap

Inserted 155 m From

 

l’

-0

  

4

 

fin

25

pstream
tatic Tap
(Rear Surface)

1

4

O

O
O

a
a

 

Bottom of Evaporator

O

Downstream

 

 

Pressure Probe

   

O o t
O O.

 

 
          
  

1
i

tmensmns ll]

- 62211311,,"2 + 7.876APem + 0.538

A a... ,

 

n
.m 9
,t
m m...
b 0
H... __
a 3 2
C Pm R
\r m A
. 0.
M _ 3
W +
4
m m
E P
H A
.3
.2
_5.
0

FIGURE 3.3 Schematic of the APE Measurement Configuration (D
(Q

 

 

 

2.25

0.75 1.25 1.5 1.75
Ape”, (in H20)

0.5

0.25

248.84 Pa).

FIGURE 3.4 Elbow Meter Calibration Curve and Fit. (Note: 1 in H20

53

 

FIGURE 3.5 Performance Measurement Experimental Configuration.

.,A.<,;

 
    

APB Measuremen . '. ‘3;
,1 .. Location 3"

 

FIGURE 3.6 Performance Static Pressure Measurement Locations.

54

 

i

\l

(a)

F"
Ur

.. ...--._.....-., .M .

O\

   
 

Medium Range

E (volts)

   
  

5 . ;
4.5 Low Range
4 _ L . -- .- . m.
0 10 20 30 40 50 60 70 80
V (m/s)
45; - . AA~ . A A , -- -
EZ— 704+ 1826 VO'22 l
40; (b) . - ' - , - / .
A "If High Range
35 -
132:8.13+7.34v‘~"33 ‘ :
N 30 , - .
”-1 Medium Range

2__ 0.41
20. E — 11.38+5.14v
15 , ___ . ~ » - AA ~ — ~
20 25 30 35 40
V11

FIGURE 3.7 Sample Hot-Wire Calibration (a) Curve and (b) Curve Fit.

55

Traverse

RenShape
Scrolled Surface

 

FIGURE 3.8 Hot-Wire Measurement Configuration.

Broaclres

 

FIGURE 3.9 Hot-Wire Probe and Traverse System for (a) Near-Impeller
Measurements and (b) Near-Surface Measurements.

56

 

 
  
  
 
  
  

Traversing
Clasp

Sleeve

Face \

Variable

_]_2 mm

   
     
   

   

Impeller

 

 

(a)

  
   
   
 

'I‘ruversing
(‘lusp

Sleeve

Impeller

 

 

(b)

FIGURE 3.10 Schematic of Hot-Wire Configurations for (a) Near-Impeller
Measurements and (b) Near-Surface Measurements.

 

 

 

 

FIGURE 3.1] Entry Port and Hot-Wire Measurement Locations. Entry Ports are
Located Along the Mid-Plane (z = 37.25 mm) at 0 = (a) 66° (b) 99° (c) 132° (d) 165°
(e) 193° (1) 220° (g) 248° and (h) 276°. Hot-Wire Measurement Locations are
Indicated by the Hash Marks.

 

Scrolled Surface 5 .

FIGURE 3.12 Static Pressure and Microphone Measurement Configuration.

58

 

FIGURE 3.13 Microphone Calibration Configuration (a) Before Insertion and
(b) During Calibration.

 

 

 

 

 

 

 

 

 

 

 

 

Seeded Flow
“Fiel d Vector Map
Laser Light- Camera / / / /
Sheet / / l l
Measurement
Plane Image 2
.;
Interrogation 3.,
Region :;,. j.
:1 .

Image 1

FIGURE 3.14 PIV Basic Principles. (Images obtained from Dantec’s website).

59

 

 

 

  

 

 

 

Region 11 Pm
- , ary Flow ': 5
Clear Plexrglas (e) ’ ”fa Direction Out l : : .

.' 2 Of Page E 1 .fi i

5 1 l 1

e: 9: 9

“I 21 E1

l 1

I l

l

.o
.-
.....
0"

 

 

 

 

 

 

FIGURE 3.15 PIV Measurement Planes. The Dashed Lines Represent the x-y
Planes while the Dotted Lines Represent the z-r Planes. The x-y Planes are located at
z = 17.25 mm, 37.25 mm, and 60.25 mm for the Rear, Mid, and Inlet Planes,
respectively. The z-r Planes are Located at 0 = 101°, 168°, and 252° for Regions 1, II,
and 111, respectively. Entry Port Locations are shown for Reference Purposes.

60

 

150

Region 11, Q = 600 rpm

 

140

130

120

Contour Map Boundary

Percent

 

llO

 

r (mm)
5
O

80

70

 

 

 

Validated

 

 

 

 

 

60

lllllllIIIIII'llIlllllrllllljllllllllIIIIIIIIIIII

 

 

 

llllllJLLlllLlLllJLllllllll‘lll

 

llll

 

50'1111

0

10 20 30 4O 50 60
z(mm)

7O

80 90

FIGURE 3.16 Representative Contour Plot of Percent Vector Validation. Data are
Cropped at Volute and Impeller Boundaries. The "L" Shaped Region of Low
Validation In the Upper Right Corner of the Volute is Due To "Burned Out" Pixels in

the Camera’s CCD Chip.

6]

4.0 Results

This chapter presents the uninstalled condition results of this study. For the pur-
poses of organization only, these results are presented on a technique by technique basis.
The first section provides data analysis methods used which were common to more than
one measurement technique. If applicable, further analysis methods are given in the sec-
tion of the individual measurement techniques. The results of the performance measure-
ments will be discussed in this chapter since they do not directly relate to the discussion of
the detailed measurements. A limited discussion of the detailed measurements is also pre-
sented in this chapter. The major discussion of the detailed measurement results is pre-

sented in Chapter 5.0.

4.1 Data Analysis

Statistics of the acquired time series data sets were used to evaluate the quantities
of interest. These statistics were used to gain information on both the mean and fluctuat-
ing characteristics of the flow field. For a given discrete time series g(t) the time mean

was defined as:

N
8=§Zam an
i=1

where N is the number of samples in the time series. The fluctuating component of the
time series was quantified with the standard deviation or root mean square (nns) value

defined by:

62

Idl—

N
~ 1 . — 2
= A -— . 4.2
g N_1X(g(‘) g) ( )
i :1
The cross-correlation function was used to determine the extent of similarity and

the time delay of a particular event between two locations. For two simultaneously sam-

pled time series (with mean values removed) the cross-correlation was calculated by [3]:

N-n
R8.8.(n) : 111—13", 2 g'im' g'2(i+")9 (4-3)

i=1

where n is the sample shift (-N S n s N). This function was normalized to give the cross-

correlation coefficient function defined by:

 

R T
R... = g“? (4.4)
182

If the two signals are identical, i.e. gl(t) = g3(t), the function is termed the auto-correla-
tion. The auto-correlation provided a measure of time-related properties of the flow at a
point, i.e. the length of a particular event or the reoccurrence of similar events.

The power spectral density function was used to resolve signal power as a function

of frequency. The one-sided auto-spectral density is defined by [3]:

AZHF
Eg(F) = ZJRglgl(r)e" ‘61, (4.5)

0

where F is frequency and t is the time lag in the auto-correlation. The discrete computa-

tions of the auto-spectral estimates were accomplished using Matlab’s pwelch function.

63

This function used an averaging record length of 2I 1 samples multiplied by a harming win-

dow with a 50% overlap of records giving a total of 320 averaged records. This results in
a relative variance errorl of 5.6%. The total area under the auto-spectral density function

over all the frequencies is the rms2 (variance) value of the time series given by:

(1"

g2 = [Eg(F)dF . (4.6)
O

4.2 Performance

The integral performance results for Q = 600, 1200, 2400, and 3000 rpm are pre-

sented in normalized form as a function of the flow coefficient ((1)), where (1) is defined as:

‘1’ = ——2-Q—— ~ (47)
(nd2/4)Ut

The performance results for the four impeller speeds are given in Figures 4.1 through 4.4.
Six dependant variables were used to characterize the performance of the uninstalled unit.
These variables include the static pressure coefficient, static efficiency, Reynolds number,
and three spatially averaged velocities. The pressure-rise performance results are given as

the static pressure coefficient defined by:

(4.8)

 

 

1. The relative variance error (%) is given by 1/ J17 - 100, where M is the total number of averaged records.

64

These values, shown in Figure 4.1, indicate a decrease in pressure-performance at lower
impeller speeds. The subscripts of "H" and "B" indicate the pressure-rise performance
results across the HVAC unit and the blower, respectively. Efficiency is a measure of out-
put to input, the output being static pressure-rise and flow rate and the input being supply

power to the blower motor. The static efficiency percentage was thus defined by:

— 100. (4.9)

_AP-Q_
‘1 60

where so (input power to motor) was defined in Equation 3.5. These results show a

decrease in efficiency with respect to 4) at lower impeller speeds, see Figure 4.2. Note,
the defined efficiency is not a direct measure of blower efficiency, rather, it is a measure of
motor\blower efficiency. Efficiency calculations typically use a direct measure of the
shaft torque x impeller speed as the reference input. However, the author was unable to
attain a direct measure of shaft torque. Hence, care should be used when assessing these
curves since they represent more than a direct result of flow dynamics.

Reynolds number (defined by Equation 3.2) and spatially averaged velocity results

are given in Figures 4.3 and 4.4, respectively. The spatially averaged velocities are

 

 

defined as:
U. = (2) , (4.10)
adj/4
“ _ Q
U1 - ndlblo (4'11)
and

65

—9 .1
U2 - 1212. (4 2)

where U1, U1 , and U2 are the spatially averaged velocities at the volute inlet, impeller

inlet and impeller outlet, respectively. Impeller tip speed (Ut) was used to normalized

these values and are presented in Figure 4.4. The excellent agreement of the curves over
the four impeller speeds tested show a confirmation of the "fan laws" (discussed in the
introduction of Chapter 3.0). These values, which are a function of the selected area (for
their definition), give an indication of the average acceleration (or deceleration) of the
flow from inlet to outlet. This observation will be brought into discussion in Chapter 5.0.
The operating points for the detailed studies are given in Table 4.1. Table 4.2

gives the operating Reynolds numbers and the normalized operating points.

TABLE 4.1 Performance: Experimental Operating Point Values

 

 

 

 

 

 

 

 

 

 

 

 

Q (rpm) 600 1200 2400 3000
Ut (m/s) 4.71 9.42 18.85 23.56
Q (m3/Hr) 56.81 131.10 278.86 360.52
APH (Pa) 5.13 24.91 112.66 175.95
APB (Pa) 16.06 70.54 285.86 463.36
p (watt) 3.81 14.55 79.98 155.74
Uinlet (m/s) 1.20 2.76 5.88 7.60
U1 (m/s) 0.70 1.61 3.42 4.43
U2 (m/s) 0.52 1.19 2.53 3.27

 

 

 

 

 

 

66

 

 

TABLE 4.2 Performance: Normalized Experimental Operating Point Values

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Q (rpm) 600 1200 2400 3000
Rec 618 1,427 3 ,03 5 3,923
(1) 0.189 0.219 0.233 0.241
WH 0.404 0.483 0.555 0.537
“’8 1.265 1.367 1.407 1.415
n“ 2.12 6.21 10.91 11.31
m. 6.64 17.65 27.69 29.79
6‘? 0.254 0.293 0.312 0.323
inlet
6"“ 0.148 0.171 0.182 0.188
6* 0.109 0.126 0.134 0.139
2

 

 

4.2.1 Inlet Study

Performance tests were conducted at 3000 rpm in order to gain insight on the
effects of altering the inlet conditions of the blower. The results of the inlet study are
given in Figures 4.5 and 4.6. These results show a decrease in performance in both pres-
sure-rise and efficiency when the inlet attachment is used. At a flow coefficient of 0.241
(the operating point used in the detailed studies) there is a decrease in performance rela-
tive to the "free inlet" of 19 and 33 percent for the "fresh air" and "recirculation" mode,
respectively. These results agree well with those of Wright, et al. [32]. In that study the
performance of a centrifugal fan for a variety of inlet flow distortions was investigated. In

their study, the inlet velocity profile was altered by adjustable vanes upstream of the inlet.

67

Their results showed a decrease in performance, anywhere from 5 to 50 percent, for all
conditions studied.

The results seem to indicate a dependence on the inlet flow direction. That is, the
direction from which flow is ducted toward the inlet (see Figure 1.8) may directly impact
the performance of the blower. These findings are not conclusive however, so further
experiments should be conducted in which the inlet geometry remains unchanged and is

rotated about the impeller axis for a number of performance tests.

4.3 Surface-Streaking

Surface-streaking tests were performed to resolve the time mean, near-surface
streamline patterns within the volute. The results are presented in a manner which makes
comparison of the 1200 rpm to 3000 rpm relatively straight forward. Figure 4.7 displays
the streak patterns along the scrolled surface for both impeller speeds. The figure provides
the locations of the eight entry ports, the intersections of the z-r plane PIV measurements,
and the locations of the cutoff (0 = 0°) and the end of the scrolled surface (0 = 290° ).
Figures 4.8 and 4.9 display the streak patterns on the inlet surface. The images from the
two impeller speeds are cropped and spliced together to make direct comparisons of the
streak patterns easier. For example, 3000 rpm is shown on the left of Figure 4.8 and 1200
rpm in shown on the right. The opposite is true for Figure 4.9. The triangles along the
scroll indicate the azimuthal locations for the z-r plane PIV measurements. Figures 4.10
and 4.11 display the rear surface streak patterns. These results are displayed in the same

manner as the inlet surface results.

68

4.4 Hot-Wire Anemometry

Point-wise velocity measurements were conducted at eighty mid-plane locations
within the volute at impeller speeds of 600 and 3000 rpm. The single sensor hot-wire

measurements resolved the x-y plane velocity magnitudes given by:

x.)

U = uf+ 0;. (4.13)

These time series data were used to statistically quantify the velocity field along the mid-
plane of the volute. Statistical convergence of the time averaged quantities revealed that
adequate sample populations were acquired. Sample convergence checks on mean and
ms values are given in Figure 4.12.

