. -me ,
y. e .1...“ J.
$3 9. 5......

.....L

. .
a... .3
. 2. ..
gum... .
.._:..m.
gm»...

 

 

2.3.5.. flux“:
7 5 .i
2. .
mi...“ 8...”... 4
n. . .
. 1:31; .. ., E...
This? 9......

5.3.9.

 

.
1.4!; .2.

3.5 ... .9.

a)... :P.I..~.. it!
2:. tJufir i:
I 1 4

. \.. J.
9...... n.

L
u...

If ,7:
.‘vvlv-Sv . z A"
.3... 5.)...

..

 

_ 9 . .E.

 

0*: Q 3;»

(3‘3”?!090/

LIBRARIES
MICHIGAN STATE UNIVERSITY
EAST LANSING, MICH 48824-1048

This is to certify that the
thesis entitled

AN EXPERIMENTAL INVESTIGATION OF THE
AERODYNAMIC SHROUD WITH AN OFF-HIGHWAY
ENGINE COOLING FAN

presented by

MICHAEL D. DUSEL

has been accepted towards fulfillment
of the requirements for the

MS. degree in

 

/ 'Major E’roiéssor’s Signature fl

— {It fit? I OS,—

Date

MSU is an Affinnative Action/Equal Opportunity Institution

Mechanical Engineering

__‘_v_ H'*" .—_.__
W

9--“

fl

__-'_—_

r O ~fi. Q'— ‘v... — v -.—~—‘-' 'V' v—
— .

-

"'—

v V"

- v v “v "'

- ‘0 v .-‘-v-“

v V’fi

 

PLACE 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

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

2/05 c:/CfiC/DateDue.lndd-p. 15

 

AX EX

SHRl

AN EXPERIMENTAL INVESTIGATION OF THE AERODYNAMIC

SHROUD WITH AN OFF -HIGHWAY ENGINE COOLING FAN

By

Michael D. Dusel

A THESIS

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

MASTER OF SCIENCE

Department of Mechanical Engineering

2005

‘\\

namic
techniq
mance ;
tion. 11
Plenum.

annular

Inleiga]
referent.
that the 1‘
Wed Io
ispmVid;
blex hoi
IOCallon a
Shroud at]

ABSTRACT

AN EXPERIMENTAL INVESTIGATION OF THE AERODYNAMIC SHROUD
WITH AN OFF -HIGHWAY ENGINE COOLING FAN

By
Michael D. Dusel

An experimental investigation of a large scale engine cooling fan with an aerody-
namic shroud has been conducted to assess the effectiveness of this active flow control
technique. The experimental study was motivated by the desire to increase the perfor-
mance and efficiency of this axial fan which is intended for use in an off-highway applica-
tion. The aerodynamic shroud utilizes a Coanda jet that is: i) delivered from a pressurized
plenum, and ii) introduced along the surface of the shroud upstream of the fan inlet. This
annular jet delivers high momentum fluid into the blade tip clearance region.

Experimental data were acquired to quantify the effect of the aerodynamic shroud.
Integral quantities including pressure rise, flow rate, and input power were measured for a
reference baseline condition and for several powered shroud conditions. The results show
that the inception of fan stall is delayed to a lower flow rate (higher pressure rise) as com-
pared to the baseline condition. Further insight into the effect of the aerodynamic shroud
is provided by detailed velocity measurements in the wake of the fan. Specifically, a dou-
ble-X hot-wire probe was used to resolve three components of velocity at a single axial
location and various radial locations. The phase averaged data show that the aerodynamic
shroud affects the “tip vortex” and the velocity field at the measurement location. Time

averaged and phase averaged in-plane PIV measurements support these data.

I
through
able. 1
Charles

funding

pie in ii
WOPlc
Laure:
Daniel
Lame

mate};

IUIUre

“'OUlc

ACKNOWLEDGEMENTS

I would like to thank my advisor, Dr. John Foss, for his guidance and support
throughout my graduate work. His dedication to his work and to his students is remark-
able. I would also like to thank my committee members, Dr. Ahmed Naguib and Dr.
Charles Petty for their valuable input during the review process. Caterpillar provided the
funding for this research and I am grateful for this as well.

I have been extremely fortunate to work with very intelligent and interesting peo-
ple in the Turbulent Shears Flow Laboratory at Michigan State University. Many of these
people contributed their time and effort during the course of this research including Al
Lawrenz, Doug Neal, Scott Treat, Kyle Bade, Richard Prevost, Aren Hellum, Amanda
Danielson, and Drew Butki. I am grateful for all their help. A special thanks to Al
Lawrenz and Doug Neal who took time to teach me things as an undergraduate that ulti-
mately lead to my interest in fluids research.

Finally, I would like to thank my family for their never ending support and my
future wife Sarah McCann for her support and motivation throughout this endeavour. I

would not have been able to accomplish what I did, had it not been for them.

iii

List of F:
List of T.
Nomencl

1.0 lntr

2.0 Exp

I)
H

I.)
I.)
I.) I.) I.) ~—<) -—4

I J
'JJ

IJ Ix)

C’J (’J

CA)
L!)
U) "U

TABLE OF CONTENTS

 

 

List of Figures ................................................................................................................... vi
List of Tables ................................................................................................................... xii
Nomenclature ................................................................................................................. xiii
1.0 Introduction -_ _- - - -- l
1.1 Motivation and Background Information ............................................................. 1

1.2 Previous Investigations of the Aerodynamic Shroud ........................................... 3

1.3 Overview of the Experimental Study ................................................................... 5

2.0 Experimental Apparatus ......................................................................................... 8
2.1 Test Fan ................................................................................................................ 8

2.2 The Axial Fan Research and Development (AFRD) Facility ............................... 8
2.2.1 AFRD Overview ......................................................................................... 8

2.2.2 Flow Rate Measurement System ................................................................ 9

2.2.2.1 Flow Meter Configuration and Measurement Technique ................ 9

2.2.2.2 Flow Meter Calibration .................................................................. 11

2.2.3 Hot-wire Traversing Mechanism .............................................................. 14

2.3 Experimental Aerodynamic Shroud ................................................................... 14
2.3.1 Prototype Design ...................................................................................... 14

2.3.2 Shroud Plenum Pressurization System ..................................................... 16

3.0 Measurement Equipment and Technique ..... -_ - - -- - "“24
3.1 Data Acquisition System .................................................................................... 24

3.2 Integral Measurements ........................................................................................ 25
3.2.1 Measurement Equipment .......................................................................... 25

3.2.1.1 Pressure Transducers ...................................................................... 25

3.2.1.2 Torque Measurement ...................................................................... 26

3.2.1.3 Optical Encoder .............................................................................. 27

3.2.2 Definition of Variables ............................................................................. 27

3.2.3 Data Acquisition and Procedure ............................................................... 31

3.3 Particle Image Velocimetry Measurements ........................................................ 33
3.3.1 PIV Equipment ......................................................................................... 33

iv

 

(J)
.3.-

 

 

 

 

 

 

 

» 4.0 Rt"

3.3.2 PIV Vector Processing ............................................................................. 34

 

 

 

3.3.3 Data Acquisition and Procedure ............................................................... 36

3.4 Hot-wire Measurements ...................................................................................... 38
3.4.1 Hot-wire Probes, Anemometers, and Thermistor ..................................... 39

3.4.2 Calibration Facility ................................................................................... 39

3.4.3 Calibration ................................................................................................ 40

3.4.4 Data Acquisition Procedure ...................................................................... 42

3.4.5 Double-X Hot-wire Processing ................................................................ 44

4.0 Results and Discussion - Integral Measurements .............. - - ...... 56
4.1 Baseline Measurements (Unpowered Shroud Condition) .................................. 56

4.2 Effect of Aerodynamic Shroud (Powered Shroud Conditions) .......................... 57

5.0 Detailed Velocity Measurements -- -- -- - . ............ - ............. -- 66
5.1 Hot-wire Measurements ...................................................................................... 67
5.1.1 Hot-wire Measurement Uncertainties ....................................................... 68

5.1.2 Phase Averaged Velocity Data ................................................................. 70

5.1.3 Phase Averaged Fluctuation Intensity (RMS) Data ................................. 72

5.1.4 Phase Averaged Vorticity Data ................................................................ 73

5.2 Particle Image Velocimetry Measurements ........................................................ 75
5.2.1 PIV Measurement Uncertainties ............................................................... 76

5.2.2 Time Averaged PIV Data ......................................................................... 77

5.2.3 Calculation of Flow Coefficient ............................................................... 81

5.2.4 Phase Averaged PIV Measurements ........................................................ 82

6.0 Summary and Conclusions -- -- - - ._...123
References ...... - ......... - - - _ -- 126

 

Figure .
Figure .'
Figure 3
Figure I
Figure 3
Figure 2
figure 2

Figure 2

F 1Sure 2.

Figure 3

Figure 1.1
Figure 2.1
Figure 2.2
Figure 2.3
Figure 2.4
Figure 2.5
Figure 2.6

Figure 2.7

Figure 2.8

Figure 2.9

Figure 3.1

Figure 3.2

Figure 3.3

Figure 3.4

Figure 3.5

Figure 3.6

Figure 3.7

Figure 3.8

LIST OF FIGURES

Schematic representation of the aerodynamic shroud concept ........................ 7
Prototype test fan (D = 0.711 m) ................................................................... 18
Schematic representation of the AFRD flow facility .................................... 18
Schematic representation of the flow metering device .................................. 19

Details of the force transducer for the flow metering device (Neal, 2002) l9
Calibration data and fit for the flow metering devices .................................. 20
Schematic representation of AFRD hot-wire traverse ................................... 21

Schematic representation of the aerodynamic shroud and plenum
configuration (Radiator not shown in top view) ............................................ 22

Radiator with static pressure tap downstream of the cooling fins ................. 23

Schematic representation of AFRD with aerodynamic shroud assembly and
pressurization system ..................................................................................... 23

Torque measurements acquired over a period of time with the fan removed
from the shaft to measure a tare value ........................................................... 48

Schematic representation of the pressure rise across the fan blades ............. 48

A schematic representation of the cross-correlation technique using in PIV
vector processing ........................................................................................... 49

A schematic representation of the interrogation window overlap used in PIV
vector processing ........................................................................................... 49

Schematic of PIV measurement regions relative to the experimental
configuration .................................................................................................. 50

TTL voltage signals from the optical encoders for PIV phase triggering taken
at a shaft speed of 20 rpm .............................................................................. 51

Illustration of the double-x hot-wire probe ................................................... 51

Hot-wire calibration facility (Top and side panel removed to show detail)..52

vi

Figure 3.9

Figure 3.10

Figure 3.11

Figure 3.13

Figure 3.13

Figure 3.14

Figure 3.15

F1glue 4.1
Figure 4.2

F1EUre 4.3

Figure 3.9

Figure 3.10

Figure 3.11

Figure 3.12

Figure 3.13

Figure 3.14

Figure 3.15

Figure 4.1
Figure 4.2

Figure 4.3

Figure 4.4

Figure 4.5

Figure 4.6

Figure 4.7
Figure 4.8

Figure 5.1

Figure 5.2

Front-view of the calibration facility. Two axes of rotation allow the hot-wire
probe to rotated in the vertical and horizontal plane ..................................... 52

A representative pressure signal from a hot-wire calibration. The high
standard deviation was reduced by fitting the curve with a 15th order

polynomial function ....................................................................................... 53

A representative image showing a fluctuating tuft. This tuft was used to find
the mean flow direction at a given radial location in the wake of the fan ..... 53

Schematic of hot-wire probe positioning device ........................................... 54

Histogram of flow angles measured by an improperly aligned hot-wire
probe .............................................................................................................. 54

Histogram of flow angles measured by a properly aligned hot-wire probe ..55

Intersection of the predicted velocities from the two wires in a single x-array
for use in the processing algorithm to find the measured velocity

and angle ........................................................................................................ 55
Baseline performance data for three independent tests ................................. 62
Baseline efficiency data for three independent tests ..................................... 62

Total system pressure rise plotted as a function of volume flow rate in
dimensionless form ........................................................................................ 63

Baseline characteristic curve showing the stall point and separate regions of
the characteristic curve .................................................................................. 63

Performance data for selected shroud conditions .......................................... 64

Two performance curves, a baseline condition and a representative powered
shroud condition, plotted with a low and high system resistance curve to

demonstrate the advantage of delaying stall .................................................. 64
Efficiency data for selected shroud conditions .............................................. 65
Performance data at showing the stall point hysteresis ................................ 65

Selected shroud conditions labeled with the operating points used for the
detailed velocity measurements (see Table 4.1 for numerical values) .......... 87

Mesh of hot-wire phase averaged data locations in r-q plane at z = 35mm
(z/Rtip=0.098), shows direction of fan rotation ............................................ 88

vii

Figure 5.3

Figure 5.4

Figure 5.8

fiwm59

Figure 5.11

F1Eure 5.13

Figure 5.3 Histograms illustrating the “sample by sample” difference between velocities
found using the pre and post calibration constants for Condition ‘A ............ 89

Figure 5.4 Histograms illustrating the “sample by sample” difference between velocities
found using the pre and post calibration constants for Condition ‘B’ ........... 90

Figure 5.5 Histograms illustrating the “sample by sample” difference between velocities
found using the pre and post calibration constants for Condition ‘C’ ........... 91

Figure 5.6 Plots of convergence for a representative data point acquired in the hot-wire

measurements ................................................................................................ 92
Figure 5.7 Phase averaged mean radial velocity for condition ‘A’ ................................ 93
Figure 5.8 Phase averaged mean radial velocity for condition ‘B’ ................................. 93
Figure 5.9 Phase averaged mean radial velocity for condition ‘C’ ................................. 94
Figure 5.10 Phase averaged mean tangential velocity for condition ‘A’ .......................... 94
Figure 5.11 Phase averaged mean tangential velocity for condition ‘B’ .......................... 95
Figure 5.12 Phase averaged mean tangential velocity for condition ‘C’ .......................... 95
Figure 5.13 Phase averaged mean axial velocity for condition ‘A’ .................................. 96
Figure 5.14 Phase averaged mean axial velocity for condition ‘B’ .................................. 96
Figure 5.15 Phase averaged mean axial velocity for condition ‘C’ .................................. 97
Figure 5.16 Phase averaged velocity magnitude for condition ‘A’ .................................. 97
Figure 5.17 Phase averaged velocity magnitude for condition ‘B’ .................................. 98
Figure 5.18 Phase averaged velocity magnitude for condition ‘C’ .................................. 98
Figure 5.19 Phase averaged RMS of radial velocity for condition ‘A’ ............................ 99
Figure 5.20 Phase averaged RMS of radial velocity for condition ‘B’ ............................ 99
Figure 5.21 Phase averaged RMS of radial velocity for condition ‘C’ .......................... 100
Figure 5.22 Phase averaged RMS of tangential velocity for condition ‘A’ ................... 100
Figure 5.23 Phase averaged RMS of tangential velocity for condition ‘B’ .................... 101

viii

Figure 5.3

Figure 5.2
Figure 5.:

Figure 5.2

Figure 5.24 Phase averaged RMS of tangential velocity for condition ‘C’ .................... 101

Figure 5.25 Phase averaged RMS of axial velocity for condition ‘A’ ........................... 102
Figure 5.26 Phase averaged RMS of axial velocity for condition ‘8’ ............................ 102
Figure 5.27 Phase averaged RMS of axial velocity for condition ‘C’ ............................ 103
Figure 5.28 Phase averaged axial vorticity for condition ‘A’ ......................................... 103
Figure 5.29 Phase averaged axial vorticity for condition ‘B’ ......................................... 104
Figure 5.30 Phase averaged axial vorticity for condition ‘C’ ......................................... 104

Figure 5.31 Representative contour plot of percent validated vectors for
condition ‘A’ ............................................................................................... 105

Figure 5.32 Time averaged PIV data: velocity magnitude in the r-z plane for
condition ‘A’ ............................................................................................... 106

Figure 5.33 Time averaged PIV data: mean radial velocity for condition ‘A’. The axial
location of the hot-wire survey is indicated on the plot. The two ‘x’ marks
denote the radial range of the measurements .............................................. 107

Figure 5.34 Comparison of the radial velocities measured by the hot-wire survey and the
two PIV regions for condition ‘A’ ............................................................... 107

Figure 5.35 Time averaged PIV data: mean axial velocity for condition ‘A’. The axial
location of the hot-wire survey is indicated on the plot. The two ‘x’ marks
denote the radial range of the measurements .............................................. 108

Figure 5.36 Comparison of the axial velocities measured by the hot-wire survey and the
two PIV regions for condition ‘A’ ............................................................... 108

Figure 5.37 Time averaged PIV data: velocity magnitude in the r-z plane for
condition ‘C’ ................................................................................................ 109

Figure 5.38 Time averaged PIV data: mean radial velocity for condition ‘C’. The axial
location of the hot-wire survey is indicated on the plot. The two ‘x’ marks

denote the radial range of the measurements .............................................. 110

Figure 5.39 Comparison of the radial velocities measured by the hot-wire survey and
the two PIV regions for condition ‘C’ ......................................................... 110

ix

Figure 54

Figure 5.4

Figure 5.4L

Figure 5.42
Figure 5.4:

Figure 5.45

Figure 54(-

F1é’ure 5.47

I
c:
p—1
(D
J
ll

Fls’tire 5

’. /r

H
('D
(Jr
J!

I
C
.1
(D
J
(J.

I
(It

Figure 5.40 Time averaged PIV data: mean axial velocity for condition ‘C’. The axial
location of the hot-wire survey is indicated on the plot. The two ‘x’ marks
denote the radial range of the measurements .............................................. 111

Figure 5.41 Comparison of the axial velocities measured by the hot-wire survey and the
two PIV regions for condition ‘C’ ............................................................... 111

Figure 5.42 Time averaged PIV data: velocity magnitude in the r-z plane for

condition ‘D’ ............................................................................................... 112
Figure 5.43 Time averaged PIV data: mean radial velocity for condition ‘D’ .............. 113
Figure 5.44 Time averaged PIV data: mean axial velocity for condition ‘D’ ............... 113

Figure 5.45 Fan and shroud configuration showing the control surface used for the
integration of the flow rate from the PIV velocity measurements .............. 114

Figure 5.46 Phase averaged PIV data: velocity magnitude for condition ‘A’ at fan blade
leading edge, note high velocity cell, 0t, and negative radial flow, B .......... 115

Figure 5.47 Phase averaged PIV data: velocity magnitude for condition ‘A’ at 5 degrees
past leading edge, note the high velocity cell, 7 .......................................... 115

Figure 5.48 Phase averaged PIV data: velocity magnitude for condition ‘A’ at 10 degrees
past leading edge, note high velocity cell, 5 ................................................ 116

Figure 5.49 Phase averaged PIV data: velocity magnitude for condition ‘A’ at 15 degrees
past leading edge ......................................................................................... 116

Figure 5.50 Phase averaged PIV data: velocity magnitude for condition ‘A’ at 20 degrees
past leading edge, note the “path” of high velocity, 1 ................................. 117

Figure 5.51 Phase averaged PIV data: velocity magnitude for condition ‘A’ at 25 degrees
past leading edge, note the “pat ” of high velocity, 1 ................................. 117

Figure 5.52 Phase averaged PIV data: velocity magnitude for condition ‘A’ at 30 degrees
past leading edge, note high velocity cell, K ............................................... 118

Figure 5.53 Phase averaged PIV data: velocity magnitude for condition ‘A’ at 35 degrees
past leading edge ......................................................................................... 118

Figure 5.54 Phase averaged PIV data: velocity magnitude for condition ‘C’ at the fan
blade leading edge, note the high velocity cell, A ....................................... 119

Figure 5.55 Phase averaged PIV data: velocity magnitude for condition ‘C’ at 5 degrees
past the leading edge ................................................................................... 119

 

 

Figure 5

Figure 5

F1 gure S

Figure 5.56 Phase averaged PIV data: velocity magnitude for condition ‘C’ at 10 degrees
past the leading edge ................................................................................... 120

Figure 5.57 Phase averaged PIV data: velocity magnitude for condition ‘C’ at 15 degrees
past the leading edge, note the recirculation region, it ................................ 120

Figure 5.58 Phase averaged PIV data: velocity magnitude for condition ‘C’ at 20 degrees
past the leading edge, note the large recirculation region, v ....................... 121

Figure 5.59 Phase averaged PIV data: velocity magnitude for condition ‘C’ at 25 degrees
past the leading edge ................................................................................... 121

Figure 5.60 Phase averaged PIV data: velocity magnitude for condition ‘C’ at 30 degrees
past the leading edge ................................................................................... 122

Figure 5.61 Phase averaged PIV data: velocity magnitude. for condition ‘C’ at 35 degrees
past the leading edge ................................................................................... 122

xi

 

Table 5.

Table 5.

Table 5.

Table 5.

Table 5.1

Table 5.2

Table 5.3

Table 5.4

LIST OF TABLES

Test Conditions for Detailed Velocity Measurements .................................. 67

Spatially averaged RMS levels for each component and each test
condition ........................................................................................................ 73

Tabulated values of the locations of the peak values for various measured
variables for both shroud conditions ............................................................. 74

Values of the dimensionless flow rate found by integration of the PIV results
compared to the measured flow rate from the integral measurements .......... 82

xii

 

Ron

.~‘\.B.r

FCCC l V er

Q
R

Roman:

A,B,n

CS

Preceiver

75 W O

'-1

ds

NOMENCLATURE

Area

Hot-wire calibration coefficient, Chapter 3 only
Discharge coefficient

Control surface

Voltage output

Correlation coefficient

Characteristic Length

Moment about point 0 in Figure 2.3
Number of samples

Atmospheric Pressure

Static pressure in upper AFRD receiver
Volume flow rate

Universal gas constant, Chapter 2 only
Radius

Temperature

Reynolds number

Velocity component

Velocity vector

Hydraulic diameter

Particle separation for PIV

Data frequency

xiii

 

(10

iii

px

2*

GrEek;

If ‘f

3)

[”1

Greek:

99>?