The statistical quantities have been normalized to make possible a direct compari-
son of the results for the two impeller speeds. The normalized mean and ms velocities are

given as:
0 = U/U, (4.14)

and

~

13* = U/Ut, (4.15)

respectively. Sample histograms of the near-impeller, mid-volute, and near-surface time
series measurements for entry port f (refer to Figure 3.11) are given in Figure 4.13. These
histograms are representative of trends in the histograms from the other entry ports as the

measurement location moves from the impeller to the scrolled surface.

69

The normalized mean and rrns velocity distributions for the ten measurement loca-
tions of each entry port are given in Figure 4.14. In order to identify spatial trends in these
data, a compact presentation similar to that devised by Humbad et al. [12] and [13] has
been applied. Specifically, the measurement domain between the impeller and scrolled
surface has been fitted with a [10 x 71] point grid (see Figure 4.15) to present this signifi-
cant amount of information in contour plots. The grid locations between entry ports were
assigned values by linear interpolation of the statistical quantities between two corre-
sponding measurement locations. For example, the obtained quantities at measurement
locations a1 and b1 were used to assign values to the nine grid locations between them.
This was repeated for the grid locations between a2 and b2, a3 and b3, ..., b1 and c1, etc.
The resulting mean and rrns normalized velocity contours for the two impeller speeds are
given in Figures 4.16 and 4.17 and Figures 4.18 and 4.19, respectively.

Power spectral density functions were computed for all acquired velocity time
series data. These spectral quantities were normalized for direct comparison of the two
operating speeds. The abscissa (frequency axis) and the ordinate (power density axis)

were normalized by:

 

60
F* = F - — 4.
Q ( 16)
and
E*(F*) - E (F)- Q (417)
U _ U 7" '
60U"

1

respectively. Sample spectra of the near-impeller, mid-volute, and near-surface time

series data for entry port f are given in Figures 4.20 (a) through (c). These spectra are rep-

70

resentative of the trends as the measurement location moves from the impeller to the
scrolled surface. Note the narrowband spectral "spikes" in Figures (a) and (b). In Figure
(3) these" spikes lie at F* = 41, 82, and 123. These correspond to the "blade passing fre-
quency" (bpf) of the impeller and its harmonics. The blade passing frequency is the rate at

which blades pass a given point, bpf is given as:

NQ
bpf — H)- . (4.18)
Normalizing this quantity gives:
bpf‘“ = N , (4.19)

where N is the number of impeller blades (N=4l for this study). The harmonic spikes do
not exist in Figure (b) and the spike at the bpf" is significantly lower, while in Figure (c)
there are no spikes. Also note the broadband spectral quantities between the range of F* =
0 and 10. From Figures (3) to (c) (impeller to scrolled surface) these broadband quantities
increase.

The above observations reveal two trends in the spectral content of the acquired
data. Specifically, a decrease in the narrowband bpf power density and the increase in the
low fiequency broadband power density as the measurement location moves from the
impeller to the scrolled surface. In order to quantify the rate at which these quantities
decrease, or increase, the power density in the specified bands have been integrated for
each time series and displayed in contour plots. The broadband normalized power was

calculated as:

71

10
1513 BB = [13; (F*)dF*, (4.20)
0

while the narrowband normalized power was calculated as:

bpf‘ +1
[153‘J bpf = j 15;; (F*)dF*. (4.21)
bpf‘ —l

The broadband results are given in Figures 4.21 and 4.22 for the two impeller speeds
tested and the narrowband results are given in Figures 4.23 and 4.24. The "drop-off" in
the bpf power with radius required the use of an exponential distribution of the contour

isolines to properly view the trends. The exponential distribution was calculated as:
C0 _"
Cn = (_) , (4.22)

where C is the isoline value and n is the isoline of interest (n = 0:1 :4; 0 = highest level).
Phase averaged statistics were computed for each time series to quantify the peri-
odic blade-to-blade flow phenomena. A portion of a sample velocity time series and the
associated (simultaneously sampled) TTL signal are shown in Figures 4.25 a and b. The
time for one complete revolution of the impeller is given by the number of samples from
rise to rise of two consecutive pulses in the TTL signal. Small cycle-to-cycle variations of

the impeller speed resulted in varying number of samples acquired per revolution. For

example, a i1 % variation in impeller speed2 would result in 328i4 samples/revolution3.

 

2. Impeller speed was held to better than :25 % of the nominal test speed for all measurement techniques in
this study, with a measurement accuracy of l.l° of impeller revolution.

3. Nominally 328 samples/revolution were acquired over 1000 impeller revolutions.

72

In order to account for this variation, the time series data for each impeller revolution were
resampled by linear interpolation at each degree (or ) of impeller rotation. The phase aver-

aged mean and rrns velocities were then calculated as:

 

 

N CV
(0)0.) = N1 Z UR(1 +(1—1).360) (4.23)
revi :l
and
N 1
~ 1 rev _ 2 2
<U>(a) = Nrev-l Z(UR(1+(i——1)-360)—(U)(a)) , (4.24)

121

respectively, where UR is the resampled time series and Nrev is the number of complete

revolutions. The resulting phase averaged mean and rms quantities from the time series of

Figure 4.25 (a) is shown in Figure 4.25 (0). Sample statistical convergence of phase aver-

aged quantities at or = 0° is shown in Figure 4.26.

Normalized Phase averaged mean velocity ( (19*) = (U)/Ut) distributions for

the ten measurement locations of entry port f are shown in Figures 4.27 and 4.28 for 600
rpm and 3000 rpm, respectively. The data shown are for the first 45 degrees of impeller
revolution. These same data are displayed as contours in Figures 4.29 and 4.30 to gain a
spatial sense of the given distributions. The corresponding normalized phase averaged

rrns quantities are shown as contours in Figures 4.31 and 4.32.

73

Note the decreasing phase mean amplitudes in Figures 4.27 and 4.28 as the mea-
surement location moves from f1 to f10. In order to gain a measure of this trend, the ms

of the phase averaged mean distributions has been calculated as:

1
359 5

1 — — 2
0phase = 359 Z (<U>(Ot)-U) 9 (4.25)
or = 0
where 0° _<_ or 5 359°. This quantity was calculated for all time series data and is pre-
sented as contours in Figures 4.33 and 4.34 for 600 rpm and 3000 rpm, respectively. The

isoline levels of these contours are distributed exponentially as given by Equation 4.22.

4.5 Scrolled Surface Static Pressure Distribution

A representative statistical convergence on the mean surface pressure is given in
Figure 4.35 showing that an adequate population has been acquired. The results of the
static pressure measurements were normalized as:

—*
—* PS

5 -1 2
2"”

 

(4.26)

and are shown in Figure 4.36 for the four impeller speeds tested.

4.6 Microphone: Fluctuating Pressure Measurements

The pressure time series data has been acquired at impeller speeds of 600, 1200,
2400, and 3000 rpm. The statistical results of these data are given in normalized form.

The results of the four impeller speeds are presented in the same figure for a given statisti-

74

cal quantity. Pressure time series were checked for statistical convergence on rrns. A
sample nns pressure convergence is shown in Figure 4.37.

The acquired pressure time series were used to quantify the fluctuating component
of pressure, i.e. the rms values, at each entry port. The calculated rrns pressure values

were normalized as:

 

P': = (4.27)
-pU.2
and are shown in Figure 4.38. Sample histograms of the pressure time series data are
shown in Figure 4.39 for entry ports a, d, and h.

Figures 4.40, 4.41, and 4.42 show sample power spectral density functions for
entry ports a, d, and h, respectively. The frequency axis is normalized as given in

Equation 4.16, while the power density axis is normalized as:

Q

:1: a1: .. _
EP;(F)—Ep;(F) 1 2
“1129111)

2 . (4.28)
The microphone drop-off in frequency response below 20 Hz is shown in these figures to
have an "insignificant effect on the total power of the signal. The equivalent normalized
frequencies at 20 Hz are F* = 2, 1, 0.5, 0.4 for the impeller speeds of 600, 1200, 2400, and
3000 rpm, respectively.

Both auto and cross-correlation coefficient functions have been calculated from
the acquired time series pressure data and are presented as a function of impeller rotation

(01 ). Sample auto-correlations are shown in Figure 4.43 for entry ports a, d, and h to show

75

'1

trends in the primary flow direction of the volute. Sample cross—correlations are given in
Figures 4.44, 4.45, and 4.46. These figures are to be read left to right and show the pro-
gression'and reduction of the peak correlation as the distance between measurement loca-
tion is increased. For example, in Figure 4.44 the cross-correlations, from left to right, are
between ports a and b, a and c, and a and d. The lag (in degrees of impeller rotation (or) ),
of the peak correlation is provided in each plot.

The peak cross-correlation lag information for the acquired pressure data is a mea-
sure of the elapsed time it takes for a given flow event to travel between measurement
locations. This information was used to calculate an average convection velocity between

consecutive entry ports, i.e. a-b, b-c, etc. The convection velocity was calculated as:

 

. _ d . 360° . 2
convection 8D (ID ()0

U (4.29)

where d6p is the distance between entry ports along the volute surface and or is the degree

p

of impeller rotation at peak correlation. These data were normalized by Ut and are pre-

sented in Figure 4.47. These results show the normalized convection speed to be on the
order of one. An initial acceleration is observed followed by a fairly constant convection

velocity and finally a deceleration toward the outlet of the volute.

4.7 Particle Image Velocimetry

Whole field velocity data were acquired along six measurement planes within the
volute, see Figure 3.15. These measurements resolved the "in-plane" components of the

velocity vector. The x-y planar measurements resolved the velocity vector given by:

76

6x)! = fix '1‘ fly, (4.30)

where ux and uy are the x and y components of the velocity vector. The z-r planar mea-

surements resolved the velocity vector given by:
u,r = G, + f... (4.31)

where uZ and u, are the z and r components of the velocity vector. Results of these mea-
surements are displayed in a stationary reference frame and are normalized by Ut.

The measurements acquired to gain time mean statistics are given in mean velocity

magnitude contours. The x-y and z-r planar magnitudes were calculated as:

2

y (4.32)

ny: L-l'i'fi

and

0,. = 105+ 03. (4.33)

respectively. Sample statistical convergence on the time mean and rrns4 values are given
in Figure 4.48. The normalized x-y plan time mean velocity magnitudes are given in
Figures 4.49, 4.50, and 4.51 for the Inlet, Mid, and Rear planes, respectively. Measure-
ments were obtained at both 600 and 1200 rpm for the z-r planes in Regions 1, II, and III.
For the two impeller speeds tested, normalized results for Region I are given in
Figures 4.52 and 4.53, Region II in Figures 4.54 and 4.55, and Region III in Figures 4.56

and 4.57.

 

4. As will be discussed. calculated rms velocities are presented in APPENDIX B.

77

Streamlines are overlaid on the given contour plots to provide the directional infor-
mation of the mean velocity vectors in the measurement plane. Streamlines are lines that

are everywhere parallel to the velocity vector and are defined by:

uxdy = 11de (4.34)

and

udr = u dz (4.35)

l. r

for the x-y and z-r planes, respectively. The displayed streamlines should not be confused
with the mean three-dimensional streamlines of the flow. That is, the displayed stream-
lines are a result of the projection of the three-dimensional velocity vector onto the mea-
surement plane, thus provide only two-dimensional information of the "in-plane"
components of velocity.

The phase averaged mean velocity magnitudes were normalized and are given in
Figures 4.58, 4.59, and 4.60 for the Inlet, Mid, and Rear planes, respectively. These x-y

plane magnitudes were calculated as:

<13...) = ./<0;’;) + (hf). (4.36)

A conditional averaging algorithm was not needed to calculate these phased quantities (as
was needed for the hot-wire phased quantities) since data acquisition was synchronized to
blade position.

The two component velocity information of the acquired PIV data allowed for the

calculation of certain components of Reynolds stress tensor in each measurement plane.

78

. ..2 -2 A
These stress components include uX , uy , and u'xu'V for the x-y plane measurements and

~2 ..2 —- . . .
uZ , ur , and u’Zu'r for the z-r plane measurements. The display of these quantities are pro-

vided in APPENDIX B.

79

 
    
  
   
 

 

   
  

 

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' .. AAA—3000 rpm \ "5,; \
Ouirniltriilrrrrlir11L111111\Y‘11 I
0 0.05 0.1 0.15 0.2 0.25 0.3

1)

FIGURE 4.1 Performance: Static Pressure Coefficient vs. Flow Coefficient.

n... n. (%)

 

 

 

    

 

: // 4:3” . “
AW 5‘ I :"8\ iii.
0:...1....1....1....K-.1..n\i

0 0.05 0.1 0.15 0.2 0.25

(1)

FIGURE 4.2 Performance: Static Efficiency

80

0.3

vs. Flow Coefficient.

 

 

 

 

 

5000 - 600 rpm .
:g ,___.__120() rpm 9(/
, ————e— 2400 rpm 0/
- A3000 rpm *4
4000 -
t x .0
.. ,A’ 29’
-- / ,Rf’
. )3 ye.’
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3000 _ ,6
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i- [w _ if ..‘1‘ 8M8»?
0,g§%.L.n..i....r....1....r....r
0 0.05 0.1 021,5 0.2 0.25 0.3
FIGURE 4.3 Performance: Reynolds Number vs. Flow Coefficient.
0.4 L- ——e—— 600 rpm Uinlct
; ~—»—~~--—4A~A 1200 rpm .
0 35 —_ AeA2400 rpm 0"
‘ : A3000 rpm ..0' '
: 0,6)”
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0.1 :- 57 flfl ’49.:
; 9 60" 1 "fl
7 65’" ”a”?
0.05 - ..
: ,6“
Oh'iilllllllllLLIIlllrlllJJIl.LALll
0 0.05 0.1 0.15 0.2 0.25 0.3

<1

FIGURE 4.4 Performance: Normalized Average Net Flow Velocity vs. Flow
Coefficient. Collapse of Curves Show Confirmation of "Fan Laws".

81

 

. ,6)“ I'OMT-‘C’Ilfgk N

1.25 -
P1.“

6* \ “
\\ HVAC
was,

0.75 :— ~\

0.5 - R l
X
Ara—Free Inlet ‘1
A Fresh-Air Mode

Ae— Recirculation Mode

W”, W3

L

0.25

IlrtTIIui

 

 

 

0 IILIJJJALIIIIIllmli [Ill

0 0.05 0.1 0.15 0.2 0.25 0.3

1)

FIGURE 4.5 Performance Inlet Study: Static Pressure vs. Flow Coefficient:
Impeller speed of 3000 rpm.