.<

Annular jet gap height

Calibration coefficient for flow metering device

Mass flow rate

Normal vector

pixel

Power

Radial direction

Non-dimensional radial position

Time delay using PIV phase positioning

Velocity component in line with probe axis measured by both x-arrays
Velocity component normal to probe axis measured by first x-array
Velocity component normal to probe axis measured by second x-array
Axial direction

Non-dimensional axial position

Pressure rise

Radial increment

Non-dimensional shroud flow rate

Angular fan speed

Phase position

Angle of the velocity vector with respect to the hot-wire probe axis

Rotation angle 1 of probe axis with respect to fan coordinates

xiv

 

J11

subscn’ Pl
aim

ei

deer

iner

lip
3hr

553

11 Efficiency

9 Tangential direction

2» Non-dimensional shroud plenum pressure
1) Kinematic viscosity

é Rotation angle 2 of probe axis with respect to fan coordinates
p Air density

1' Torque

(1) Flow coefficient

x Non-dimensional shroud power

1|! Pressure coefficient

a) Vorticity

Subscripts:

atm Atmospheric

ci Calibration inlet

decr Decreasing

incr Increasing

n North

r Radial direction

5 South

tip Fan blade tip

shr Shroud

sys System, encompassing test fan and aerodynamic shroud

XV

 

Aeron } m
AFRD
CCD
HWA

P1 V

R\lS

Acronyms:

AFRD

CCD

HWA

PIV

RMS

Axial direction

Tangential direction

Axial Fan Research and Development Facility
Charge Coupled Device

Hot-wire Anemometry

Particle Image Velocimetry

Root mean squared

xvi

 

 

1.0

1.1 .\l(

in cons!
siderablu
S} stem.
C001lllg l
“ram-arr
eV01Ve a;
dissipate
and 011 Ct
01 the rar
[116 Cngm
“5411116 0U
as redU'Cir
ln
C001ing {L1
exam Fesu

Stud" is a -

U

1.0 Introduction

1.1 Motivation and Background Information

The engine cooling fan used in a typical “off-highway” application, such as found
in construction or agricultural equipment, can be quite large in diameter, consume a con-
siderable amount of power, and contribute a measurable amount of noise to the overall
system. The cooling fan in this type of application is fully responsible for the convective
cooling of the radiator when the engine is operating, unlike a passenger vehicle in which a
“ram-air” effect is present when the vehicle is moving. As these engines continue to
evolve and the power output increases, there is a greater demand on the cooling system to
dissipate the added thermal load. Moreover, the addition of air conditioning condensers
and oil coolers adds to the system resistance. One obvious solution is to increase the size
of the radiator and cooling fan. However, this has a negative effect on the efficiency of
the engine because the larger coolant pump and fan will have a parasitic effect on the
usable output power. Thus, it is apparent that any effort to optimize the cooling fan, such
as reducing power input or increasing mass flow rate through the fan, is justified.

In response to this motivation, an experimental investigation of an axial flow type
cooling fan intended for use in an off-highway application has been conducted and the rel-
evant results and analysis are reported in this thesis. The defining characteristic of this
study is a unique aerodynamic shroud that is used with this cooling fan. The shroud con-
sists of a pressurized plenum that produces an annular jet of air which is injected along the
surface of a curved shroud upstream of the fan inlet. This Coanda jet is designed to add
high momentum flow to the blade tip region and weaken, or diminish, the reverse flow

(known as the tip leakage flow) and the resulting tip leakage vortex. A schematic of the

 

aerodynurr
local press
aerodynum
uted to a la
imemal stu
1.2.

Aui
amount of
representat
ance and 11
SIUdres is a
fan: AXia'.
mcmufam

dellVEr. \

aerodynamic shroud concept is presented in Figure 1.1. The reverse flow is caused by a
local pressure differential between the pressure and suction sides of the fan blade. The
aerodynamic shroud was originally conceived to prevent the strong reverse flow, attrib-
uted to a large tip clearance, in the blade tip region of a light truck cooling fan. An exper-
imental study was derived from this motivation and a detailed review follows in Section
1.2.

Automotive engine cooling fans have been the focus of interest in a limited
amount of literature, Baranski (1974), Mellin (1980), and Morris and Foss (2001) provide
representative publications. These studies have recognized the significance of tip clear-
ance and the effect on the performance and efficiency of the cooling fan. A wider range of
studies is available when considering the cooling fan in the larger context as an axial flow
fan. Axial flow fans are found in numerous applications, perhaps because of their case of
manufacture and the relatively low pressure rise and high flow rates which they can
deliver. Most of the recent studies documented in the available literature are directed at
understanding the source of aeroacoustical noise emitted from a rotating fan. Some recent
examples of this include Fukano and Jang (2003), Sorensen (2001), and Wu et a1. (1998).
These studies often investigate different blading schemes or a change in the boundary con-
dition around the fan.

A few studies, in particular, have used different shrouds to obtain enhanced perfor-
mance or lower noise generation of the fan. Since the vortical flow field around the tip
clearance is known to be a source of loss and production of noise, the fan shroud is often
the focus. For instance, Longhouse (1978) stated that tip clearances on the order of 3-4%

of blade chord can contribute up to 15 dB in overall noise levels throughout the operating

 

range of

mate 10'
ing-ring

was deCi
reduce 11
Hull) shr
designed
reverse fl
2001). T

Controllin

1.2 Preyi

Th
Mel}; larg
Clearance 1
Cooling far
firm [0 1m.
“1111 3 didn-
sueh as fun
meauremer

shroud mm

1’0
J ' Y
u r“It‘s.

bl the f

an gr]

[£1111er ”l d

range of the fan. Furthermore, he stated that large tip clearances can result in an approxi-
mate 10% drop in peak efficiency. In response, Longhouse introduced a contoured rotat-
ing-ring shroud that is attached to the fan blade tips. The author found that the noise level
was decreased by as much as 12dB. Another study by Jang et al. (2001) attempts to
reduce the production of noise by controlling the vortical flow field near the tip of a par-
tially shrouded propeller fan using an optimized shape shroud. In particular, the newly
designed shroud attempts to “hinder the development of the tip vortex by diminishing a
reverse flow caused by the tip vortex between the rotor tip and the shroud” (Jang et al.,
2001). Thus, the modification and optimization of the fan shroud is not a new method of

controlling the flow field around a fan.

1.2 Previous Investigations of the Aerodynamic Shroud

The original development of the aerodynamic shroud was motivated by the rela-
tively large tip clearance of cooling fans in light and medium duty trucks. The large tip
clearance is necessary because of the large relative motions between the engine mounted
cooling fan and the chassis mounted shroud and radiator (Morris 1997). Morris was the
first to investigate the efficacy of the aerodynamic shroud on an automotive cooling fan
with a diameter of 457 mm and a tip clearance of 25 mm. Integral flow measurements
such as fan input power and volume flow rate were acquired in addition detailed velocity
measurements in the wake of the fan. These data demonstrated that the aerodynamic
shroud increased performance at higher flow rates but decreased performance at lower
flow rates. Because of this result, it was concluded that system resistance, characterized
by the fan grill, radiator, and air-conditioning condenser in an automobile, plays an impor-

tant role in determining the efficiency of the aerodynamic shroud and fan combination. In

 

 

pam'cu
allows
tire 0b
larjet c

an addi

immers:

fan blad

fan [0 pr

Strong 1‘21

an axial (
stud} 1M
radial fic
Shroud Vt.
50

[Ural builc
dynamic g

10”“ Sx'sr

particular, a large system resistance lessens the efficiency and a smaller system resistance
allows the aerodynamic shroud to increase the efficiency of the fan. Furthermore, qualita-
tive observations using a tuft in the wake of the fan showed that the high momentum annu-
lar jet changed the reverse flow in the tip clearance area to a positive axial flow providing
an additional contribution to the net flow rate through the fan.

In the original investigation by Morris (1997), the cooling fan was partially
immersed in the shroud. In particular, the shroud covered only the upper portion of the
fan blades and had a sharp edge outlet. The partially shrouded configuration allowed the
fan to produce static pressures even throughout the stall range, which is characterized by
strong radial flow. However, when the aerodynamic shroud was used, it forced the flow in
an axial direction explaining the decrease in performance at the lower flow rates. In a later
study (Morris and Foss, 2001), the authors used a rounded outlet that accommodated this
radial flow and found that the enhanced performance of the fan with the aerodynamic
shroud was extended into even lower flow rates.

Neal (2002) studied the effectiveness of the aerodynamic shroud with an agricul-
tural building ventilation fan. The observation from the Morris (1997) study that the aero-
dynamic shroud lead to increased performance and efficiencies with higher flow rates and
lower system resistance naturally suited the use of the aerodynamic shroud with agricul-
tural ventilation fan systems. Agricultural building ventilation fans are characterized by a
low system resistance (only the protective cage and motor on the downstream side and
streamlined shutters on the upstream side add to the resistance) and a high volume flow
rate (Neal 2002). The propeller fans tested were used in unison with a diffuser cone on the

downstream side. In these studies, the aerodynamic shroud allowed the use of a wider dif-

 

fuser C(
toa lou
ber of (
gated. 1
inereasin
T

mance gr
not all pr
ments. 1t.
namic Sim

ac”1‘1.me

L3 Ot'en'i
The
C‘:’“fié’ururi‘ci
mUCh differe
design “1111 C

on the rmignir

fuser cone by allowing the flow to stay attached along the walls of the diffuser. This lead
to a lower velocity at the exit of the diffuser and a more uniform velocity profile. A num-
ber of different blade designs, shroud configurations, and diffuser cones were investi-
gated. The results show that the aerodynamic shroud increased efficiency by 15% while
increasing flow rates by over 35% for the optimal propeller fan of those investigated.

The fact that different blade designs resulted in much different fan/shroud perfor-
mance gives evidence that the aerodynamic shroud is sensitive to blade geometries and
not all production fans can benefit from the active shroud without necessary enhance-
ments. It also supports the suggestions of Neal (2002) and Morris (1997) that the aerody-
namic shroud can benefit from a blade design specifically designed for use with the

aerodynamic shroud.

1.3 Overview of the Experimental Study

The present study intends to investigate the aerodynamic shroud in a similar test
configuration as the Morris (1997) study, however, the fan and boundary conditions are
much different. The previous study effectively demonstrated the efficacy of the shroud
design with one particular automotive cooling fan and the results were strongly dependent
on the magnitude of the tip clearance. This study further investigates the feasibility of the
aerodynamic shroud in a different application and extends the general knowledge of the
how the aerodynamic shroud affects the performance of a fan.

The experimental program was conducted in two separate portions and will be pre-
sented as such. The first portion contains overall performance and efficiency measure-
ments for one baseline test condition and several powered shroud conditions. The

separate aerodynamic shroud conditions are characterized by a “shroud power,” defined as

 

the pro
\‘ariabl:
portion
the ts a1.
techniqL
velocity
aged and
ues of ti
dimensior
additional
methods u
Th:
imental 8p]
moron-p6 ,
e‘iullmtent
integral me;
We derail.

“0m all of I,

the product of shroud pressure and shroud volume flow rate, and it will be shown that this
variable is important in defining the contribution of the aerodynamic shroud. The second
portion of the experimental program contains a detailed investigation of the flow field in
the wake of the fan. Two different techniques were used to accomplish this. The first
technique used a double-X hot-wire probe to simultaneously resolve three components of
velocity at a singular axial location downstream of the fan. These results were phase aver-
aged and are presented as contour plots. These data include the mean and fluctuating val-
ues of the velocity, as well as the axial vorticity. The second technique uses two-
dimensional PIV measurements to present a planar view of the flow field and provide
additional analysis of the radial and axial velocities. Time averaged and phase averaged
methods were used and each are presented as contour plots.

The thesis structure is defined as follows. First, a detailed description of the exper-
imental apparatus is presented in Section 2.0. This includes the test fan, test facility, and
prototype aerodynamic shroud. This is followed by descriptions of the measurement
equipment and techniques in Section 3.0. The experimental results and analysis of the
integral measurements are presented in Section 4.0 followed by the results and discussion
of the detailed velocity measurements in Section 5.0. Finally, the main conclusions drawn

from all of the results are presented in Section 6.0.

      
   
 

Inlet Flow Direction
Jet Gap (g)
///////////,,
Jet Thickness (I;

Test Fan .
\‘

111""

I
Tip Clearance —..l E—

Dtip

 

 

 

17/

 

hub

-b-

Figure 1.1 Schematic representation of the aerodynamic shroud concept

 

2.0 I

2.1 Test

T
imental i
hub to ti;
respect to
trailing er

immersior

2.2 The A
All
The AFRII
Coming fan
F055 61 al.,
and inpm p,
try and Pan;

563110113 deS£

2.0 Experimental Apparatus

2.1 Test Fan

The 711 mm prototype fan pictured in Figure 2.1 was the test fan used in all exper-
imental investigations. The nine molded plastic blades were attached to a steel hub with a
hub to tip ratio of approximately 0.37. The blades had a constant pitch (chord angle with
respect to plane of rotation) of 34° and a single thickness of 4.5 mm from leading edge to
trailing edge. The chord length was 157 mm at the tip and the projected width (total

immersion height of the fan blade) was 100 mm.

2.2 The Axial Fan Research and Development (AFRD) Facility

All test fan data contained in this thesis were acquired in the AFRD flow facility.
The AFRD facility was originally designed for the testing and research of automotive
cooling fans, although it has been used for agricultural ventilation fan research as well (see
Foss et al., 2001). Integral flow quantities (fan or system pressure rise, volume flow rate,
and input power) as well as detailed velocity measurements including hot-wire anemome-
try and particle image velocimetry, can be acquired in the AFRD facility. The following

sections describe the AFRD facility in more detail along with the relevant calibrations.

2.2.1 AFRD Overview

A schematic representation of the test facility is shown in Figure 2.2. The AFRD
is comprised of two sections: an upper receiver (C) in which the test fan is mounted, and a
lower receiver (H) in which the flow rate is measured. The upper receiver is approxi-
mately 2 min height and 3 m in width and depth. The lower receiver is 1 m in height with

the same width and depth dimensions. Air is moved through the facility by both the test

 

 

fan (A)
directec
prime r
SQA-3C
outlet is
the flow

fan.

fan (A) and the prime mover (N). Air is drawn into the upper receiver from the laboratory,
directed through a pair of inlet nozzles (L) into the lower receiver, and moved through the
prime mover back to the laboratory. A large centrifugal blower (Chicago Blower Co.
SQA-36.5) is used as the prime mover. Its inlet is located in the lower receiver (1) and its
outlet is equipped with a computer controlled throttle (M). This throttle is used to adjust
the flow rate through the AFRD, consequently adjusting the pressure rise across the test
fan.

The vertical drive shaft is located in the center of the upper receiver and supported
by two bearings (D) on a structural shaft support. The upper bearing is approximately
0.68 m from the fan exit plane to prevent obstruction of the exiting flow. The fan is driven
by a Reliance Electric 15 hp motor (G) and a 2:1 belt drive pulley system so that the rota-

tional speed of the fan is twice that of the motor.

2.2.2 Flow Rate Measurement System

A unique metering device is used to obtain an integral measurement of the mass
flow rate through the AFRD facility (Morris et al., 2001). The measurement device con-
sists of a large turning passage in the lower receiver that redirects the downward flow
from the upper receiver to the horizontal direction. In redirecting the flow, a moment of
momentum flux is created at the pivot point. The significance of this action in measuring
the mass flow rate and a detailed explanation concerning the measurement apparatus will
be provided in subsection 2.2.2.1 followed by an explanation of the calibration method in

subsection 2.2.2.2.

2.2.2.1 Flow Meter Configuration and Measurement Technique

 

 

 

the tun
TWO 1dr
ncalde
prevent

tieal no;

of 0.406
SUP-POT!

anchorec
momentt
Point (P,
motion.

Figure 2..

expefienc

Figure 2.3 illustrates the configuration of the flow metering device. The inlet to
the turning vane from the upper receiver consists of a 1/4 circle of radius 0.406 m (T).
TWO identical inlets and turning vanes exist on either side of the AFRD provide a symmet-
rical delivery of air to the lower receiver. The inlets in the upper receiver are filtered to
prevent drafts from influencing the flow into the turning vanes. Each inlet leads to a ellip-
tical nozzle with a 4:1 area contraction ratio (V).

The turning vane itself is created from a curved aluminum sheet (Q) with a radius
of 0.406 m and a span of 1.6 m. It is supported by a sharpened point resting on a divoted
support (0) on either side of the span. These points act as a low friction pivot and are
anchored directly to the ground isolating them from AFRD vibrations. The moment of
momentum flux created at the pivot point (0) is balanced by the moment of the force at
point (P) which supports the lower portion of the turning vane and resists rotational
motion. This force is measured with a ring type strain gauge assembly as illustrated in
Figure 2.4. The strain gauge assembly outputs a voltage proportional to the tensile force
experienced by the ring.

Since the force (Fp) can be measured, the control volume formulation of the
moment of momentum flux can then used to compute the mass flow rate, iii. The follow-
ing analysis will explain this in further detail. Using the control volume shown in Figure
2.3 that extends from nozzle outlet to the outlet of the turning vane along the span of the
turning vane, the expression for the momentum of momentum flux as derived in Potter
and Foss (1982) is given by:

2M0 = ;0_Pxfi,. = jppxr‘q(r‘;.;,)dA (2.1)

CS

10

 

.w'.

 

 

Folloui

less usir

A consta'

the upstrr
measured

Upon the s

the equatir

T116 C0€ffic

”“10" 2.2.;

t-oltage OUIPI
mplaced the 1:
ettiptieal shape

elated the 1nCo

Following a simplification and reduction of variables, the integral can be made dimension-

less using a characteristic length (L) and velocity (U) (Morris et al., 2001):

U 2

e0_,,xi“,. = —pL3U21[£x:/](:/-fz)dfi. (2.2)
CS L

A constant (cl ) will replace the dimensionless integral if it is not a function of flow rate or

the upstream velocity profile. According to Morris et a1. (2001), “The magnitude of the

. . . 2
measured force can be wrrtten rn terms of the mass flux m = csz U , where c2 depends

upon the selection of U and L.” With the cross-product (7‘0 _ P x FF) written as l - F P

the equation simplifies to

 

C
FP = [ 1 ]m2 = kmz. (2.3)
2
pczlL

The coefficient, k, was evaluated by the use of a calibration procedure discussed in sub-

section 2.2.2.2. From this equation, and given the measured force, F P , the mass flow rate,

if: , can be reliably measured.

2.2.2.2 Flow Meter Calibration

Calibration of the force transducers was necessary to compute a flow rate from the
voltage output. A known flow rate was produced through a planar contraction which
replaced the test fan at the inlet to the upper receiver. The calibrated inlet consisted of an
elliptical shaped nozzle with a 6.3:1 contraction ratio. The relatively high area ratio accel-

erated the incoming flow ensuring a thin boundary layer at the exit and a Cd value close to

11

 

 

1.0. The
a previou

Tl
flon rate

numbers 1

“here ('6
the kinem
increasing

Th
from 150‘
regjolls of
mem thick
fieiem. T}
the Scale n

Th:
hence the ‘

“here the s

1.0. The evaluation of C d = C d(Re) was performed through two separate procedures in
a previous study by Neal (2002) and is described in the following paragraphs.

The first procedure involved a 3/16 scale model study. By using a known mass
flow rate from a sonic nozzle test stand, the Cd value was found for a range of Reynolds

numbers between approximately 50,000 and 150,000 with Reynolds number is defined as:

Re ___ UCldH
V

 

(2.4)

where U a. is the velocity at the contraction inlet, d H is the hydraulic diameter, and v is

the kinematic viscosity of the laboratory air. From this study, the Cd values increase with
increasing Reynolds number to Cd=0.980 for Re=150,000.

The second procedure involved a hotwire study with a Reynolds number range
from 150,000 to approximately 775,000. The hotwire survey traversed the near wall
regions of the exit of the contraction on both the length and width sections. The displace-
ment thickness found from this investigation was then used to estimate the discharge coef-
ficient. The Cd value found from this procedure was also 0.980 confirming the results of
the scale model study.

The maximum Reynolds number for the present study was approximately 800,000,
hence the use of a Cd value of 0.980 was justified. Given the Cd value of the planar con-

traction, the mass flow rate through the AFRD facility was found by:

 

. 2
m : pAciCdA/6(Patm—Preceiver) (2‘5)

where the subscript “c1” refers to the contraction inlet.

12

 

 

 

A
the prime
pressure ti
The calibr

transfer f u

1n equatio
transducer
flow rates
The Calrbr
include th
ment devil

Sir
flow mEte

to:

A range of flow rates was induced through the calibrated contraction by throttling
the prime mover. The voltage output from both force transducers and from the differential
pressure transducer were recorded by an AID acquisition system described in Section 3.1.
The calibration data were plotted and fitted with a linear least squares fit. The resulting

transfer functions for the North and South flow meters were:

n 2.7816 [En-offset (kg/s) (2.6)

ms = 2.9214 /Es—0ffset (kg/s) (2.7)

5.
II

In equations 2.6 and 2.7, the term ‘offset ’ denotes the “pre-load” voltages from the force
transducers. The pre-load was necessary to avoid fully unloading the strain gauges at low
flow rates due to the vibrations in the system. The fitted data are shown in Figure 2.5.
The calibration data were acquired in two separate trials. Correspondingly, the plots
include the results from both trials to qualitatively show the repeatability of the measure-
ment devices.

Since two independent measurements of flow rate were acquired from the two
flow meters but only one integral measurement is required, they were averaged according
to:

mn+ms

measured " 2

th (2.8)

where the subscripts ‘n ’ and ‘s ’ denote the “north” and “south” flow meters. Although

the mass flow rate is the actual quantity measured, the data are presented in the results

with the volume flow rate, Q , as given by:

11':
_ measured
Qmeasured — p (2'9)

13

 

 

The diffe

ically no

2.2.3 Ho
A
probe dot:
the traters
steel ring I
tional mot
box. Rad
“$40241
mounted 0
ing in this .
”3161 Was
“"3 Probe
faced “11h
m a CUStor

ExPEri
23'] Pr 01(

The

2.7 The d1

The differences in the measured volume flow rates between the two flow meters were typ-

ically no greater than 1.2%.

2.2.3 Hot-wire Traversing Mechanism

A stepper motor controlled traverse system allowed movement of the hotwire
probe downstream of the fan in the r, 6, and z directions. A schematic representation of
the traverse is shown in Figure 2.6. The entire traverse was mounted on a 1.52 m diameter
steel ring that surrounded the fan allowing it to be moved in the azimuthal direction. Rota-
tional motion of the steel ring was provided by a 6:1 pulley system and a 100:1 worm gear
box. Radial motion of the hotwire probe was provided by a linear slide (Velmex Inc.
#MB4024BJ-S4-20) with a maximum horizontal travel of 51 cm. Two linear slides were
mounted on the steel ring (180° apart) but only one was necessary for the probe position-
ing in this study. A vertical linear slide (Velmex Inc. #MB4024BJ-S4-16) with 40.5 cm of
travel was mounted to each horizontal linear slide to provide axial movement of the hot-
wire probe. All stepper motors were controlled by a Velmex #NF90-3 controller inter-
faced with a PC from which stepping commands were sent. The commands were written
in a custom Matlab code. The manufacturer supplied accuracy of the linear slide was

0.05mm/meter.

2.3 Experimental Aerodynamic Shroud
2.3.1 Prototype Design

The experimental aerodynamic shroud assembly is shown schematically in Figure
2.7. The different parts of the complete system include the radiator, the shroud plenum,

the fixed venturi shroud and the auxiliary pressurization source. A radiator was included

14

 

 

to mimic
condition
mance of

T1
Modine \
directly c
sealed of
Upstream
static pre;
tor and p;

21tor and t

to mimic an installed condition. It was important to include the pressure drop and flow
conditioning associated with the radiator because of the possible influence on the perfor-
mance of the cooling fan.