 

  

30 F {rake—HRS-
_ /B/ 7': III,“ I“ ..
L l/ifb/Qx-MN m \
.— _ , *9\ II-‘K
25 - % Blower 8‘0»
A. : a
\ *- 6’5" , _» ._ '3’» :El
&/ 20 - fllfi‘fl\fl\
m T I‘KRIYr
57 ~ PK
5‘: 15 -
C j
10 h-
5 f' Ara—Free Inlet .
. AAA—A Fresh-Air Mode .\
” —e—— Recirculation Mode K
01111.1;1111111'11111111lir‘xrl

 

 

 

0.05 0.1 0.15 0.2 0.25 0.3

1)

FIGURE 4.6 Performance Inlet Study: Static Efficiency vs. Flow Coefficient:
Impeller speed of 3000 rpm.

Ct

82

A

H cgwum : 9:va E enscwoz

:2 w. ..nx.r . . ..r

:Q .505 .CmEcm

28825
.505 E5

 

Scrolled Surface.

ing:

.7 Surface-Streak'

FIGURE 4

83

Q = 1200 rpm

12 = 3000 rpm

A Indicates z-r PIV
Measurement Plane

1' Bifurcation
‘ Line

I

A Indicates z—r PIV my .
Measurement Plane is.

   

  
 
  
  

FIGURE 4.9 Surface-Streaking: Inlet Surface (Images Cropped for Comparison).

84

 

Q = 3000 rpm V Q = 1200 rpm

,5 ,Cutof is -
.t > "‘zv ' 'I ‘ "9 .v . "

A Indicates z—r PIV
Measurement Plane

    

Bifurcationé‘ ,

 

FIGURE 4.11 Surface-Streaking: Rear Surface (Images Cropped for Comparison).

85

A Sn

W
U=6A7mm

    

fl

% Difference from Mean
:3
M

0i
i
- l , egg --_. _‘
0'50 l 2 3
NumberofSamples x105
E 4.’ " "—6003;
E 3; 6:1.17m/s;
E l u
“- 2
8
C:
2 l
=93 l
'5 0
5°
'10 i *2 a *3

Number ofSamples x l0S

 

_ __,.,__ _

7*, __Y_._..

3000 rpm
U = 33.02 m/s l

.o .0
hO‘

 
 
  

 

  

g
Q)
2
E
g 0.2
8
5 0
:5:
'5—02
b\b
-0.4 -- M, -~-———- -
O . l 2 3
Number of Samples x 105
3 ,, ,, m -M
E 3000 rpm
E 2 ii = 6.47 m/s
2
“- l
8
C:
8 0
:2
5 -|
a“

7 g , __, .
“o l 2 3
Number of Samples x 105

FIGURE 4.12 Hot-Wire: Time Averaged Mean and rms Velocity Statistical
Convergence for Location g1.

86

O

 

 

600 rpm 7
8‘ —e—— 3090 rpm
a. 6-
8
CU
m 4
s
2,
(a)
O .__ .
0 0.5
10 i
8
ii 6 l
e
a, ,
Cl) 4;
be
2‘
O ":';-.-::1’;?,iiE¥-— - -—~— -
0 2 2.5
"——- 60??me
“*5- 3099-:Pgl
1 5 2 fl 2 5

 

FIGURE 4.13 Hot-wire: Normalized Velocity Histograms for Measurement
Locations (a) f1 (b) f5 and (c) f10.

87

    

 

——B— 600 rpm
~—9—- 3000 rpm

0.5
~ 0.25 "N
U. E
O ,

al-lO bl-lO cl-lO

 

O

\l

U”:
lllll'llllllIjlllllllJoI"

 

 

 

 

 

 

 

 

ix
.
t
I
I
\ D
.Q‘\
.
I-i
k,/

MK;

d1-10 el-lO f1-10 gl-lO 111-10

Measurement Location

FIGURE 4.14 Hot-Wire: Normalized Velocity Distributions.

88

y (mm)

100

50
O
Impeller
-50
-150 -100 -50 O 50 100
' x (mm)

FIGURE 4.15 Hot-Wire: Mesh Positions for Interpolated Values.

89

   

150

y (mm

y (mm)

 

 

14

150
x (mm)

FIGURE 4.16 Hot-Wire: Normalized Mean Velocity Contour: 600 rpm.

I o = 3000 rpm

 

 

 

x (mm)

FIGURE 4.17 Hot-Wire: Normalized Mean Velocity Contour: 3000 rpm.

90

y (mm

’_ Q=600rpm

100

50

y (mm)

 

 

-50
-150 -100

 

x (mm)

FIGURE 4.18 Hot-Wire: Normalized rms Velocity Contour: 600 rpm.

: n =3000 rpm

 

 

 

x (mm)

FIGURE 4.19 Hot-Wire: Normalized rms Velocity Contour: 3000 rpm.

91

 

 

10" ; .
I s . - 600 rpm
. i , 3000 rpm
10‘?
IO" {- t l
. 3 Er...“ ‘ MLM ”—1 I
“-1 10 4 " mwrm W~M~w«i‘— . .
r- 7, ii 4" ’l‘IM k!‘;‘y"m.'-,~‘W 5"": “Vi
105r
y (a)
10-0 1 1 l 1

l l l l
0 20 40 60 80* I00 120 I40 160
F

 

 

 

 

10'l r-
t — 600 rpm
. -- 3000 rpm
10" -
10" r l
a 3 p _—.l'i
LL] ‘ “WNW ‘J I".
1 0‘4 _ ‘*.Vp)’v;:ff?ic-1Y‘T?M\
: »,'~'~\T‘Mf.mw.w ~
V . 1 4r W W Wmt‘lemmggm
10“ —
(b)
1 04‘ l l l I l l l l
0 20 40 6O 80. I00 120 I40 I60
F
10" :—
~ 600 rpm
1 3000 rpm
10”r
l
h
10 ‘ r- "r
o D ” “~‘
m 4 ~ \\
10 I' iniix‘uf‘fmm
: Wok-
_ . 3.3.(‘fifiru my ‘
10-5 I'" \ mug... 9*.“ W‘hig'. “my” ., .-r ”V.” ,L “was: i-w“
I (C) “""‘I
l 0'0 l l A I I l I L
0 20 40 60 80* l()() I 20 I40 160

FIGURE 4.20 Hot-Wire: Normalized Velocity Power Density Spectra for
Measurement Locations (a) fl (b) f5 and (c) f10. Arrows in Figures a and b Indicate
the Spectral Peaks at 600 rpm.

92

y (mm)

 

x (mm)

 

J‘EU BB

1 . 10E-02
9.00E-03

-. 7.00E-03

5.00E-03
3 .00E-03

FIGURE 4.21 Hot-Wire: Normalized "Broadband" Energy Contour: 600 rpm.

: Q=3000rpm .

y (mm)

 

x (mm)

 

J’EU BB

1 . 10E-02
9.00E-03

7.00E—O3
7' 50013—03

3.00E—03

FIGURE 4.22 Hot-Wire: Normalized "Broadband" Energy Contour: 3000 rpm.

93

y (mm)

y (mm)

 

 

.
IE U bpf

1.00E-02
35013-03
~ 1.2213-03
4.29E-04
1.50E—04

 

 

 

 

I I I
150
x (mm)
FIGURE 4.23 Hot-Wire: Normalized "bpf" Energy Contour: 600 rpm.
I Q = 3000 rpm . .

i Eu bpf
1.00E-02
3.50E—03
1.22E-03
4.29E-04
1.50E-04

Impeller
I I LI 6L1 14L l_l
-100 150

x (mm)

FIGURE 4.24 Hot-Wire: Normalized "bpf" Energy Contour: 3000 rpm.

94

 

v.» A U!
o g o
I
I
I
I
I

to
ES?

Velocity Signal U

0

Acquired Time Series (m/s)

A
9.)
V

0 , , , r .; _ __
l ()0 200 300 400 500 600 700 800
Sample #
6 I
5 I " I'I I" I

A

 

 

 

 

 

 

 

 

 

‘ TTL Signal (volts)
N
I

 

  
 

 

(b) .

-1 7 . 7 7 7 . ——.....-—— we

I 00 200 300 400 500 600 700 800
Sample#

50 ~ ~ ~ , ~ ~ a . ;
73 Phse Mean <U> I
E40 ~ .

'5

Eu 7

2}; 30+

1::

Q)

DD

3 20'

0)

> .

< I

g 10I -
a I (C) Phase rms <U>

  

 

‘LfiLA— ,WL

0 50 ' 7300 f i 130 if 200 ‘250 30 350
(I (degrees)

FIGURE 4.25 Hot-wire: Sample (a) Time Series, (b) TTL Signal, and (c) Phase
Averaged Quantities. Data are Taken From Measurement Location g1, 3000 rpm.

95

% Difference from Mean

% Difference from rms

¢
VA

800

 

7 600 rpm
0.5, <U> = 6.69 m/s
0.
2400 600 700 800 900 I 000
Number of Samples
2‘ 97500 Fpm
<U> = 1.00 m/s
0.
-2.

600 700 800 900 1000

Number of Samples

C/"r Difference from Mean

% Difference from rms

".400

3000—mp?
<U> = 34.59 m/s

 

70.5 I

I

600 700 800 900 1000
Number of Samples
5 . 7 if“
3000 rpm
4 <0>=4J4mm

    

0 *

600 700 800 700 1000

Number of Samples

FIGURE 4.26 Hot-Wire: Phase Averaged Mean and rms Velocity Statistical
Convergence for Location g1 at CL = 0° (TTL Signal Rise).

96

  
 
 
   

W 600’rpm—I:T
. —~— f2

7.... f3
.. I —«-— f4
.. + f5

—e—f6

 

 

, ,77‘-~.~.V|r , p i
->
0.9-

_‘

 

0 5 7910‘ 1'5 20’ 25 30 ’35 40 ’45
or (degrees)

 

FIGURE 4.27 Hot-Wire: Normalized Phase Averaged Mean Velocities for Entry
Port f: 600 rpm.

    

. 300rpm—T—e— f1—

0 '5’ 10 ”'15 20 25 30 35 _"'
a (degrees)

FIGURE 4.28 Hot-Wire: Normalized Phase Averaged Mean Velocities for Entry
Port f: 3000 rpm.

97

<U>
1.40
1.27
1.15
1.02
0.90

5f (mm)

 

20
(1 (degrees)

 

FIGURE 4.29 Hot-Wire: Contour of Normalized Phase Averaged Mean Velocities
for Entry Port f: 600 rpm. (8f = r - r; + 2)

Q = 3000 rpm _.
<U >
1.40
1.27
1.15
1.02
0.90

5f (mm)

 

(1 (degrees)

FIGURE 4.30 Hot-Wire: Contour of Normalized Phase Averaged Mean Velocities
for Entry Port f: 3000 rpm. (6f = r - r2 + 2)

98

sol? Q=600 rpm ~,

 

0J4

40 053

0J2

0J0

”‘30 009
a
Si
“-4
0°

    

00 10 n 20 30 i 40
(1 (degrees)

FIGURE 4.31 Hot-Wire: Contour of Normalized Phase Averaged rms Velocities for
Entry Port f: 600 rpm. (6f = r - r2 + 2)

50

    

Q = 3000 rpm ~.
.. _ <U >
‘ 0J4
40 0.13
0.12
0.10
A 30 0.09
E
ii
%
20’
10‘
0 ( , , - a" 7 .
0 10 20 30 40
0. (degrees)

FIGURE 4.32 Hot-Wire: Contour of Normalized Phase Averaged rms Velocities for
Entry Port f: 3000 rpm. (6f = r - r; + 2)

99

/U,

ophme

  
  

   

50

y (mm)

Impeller/

 

_50 I I

 

x (mm)

FIGURE 4.33 Hot-Wire: rms of Phase Averaged Velocity Normalized by Tip
Velocity: 600 rpm.

y (mm)

 

 

Q=3000 m ,
I- rp opium/Ul
1.00E-01
100 5.00E-02
2.50E-02
1.25E—02
6.25E-03
50
0.
Impeller/I .
_50 l 1 I l I l 1
—150 -100 150

x (mm)

FIGURE 4.34 Hot-Wire: rms of Phase Averaged Velocity Normalized by Tip
Velocity: 3000 rpm.

100

600 rpm

 

% Difference from Mean

 

' 0 1000 2000 3000 4000
Number of Samples

FIGURE 4.35 Static Pressure: Representative Static Pressure Statistical
Convergence on Mean for Entry Port c.

1.4

1.2

 

   
 

IIIIIIIIIIIIII

 

 

 

* w 0.8
In.
0.6 l
i —-—e——- 600 rpm
: 1200 rpm
0'4 T ~~——e— 2400 rpm
; -----—-—<>—-——— 3000 rpm
0.2 L'
0 b l l J l L l I
a b c d e f g h
Entry Port

FIGURE 4.36 Static Pressure: Normalized Static Pressure Distributions, Lines are
Plotted to Show Connectivity.

101

3000 rpm
(5
g 5
4
:3
Q)
B
fa; 2
a
b" 1
0
_] '
0 0.5 I 1.5 2 2.5 3
Number of Samples x 105

FIGURE 4.37 Microphone: Sample Fluctuating Pressure Statistical Convergence on
rms for Entry Port (1.

0.06

 

  
   
 
 
 

600 rpm
---..77___.. 1200 rpm
7777—«9— 3000 rpm

0.055

0.05

0.045

0.04

IIII'IIIIFIIIFTII'IIIT'

 

 

  

IIIII'I

0.025

I
I

IX
I

0.02

TII'II

  

 

 

0.015 _
0.01 E-
0005 E-
0 3 L l I l l l l
a b c d e f g h
Entry Port

FIGURE 4.38 Microphone: Normalized Fluctuating Pressure Distributions, Lines
are Plotted to Show Connectivity.

102

I ~54 600 rpm

 

 

 

 

‘ + 1200 rpm
, P
20 Entry on a _e_ 2400 rpm
i3 15: —6— 3000 rpm
0.
E
N
‘2 10
5 ;
SI
0 ' t’
-0.5 -0.4 0.2 0.3 0.4
P';
25 , ,
--E- 600 rpm
20 Et P td +1200'Pm
” 'y or 7% -e— 2400 rpm
33 15 J3, wom
Q.
E
N
‘2 l0
5
5
0
—0.5 -0.4 -0.3
PT;
25 , ,7,
+ 600 rpm
20‘ Entry Port h + 1200 rpm
—e— 2400 rpm
3 + 3000 rpm I
315‘ i _.
E .
{U I
‘2 10"
E
5.
0 ‘:/'-.'.—‘ ;. . " t :1" Hit". :.:." ::: 2::
-0.5 -0.4 -0.3 -0.2 -0.1 0 0.1 0.2 0.3 0.4

FIGURE 4.39 Microphone: Normalized Pressure Histograms for Entry Ports a, d,
and h.

103

-4
10

 

 

() 20 40 60 80 100 I20 140 160

F*

1200 rpm i
I

      

 

 

 

ufio‘I
0 20 40 00 80 I00 I20 ‘140 160
F*
-4 ,, - # -——. .
'0 2400 rpm I
I
*, 2;, -j $44.“;
“Profit
0 20 240 60 ’80 I00 I20 I40 i60
F*
74 — ~~ »- --—— I
'0 3000rpn1I
I
~ _* LI-” ..n-nn_ -_ I
-x» ”r " "'"' ' " - WV iv f" L'“~‘~““.: 2 ~'
a , wI
Lu10 - =
0 20 WP 40 60 80 I00 120 140 7' 160
F*

FIGURE 4.40 Microphone: Normalized Pressure Power Spectrums, Entry Port a.

104

 

 

0 20 40 00 80 I00 120 'i40 160

F*

'0 1200rpn1I

  
  

 

 

 

15’;
LL110
0 20 40 00 80 I00 120 1407i60
F*
-4 _ 7— . 7 .
10 2400 rpm
*7):
”310‘-
0 20 40 00 80 I00 I20 140 160
F*
'0-4 _ _ . 7 __ .. I .
3000rpnr,
*2]:
Ln“10'6
0 20 7 40 60 80 I00 120 140 100
F3:

FIGURE 4.41 Microphone: Normalized Pressure Power Spectrums, Entry Port (1.

105

 

0 20 40 7 00 so I00 120 140 160
F*

 

 

0 20 _“ 40 00 80 I00 120 ‘7140 160
F*

-4 . . __. __ .
‘0 ' 2400rpnii

 

0 20 If 40 f 60 80 I00 120 7140' igo

 

 

F3:
3000 rpm i
m 7 FUD
0 20 If 40 I f 60 80 I00 I 120 I40 160

F2»:
l

FIGURE 4.42 Microphone: Normalized Pressure Power Spectrums, Entry Port h.

106

 

 

 

 

 

 

 

 

 

   

 

 

 

 

0 8 600 rpm 0 8 600 rpm 600 rpm
06 0.6
g 0.4{ g 0.4‘
3‘ 0.2I 7 3‘ 0.2
“Ire” °IF —
-0.2~ -0.2
I.I70 I.I5O 1.300,
.04 - - 7 -04 7 -o.4 - -- - --———- -- 7 7a
0 10 20 30 40 0 1o 20 30 40 0 10 20 30 40
(\ (degrees) (\ (degrees) (r (degrees)
0 8 1200 rpm . 1200 rpm 1200 rpm
0.6
g 04'r
3‘ 0.27 .
0* W —— I.
70.2 - . I
I050 L350 1.390 '
-04 7 -o.4 7 7 -0.4 -—- 7 ~-- - _.
o 10 20 30 40 o 10 20 30 40 0 10 20 30 40
(\ (degrees) (\ (degrees) (1 (degrees)
, 7 , I, 7 L _ __ _
0.8 -400 rpm 1 0‘8 -400 rpm 018 2400 rpm
06‘ . 0.6
I
g 0.4I 7 g 0.4
34 027 - 3‘ 0.2
o. 0
-0.27 -02 -0.2 <
0720 L690 1.440
70.4 7 7 7 04 -0.4 - - “—7- 77
o 10 20 30 40 o 10 20 30 4o 0 10 20 30 4o
(\ (degrees) (\ (degrees) (X (degrees)
0 8 3000 rpm 0 8 3000 rpm
06; 0.6
g 0.4) ~ g 0.4
& 02+wa a: 0.2
0} . , 0
-02. 0.690 f -0.2 1.900
-0.4 7 —0.4 - -0_4 - 7 777* 7 -
0 10 20 30 40 0 10 20 30 40 0 10 20 30 4o
(\ (degrees) (\ (degrees) (\ (degrees)

FIGURE 4.43 Microphone: Normalized Pressure Auto-Correlations for Entry Ports
a, d, and h (R*(or)=0.5 Given at Bottom Right of Each Plot).

107

0.2I , 7 0.2 - 0.2 rrrrrr

 

 

 

000 rpm I 000 rpm 600 rpm
017 0.1 0.1
I ‘ '
a I I [I a a ‘
* 0w * - p * OWW * 0W
a: I I 06 a: 7
01- ‘ 70.1 70.1 I
42° 80° 109°
02 7777 7 I -0.2 7 70.2 77 - -- 77
0 50 100 150 200 o 50 100 150 200 0 50 100 150 200
(1' (degrees) (\ (degrees) 01’ (degrees) .
0.2 7 0.2 0.2 77 7 7 7777
l200 rpm 1200 rpm 1200 rpm I
0'“ 7 0-1 0.1 I
..D U . "U -
CU fl 1
* 0 , * 0W :3 0- -
o: I ‘ 05 ad
0.1; I 70.1 70.17 I
, 400 ‘ 85° 115° I
02:7 7 7 I 70.2 77 70.2 7 7 77777777 777
0 50 100 150 200 0 50 100 150 200 0 50 100 150 200
(1 (degrees) (.1 (degrees) ()1 (degrees)
0.2 . 0.2 77 0.2 77-7 777-777. I
2400 rpm 2400 rpm 2400 rme
0.1 01 7
a ,1 PMMW' , a i I
* 0 I * 0W
D4 04
70.1I i 70.17 ‘ 70.17 7
. 37° 81° T 108° I
02 I 7 70.2 7 7 70.2 77 7777—7 7 7
0 50 100 150 200 0 50 100 150 200 0 50 100 150 200
('1‘ (degrees) (1' (degrees) 01 (degrees)
0.2 i 7 7 7 0.2 7 7 _ 0.2 7 77 777 7777
7 3000 rpm; 3000 rme 3000 rpm!
‘ .
0.1 i I 0.1 7 01 .
. I
..D U '0 I
N a .
.7 0W .7 0W 5 0W
a: a: l a:
70.17 E 7, 70.1 7 701- I
- 37° 231° 130° |
02' 7- - 7 77 70.2 7 70,2 77 77 777 77- 77 7
0 50 100 150 200 0 50 100 150 200 0 50 100 150 200
(X (degrees) (1* (degrees) 01’ (degrees)

FIGURE 4.44 Microphone: Normalized Pressure Cross-Correlations. Progression
of Peak Correlation is Shown (Left to Right) from Entry Port a to b, a to c and a to d.

108

0.2 - 7 0.2 02 777. . _

 

600 rpm (500 rpm 600 rpm ‘
0.17 0.1 0.1
'O I 0 u— i
a: a: a: E
-0.17 01 701
40° 78° 121°
02 777 -0.2 - 02 ——-7—777 7 ——-- 7
0 50 100 150 200 0 50 100 150 200 0 50 100 150 200
(1 (degrees) (1 (degrees) ()1 (degrees)
0.2 7 0.2 0.2 7777 7777 7 7 7
l200 rpm l200 rpm 1200 rpm
0.1 - 0.1 01
-o l o - I
*0 0AM ;’ 0W ,3 CW
Dd CZ M .
70.1. - -o.1 701
43° 81° 122°
02 7 7 - - 70.2 70.2 7 77 7777 77
0 50 100 150 200 O 50 100 150 200 0 50 100 150 200
(r (degrees) (11 (degrees) ()1 (degrees)
02. 7 7 0.2 0.2 7 777777 77777 7
2400 rpm 2400 rpm 2400 rpm
01

o .
R* cf

0

l

 

-o.1;» 01- 70.1

 

43° 82° 128°
02‘ 7 - 02 -0.2 77-7-77 7 7— 7
o 50 100 150 200 0 50 100 150 200 o 50 100 150 200
(\ (degrees) 0 (degrees) 01 (degrees)
0.2 0.2 7 0.2 777 77 --

3000rpn1 3000rpn1 * ‘“3666IprI

   

-0.1 ‘- 7 -0.1 01
42° 85° 131°
702 7r7 - 7 7 70.2 7 70.2 7 —7 7- 77
0 50 100 150 200 O 50 100 150 200 0 50 100 150 200
()1 (degrees) (‘x' (degrees) ()1 (degrees)

FIGURE 4.45 Microphone: Normalized Pressure Cross-Correlations. Progression
of Peak Correlation is Shown (Left to Right) from Entry Port c to d, c to e and c to f.

109

 

 

 

 

0.2. - 0.2 g 0.2 7 7v7
600 rpm 600 rpm 600 rpm
017 0.1. 0.1 I
I
H— on .1: I
Q) Q) <1)
* 0W * 0W * 0mm
M a: M
‘0" ' -0-1 70.1
' 430 89° 144°
.02 7 70.2 702 777777777777— 7
0 50 100 150 200 0 50 100 150 200 0 50 100 150 200
(\ (degrees) (\ (degrees) (X (degrees)
02. 777 7 0.2 0.2 77 7‘7
I200 rpm. 1200 rpm 1200 rpm
01 0.1
8’ '5 I
* 0W * 0W
M M
70.1 70.1?
I 42° 85° 134° I
0.2 I 7 77 -0.2 70.2 7777 77 7
0 50 100 150 200 0 50 100 150 200 0 50 100 150 200
(1 (degrees) (1' (degrees) (X (degrees)
0.2 7 7 0.2 0.2 7 77 777-777 777 7
2400 rpm 2400 rpm 2400 rpm
0.1
‘ .
0W CW
701
83° 127°
....7..- - . .. . --. , _‘ - . -- .- .02 . __+h.____-L. ___._- ,
o 50 100 150 200 0 50 100 150 200 0 50 100 150 200
(1" (degrees) (1’ (degrees) (X (degrees)
0.2 02 0.2 7 77 7 7 7
3000 rpm. 3000 rpm 3000 rpm
01 0.1
‘0—
Q)
9(-
:24
-01 ~01
82° 130°
702% - 7 70.2 7 70.2 7 777 7777 I
0 50 100 150 200 0 50 100 150 200 0 50 100 150 200

(1' (degrees) (\’ (degrees)

(X (degrees)

FIGURE 4.46 Microphone: Normalized Pressure Cross-Correlations. Progression
of Peak Correlation is shown (Left to Right) from Entry Port e to f, e to g and e to h.

110

 

 

 

1..) :-
12 E-
1.15—
1%-
DH 0.9 E-
\ 0.8 E—
a-
8 E
“>906 I:-
g :
005 :-
D 0.4:- ———-B— 600rpm
03; ----—~é~—-———- 1200rpm
E ——C—>— 2400 rpm
02;” ———<+— 3000rpm
0.1:-
0 3 I l I L I l I
a b C d e f g h
Entry Port

FIGURE 4.47 Microphone: Normalized Convection Velocity. Values are Based on
the Lag Time of the peak in the Cross-Correlations and the Separation Distance
Between Entry Ports a-b, b-c, c-d, d-e, e-f, f-g, and g-h. Lines are Plotted to Show
Connectivity.

111

 

3, - -
C‘.
G
d.)
2
E
E
“—4
d)
U
i:
8
g
D
{S
100 200 300 400 500
Number of Samples
3 _
2. U? = 1.8] m/s
1
0

  
  

% Difference from rms

100' 200* 300 4110 500
Number of Samples

C/r Difference from Mean

% Difference from rms

 

100 200 __”3001 460 500
Number of Samples

 

100 2110 @350“ 4110 500
Number of Samples

FIGURE 4.48 PIV: Time Averaged Mean and rms Velocity Statistical Convergence
for Region 11 (z = 43.3 mm, r = 103.9 mm) at Q = 600 rpm (97 Percent Validation).

112

 
   

 

 

 

 

- Inlet Plane
_ Q =600 rpm
100 -
E L
E 50
>~ I
0 —--H -
_ h
Impeller
60”. 1.. 1. .. 1. .. .1. .. .1 .. . ..I
-150 -100 -50 O 50 100 150

x (mm)

FIGURE 4.49 PIV: Normalized Velocity Magnitude Contour with Streamlines: Inlet
Plane (x-y) 600 rpm (See Figure 3.15 for Plane Definitions).

   
   

 

 

 

 

_ MidPlane
_ 9 =600 rpm
100 —
a L
g 50 ..
>~ :
0 —-fi
_ h , .
_ . |
” Impeller '
60”...1...1...-1....1../i....n
-150 -100 -50 0 50 100 150

x (mm)

FIGURE 4.50 PIV: Normalized Velocity Magnitude Contour with Streamlines: Mid
Plane (x-y) 600 rpm.

113

17

l

100

11'

I

I

50—

y (mm)

 

 

 

 

 

I" Impeller
60” ..1...1....1.. . 1
-150 -100 -50 0 50 100 150
x(mm)

FIGURE 4.5] PIV: Normalized Velocity Magnitude Contour with Streamlines: Rear
Plane (x-y) 600 rpm.

114

IlllllllllIFTlT

9?
”03.
o

s
)—

 

 

h!lll'IllIlIlTllIlllJlllI

 

 

lllllilllllIIIIIIIIIIIJIIIIIIJIIIIIJIIIIII

 

 

0 10 20 30 40 50 60 70 80
2 (mm)

FIGURE 4.52 PIV: Normalized Velocity Magnitude Contour with Streamlines:
Region I (z-r) 600 rpm (See Figure 3.15 for Region Definitions).

llll’llllll‘lj—I’

Region I
Q =1200 rpm

 

ljrlllllllllllJllIllrlll

 

 

 

 

llllllllllllllllllllllLLllIJ-J-JIIIJllll l

 

0 10 20 3O 40 50 6O 70 80
2 (mm)

FIGURE 4.53 PIV: Normalized Velocity Magnitude Contour with Streamlines:

Region I (z-r) 1200 rpm.