The radiator used in the experimental study was a production model radiator from
Modine with an effective cooling area of 0.83 m by 0.83 m. The radiator was placed
directly on top of the shroud plenum and the area around the sides of the radiator was
sealed off to prevent air from leaking into the area downstream of the radiator and
upstream of the fan inlet. The downstream side of the radiator frame was equipped with a
static pressure tap to measure the pressure drop across the restrictive elements. The radia-
tor and pressure tap location is pictured in Figure 2.8. The axial distance between the radi-
ator and the leading edge of the fan was approximately 140 mm.

The shroud plenum was built in two sections. The lower section is fixed and the
upper section moves vertically to adjust the gap height from which the jet originates. An
axisymmetric, contoured carbon-fiber piece with a sharp trailing edge is attached to the
upper shroud section to create the thin wall jet from the shroud plenum. The contour is
designed to provide a smooth inlet to the fan and to act as a nozzle to accelerate the air
exiting the plenum.

The geometric dimensions of the aerodynamic shroud were chosen based on the
fan diameter, the tip clearance, and the typical dimensions of a radiator for this fan diame-
ter. To replicate a field application, the design of the aerodynamic shroud was dependent
on representative spatial constraints of an underhood installation. For this reason, the
shroud plenum was made much smaller relative to the fan diameter than the previous aero-

dynamic shroud investigations of Morris (1997) and Neal (2002). Specifically, the shroud

15

 

 

plenum v
the plenur
A calibrat

Tl
tip cleara‘
mill with
previous 1
1.2. The

the shrou

plenum was 0.891 m square, or 1.25 times the diameter of the fan. The upper section of
the plenum rested on four 1/4 inch threaded rods which were used to adjust the gap height.
A calibrated “wedge” was used to set this height to the desired value.

The inside diameter of the venturi shroud was 729 mm corresponding to a 9 mm
tip clearance for the 711 mm fan. The shroud was fabricated from wood on a five-axis
mill with a curved inlet and outlet. The use of the curved outlet was motivated from the
previous research of Morris (2001). Those findings are discussed in more detail in section
1.2. The upper section of the venturi shroud had a thinner cross section to accommodate
the shroud plenum and to allow the coverplate to be attached to the venturi shroud (see
Figure 2.7). The shroud plenum was constructed on the coverplate with a 22 mm gap
between the venturi shroud and the plenum sidewalls at the four quadrants. This gap was
sufficient to allow a uniform axisymmetric jet to exit the plenum. Four inlets with a diam-
eter of 89 mm each are located in the four comers of the shroud plenum to allow for the air

delivery as shown in Figure 2.7.

2.3.2 Shroud Plenum Pressurization System

The plenum was pressurized by injecting air into the four inlets on the underside of
the plenum. Figure 2.9 illustrates the piping system and auxiliary blower arrangement in
relation to the testing facility. Air was delivered by a centrifugal blower (Spencer Turbo-
Compressor, Lot NO. 37215) through 3” PVC piping. The piping entered the plenum
from the upper receiver. This arrangement differed from the experimental configurations
of Morris (1997) and Neal (2002) because of the presence of a radiator upstream of the fan

inlet.

16

 

centrifu,
the inlet

T2116 \\ 213

where A,

nozzle. Tl

 

the pipin
used to rr
the shrou

F316 “'35 I

“here

of P
Shroud.

The flow to the shroud was adjusted with a butterfly valve on the outlet side of the
centrifugal blower. A Smith and Wang nozzle with a known discharge coefficient, C d , on
the inlet side metered the flow (see Figure 2.9). Using this method the shroud mass flow

rate was found according to

 

mshroud : pathnozzleCdA/(Patm _ Pnozzle) (2°10)
where Amaze and mee is the area and static pressure at the throat of the Smith and Wang

nozzle, respectively. Since the air flow underwent viscous heating through the blower and
the piping system, a thermocouple (Omega Engineering Inc., model 670, Type ‘T’) was
used to measure the air temperature in the shroud. A small hole was drilled into the top of
the shroud and the thermocouple was suspended into this hole. The shroud volume flow

rate was then adjusted for each data point according to the measured temperature:

 

_ mshroud
Qshroud - p (2'11)
shroud
where
P +P
hroud atm
ph d: 3 (an)
3 ma RTshroud

A static pressure tap in the bottom on the shroud plenum allowed the measurement

Of Pshroud°

l7

 

Figure 2.1 Prototype test fan (D = 0.711 m)

A B7

 

 

 

 

 

 

  

/
H

A Test Fan H Lower Receiver

B Pressure Tap I Inlet to Prime Mover
C Upper Receiver .1 Force Transducer

D Drive Shaft Supports K Turning Vane

E Traverse L Nozzle

F Torque Sensor M Throttle Plate

G Drive Motor N Prime Mover

Figure 2.2 Schematic representation of the AFRD flow facility

18

 

 

 

 

 

Low Friction Pivot

Force Transducer
Turning Vane

Counter Balancing Spring
Lower Receiver Floor
Inlet Filter

Upper Receiver Floor
Elliptical Nozzle

 

<C—1W7UO'UO

 

 

 

 

 

 

12 Volts

 

E (output)

 

 

Force Load

Figure 2.4 Details of the force transducer for the flow metering device (Neal, 2002)

19

(a

(in) ii

Ki

\ .

- .131”?

4 . 0")

rant: 8:: s2... 3...:

a).
.N.
1-

1.041

1,! . it?)

I:

Mass Flow Rate (kg/s)

Mass Flow Rate (kg/s)

 

8.00 r

7.00 *—

   

6.00 ”—
m, = 2.7816*(En - offset)”

 

5.00 _

4.00 t—‘
F

3.00 *—

 

  

2 '00 b I Trial I

0 Trial 2
Least Squares F it

 

 

1.00 *—

 

 

   

 

 

L

 

 

0.00 l l 1 l l l l l l l l l l l m l L l l 1 l J
0.00 0.50 1.00 1.50 2.00 2.50

(En - offset) 0'5

a) North Flow Meter Calibration

3.00

 

8.00

T

l

7.00

6.00

TFTI

m, = 2.9214*(Es - offset)“
5.00

r l

4.00

l

3.00 *—

    

 

   

2 '00 _ I Trial I

0 Trial 2

1.00 *- Least Squares Fit

 

 

 

 

  

l

0.00 l l l L l l l 4 l l l l l l L l l l

   

 

l

 

 

0.00 0.50 1.00 1.50 2.00

(rt:s - offset) 0'5
b) South Flow Meter Calibration

Figure 2.5 Calibration data and fit for the flow metering devices

20

2.50

 

Fl

0007.“?

 

Azimuthal Traverse Ring
20" Horizontal Slide
Stepping Motor

Vertical Guide

‘I-

IQ'UU'!

1'
i'

 

 

Ii '
I 1:
l1 .

A
B
1— I
_-_
—~ I
u I D
.le
F I

16" Vertical Slide
Linear Bearings
Probe Mount
Cross Supports

Figure 2.6 Schematic representation of AFRD hotwire traverse

21

 

 

 

 

 

Vent“

AFRD

Shroud Plenum (‘overplate/Base
of Plenum ;

 

 

 

.I-Is"ll
III-u-IID

 

 

 

 

Upper Contoured
Shroud Section

Air Delivery lnlets

Gap adjustment
mechanism

    

Radiator

 

Venturi Shroud

Air Delivery lnlets

 

Figure 2.7 Schematic representation of the aerodynamic shroud and plenum

configuration (Radiator not shown in top view)

22

 

 

Figure

 

DOW}

Figtlre 2.9 S

Static Pressure

. 9/ Tap Location

  
 

    
 

31..

Figure 2.8 Ra ator with static pressur

  

ap dow

 

tream of the cooling fins

A

    

A Aerodynamic Shroud Assembly E Prime Mover

B Upper Receiver F Spencer Blower Throttle
C Shroud Piping G Smith and Wang Nozzle
D Prime Mover Throttle H Spencer Shroud Blower

Figure 2.9 Schematic representation of AFRD with aerodynamic shroud assembly
and pressurization system

23

 

 

3.0 M

Tl
obtain the
the first h;
out an arm
in Section
investigate
program,
wake of tilt
about the p
calibration
“1th usec

tion 34,

3.1 Data A

All 1

3.0 Measurement Equipment and Technique

The details of the measurement equipment, techniques, and processing used to
obtain the results are presented in this chapter. Integral measurements were acquired for
the first half of the experimental program to characterize fan performance with and with-
out an aerodynamic shroud. The quantities measured and the methods used are described
in Section 3.2. The information gathered from these measurements was used to further
investigate the mechanistic features of the fan flow in the second half of the experimental
program. This was accomplished by acquiring detailed velocity measurements in the
wake of the fan using particle image velocimetry (PIV) and hot-wire anemometry. Details
about the PIV system and vector processing are given in section 3.3. Finally, the hot-wire
calibration method and technique using the double-X probe as well as the processing algo-
rithms used to simultaneously resolve three components of velocity are presented in sec-

tion 3.4.

3.1 Data Acquisition System

All integral and hot-wire data acquired in this thesis were sampled and recorded
using a 16 bit, 8 channel Iotech Wavebook/516 AID acquisition module accompanied by
an 8 channel expansion module. The AID system had a range of 3510 volts with a resolu-
tion of 0.3 mV. Measurements were acquired as “sample-and-hold” meaning that all chan-
nels of data were acquired and recorded simultaneously. The data were saved in a binary

format and transferred to a PC for processing using custom Matlab programs.

24

 

 

3.2 In

overall
efficier
specif}
are use
shroud
The eqt
are criti
definitic

dure use

3.2.1 M

3.2.1.1 1

3.2 Integral Measurements

In engineering practice, it is common to use integral quantities to examine the
overall performance of a fan, blower, or compressor. Namely, the pressure rise AP, and
efficiency n , as a function of volume flow rate Q , are significant variables to consider in
specifying an air moving device for a particular application. In this study, these variables
are used to compare the baseline performance of the cooling fan without the aerodynamic
shroud to the performance enhancement (or degradation) with the aerodynamic shroud.
The equipment used to measure these quantities and produce the performance curves that
are critical in assessing the effect of the aerodynamic shroud is given in Section 3.2.1. The
definition of the variables presented in the results is given in Section 3.2.2. and the proce-

dure used to acquire the integral measurements is described in Section 3.2.3.

3.2.1 Measurement Equipment
3.2.1.1 Pressure Transducers

Four separate pressure measurements were necessary to fully characterize the per-
formance of the fan and shroud combination. The static pressure in the upper receiver of
the AFRD was measured with a 10 Torr MKS Baratron pressure transducer, model 398
HD head unit with a Type 270B signal conditioner. The static pressure at the downstream
face of the radiator was measured to acquire both the pressure drop across the radiator and
the pressure rise across the fan (see section 3.2.2). This pressure was measured using a
Validyne model DP15-20 with a range of 0-868 Pa. A Validyne DP10-22 head unit was
used to measure shroud pressure when the shroud pressure was under 1245 Pa. For pres-
sures greater than this value, a Validyne model DP15-26 was employed. The three Vali-

dynes were each used in conjunction with a model CD15 carrier demodulator. A fourth

25

 

 

pressu
inv’OlV
anache
pressur
tude of
Type 2

transdu

Torquem
directly It
accuracy
ean'ty of .-
to of :01
from bend:
Ulbute 10 m
The
Cision mach;
Jew arm to
tioner and Cor,
dueed an 0Ut
Dr
A (are n

1h? fl’lCllOn {mm

pressure was necessary to obtain the flow rate through the shroud flow system which
involved measuring the static pressure at the throat section of the Smith and Wang nozzle
attached to the inlet of the auxiliary shroud blower as described in Section 2.3.2. This
pressure was measured by a 10 Torr or 100 Torr MKS Baratron (depending on the magni-
tude of the pressure), model 398 I-ID head unit with a Type 270 signal conditioner and a
Type 274 head selector. The reported accuracy of the Baratron and Validyne pressure

transducers is i0.08 % and i025 % respectively.