115

 

 

 

 

14o _—
130 :—
120 :—
110 _— [7;
E 1-
5100 L 0.40
:1 : 0.33
: 0.25
90 :‘ 0.17
: 0.10
80 —
70 :—
60:111llllllllllllllllllllijjllllllJllllllIlllll

 

 

0 10 20 30 4O 50 60 7O 80
2 (mm)

FIGURE 4.54 PIV: Normalized Velocity Magnitude Contour with Streamlines:

 

 

 

 

 

 

Region [I (z-r) 600 rpm.
140 r
p
t
130 L'
C
120 E
110 ~— —.
A : U2.
E : *
5.100 _— 0.40
T: _ 0.33
i 0.25
90;” 0.17
2 0.10
80 r
70 {—
60:11lJllllllllllllllll1lllllllllllJllllrllllllll

0 10 20 30 40 50 6O 70 80
2 (mm)

FIGURE 4.55 PIV: Normalized Velocity Magnitude Contour with Streamlines:
Region II (z-r) 1200 rpm.

116

 

_ITI7]IIIIIIIIIIIIIIIIIlllllIlllIII‘lfi

 

 

 

 

 

 

llllllllll l l l l J 1 I l 1 ll

 

O 20 40 60 80
2 (mm)

FIGURE 4.56 PIV: Normalized Velocity Magnitude Contour with Streamlines:

I40

70

60

Region III (z-r) 600 rpm.

E Region III
: Q = 1200 rpm

 

: r

   

 

 

_|llllllll'IIllllllI'llll'llIIlIITll

 

IALIIIIIIIIIIIIJI I ll

 

20 40 60 80
2 (mm)

0

FIGURE 4.57 PIV: Normalized Velocity Magnitude Contour with Streamlines:

Region III (z-r) 1200 rpm.

117

// l
Inlet Plane .

Q = 600 rpm

      
      
  

.......

75

'1111

50

y (mm

25 ......

 

 

-75 -50 -25 0 25 50 75
x (mm)

FIGURE 4.58 PIV: Normalized Phase Averaged Velocity Magnitude Contour with
Streamlines: Inlet Plane (x-y) 600 rpm.

//
Mid Plane

Q = 600 rpm

50

11111111

y (mm)

25

 

 

 

   

-7 -50 -25

x (mm)

FIGURE 4.59 PIV: Normalized Phase Averaged Velocity Magnitude Contour with
Streamlines: Mid Plane (x-y) 600 rpm.

118

//

Rear Plane
Q = 600 rpm

     
    

111‘111N11

y (mm

 

 

111111111111111111111 “‘74“...-
-100 -75 -50 -25 O 25 50 75 100
x(mm)

FIGURE 4.60 PIV: Normalized Phase Averaged Velocity Magnitude Contour with
Streamlines: Rear Plane (x-y) 600 rpm.

119

5.0 Discussion

The discussion of the flowfield within the volute is presented in this chapter on a
flow feature basis. Specifically, the flowfield is subdivided into distinctive or salient fea-
tures that were revealed by the statistical analysis of the acquired time series data. These
features are primarily quantified through the integration of results from multiple measure-
ment techniques. The following sections present each of these salient features in an order

which provide an increasing depth of understanding on the overall flowfield complexity.

5.1 Primary Flow

The primary purpose of the volute is to collect the flow discharging from the
impeller, direct the flow to the blower outlet, and to convert the kinetic energy of the high
speed fluid discharging from the impeller into flow work (pressure/density). Ideally the
flow discharging the impeller would adhere to flow behavior of a source vortex, as dis-
cussed in Section 1.1. Conditions of this source vortex are streamlines with increasing

radius with respect to azimuthal location, a decreasing velocity magnitude in the radial

direction, and, for a given radius, IUI = constant.

The results that show the closest approximation to this ideal flow behavior are the
time mean quantities along the mid—plane of the volute. The hot-wire results in
Figures 4.16 and 4.17 show a radial decrease in velocity magnitude, however, the distribu-
tion is not axisymmetric. This observation is also true for the PIV results of Figure 4.50.
Qualitatively the non-dimensional velocity results of the two techniques show good agree-
ment except for the near impeller PIV results in Region I where high velocity magnitude

levels extend farther into the volute. This may be attributed to the inertia of the seed parti-

120

cles leaving the impeller in this region. Quantitatively the PIV results are slightly lower in
magnitude than the hot-wire results which is likely due to the ability of the seeding to
accurately track the motion of the air given the centrifugal nature of the flow. From the
impeller to the scrolled surface the PIV results show a velocity range from about 1.3 to 0.9

Ut, while the hot-wire results show a velocity range from about 1.5 to 0.9 Ut.

Overall the hot-wire results of the two impeller speeds tested show good agree-
ment. However, a feature which is more pronounced at 3000 rpm is an increase in veloc-
ity magnitude near the scrolled surface along entry ports a and b. This feature is smoothed
by the contour plots, but by examining the distributions of entry ports a and b in
Figure 4.14 one can clearly see this feature. The likely cause of the velocity increase near
the surface in this region is the recirculation of high speed flow past the cutoff of the
volute.

The mid-plane streamline patterns near the impeller show a close resemblance to
those of a source vortex, see Figure 4.50. This figure shows the streamlines diverge upon
exhausting from the impeller and then converge toward the scrolled surface. In a two-
dimensional sense diverging streamlines indicate flow deceleration, while converging
streamlines indicate flow acceleration. The contour levels in this figure show deceleration
in the streamline direction. This observation is consistent with the initial streamline
behavior, however, it is not consistent with the streamline behavior near the scrolled sur-
face. The only explanation for this inconsistency is that there must be a significant axial
flow component in this region. That is, along the mid-plane, the flow exhausting from the
impeller initially shows the character of a radial-tangential flow and gains a significant

axial component near the scrolled surface.

121

In Region II of Figure 4.50 a relatively large deceleration in seen in the azimuthal
direction between entry ports (1 and e. This feature is also seen in the hot-wire results of
Figures 4.16 and 4.17 but is less evident due to the nonexisting measurement points
between the entry ports. This deceleration may indicate a relatively large adverse pressure
gradient in the same region. However, Figure 4.36 shows a relatively smooth adverse
pressure distribution from entry port a to h without a sudden increase from entry ports d to
e that would indicate a large flow deceleration in this region. Coupled with the observa-
tion of converging streamlines, the relatively large deceleration indicates a relatively large
axial flow component in this Region. The details of Region II have been implied by the

general discussion of the primary flow in the mid-plane.

5.2 Double Vortex (Secondary Flow)

The discussion above has given a strong indication of a secondary flow in the axial
direction. This secondary flow is readily seen by observing the z-r plane PIV results in
Figures 4.52 through 4.57. As seen in these figures, 3 pair of counter rotating vortices are
generated which are superimposed on the primary flow. The flow is shown to mainly
enter the volute radially from the mid-span of the impeller and to impinge on the scrolled
surface near the mid-plane of the volute. The flow is then directed axially to both the inlet

and hub surfaces and back radially toward the impeller along these surfaces. The second-

ary flow pattern near the hub surface of the volute has (00 > 0 , while the flow pattern near

the inlet surface has (00 < 0. These structures are not symmetric about the volute mid-

plane. The vortex structure near the inlet surface is shown to occupy a larger cross-sec-

tional area of the volute and has higher flow velocities.

122

These structures can also be inferred from the streamline patterns of the inlet, mid
and rear-plane PIV results of Figures 4.49 through 4.51. As discussed in the previous
section, the convergence of the mid-plane PIV streamlines coupled with the deceleration
in velocity magnitude suggests an axial flow component near the scrolled surface. As the
plane of observation moves to the inlet and rear planes (Figures 4.49 and 4.51) the veloc-
ity magnitudes are notably lower and the streamlines show the flow has a reduced radial
component. The the radial velocity components of Figure 4.49 are not only reduced rela-
tive to the mid-plane, they are negative, i.e. toward the impeller.

As discussed in Section 4.7, the streamline patterns of the PIV results provide only
two-dimensional information about the "in-plane" velocity vector. The time mean, near-
surface (or limiting) streamline patterns given by the surface-streaking results can be used
to gain a three-dimensional understanding of the streamline patterns within the volute. As
discussed above, the streamline patterns show that the flow originating from the impeller
impinges on the scrolled surface near the mid-plane of the volute and is directed toward
the inlet and rear surfaces. This observation can also be made from Figure 4.7. The
results of the 3000 rpm condition clearly show that the limiting streamlines diverge toward
the inlet and hub surfaces nominally along the mid-plane of the scrolled surface. Similar
patterns are evident for the 1200 rpm results of Figure 4.7, however, they are not as clearly
resolved given the low shearing effects of the flow at this lower impeller speed. At the
inlet (Figures 4.8 and 4.9) and rear surfaces (Figures 4.10 and 4.11) the streamlines are
shown to have a strong negative radial component toward the impeller. Close to the cutoff
at the outlet of the blower, see Figure 4.9, the radial component becomes even more sig-

nificant indicating a strong pressure-differential from the blower outlet to the cutoff.

123

The stream-wise vortex motions described in this section are similar to the flow
patterns of a curved channel flow. The bulk flow in a curved channel establishes a radial
pressure gradient such that the pressure at the outer radius is greater than the inner radius.

This pressure gradient results from the Euler-n equation, given by:

IJ

0P _ pV
g " R 9 (5'1)

where P is pressure, n is the streamline outward normal direction, and R is the radius of
streamline curvature. Viscous effects cause the flow along the walls to have a lower
velocity than that of the bulk flow; hence, the momentum of this fluid is not sufficient to
balance the pressure gradient. Consequently, a secondary flow is established in which the
fluid near the wall responds to the radial pressure gradient by moving from the outer
radius to the inner radius along the periphery of the curved channel [25]. That is, the
streamlines now form a helical pattern. This leads directly to the surface streaking obser-
vations: the streamline divergence along the volute mid-plane toward the inlet and rear
surfaces, and then from the volute scrolled surface toward the impeller.

In a study by Widener ate al. [29], thee-dimensional measurements of the mean
velocity field were made in a square—cross-sectional, 90 degree duct flow. The secondary
flow pattern results at the exit plane of the curved duct (CD) are compared below to that of
the volute (V) PIV z-r plane results (Figures 4.52 through 4.57).

1. CD: The vortex structures were symmetric about the mid-plane in both cross-sectional
area and velocity. The structures are centered at 0.2 height (radial) location and 0.25,
0.75 width (axial) location.

V: The vortex structures are not symmetric about the mid-plane. As discussed above,
the vortex structure near the inlet surface is shown to occupy a larger cross-sectional

area of the volute and has higher flow velocities. The location of the vortex centers are
shifted in the positive radial direction from Region I to 111. The approximate radial

124

position of the vortex centers are (given in cross-sectional height) 0.33, 0.55, and 0.67
for Regions 1, II, and 111, respectively.

2. CD: The largest secondary velocity magnitudes are located along the inner radius and

are approximately 40% of the center-line velocity at the same radial location.

V: The largest secondary velocity magnitudes are located at the center of the cross sec-

tion and along the volute surface (outer radius) and are, interestingly enough, also

approximately 40% of the center-line velocity at the same radial location. However,

the area over which the high magnitudes exist in the volute is larger and more evenly

distributed than that of the curved duct.
In summary of these comparisons, the curved duet flow is a useful reference for the sec-
ondary flow behavior, but is clearly only a related (not a similar) flow environment. The
two main differences being the increasing cross-sectional area and the additional flow
being added to the volute as the point of observation moves in the azimuthal direction.
The addition of flow at the inner radius of the volute (i.e. from the impeller) is the major
cause of the differing flow behaviors. The flow entering the volute is delivered mainly
from the mid-span of the impeller with a significant radial component. This radial flow
component then acts to "pump" the vortex motions giving them additional strength and
also "pushes" the vortex centers of rotation toward the outer radius of the volute (i.e.
toward the scrolled surface).

Reynolds number dependence of the secondary flow can be inferred by comparing

the PIV results of the two impeller speeds tested. In Figures 4.52 through 4.57 the 600
and 1200 rpm PIV results for each Region are displayed sequentially for ease of compari-
son. The results show excellent agreement in both velocity magnitude and flow direction,
however, there is some difference in the location at which the radial flow diverges on the
scrolled surface. This location of divergence has also been observed from the surface-

streaking discussion of Figure 4.7. In Region I at 600 rpm the divergence point lies closer

to the inlet surface, by 6mm, than at 1200 rpm. The opposite is true for Region III, where

125

for 600 rpm, the divergence point lies closer to the hub surface by 2 mm. Region II shows
no measurable difference in the location of this point. The shift in this location gives an
indication that the axial flow strength at the inlet of the volute is a function of both azi-
muthal location and impeller speed.

The Reynolds number dependence of the secondary flow can also be inferred by
comparing the surface-streaking results of the 1200 and 3000 rpm impeller speeds. In
Figures 4.8 through 4.11 the 1200 and 3000 rpm streaking results are cropped and dis-
played side by side for ease of comparison. The results show similar patterns between the
two impeller speeds for both the inlet and rear surfaces. A direct comparison of the rear
surface streaking results (Figures 4.10 and 4.11) can not be made since the irregular pat-
tern indicates that the streaking solution for the 1200 rpm condition did not conform to the
near surface streamline patterns. However, the results of the inlet surface (Figures 4.8
and 4.9) are quite clear and a direct comparison is easily made. By observing the stream-
line patterns where the images are spliced, one can see a difference in the angle of the
streaking patterns. The 1200 rpm results show a stronger radial component with respect to
the primary flow direction than the 3000 pm results. That is, the strength of the secondary
flow with respect to the primary flow is stronger at the lower impeller speed. Given that
the reynolds number is a measure of inertial forces to viscous forces, the argument can be
made that viscous forces have an increasing effect on the flow as the impeller speed is
reduced.-

The irregular static pressure distribution along the mid-plane as seen in
Figure 4.36 may be explained by examining 3000 rpm streamline patterns of Figure 4.7.