3.2.1.2 Torque Measurement

A torque sensor from S. Himmelstein and Co. (MCRT 3120TN(2-2)NN Pulley
Torquemeter) was used to measure the torque applied to the drive shaft. It was connected
directly to the lower end of the shaft below the pulley and had a range of i200 in.lb. The
accuracy of the torque readings were supplied by the manufacturer and included a nonlin-
earity of :02 in.lb. (0.1% of full scale), a hysteresis of _+_0.2 in.lb., and a non repeatabil-
ity of $.01 in.lb. It was equipped with a strain gauge bridge designed to cancel signals
from bending and thrust loads. Hence, any bending load from belt tension does not con-
tribute to the overall measurement.

The torque sensor was calibrated prior to experimental data being acquired. A pre-
cision machined lever arm was fit directly on the torque sensor. A load was applied to the
lever arm to produce a known torque. This torque was then entered in the signal condi-
tioner and controller module (S. Himmelstein and Co., 700 Series) so that the module pro-
duced an output of +/- 10 volts directly proportional to the full scale torque range.

A tare measurement was taken prior to testing to account for the torque applied by

the friction from the bearings. The friction is a function of grease viscosity, and in turn, a

26

 

 

 

fiuncfio
before

fOrapp
measun

ingthe'

approxir
measuret

totesfing

function of temperature. In this manner, the bearings must reach a steady state condition
before a constant tare value can be found. To find a proper tare value, the shaft was run
for approximately 2 hours at the testing speed with the fan removed. A plot of this torque
measurement is shown in Figure 3.1. The measured torque was then adjusted by subtract-

ing the tare value

1: T (3.1)

actual = measured" Trare °
It is evident from the plot that the tare decreases with an exponential decay to
approximately 5.5 in-lb. After 45 rrrinutes, the torque was within 0.3 in.lb. of the torque

measured at 150 nrinutes. Consequently, the fan was operated for at least 45 minutes prior

to testing with 1: = 5.5 in-lb.

tare

3.2.1.3 Optical Encoder

An optical encoder was used to accurately set, monitor, and record the fan rota-
tional speed. The optical encoder used for the integral measurements consisted of an
infrared interrupt circuit (Omron EE-SG3, slot sensing method) and a disk with 60 evenly
spaced holes attached to the fan shaft. Using a Schnritt trigger (Fairchild Semiconductor
4093B), a TTL signal was produced (0 or 5 volts depending on the location of an individ-
ual hole in relation to the infrared sensor). A time series of the measured voltages was
recorded during data acquisition and subsequently processed to deterrrrine an accurate

rotational speed.

3.2.2 Definition of Variables
One of the most common ways to characterize the performance of a fan is in terms

of the pressure rise (AP) as a function of flow rate (Q). The measurement of Q is

27

 

 

 

descril
into ac
tive of
air by i
is adde
quantity

vant 110

AP is (1.
blades ar
ciently “;
HOWEVer.

C1€3f1)'(jc

Where p
r1

recelter

described in section 2.2.2, however, the added flow from the shroud plenum must be taken
into account in a powered shroud condition. It is important to note that a primary objec-
tive of the aerodynamic shroud is to increase the transfer of thermal energy to the ambient
air by increasing the amount of convected air through the radiator. Since the shroud flow
is added downstream of the radiator, this amount of flow must be subtracted from the
quantity measured by the flow meters. Hence, in the powered shroud condition the rele-
vant flow rate is
Qfan = Qtotal— Qshroud ' (3'2)
The definition of AP must also be clarified according to the application. Usually,
AP is defined as the static pressure differential between the downstream side of the fan
blades and the upstream approach flow. In the absence of restrictive elements and suffi-
ciently “far upstream”, the upstream static pressure is simply the quiescent atmosphere.
However, when restrictions are present (e.g., a radiator) the pressure differential must be
clearly defined. In the present investigations, the pressure rise, AP , is defined as:

AP P

fa (3.3)

receiver _ radiator

n=P

where P is the static pressure measured just downstream of the radiator and

radiator

P is the static pressure in the AFRD upper receiver. Figure 3.2 is presented to

receiver

clarify this configuration. Both P and P were measured independently of

radiator receiver

each other and referenced to atmospheric pressure. This method was used so that the pres-
sure drop across the radiator and the total system pressure rise could be measured as well.

It is typical in turbomachinery literature to represent the aforementioned variables
in dimensionless form. Morris (1997) performed a dimensional analysis in the previous

aerodynamic shroud investigations and the resulting dimensionless variables will be used

28

 

 

 

 

in this
the tip
effects
number

(Barans

and the 1

“ here A

shroud ar

cally Simi
relationsh
Cffects are

An
0f Voiume
CiEncy is u:

“16 p0“€r 11

“here

in this study as well. The length and velocity scales used were the fan diameter (D) and

the tip speed (U ). Because the velocity and pressure fields are insensitive to viscous

tip
effects and governed solely by inertial effects and boundary conditions, the Reynolds

number is not considered a significant parameter. This is customary in many fan studies

(Baranski, 1997). Hence, the pressure coefficient is defined as:

\t! = —-’i"—” (3.4)

and the flow coefficient is defined as

/A
(1) _ Qtan [low (3.5)

- U tip
where Aflow = $0251" — Dzhub) represents the effective annular flow area between
shroud and the fan hub.

These non-dimensional variables are especially useful when comparing geometri-
cally similar fans of different diameters or operating speeds. In those cases, the functional
relationship, it! = tum), will remain constant. The limiting case is when compressibility
effects are present.

Another integral quantity of considerable interest is the fan efficiency as a function
of volume flow rate. Many definitions of efficiency exist. For this study, a static effi-
ciency is used where it is defined as the ratio of energy input to the fluid in unit time over
the power input to the fan. In terms of the relevant variables, it is defined as

n = M (3,6)
50 fan

where

29

 

and t
This (1
from t.

21556551

accomr
The rea
in effici

0356:. the

Where

In equatio
air dEHVer
based 0n it
Its Csu‘mate
C0uld be USr
mine the aCtr

was dictaled

sofa" = 1: - $2 (3.7)

and T is the shaft torque that is required to rotate the fan at £2 radians/second.

This definition of power is instructive because it is independent of the mechanical losses
from the motor, belts, and bearings in the experimental configuration. Therefore, it is an
assessment of the fan only.

In the active shroud condition, a modified definition of efficiency is necessary to
accommodate the added power consumption associated with the aerodynamic shroud.
The reader is reminded that, in using the aerodynamic shroud, it is desired to realize a gain
in efficiency greater than the additional power needed to pressurize the shroud. In this

case, the efficiency is defined as:

 

 

Tl = APfaanan (3.8)
safari + sashroud
where
AP -
Joshroud : shroud Qshroud. (3.9)

1{shroud

In equation 3.9, the term 11 d is present to account for the inefficiencies of a shroud

shrou
air delivery system. An estimated value 11 shroud = 0.7 was used for the calculations
based on the previous aerodynamic shroud work of Morris (1997) and Neal (2002). Mor-
ris estimated this number from typical operating efficiencies of centrifugal blowers that
could be used for a practical aerodynamic shroud application. It was not useful to deter-

mine the actual efficiency of the experimental aerodynamic shroud since the configuration

was dictated by the constraints of the experimental setup and available hardware.

30

 

 

shrou
cal 1y.
shrouc

plenur

and

The func
The com!-
Variable r
Sionless s,

“‘0- The

It is clear that the additional boundary conditions imposed by the aerodynamic
shroud modify the functional dependence of the flow coefficient and efficiency. Specifi-
cally, the pressure coefficient and efficiency become a function of the flow coefficient,
shroud plenum pressure, and the shroud flow rate. In non-dimensional terms the shroud

plenum pressure and shroud flow rate are:

AP
A = _£’_’_’2“_4 (3_10)
viii-p
and
Q
A = (Tibia—51. (3.11)
tip ow

The functional relationships are then modified to t1! = w(¢, I», A) and r1 = 71(4), A, A).
The complexity of the data analysis is greatly increased using three variables. Thus, a new
variable representing a dimensionless shroud power is created by combining the dimen-
sionless shroud power and shroud flow rate, therefore, reducing the number of variables to

two. The definition adapted from Morris (1997) is:

= APshrrgudQshroud. (3.12)
pH A

tip ow

 

X

This variable simplifies the data analysis and it provides a useful method of characterizing

the contribution of the shroud. Specifically, using x, the 1.1! and 1] functions can be

cxnressed as v = Wt. x) and n = n(¢. x).

3.2.3 Data Acquisition and Procedure
The equipment described in Section 3.2.1 was used to acquire the integral quanti-

ties of interest. All channels of data were acquired for 10 seconds at a sampling rate of 10

31

 

kHZ.T

acquirer

tion. Tr
during 1
shroud ]
delivere
tests. tht
other. 1
tnovedj
COmplet
repeatec
rate COn
“5 data
Characte
TEVerSec

Sure rise

above.

expefim
gap heig
refers {C

kHz. The main portion of the integral data and the results presented in this thesis were
acquired at a fan speed of 950 rpm.

A characteristic curve, or performance curve, was produced for each shroud condi-
tion. To clarify, the term “shroud condition” refers to the state of the aerodynamic shroud
during the test. The “shroud oft” condition means that no auxiliary air is supplied to the
shroud plenum. A “powered shroud” or “active shroud” condition implies that air is being
delivered to the shroud creating an annular jet upstream of the fan inlet. For all integral
tests, the procedure involved acquiring data from one extreme of the operating range to the
other. For instance, the test would start at a very low pressure rise. The throttle was
moved in small increments, acquiring data at each increment, until the AFRD throttle was
completely closed to create a “full shut-oft” condition (largest 111). This procedure was
repeated going from the full shut-off condition back to the low pressure rise, high flow
rate condition to test for repeatability in the measurement. Approximately 30 data points
(15 data points in each direction) were acquired over the course of the experiment to fully
characterize the performance curve. To test the effects of hysteresis, this method was
reversed for some tests, starting from full shut-off, increasing flow rate until a low pres-
sure rise, and returning back to the starting pressure rise condition.

The procedure for a “shroud off” condition was relatively simple as explained
above. Conversely, the powered shroud conditions presented an added complexity in the
experimental variables. All of the powered shroud integral data consisted of a constant
gap height and shroud pressure for the duration of the experiment. The gap height, g,
refers to the space between the annular jet nozzle and the fixed venturi shroud surface.

Three different gap heights were used including a 2 mm, 3 mm, and 4 mm gap. Further-

32

 

more. ft

SUTCS T31

3.3 Par

ate a vel
used to

Wake of
ment use
cOmpute

EXpen' me

3.3.1 PI‘

T1
Computer

generator.

more, four different shroud pressures were tested for each gap height. The shroud pres-

sures ranged from nominally 250 Pa to 1500 Pa depending on the gap height.