It is reasonable to assume the axial location of the highest static pressure would lie at the

126

point of streamline divergence for any given azimuthal location. The entry port locations
for the pressure measurements are indicated by the overlaid circles along the mid-plane of
Figure 4.7. The point of streamline divergence with respect to azimuthal location is
shown to "wander" about the these measurement locations. The irregular increase in the
mid-plane pressure distribution is a logical result of the observed "wander".

Given that the "desired" flow direction is that of the primary flow direction, the
secondary flow, i.e. double vertex, is seen to be a significant source of excess kinetic
energy whose subsequent dissipation will contribute to the overall losses in this flow sys-
tem. This statement is supported by the observation that secondary flow velocity magni-
tudes are up to 40% of the primary flow velocity magnitudes. A number of other
undesirable effects on blower performance can be associated with the secondary flow.
Two of these effects are directly measurable and are discussed in the next two sections.
Two other effects on blower performance are discussed below which are simply inferred
from the observed flow patterns of the r-z plane PIV results.

The streamline patterns (Figures 4.52 through 4.57) show the existence of a strong
negative radial flow along the side surfaces of the volute. In Regions I (Figures 4.52
and 4.53) and 11 (Figures 4.54 and 4.55) the negative radial components are shown to
extend all the way to the impeller outlet. That is, the flow is shown to re-enter the impeller
near the inlet and rear surfaces of the volute. The blades of the impeller would then have
an "egg-beater" effect on the re-entering flow. It is a reasonable assumption that this
effect would have two negative associated aspects: the dissipation of energy and the pro-

duction of noise. The dissipation of energy would have a direct effect on the efficiency of

127

the blower while the noise production could be a significant contributor to the broadband

acoustics of the blower.

5.3 Active Impeller Width

The double vortex structures have been shown to be a significant source of ineffi-
ciencies, given that these structures are sustained by the transfer of energy from the pri-
mary ("desired") flow. The negative radial component of these vortex motions are shown
in this section to have additional undesirable effects on the performance characteristics of
the blower.

The streamline patterns of Figures 4.52 through 4.57 show the mid—span of the
impeller to be the main contributor of mass flux to the volute. That is, along the mid-span
of the impeller there is a considerable radial flow into the volute. As the point of observa-
tion along the impeller outlet moves toward the volute side surfaces the radial component
is shown to decrease. In fact, in Regions 1 and II, the inlet and rear planes are shown to
have a negative radial component, i.e. flew back into the impeller. This observation sug-
gests that only a portion of the impeller width is active in adding mass to the volute.

In order to gain a measure of the impeller active width, one needs to examine the
radial velocity profile at the exit of the impeller. Specifically, the width of the profile
which has a positive component, i.e. into the volute, must be measured and compared to

the fill] width of the impeller outlet (b2). This measure was easily obtained by plotting the

radial velocity component with one istach at 0 m/s. see Figure 5.1. An istach at this level

reveals the transition points of the flow from positive to negative. Similar measurements

128

were obtained for all three Regions at both impeller speeds tested. These measures were

calculated as percentage of impeller outlet width and are given in Table 5.1.

 

 

 

 

140*— .
rRegronII
.Q=600rpm

120-

.

E .—

8100—

V .—

L-

80-

60~11 Illllllllllllllll'llltilllllllllllllllll

 

 

 

0 10 20 30 40 50 60 70 80
2 (mm)

FIGURE 5.1 PIV: Sample Radial Velocity Contour with one isotach at 0 m/s
Showing the Distribution of Positive and Negative Values at the Impeller Outlet.

TABLE 5.1 PIV: Impeller Active Width, i.e. Perce_ntage of Impeller Outlet Width
(b2) With a Positive Radial Velocity Profile (ur(H g z s H + b2, r2) > 0).

 

 

 

 

 

 

‘1 Region I II 111 J‘
|| 600 rpm 76.9% 63.1% 100% 1|
1200 rpm 75.4% 63.1% 100%

II II

 

129

Care must be taken when assessing the results given in Table 5.1 as they are not a
"true" measure of the impeller outlet active width. That is, contributions to the positive
radial velocity profile are not only from the addition of mass to the volute. A portion of
the positive profile is the result of the recirculation of the mass already within the volute.
Thus, the "true" active width is smaller than what is reported in Table 5.1 (at least for
Regions I and II).

The results given in Table 5.1 show that the active width of the impeller is a func-
tion of azimuthal location. The results also show the active width is not a function of
impeller speed, at least not for the two low impeller speeds tested. Region II is shown to
have the lowest percentage of impeller active width. This may be a direct result of having
the largest secondary flow velocities. Region III is shown to have 100% active impeller
outlet width. Upon inspection of Figures 4.52 through 4.57 it is seen that the vortex struc-
tures of Region III are removed from the impeller. That is, the vortex structures of Region
III do not "crowd" the impeller as is observed in Regions I and II. This observation
strongly implies that the vortex structures significantly effect the flow exhausting from the
impeller.

In the study by Raj [26], smoke visualization was used to investigate the active
impeller width at the inlet of the impeller. In that study, Raj found that the extent of sepa-
ration, see Figure 1.4, was greatest just past the cutoff and minimum in the area near the
outlet region of the volute. That is, the active impeller width increases width azimuthal
location. Although the results of the current study do not show a direct correlation

between active impeller width and azimuthal location, it is interesting to note that the loca-

130

tion of the most efficient use of the impeller width lies close to the outlet region of the
volute.

Past investigations, e.g. [7,26, 27], have based the measure of active impeller
width on the flow separation zone at the inlet of the impeller. A common assumption of

these investigations is that the separation zone is a result of the inability of the axial flow

entering the volute inlet (see Figure 1.4) to make the sharp 90° turn to a radial direction
toward the impeller inlet. The current investigation has shown the active width at the out-
let of the impeller is directly related to the double vortex structures within the volute. It
may be reasonable to assume from these observations that the flow structure within the
volute also affects the active width at the inlet of the impeller and that the separation zone
at the inlet of the impeller affects the active width at the outlet of the impeller. That is,
there may be a coupled effect of the flow structures at the inlet and outlet of the impeller
which effect the use of the impeller width at both radial locations. Further experiments
should be conducted at multiple azimuthal locations which investigates the correlation of
the active widths between the impeller inlet and outlet to see if these structures do indeed
have a coupled effect.

Eck [7] suggests that the flow from the inlet of the volute (defined as "inlet" in
Figure 1.10) to the inlet of the impeller should be accelerated to reduce the extent over
which the separated region exists. For the current study, Figure 4.4 shows that the flow is
decelerated from the volute inlet to the impeller inlet by a factor of 1.72. This large decel-

eration would suggest that a substantial separation zone exists at the inlet of the impeller.

Hence, according to Eek, the impeller width should be reduced to below 1.72'I of the cur-

rent width in order combat the separation at the inlet.

131

5.4 Separated flow and Recirculation

By observing the 3000 rpm results of Figure 4.7 the limiting streamline patterns

abruptly end at the indicated bifurcation linesI before they reach the sides of the scrolled
surface. The streamline patterns then resume inside the bifurcation lines along the inlet
and rear surfaces, see Figures 4.8 through 4.11. These observations indicate the secondary
flow separates from the volute as it makes the turn from the scrolled surface to the inlet
and rear surfaces.

A confirmation of the flow separation at the comers of the volute, as indicated by
the streaking results, can be made by observing the streamline patterns of Figure 5.2. This
figure is a zoomed view of the PIV results from Region III showing the details of the flow
patterns near the comer of the scrolled and inlet surfaces. The streamline patterns show
the secondary flow separating from the scrolled surface as it approaches the inlet surface.
Also shown in this figure is the resulting recirculation zone within the boundary of the
separated flow. This recirculation is yet another source of losses associated with the sec-
ondary flow. It is noted here that the production blower had a slight fillet (i.e. radial cur-
vature) at the comers of the volute which may alter or have an effect on the position of the

separation point.

 

 

1. Bifurcation Line: A line which divides two regions of flow.

132

 

 

 

 

 

 

 

 

 

 

r40 3- Region III
I Q = 600 rpm

135 - _ _
E _ Recrrculatron
E 130 ’_ Stagnation Point Separation Point /
i... C I I

125 .7- 1 g - g f /

nfif \ x

l l l I I l l l L l
40 60
2 (mm)

FIGURE 5.2 PIV: loomed View of Region III z-r Plane. The Streamlines Reveal a
Separation of the Secondary Flow and a Region of Recirculation.

The existence of the separation is a result of the abrupt change in geometry as the
flow moves along the scrolled surface to the inlet and rear surfaces. In order to follow the
geometry at the comer of the volute, the flow would be required to turn at an increasingly

smaller radius of curvature and hence an increasingly larger pressure gradient. That is,
. . . 6P
according to the Euler-n equatron (Equation 5.1), for V i 0, 5; —-> 00 as R -) 0. The

flow field responds to this "unreasonable" requirement by forming its own boundary for
which the required pressure gradients are available [25].

The bifurcation lines of Figure 4.7 also indicate a poor transition from the end of
the scrolled surface 0 = 290° to the outlet of the volute. It has already been established

that the bifurcation lines indicate the point of separation (or reattachment) of the flow

from the‘ surface of the volute. As seen in this figure the bifurcation lines along the sides

of the scrolled surface join together just after 0 = 290°. This indicates the flow in the

133

primary flow direction separates from the scrolled surface at the volute outlet which sug-

gests poor diffusive properties at the outlet section of the blower.

5.5 Jet-Wake

Upon inspection of the phase averaged velocity magnitudes of Figures 4.58
through 4.60 one can clearly see the existence of the "jet-wake" flow pattern at the exit of
the impeller. The jet-wake structures are shown to consist of a high velocity "jet" near the
blade pressure surface and a low velocity "wake" near the blade suction surface. The
observations that were made on the time averaged quantities concerning the axial and azi-
muthal velocity distributions and streamline patterns can also be made for these phase
averaged quantities. Therefore, only observations specific to the phase averaged quanti-
ties along the mid-plane are presented in this section.

As seen in the mid-plane results of Figure 4.59, the profiles of the jet-wake struc-
tures change with azimuthal location. The first observation is the extent to which the high
velocity jet structure fills the blade gap in each Region. As seen in this figure, Region I
fills the largest area of the blade gap, followed by Region III, and finally Region II. The
second observation is the extent to which the jet structures extend into the volute. This
measure follows the same order as before with the jet structures extending the furthest into
the volute in Region I, followed by Region III, and finally Region 11. These observations
reveal that the flow in the mid-plane of Region I is the strongest and most uniform fol-
lowed by Region 111. That is, along the mid-plane of Region I, the flow through the blade
gaps has the highest flow rate and the flattest distribution.

In a study by Kawahashi et a1. [15], PIV measurements were acquired to analyze

the flow through blade passage of a FC automotive HVAC blower. The results of that

134

study showed the existence of a separation bubble on the suction side of the blade. The
separation bubble was formed by flow separation at the leading edge of the blade and reat-
tachment near the trailing edge. The size of the separation bubble was shown to be a fianc-
tion of both azimuthal and axial location. This indicates that the flow angle at the inlet of
the impeller is also a function of both azimuthal and axial location, since separation at the
leading edge is due to poor inlet flow angles (i.e. high angles of attack). The separation
bubble was also shown to have a direct impact on the flow through the blade passage.
Specifically, a large separation bubble resulted in low velOcity magnitudes through the
blade passage and large wakes, i.e low "blade gap filling" by the jet structure. Kawahashi
et al. also noted the unsteadiness of the separation bubbles and considered the fluctuations
of these to be a fundamental mechanism of noise. These results indicate Region II of the
current study has considerable blade passage blockage due to a relatively large separation
bubbles on the suction side of the blades and therefore high angles of attack at the blade
inlets. These results also indicate Region 11 may have considerable contributions to the
turbulence of the blower due to the unsteady nature of the separation bubble.

The unsteady separation bubble may also have a direct effect on the velocity fluc-
tuation levels of the flow through the impeller gaps. This observation has been inferred
based on a comparison of the phase mean values to the phase nns values of Figure 4.25c.
Upon closer inspection of this figure, see Figure 5.3, it is seen that the highest rrns levels
are along the rising edge of the mean values, while the lowest rrns values are along the
falling edge of the mean values. Given that the flow through an impeller gap is bounded
by the pressure side of an impeller blade and (based on the results by Kawahashi et al.

[15]) an unsteady separation bubble along the suction side of a blades, it is reasonable to

135

assume that the high mis values along the rising edge of the mean jet velocity profile are
associated with the shear layer and possibly the fluctuations of the separation bubble. The
global maximum in the rms values are shown to be along the rising edge of the mean val-
ues at about 60% of the minimum-to-maximum velocity difference, while the global min-
imum in rrns values lies near the minimum mean values. The rms values also exhibit a
local minimum near the peak of the mean values followed by a local maximum at about
60% of the min-to-max variation along the falling edge of the mean values. The observa-
tion concerning the global minimum of the rrns values lying near the minimum mean val-
ues (i.e. in the wake region) is in direct contrast to the results of Raj [26]. In that study,
Raj reported the wake fluctuations to be high compared to those of the jet region fluctua-

tions. No direct reasoning has been made for these conflicting results.

50 7 7 7

/\

45'

1240"
E ,
:35.1
m 1
C:
60

 
  

<U>

*— N N U.)
U‘ c “ft-O

Phase Averaged Si

9

 

 

 

 

 

5 1 <U>

0 - .- . . .-. - . ..
40 45 50 55
0’ (degrees)

FIGURE 5.3 Sample Phase Averaged Quantities from Figure 4.25c.