3.3 Particle Image Velocimetry Measurements

Particle image velocimetry (PIV) is an instantaneous measurement used to evalu-
ate a velocity field across a planar area. In this study, a two-component PIV system was
used to measure the axial and radial components of velocity in the outer region of the
wake of the fan. Both time averaged and phase averaged data were acquired. The equip-
ment used in the PIV measurements is described in Section 3.3.1. The technique used to
compute and validate vectors is described in Section 3.3.2 followed by details of the

experimental configuration and procedure in Section 3.3.3.

3.3.1 PIV Equipment

The PIV system consisted of a CCD (charge coupled device) camera, a laser, a
computer with the main processing unit and necessary software, and an atornizing aerosol
generator.

The 2k x 2k pixel, 12-bit camera was a Flowmaster 28 model with a Kodak KAI
4000 sensor capable of 15 single frames per second. A 532 nm band pass filter was used
in front of the lens to filter out ambient light.

A New Wave Research Minilase III, NszAG dual-head 10 Hz laser was used for
all PIV measurements. The laser was equipped with light-sheet optics to expand the laser
beam into a planar sheet. The manufacturer specified the output of the laser to be 50 m] at

a wavelength of 532 nm with a pulse width of 5-7 ns.

33

 

 

ing Unit 1
data. La'
ment. T1
pulse as v
used for 3
Th

H vvith an
octyl sebar
0.21mi am
A s

“‘35 used f(
a fBflective
“'3 used to
mp the “log

averaged d3:

In thr:
mOde. The n
is perfotmed (
tion of the pro

lcal SlZe 0f the

A computer equipped with dual Intel Xeon processors and a Programmable Tim-
ing Unit (PTU) card was used to control the PIV system as well as acquire and process the
data. LaVrsion’s proprietary software, DaVis, was the main interface with the PIV equip-
ment. The software allowed the synchronization of the camera exposure with the laser
pulse as well as the control of the many other laser and camera settings. DaVis was also
used for all processing and post-processing of the acquired images.

The flow was seeded with a Topas Atorrrizer Aerosol Generator, model ATM 210/
H with an adjustable particle production rate. The aerosol fluid used in the generator was
octyl sebacate (DEHS) and the manufacturer reported particle sizes generated are between
0.2p.m and 0.3 um.

A slightly different configuration of the optical encoder described in Section 3.2.1
was used for the particle image velocimetry measurements. Instead of an interrupt circuit,
a reflective type receiver/transmitter (Omron EE-SFS, phototransistor sensing method)
was used to output a TTL level signal. A narrow strip of a reflective material was used to
trip the “logic high” once every revolution. This provided the trigger signal for the phase

averaged data that will be described in Section 3.3.3.

3.3.2 PIV Vector Processing

In this study, the CCD camera records images in the double frame/double exposure
mode. The images are divided into separate interrogation regions and a cross—correlation
is performed on each region to resolve a single vector (see Figure 3.3). The spatial resolu-
tion of the processed data is dictated by the number of interrogation regions and the phys-

ical size of the image area.

34

 

 

An
els for eacl
tion is the
than half I
reduced. T

longer be c

where ds is

desired inter

An interrogation region is composed of a number of pixels. As the number of pix-
els for each interrogation region decreases, the spatial resolution is increased. The lirnita-
tion is the magnitude of particle shift. In particular, if the mean particle shift is greater
than half the size of the interrogation window, then the accuracy of the correlation is
reduced. This is because some particles move outside the interrogation region and can no

longer be correlated. The optimum particle shift is given by Keane and Adrian (1990):

l
O.lpx<ds<ZdImWin (3.13)

where ds is the separation of the particle images on the CCD, and d represents the

intWin
desired interrogation window size in pixels (designated as px.) (LaVrsion GmbH, 2002).
An adaptive multi-pass feature in the DaVis software was used in the PIV process-
ing to manage the issue of particle shift and optimize the size of the interrogation window.
Using multiple passes, the accuracy of the correlation is increased as explained by the fol-
lowing. During the first pass, a reference vector is calculated for each interrogation win-
dow. In the next pass, the interrogation window is half the size of the first interrogation
window and shifted according to the magnitude and direction of the reference vector. The
number of vectors calculated with each pass is increased, therefore, increasing the spatial
resolution without sacrificing correlations. Moreover, a window overlap can be used to
further increase the spatial resolution. An example of a 50% window overlap is shown in
Figure 3.4. The PIV measurements acquired in this study use a total of four passes start-
ing with an initial pass of 128x128 pixels and a 50% overlap. The final window size is
32x32 pixels with a 50% overlap to result in a total of 16,384 vectors for each image.

Validation schemes were used to remove spurious vectors resulting from the pro-

cessing. Spurious vectors arise from a number of undesirable circumstances including

35

 

 

noise. p001
are availab
study was
0nd hi ghes
study. if th.
is to prever
a “median
vectors anc
ables to set
magnitude r
less than thr
at least three
or the vector
remove as m

CedUre Was IT

333 Data A

The prl

directions near

noise, poor seeding, or high velocity gradients. A great number of validation techniques
are available in the DaVis software package. The first validation technique used in this
study was a “peak ratio” validation. The peak ratio is the ratio between the first and sec-
ond highest correlation peaks. The criterion can be set by the user, and in the case of this
study, if the peak ratio was less than a value of 1.2 the vector was tagged as spurious. This
is to prevent any ambiguity of the correct shift. A second validation technique was termed
a “median filter”. This involved computing a median vector from the eight surrounding
vectors and comparing that value to the middle vector. Again, there are user defined vari-
ables to set the criteria for this action. In this study, the middle vector was removed if its
magnitude was greater than three times the rrns of the neighboring vectors or if there were
less than three neighboring vectors. In addition to this, a vector was inserted if there were
at least three good vectors surrounding it. The inserted vector is either the original vector
or the vector from the second, third, or fourth lower correlation peaks. It is important to
remove as many spurious vectors as possible when computing averages. Hence, this pro-

cedure was followed in the present investigations.

3.3.3 Data Acquisition and Procedure

The present study seeks to resolve the velocity components in the axial and radial
directions near the outer portion of the fan exit plane. All PIV data were acquired at 950
rpm. Due to the limited size of the measurement field, two separate measurement planes
were used in the PIV experiments. A schematic representation of the measurement loca-
tions is shown in Figure 3.5 and denoted as “region 1” and “region 2”. As shown, region
2 overlapped region 1 by approximately 12 mm. Due to the differences in placement of

the camera, region 2 is a larger field of view, or smaller magnification, than that of region

36

 

1. Hence
lated ever
Ti
ditions. l
processed
the veloci
The phase
phase trig
control 01
It
adJ'USt the
\‘idual f'd‘
110n_ T1]:
the insra,
lowing
CUCOder
Were the
rePT'ESen
Check at
9‘38. N,

1. Hence, region 1 has a slightly greater spatial resolution. Specifically, a vector is calcu-
lated every 0.83 mm in region 1 compared to 0.96 mm in region 2.

Time averaged and phase averaged data were acquired for the selected shroud con-
ditions. For the time averaged data, approximately 200 vector maps were recorded and
processed in both regions of interest. Phase averaged data were also acquired to capture
the velocity field at eight different phase positions between two blades of the rotating fan.
The phase averaged acquisition system used the TTL output on the optical encoder as the
phase trigger. A “logic high” from the TTL signal used with a delay time allowed full
control of the phase plane in which data were acquired.

It was instructive to know the position of the fan at the first phase plane and then
adjust the phase positions in equal increments to capture the dynamics between two indi-
vidual fan blades. The leading edge of the fan blade was chosen as the first phase posi-
tion. Thus, it was necessary to find the proper delay time between the trigger signal and
the instant that the leading edge crossed the PIV laser plane. This is described by the fol-
lowing. The optical encoder on the fan shaft was used in conjunction with an optical
encoder that was aligned with the leading edge of the fan blade. The two output signals
were then acquired while the fan was run at various speeds from 20 rpm to 130 rpm. A
representative output signal is shown in Figure 3.6. The various speeds provided an error
check and a larger sample from which to average. From these data, the number of sam-
ples, N, between the rising edges of the two signals could be found. Since the data fre-

quency, f, and the speed of the test, to , were known, the phase lag, 0t , was found by:

NO.)
-I- _ 3.14

37

 

 

where (I 1

speed. Q .

where Q is

T0:
was increas
leading edg
ments. Bec

100 images

3.4 Hot-wit

HOl-t
nature of the
SimUItaneOUS
downstream (
3\‘eraged USir.
continuous sa.
that “fire nor a
UfatiOn and CC]
new). {aanatc
it

let hOt‘ere 1

3.4.210110Wed b

where or is in radians. With this information the proper delay time, rd, for a given fan

speed, 82 , is simply:

(3.15)

{019

where Q is in radians/second.

To acquire data at the different phase positions between the blades, the delay time
was increased according to the increase in phase lag. Since there were 40° between two
leading edges, data were acquired from 0° (leading edge) to 35° in five degree incre-
ments. Because of the size and number of files involved in the phase averaged data, only

100 images were recorded for each phase angle.

3.4 Hot-wire Measurements

Hot-wire anemometry data were acquired to further evaluate the instantaneous
nature of the flow field in the wake of the fan. A double-X hot-wire probe was used to
simultaneously measure three components of velocity at several radial locations just
downstream of the fan exit plane for the selected shroud conditions. The data were phase
averaged using the signal from the optical encoder. Since the hot-wire measurement is a
continuous sample of a the velocity field, these results give detailed statistical information
that were not available from the other measurement methods. The hot-wire sensor config-
uration and equipment is described in Section 3.4.1. The hot-wires were calibrated in a
newly fabricated calibration unit designed specifically for the purpose of calibrating dou-
ble-x hot-wire probes. A detailed description of this calibration unit is given in Section

3.4.2 followed by the calibration method in Section 3.4.3. The actual method of acquiring

38

 

 

the data is

tion of the

3.4.1 Hot-
The

of two x-vv
a i45° ang
seen in Fig!
active sensi
and soldere.
nector so [h
Wires were I
I)‘pical noise
A the
achlSltlon 1C
[hi‘nnistm ha

FCSPOnSe Of u

34‘2 Callbra

the data is described in Section 3.4.4. Finally, the processing algorithm and the calcula-

tion of the statistical quantities of interest is given in Section 3.4.5.

3.4.1 Hot-wire Probes, Anemometers, and Thermistor

The double-X hot—wire probe consisted of four tungsten wires in an arrangement
of two x-wire pairs rotated 90° to each other. The wires on a single x-array are oriented at
a i45° angle to the probe axis and spaced 1 mm apart. An illustration of this probe can be
seen in Figure 3.7. The four separate 5 um tungsten wires were 3 mm long with a 1 mm
active sensing length in the center. The two ends of the wire were copper plated to 30 um
and soldered to stainless steel broaches. These broaches were then wired to a BNC con-
nector so the output signal could be recorded by the data acquisition device. .The hot-
wires were driven by four DISA type 55M10 constant-temperature anemometers with a
typical noise level of 2 mV RMS (Morris, 1997).

A thermistor was placed in the vicinity of the hot-wire during calibration and data
acquisition to compensate for ambient temperature fluctuations between data points. The
thermistor had an accuracy of i0.2 K, a sensitivity of 2.03 K/Kohm, and a frequency

response of 10 Hz at 293 K (Bohl, 1996).