136

The jet-wake structures are undesirable since their large velocity gradients add sig-
nificantly to the turbulence levels and therefore the noise levels of the HVAC system.
This non-uniform velocity profile also provides excessive kinetic energy beyond that
associated with the desired mass flux. By dissipating this excess kinetic energy before the
air flow interacts with the downstream surfaces (i.e volute surfaces and evaporator) the
flow will have obtained the maximum exchange of momentum from the jet structures.
That is, the flow will have obtained the maximum amount of flow energy without further
loss to dissipation at the surfaces. Hence, a high rate of dissipation or mixing into the bulk
flow of the volute is desirable. A measure of the rate at which the jet-wake structures dis-
sipate in the volute is given in Figures 4.33 and 4.34 for the 600 and 3000 rpm speeds,
respectively. A comparison of the two impeller speeds tested shows good agreement,
however, the 3000 rpm results show slightly higher values. The results show that rms val-
ues (as given in Equation 4.25), which describe the spatial distribution of the jet-wake pro-

files, dissipate rapidly within the volute. Specifically. the rms value of these profiles fall

off exponentially in the radial direction to 12.5% of the highest contour level at roughly2 r/

r2 = 1.28. In a study by Miner et al. [21], laser velocimeter measurements of the flow field
within the volute of a centrifugal flow pump were acquired and analyzed. In that study
they reported that the jet-wake structures were also dissipated rapidly into the volute and
were absent for values of r/r2 = 1.2 and higher. Although the Miner et al. study was per-
formed on a centrifugal pump, their result is in good agreement with the centrifugal

blower results of present study.

 

2. Based on the average radial location ofo /Ut = I251: — ()2 at 0 = 90°, 180°, 270°.

phase

137

5.6 Flow Field Fluctuation Intensity

In general, turbulent fluctuations have unwanted effects on the performance char-
acteristics of turbo-machinery. Two of the unwanted effects are energy dissipation (i.e.
losses) and broadband noise generation. Energy dissipation is generally accepted to take

place at the end of the energy cascade. This energy cascade is described as the successive

transfer of turbulent kinetic energy to smaller and smaller eddies3 until, at the smallest
scales, viscosity is effective in dissipating the kinetic energy into thermal energy. Aeroa-

coustic broadband noise generation has no universal theory on its production but it is gen-

erally accepted to be generated by vertical motions4 (i.e. eddies) in the flow field. It is
therefore the objective of this section to identify regions of high turbulence levels within
the volute.

A first measure of the turbulence levels within the volute is given by the mid-plane
time averaged rrns velocity contours in Figures 4.18 and 4.19. These rms values are of the
order of 13% of the mean values, see Figures 5.4 and 5.5. Three distinctive regions of
high rms values exist in Figures 4.18 and 4.19. The first region is the "nearly" axisymmet-
ric band that surrounds the impeller (A81). The second region is the tongue-like band
(TLB) that spans from approximately entry port b to entry port (1 along the mid-height of

the volute between the impeller and scrolled surface. The TLB is clearly defined by the

isoline at 13* = 0.16 for the 600 rpm condition (Figure 4.18), however, at 3000 rpm

(Figure 4.19) the TBL is not as clearly defined. Although a clearly defined TLB is not

 

3. An eddy has no precise definition, but is conceived to be coherent turbulent motion localized within a
region of size L. A region occupied by a large eddy can also contain smaller eddies.

4. Based on Sir James Lighthill’s theory on the mechanisms of aerodynamic sound generation.

138

present at 3000 rpm, high rms are observed in this region and shall be referred to as the
TLB. The third region is the thin band (ThB) near the scrolled surface that starts from
roughly entry port d and ends between entry ports f and g. Although these time averaged
rrns values give an indication of the turbulence levels within the volute, care must be used
when assessing them since they also give a measure of the periodic blade passing fre-
quency flow fluctuations from the impeller.

In order to account for the inability of the time averaged rms values to distinguish
between turbulent and periodic fluctuations, one must inspect the spectral contributions to
the fluctuations over specified frequency bands. Figures 4.21 and 4.22 reveal that the
TLB and Th8 have their major contributions from turbulent fluctuations, while
Figures 4.23 and 4.24 reveal that the ABI has its major contributions from the periodic
fluctuations of the flow exiting from the impeller.

It is interesting to note that the TLB falls within Region II, where the phase aver-
aged PIV results of Figure 4.59 showed a considerable wake at the exit of the impeller. It
had been inferred in Section 5.5 that the large wake regions were caused by large fluctuat-
ing separation bubbles on the suction side of the blades and thus had considerable contri-
butions to the turbulence of the blower. This observation suggests that relatively poor
impellerflow conditions exist along this TLB. It is also interesting to note that the peri-
odic contributions to the flow field’s fluctuation intensity (Figures 4.23 and 4.24) fall off
exponentially in the radial direction. This observation shows a high correlation with the
rate at which the jet-wake structures dissipate in the volute, see Figures 4.33 and 4.34.

The fluctuating pressure results, see Figure 4.38, show that the fluctuation levels

start off low near entry port a, increase to maximum levels around entry ports (1 through f,

139

and then fall off toward entry port h. The fluctuating pressure results between entry ports
(I through f are in good agreement with the ThB of high broadband fluctuating velocities
observed in Figures 4.21 and 4.22. This suggests that these pressure fluctuations are
caused by the flow field turbulence and not the periodic flow fluctuations associated with
the impeller. Upon inspection of the pressure spectra of entry port d (Figure 4.41), the
periodic fluctuations (at F *=41) are shown to indeed have little contribution to the overall
fluctuating content. The periodic fluctuations at entry ports a and h, see Figures 4.40
and 4.42, are shown to be much more evident than at entry port (1 but the major contribu-

tions to the overall fluctuating content is still dominated by the broadband components.

100 —

50'—

y (mm)

   

 

 

O - 3.
h
: Impeller
_50 1 l J 1 i 1
-150 -100 -50 0 50 100 150

x (mm)

FIGURE 5.4 Hot Wire: Ratio of rms to Mean Velocities: 600rpm.

140

: r2 =3000 rpm

1' (mm)

 

 

 

x (mm)

FIGURE 5.5 Hot Wire: Ratio of rms to Mean Velocities: 3000rpm.

141

6.0 Summary and Conclusions

The flow field within the volute of an automotive HVAC blower has been mea-
sured using multiple experimental techniques. The following observations and conclu-
sions, based on the statistical analysis of the acquired data, are made on the basis of these
measurements.

1. The direction from which flow is ducted to the volute inlet impacts the perfor-
mance of the blower. It is speculated that the direction in which the flow enters the volute
inlet effects the flows angle of attack at the inlet of the impeller blades. The angle of
attack is an important parameter as it effects the extent to which separation occurs on the
suction side of the blades and thus the mass flux through the blade gaps. From measure-
ments of the flow field within the volute, it has been inferred that the angle of attack is a
function of both axial and azimuthal location. Hence, an optimum inlet flow direction is
expected to exist which least negatively effects the performance of the blower.

2. The flow within the volute is highly three-dimensional. A secondary flow
develops within the volute that is superimposed on the primary flow. The secondary flow
is a pair of counter rotating vorticies (commonly called the "double vortex") which exhib-
its velocity magnitudes up to 40% of that of the primary flow. The existence of the double
vortex is believed to be induced by the curvature of the volute flow path but it is "ener-
gized" by the high momentum fluid exhausting from the impeller mid-plane. This second-
ary flow is seen to be a significant source of excess kinetic energy whose consequent
dissipation will contribute to the overall losses in this flow system.

3. The contribution to the flow from the impeller is the highest near the mid-plane

and drops off as the plane of observation moves axially toward the inlet and rear surfaces

142

of the volute. In fact, depending on azimuthal location, flow re-circulation (i.e. flow back

into the impeller or fir < 0) has been observed near these outer surfaces of the volute.

This observation shows that only a portion of the impeller width is active in adding mass

to the volute. If the secondary flow velocities of the double vortex result in ur < 0 at r=r2,

then the effective width of the impeller is less than the full impeller width. This phenom-
enon was observed in Regions 1 and 11, see Figures 4.52 through 4.55.

4. The double vortex structure has three other negative effects on the performance
characteristics of the blower. The first effect was directly observed. It is the flow separa-
tion along the comers of the volute where the scrolled surface joins the inlet and hub sur-
faces. The separation along these comers cause regions of recirculation which add to the

losses of the blower. The second and third effects were inferred based on the re-entrance
(a, < 0) of flow into the impeller. The impeller blades would likely have an "egg-beater"

effect on the flow re-entering the impeller and thus dissipate the energy of the flow and
produce broadband noise.

5. Phase resolved measurements show the presence of a periodic jet-wake struc-
ture at the exit of the impeller. These jet-wake structures exhibit large velocity gradients
that "die-out" exponentially (see Equation 4.22) in the radial direction. The largest fluctu-
ations associated with the jet-wake structures were found to be along the rising edges of
jet profiles (i.e. near the suction surface of the impeller blades). A literature reference [15]
gave the insight that these large fluctuations are due to the fluctuating nature of the separa-

tion on the suction side of the impeller blades.

143

The size and strength of the jet region was found to be a function of azimuthal
location. Here, size is defined as the extent to which the jet structure fills the blade gap
and strength is defined as the extent to which the jet extends into the volute. The size of
the jet was correlated to the jet strength, where a relatively large jet also had a relatively
high strength. It has been inferred that the size and strength of the jet is dependant on the
amount of separation on the blade suction surface, where a large separation results in a
small-weak jet.

6. The flow within the volute has significant fluctuating velocity levels with rms
velocities on the order of 13% of the mean velocities. The fluctuating components of
velocity were decomposed to distinguish between the turbulent and the periodic contribu-
tions. The periodic contributions (caused by the periodic nature of the jet-wake structures)
were shown to be dominant near the impeller and to "die-out" exponentially (see
Equation 4.22) in the radial direction. A "tongue-like" band of high turbulence level fluid
was observed in a region where large wakes (i.e. small jets) were present. It is inferred
that this band of high turbulence is a result of large fluctuating separation bubbles along
the suction sides of the blades in this region. This suggests that the turbulence levels
within the volute have major contributions from the flow through the impeller. This also
suggests a possible region of significant broadband noise generation.

7. Results acquired at multiple impeller speeds indicate that the flow field has a

weak Reynolds number dependence and that it scales on blade tip speed (Ut). This con-

clusion was made by the comparison of the normalized results which showed good agree-

ment at all impeller speeds tested. Thus, analytical and computational investigations of

144

this (or a similar) blower could be carried out at a given impeller speed and scaled by Ut to

obtain the flow properties at any other desired impeller speed with reasonable reliability.
8. The results of this study have shown a strong impeller\volute interaction. That
is, the flow field within the volute strongly effects the flow through the impeller and the
flow through the impeller strongly effects the flow field within the volute. However, a
feature which seems to dominant the negative characteristics of both the flow within the
volute and the flow through the impeller appears to be a fundamental property of the
volute flow field. This feature is that of the double vortex. This observation suggests that
changes in the volute housing which limit the strength of the double vortex motions could
prove to be a significant step toward a high performance, low noise blower. It is. recog-

nized that this may require a complete redesign of the housing.

145

TABLE A.1 Operating Conditions of the Instrument Panel (Installed Condition) at

APPENDIX A

the Highest Blower Control Setting.

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Mode T825313?” APE (pa) Q (m3/min) rpm
Panel Lo-Lo 90.27 6.181 3076
Lo-Hi 84.21 5.375 3157

Hi-Hi 72.17 4.184 3335

Hi-Lo 78.26 4.762 3220

Bi-Level Lo-Lo 90.84 6.274 3074
1.0-Hi 83.64 5.363 3172

Hi-Hi 70.81 4.138 3346

Hi-Lo 78.99 4.937 3236

Floor Lo-Lo 75.74 4.220 3355
Lo-Hi 64.03 3.681 3400

Hi-Hi 54.98 3.299 3473

Hi-Lo 64.97 3.763 3447

Defrost/Floor Lo-Lo 81.78 4.863 3236
Lo-Hi 73.73 4.156 3303

Hi-Hi 58.83 3.516 3402

Hi-Lo 73.37 4.229 3292

Defrost Lo-Lo 77.80 4.387 3221
Lo-Hi 68.55 3.899 3310

Hi-Hi 55.74 3.346 3395

Hi-Lo 67.04 3.879 3298

 

 

146

 

 

APE (Pa)

90

80

70

60

50

40

30

20

 

Temperature Setting Lo-Lol

 

  

-- -,

“" Bi-Level

 

 

 

' / r ‘ Floor 3
A, '/ - "7"“ Defrost/Floor» 1
‘e'Defrost

1

1 3 4 5 6

Q (mi/min)

FIGURE A.1 Installed Condition Evaporator Load Curves (Temperature Setting

APE (Pa)

90

80

70

60

50

40

3O

20

 

Temperature Setting Lo-Hil

 

Lo-Lo).

 

" Bi-Level
Floor
"7 Defrost/F loor
‘9‘ Defrost

 

1

 

3 4 5 6

Q (mi/min)

FIGURE A.2 Installed Condition Evaporator Load Curves (Temperature Setting

Lo-Hi).

147

90 - Temperature Setting Hi-Hil . #777 i

 

7. _ A /
6, - i . ,/ _ _

5. . w /

3O 7 / + Panel
/ ‘- Bi-Levcl f
20 f ‘ ‘ Floor
1“ " Defrost/Floor 1
10 0’) I 1 1 ‘9‘ Defrost 71

O __ ...--_.i__._... 7
0 l 2 3 4 5 6

APE (Pa)

 

 

 

 

 

Q (1113/min)

FIGURE A.3 Installed Condition Evaporator Load Curves (Temperature Setting

 

 

 

 

 

 

 

Iii-Hi).
90 - Temperature Setting Hi-Lol -
80 - 5
i .
70 l t
60 . 77 71
a 1 '
9.": 50 ‘ 7| -
LI
0.
<1 40
30 + Panel
“'— Bi-Level
20 Floor 1
"““‘ Defrost/Floor 1
10 '9’ Defrost
T
0 7 7 5
0 l 2 3 4 5 6

Q (ml/min)

FIGURE A.4 Installed Condition Evaporator Load Curves (Temperature Setting
Hi-Lo).

148

y (mm)

y (mm)

L...

—

100

50

 

APPENDIX B

Inlet Plane ~
Q = 600 rpm ux / Ut
0.20
0.15
0.10
0.05

0.00

 

Impeller

 

 

 

 

  
 
 
 
  

 

 

 

 

_50 l l l I l l l 1 l I
-150 -100 -50 0 50 100 150
x (mm)
FIGURE B.1 Normalized ux rms Velocity Contour: Inlet Plane 600 rpm.
Mid Plane ~
1 Q = 600 rpm ux / Ut
— 0.20
100 — 0.15
— 0.10
_ 0.05
0.00
50 —
0 — --1
7 h
1—
p
_50 l a l l_l
-150 -100 -50 0 50 100 150
' x (mm)

FIGURE B.2 Normalized ux rms Velocity Contour: Mid Plane 600 rpm.

149

100

Rear Plane
Q = 600 rpm

   

 

 

 

 

  

 

E
E 50
>‘ U.
0
-50 I
-150 -100 -50 0 50 100 150
x(mm)
FIGURE B.3 Normalized ux rms Velocity Contour: Rear Plane 600 rpm.
Inlet Plane Hy /Ur
_ - 0.20
100 - 0.15
— 0.10
- 0.05
r 0.00
e ‘_
E, 50
e i
0 —
_ h
‘ Impeller
-50 — 1_1 14 1 1 I
-150 ~100 -50 0 50 100 150
x(mm)

FIGURE B.4 Normalized uy rms Velocity Contour: Inlet Plane 600 rpm.

150

y (mm)

y (mm)

Mid Plane , -" ~
(2 = 600 rpm “y / U:

 

 

 

FIGURE B.5 Normalized uy rrns Velocity Contour: Mid Plane 600 rpm.

Rear Plane

    

100 —

50' —

 

750 ’
-150 7100 750 0 50 100 150

FIGURE B.6 Normalized uy rms Velocity Contour: Rear Plane 600 rpm.

151

y(mm)

   
     

 

 

 

 

 

 

- Inlet Plane _r_T_/ U2
- Q = 600 rpm u xu y t
— 0.010
100 _ 0.005
_ 0.000
_ -0.005
— -0.010
50 -
0-—-1
f" h
’ Impeller
_50 Pl 1 1 l l L l l l J l l |
-150 -100

 

FIGURE B.7 Normalized Reynolds Shear Stress Contour: Inlet Plane 600 rpm.

100

50

y(mm)

-50

-150

Mid Plane

- —T—. 2
u xu y / Ut
0.010
0.005
0.000
-0.005

-0.010

 

I
° .‘ca.
, ., '°-.'.'
0

 

 

 

 

 

 

T h
” Impeller
1 I 1 1 i 1 1 I
-100 -50 0 50 100
x (mm)

FIGURE B.8 Normalized Reynolds Shear Stress Contour: Mid Plane 600 rpm.

152

1 Rear Plane
Q = 600 rpm

 
   

100

7.x: .7:

50

y (mm)

 

x(mm)

FIGURE B.9 Normalized Reynolds Shear Stress Contour: Rear Plane 600 rpm.

140

 

 

 

 

 

 

 

:—ReionI
ZQ=600 m
130T 111 I 1
120:— l I I
E ' I '
110_
a E
5100:
.. :
90:—
80:-
70:-
E 1_1_I Ii 1 11 1 I r I
60 1141_L lj_ll 1411 IillllllIJllllllrll
010 20 30 40 50 60 70 80
z(mm)

FIGURE B.10 Normalized uz rms Velocity Contour: Region I 600 rpm.

153

 

 

 

 

 

 

 

E. Region II
: Q = 600 m
120 E- l | I
110 :— E /Ut
E I i
8100 .- 0.16
7: - 0.14
: 0.12
90 :— 0.10
: 0.08
80 _—
70 :—
60 _1 1

0 10

FIGURE B.11 Normalized uz rrns Velocity Contour: Region II 600 rpm.

140

 

 

 

 

 

F Region III
: Q = 600 m
.30 .- r‘1 . a
120 :-
AllO ;- fiz/Ui
E : ;
E100 :- V" 31146;
H 90 Z_ " 0.12
: 0.10
: 0.08
80 _—
70 :— 1 | l
E I
60 I I I l I I 1 II I I I l 1 I | l l I

 

 

 

O 20

40 60 80
2 (mm)

FIGURE B.12 Normalized uz rms Velocity Contour: Region HI 600 rpm.

154

Region I
Q = 600 rqm

.—
A
O

Irrrlrrrrln—i—WT

 

  

 

llllllllllllllllllllllll

 

 

III-11111 IJJIIIIIIIIIIlllIlIIlJllllIlllIIll
0 10 20 3O 40 50 60 70 80
2 (mm)

 

O\
0

FIGURE B.13 Normalized ur rms Velocity Contour: Region I 600 rpm.

7? Region II
T Q = 600 rpm

...1
b)
O

111'

 

11111111'1IIIIIIIIIIITI‘II

 

 

\l
O
llll

I

 

 

I l I | I 1 I I I I I I I I I l 11 I I I I I I I I I I J I 1 I 1 I I ' L
0 10 20 30 40 50 60 70 80
2 (mm)

0
O

 

FIGURE B.14 Normalized ur rms Velocity Contour: Region II 600 rpm.

155

Region 111
Q = 600 rqm

 

 

 

    

 

_l|IlllllIllIlllllIIllllIllIlllllllllllll

 

 

I I I I I I I I I 1 I I I I I J l I
O 20 40 6O 80
2 (mm)

 

FIGURE B.15 Normalized ur rms Velocity Contour: Region III 600 rpm.

140 .
e ron I

= 600 ram

.—n
(a)
O
Illllllllllllll

 

 

IIIIIIlllllllllllIllllll

 

 

 

1LLIIIIILIJ_IIllllIIIIIIILLIIIJIIIlIIIllIII
0 10 20 30 40 50 6O 70 80

05
O

 

FIGURE B.16 Normalized Reynolds Shear Stress Contour: Region I 600 rpm.

156

 

 

 

r(mm)
5
O
_IIII‘IIIII'IIII'IIIII’IIIllllll'IllllllIIT

I I IIIIIIIliJ-IIIIIIIIIIIIIIIIIIJIIIIIIIIIIII
0 10 20 30 40 50 60 70 80
z(mm)

 

 

 

FIGURE B.17 Normalized Reynolds Shear Stress Contour: Region II 600 rpm.

Region III
- Q = 600 rqm

 

      

 

 

r(mm)
—‘
8 o
_llllllllllllll'llllllllITIIlllll11'

 

 

90
80
7o 1 | 1
60 I l I L I I I l I I I I I 1 J l l I l I
0 20 40 60 80
2 (mm)

FIGURE B.18 Normalized Reynolds Shear Stress Contour: Region III 600 rpm.

157

DE?
"03.
O
S
D—t

 

 

    

lllllllIIIIlllllllllIllI IITIIIIIIIIIIII

lIlIIIIIJ-IlllIIIIJIIILIIIIIIIIJIIIIIIILJ
0 10 20 30 40 50 60 7O 80
z(mm)

I

 

 

 

 

0‘
0

FIGURE B.19 Normalized uz rms Velocity Contourz: Region I 1200 rpm.

 

 

 

4 - .
l 0: Region II
: Q=1200 m
120 :— 1 | I
110 :' u /U‘
E100 :_ 0.16
r : 3'13
90 :— 0.10
: 0.08
80 _-
70 {-
E I
6011 J lIIIIIJIIIIIJ-lIIIIIIIIIIIIIIIII-l-LIIIIIIII

 

 

 

0 10 20 30 4O 50 60 70 80
2 (mm)

FIGURE B.20 Normalized uz rms Velocity Contour: Region II 1200 rpm.

158

0 7 Region III

130:_Q=12001‘pm

 

   

 

 

70’. 1 I I

 

 

60 "1 1 I 1 I 1 | 1 11 I 1 . I 1 17 1 1 I 1
0 20 40 60 80
2 (mm)

 

FIGURE B.21 Normalized uz rms Velocity Contour: Region III 1200 rpm.

 

 

 

140— .
: Re ronI
: Q=1200 m
I3 _- 1.1) I 1
120:— I I I
E 1 1 1
A110: ur/U‘
E : .
51007 I: 0.16
7 : -- 31:
9°:— 1 0.10
3 0.08
80:-
70:—
E I
6011llIIIlllIlIIIIIIlIlIlllllllllJlllllIlIJlll

 

 

 

 

0 10 20 30 40 50 6O 70 80
2 (mm)

FIGURE B.22 Normalized ur rms Velocity Contour: Region I 1200 rpm.

159

 

 

 

r40- .
i Region II
: Q=1200 m
120:— I 1 I
110:— fir/Ut
E : . .
8100,— . 0.16
T: : 0.14
7 0.12
90:” 0.10
: 0.08
80_—
70:—
E I
6041-1llIllllIIIlIIIIIIIIIIIIIIIIIJIIIIIIIIIII

 

 

 

0 10 20 30 40 50 60 70 80
2 (mm)

FIGURE B.23 Normalized ur rms Velocity Contour: Region II 1200 rpm.

140 e ion III

= 1200 1pm

107°

 

fir/U,

 

-—o
—I
O
IllllllllllllfIll'llIlllIlllIllll’

 

 

 

O\ \l
O O
__lllllll

 

 

I I I I I 1 I I I J l I 1 l I l
0 20 40 60 80
2 (mm)

 

FIGURE B.24 Normalized ur rms Velocity Contour: Region III 1200 rpm.

160

Region I
Q =1200 rpm

H

b)

C
lllI'lIlI'lllll

 

 

 

llll'lllIlIllllIlIIlllIl

 

 

1 I I l I III I l l I l I I I III I 1 l I | I I I I I J l 1 I 1 I I I I l I LI
0 10 20 30 40 50 6O 70 80
z (m)

 

O\
O

FIGURE B.25 Normalized Reynolds Shear Stress Contour: Region I 1200 rpm.

 

 

 

 

140 — .
: Region II
: Q = 1200 m
120 :— I l I
110 3- “_ 2
A : u'zu'r / Ut
a :
8100 r 0.008
:4 - 0.004
: 0.000
90 3' 0004
: ~0.008
80 :-
70 :- I | I
h l
60 I I I l l I l l l l I l I I l [J l LII I I I I IJ I I I l I l I I I I I 1 I I

 

 

 

0 10 20 30 40 50 60 70 80
2 (mm)

FIGURE B.26 Normalized Reynolds Shear Stress Contour: Region II 1200 rpm.

161

 

 

 

 

 

 

 

 

 

I; Region III
: Q = 1200 m

130 I‘ Ip I

120 f—

110; __u'zu'r/Uf
E : I
awo- V 333:
H 90 :_ . I 0.000

: -0.004
; -0.008

80 _-

70 :-

60 :I I I II I I | I II I ‘ I I I J I I I I I_L

0 20 40 60 80
2 (mm)

FIGURE B.27 Normalized Reynolds Shear Stress Contour: Region III 1200 rpm.

E/Inggane '5: , A \\

y (mm

 

 

 

-75 -50 -25 0 25 50 75
x (mm)

FIGURE B.28 Normalized Phase Averaged ux rms Velocity Contour: Inlet Plane
600 rpm.

162

   

 

 

 

 

 

 
 
     

 

 

l
r.
A 50 -
:1 ..........
25 ........
O
-75 -50 -25 0 25 50 75
x (mm)

FIGURE B.29 Normalized Phase Averaged ux rms Velocity Contour: Mid Plane
‘ 600 rpm.

* //
C R r Plane
- Q = 600 rpm

 

 

 

 

 

 

y (mm)

 

  

IIIIIIIIIIIIIIIIIIIIIIJ.
-100 -75 -50 -25 O 25 50 75 100
x(mm)

  

FIGURE B.30 Normalized Phase Averaged ux rms Velocity Contour: Rear Plane
600 rpm.

163

//

Inlet Plane
Q = 600 rpm

 
 

75

50

llllllll

 
 
     

y (mm

h..

25 ......

 

 

-75 -50 -25 O 25 50 75

FIGURE B.3] Normalized Phase Averaged uy rms Velocity Contour: Inlet Plane
600 rpm.

// I
Mid Plane .
Q = 600 rpm

50

IIITIII

 
   

y (mm)

25 ......

 

  

IIIIIIJIIIIIIIIIIIIII -.
-50 -25 O 25 50 75

X (min)

FIGURE B.32 Normalized Phase Averaged uy rms Velocity Contour: Mid Plane
600 rpm.

164

//

Rear Plane
Q = 600 rpm

    
  

75

 

 

0.20
0.18
" 0.16
0.14
0.12

 

 

 

IIIIIIIVII

 

 

   

IIIIII IIIIIIIIIIIIIILIIIIII .
-1 OO -75 -50 -25 O 25 50 75 100
x(mm)

FIGURE B.33 Normalized Phase Averaged uy rms Velocity Contour: Rear Plane
600 rpm.

-
0
. 0
u
n
o
I A ‘
.. .. ..
Z '72- t; . ,
- pg. '1. ~' 2. .
. v, “I ."'.'.‘.
. '4‘“ . ‘I ‘ '.
a .‘ ..: I f. ,
‘ 0" [5
U . ,I . ‘

Inlet Plane
Q = 600 rpm

   
 

75

 

 

0.01
0.01
0.00

-0.00

. -001

25 ........... . -0.01

 

 

 

50

IIIIIIIW

y (mm)

M

 

 

-75 -50 -25 O 25 50 75
x (mm)

FIGURE B.34 Normalized Phase Averaged Reynolds Shear Stress Contour: Inlet
Plane 600 rpm.

165

//
Mid Plane

Q = 600 rpm

 
    

75

 

 

 

 

 

50

 
    

Fill'ITIT

y (mm)

25

......

 

     

I ' '
I» g, .’

 

 

D/' . l
W11111111111..11II..11 1
-75 -50 -25 O 25 50 75

X (turn)

FIGURE B.35 Normalized Phase Averaged Reynolds Shear Stress Contour: Mid

 

 

 

 

 

 

-0.00

-0.01
’ -0.01

Plane 600 rpm.

E // ‘, ' ~

_ Rear Plane ‘_ 41‘ .

— o = 600 \ .
75 / rpm " . ' / ”12>

_ h 1 w 1

_ 1 '.) 0.01

a 0.01
50 :- ’ 0.00

I—

y (mm

25 ................

 

        

1111I1 I1111l111111111|1111l11
-100 —75 -50 -25 0 25 50 75 100
x(mm)

FIGURE B.36 Normalized Phase Averaged Reynolds Shear Stress Contour: Rear
Plane 600 rpm.

166

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170

IIIIIIIIIIIIIIIIIIIIIIIIIIIIIII

11][1121]]1]12[]11[[1[111