3.4.2 Calibration Facility

The hot-wire calibration facility allowed the double-X hot-wire probe to be cali-
brated over a range of velocities and angles. An illustration of the calibration facility is
shown in Figure 3.8. A centrifugal blower was used to pressurize a small plenum creating
a velocity at the outlet of the calibration unit. Multiple screens just downstream of the ple-

num inlet (blower outlet) were used to dampen large pressure and velocity fluctuations

39

 

 

and inhere
straight an
The hot-vv
A brief st
ensuring t1
Tht
axes as see
ing or re-p
#NF90-3 or
The control
515131)1 ng cor
the desired a
questionable
I010 V011 ran
10 the hot-W11
angle. This Pl
angle on borh l
31 each angle “
age was FCCOrd

angle aleach Ca

3.4.3 Cal", Pa 110

A quasi-“

and inherent unsteadiness. The smooth converging section preceding the outlet ensured a
straight and steady outlet flow. The rectangular outlet was 2.5 cm wide by 21.5 cm long.
The hot-wire was aligned with the center of the outlet just downstream from the exit plane.
A brief study showed that the boundary layer thickness at the exit plane was minimal
ensuring that the hot-wire was in the inviscid core.

The calibration facility had the ability to rotate the hot-wire around two different
axes as seen in Figure 3.9. This allowed both x-wire arrays to be calibrated without mov-
ing or re-positioning the hot-wire probe. Two stepper motors controlled by a Velmex
#NF90-3 controller were used to rotate the probe in the horizontal and vertical planes.
The controller was interfaced with a computer running a custom Matlab code from which
stepping commands were sent. A specific amount of steps were sent to move the probe to
the desired angle. However, the accuracy in determining the angle from this method was
questionable because of skipped or rrriscounted steps. Therefore, a potentiometer with a 0
to 10 volt range was attached to the output shaft. The potentiometers were calibrated prior
to the hot-wire calibration such that the output voltage was directly proportional to the
angle. This process involved manually moving each traverse to a physical stop at a known
angle on both the positive and negative extremes of the angle range. The voltage recorded
at each angle was then used to find a linear fit. During calibration, the potentiometer volt-
age was recorded along with the hot-wire anemometer voltage to accurately resolve the

angle at each calibration point.

3.4.3 Calibration
A quasi-steady method was used to calibrate the double-X hot-wire probe over a

range of i36° in approximately 6° increments for each x-array. The first x-array was

40

 

calibrated
in the hon
calibratior
traverse b
This resul
each an glc

s over 3 4(
with a 10 '
noulli equa
the unsteac
Bernoulli Ct
several steat
Confirming r
The v

UliSIOr Were a

data “Ere user

where E IS the

IS the pitch an 91.

!

Ofm .
e tam.

n

a .
lgonthm raga/(CC

of . .

”100 .
. Despite Sui

calibrated through each of the 13 angles using the horizontal traverse by rotating the probe
in the horizontal plane about the vertical axis of rotation. Immediately following the first
calibration, the second x-array was calibrated through the same angles using the vertical
traverse by rotating the probe in the vertical plane about the horizontal axis of rotation.
This resulted in a total of 26 calibration files for a full double-X probe calibration. At
each angle, the flow speed was increased from a near zero velocity to approximately 30 ml
5 over a 40 second period. The static pressure in the calibration unit plenum was measured
with a 10 Torr Baratron pressure transducer described in Section 3.2.1.1. Using the Ber-
noulli equation, the velocity was calculated for each sample acquired. It was assumed that
the unsteady effects from varying the velocity were negligible justifying the use of the
Bernoulli equation. As noted in Morris (2002), the assumption was tested by performing
several steady-state calibrations over the same velocity range as a quasi-steady calibration
confirming that the differences were not measurable.

The voltage signals from the pressure transducer, hot-wire anemometers, and ther-
mistor were acquired during the 40 second period at a rate of 1 kHz. For each sensor, the

data were used to fit the equation

52 = An) + 8000"”) (3.16)
where E is the voltage from the hot-wire anemometer, Q is the measured velocity, and y
is the pitch angle of the probe. The best fit equation was found by iterating through values
of the ‘n ’ variable until the minimum standard deviation was found. This processing
algorithm resulted in 13 separate transfer functions for each of the four sensors. Because
of a noisy pressure signal, typical standard deviations of approximately 0.3 m/s were com-

mon. Despite substantial flow conditioning downstream of the blower (upstream of the

41

 

 

calibratio
tions on t1
ters and 1
signal no
higher vei
the mean -
with the p
same trent
15th order
C211 standart
deviations .-
tions.
Beca
1361016 and 21
the calibratic
rePC'ated. A(
to Process the
110m {be P F€~C
this thesis and

the ”suits of m.

3.4.4 Data ‘4th

calibration unit outlet), the high standard deviation may have been due to pressure fluctua-
tions on the outlet side of the blower. Further investigations concluded that the anemome-
ters and pressure transducers were not responsible for this noise. The noise from the
signal was clearly fluctuating through a mean value and was much more prevalent at the
higher velocities. To remedy this problem, a 15th order polynomial curve was fit through
the mean of the pressure signal as a smoothing function. A representative pressure signal
with the polynomial fit is shown in Figure 3.10. Since the hot-wire signals displayed the
same trend as the pressure signal, each of the four hot-wire signals were also fit with a
15th order polynorrrial smoothing function. The combination of these actions led to typi-
cal standard deviations of approximately 0.02 ans which is representative of the standard
deviations achieved with the more time consuming discrete point, steady state calibra-
tions.

Because of the nature of hot-wire “drift” over time, a calibration was performed
before and after each experiment termed the “pre—calibration” and “post-calibration”. If
the calibration did not agree well, the data were discarded and the experiment was
repeated. A common method of checking for disagreement between the calibrations was
to process the post-calibration files with the calibration information (A, B, and n variables)
from the pre-calibration. Calibration drift did become an issue with the data presented in
this thesis and the uncertainties predicted as a result of drift will be presented along with

the results of the each experiment.

3.4.4 Data Acquisition Procedure

42

 

 

Ho
mm (r/R
trailing ed;

Pre
digital can
each of the
the hot-vvir
investigate
resentati ve I
hot-wire prc
device mour
angle to be a.

the probe P05

Hot-wire measurements were acquired for 17 radial locations, in increments of 10
m (r/ R = 0.028 ), at an axial distance of 35 mm (z/ R = 0.098 ) downstream from the
trailing edge of the shroud for the selected shroud and operating conditions.

Preliminary tuft observations were made to estimate the mean flow direction. A
digital camera configured with a long exposure time captured the fluctuating tuft image at
each of the 17 locations across the fan exit plane for the shroud conditions tested during
the hot-wire data acquisition. The camera was situated at two different orientations to
investigate the fluctuations in the axial-radial plane and the radial-tangential plane. A rep-
resentative tuft image is shown in Figure 3.11. This information was then used to align the
hot-wire probe with the mean flow direction. The probe was aligned with a positioning
device mounted on the hot-wire traverse described in Section 2.2.3 allowing the pitch
angle to be adjusted by rotating the probe around two axes. A schematic representation of
the probe positioning device is shown in Figure 3.12.

Hot-wire measurements were acquired for 60 seconds at a sampling rate of 11.4
kHz resulting in a total of 684,000 samples per radial location with the double-X hot-wire
probe. This sampling rate corresponded to one sample every 05° of fan rotation for a fan
speed of 950 rpm. The x-array hot-wire configuration can only measure angles up to
approximately 21:36° since any angles greater than the calibration limits cannot be reliably
resolved. Therefore, it was important to check the alignment of the probe during data
acquisition. The processing algorithm resolves speed and angle from each x-array as
explained in the following section. To check the alignment of the probe, histograms of the
angles found from each x-array were output from a modified version of the main process-

ing program. It was desired to center the histogram around a zero angle so that the probe

43

 

 

was alignel
data tO g6“
Provided a
improperb
[jv'ely- If t1
of the Prob

3.4.5 Dout‘
The
and angle fr
double-x 110?
one Speed/an
The t
From the tit?-

(QI‘QL‘) are f‘

n) correspond

for each sens

i=1.2,...,13 (M

intersection of
points to the in
we intensive

Samples for eacl

was aligned with the mean flow direction. The modified version processed one second of
data to generate these histograms. Although the data were not converged, the histograms
provided a good representation of the actual angles. Representative histograms of an
improperly and a properly aligned probe are shown in Figure 3.13 and Figure 3.14, respec-
tively. If the histogram was apparently skewed or truncated, the corresponding pitch angle
of the probe was corrected and the data point was repeated as necessary to produce a cen-

tered histogram.

3.4.5 Double-X Hot-wire Processing

The goal of the main processing algorithm was to solve for the instantaneous speed
and angle from each x-array resulting in two separate speed/angle pairs for the complete
double-x hot-wire probe. The following description details the method followed to find
one speed/angle pair. The process was simply repeated for the other x-array.

The voltages from the two sensors of the x-array will be denoted as E I and E2.
From the time series of recorded voltage pairs, (E 1,E2), the two predicted velocities
(Q ,,Q2) are found using equation (3.16) for all 13 sets of calibration coefficients (11,3, and
n) corresponding to each of the 13 calibration angles. This results in 13 separate velocities
for each sensor. That is, equation (3.16) is solved for Q1(E 1.)?) and Q2(E2,y,-) where
i=1,2,...,13 (Morris, 2002). The measured speed and angle, (Q,y), from the x-array is the
intersection of the two curves, and is found by interpolating between the three closest
points to the intersection using a polynomial fit (see Figure 3.15). The processing can be
quite intensive since this routine is performed for each sample out of the 684,000 total

samples for each x-arrays of the double-X hot-wire probe.

 

The 1
ponents of 1
probe axis \
denoted as
uon.the ve
array will b
redundant n
nate system.
tioned in a s
necessary to
(r: 9' 3) “he
with the t‘Ota

transformatio-

 

The speed and angle pair found from each x-array were decomposed into two com-
ponents of velocity. For purpose of discussion, the velocity component in line with the
probe axis will be denoted as 141 and the velocity vector normal to the probe axis will be
denoted as v where the subscript ‘1’ denotes the first x-array. Continuing with this nota-
tion, the velocity components in the axial and normal probe directions in the second x-
array will be denoted as u2 and w. Since the velocity components ul and u2 are a
redundant measurement, this results in three components of velocity in the probe coordi-
nate system. However, these velocity components are relative to the probe which is posi-
tioned in a spherical coordinate system using the device in Figure 3.12. Therefore, it was
necessary to transform the velocity components to the cylindrical fan coordinate system
(r, 0, z) where r is positive outward from the center of rotation, 9 is positive tangential
with the rotation of the fan, and z is positive downward from the fan exit plane. The

transformation equations are given as follows:

Ur = uxsinCcosé-vcochosé-wsinfi (3.17)
U9 = uxsinCsinfi—vcosfisin§+wsinCcosl‘, (3.18)
UZ = uxcos§+vsin§ (3.19)

where the angles C and g are the angles from the probe positioning device, and the sub-
script ‘x’ can be from either x-array or an average of the two.

Post-processing of the initial speed and angle calculations were performed to
obtain the desired velocity data and statistical quantities. These include: the time mean
and rms of the velocity components, the phase averaged mean and rms, and axial vorticity.

The time mean of the velocity components was calculated according to

45

 

and the rr

The velocr
Phase ave:
velocity fit
not be seer

nents “Em

and the phas

U(r,e,z) = 2 U, (3.20)
= l

l
N
i
and the rms of the velocity components was given by
l
.. l N 2 2
U(r,9,z) = N——l Z (Ui—U) . (3.21)

i = l
The velocity data were phase averaged using the phase trigger described in Section 3.3.3.
Phase averaged data provides a view of the blade-to-blade dynamics and the resulting
velocity field. It can be used to discern distinguishing features of the flow field that can-

not be seen from the time mean information. The phase averaged mean velocity compo-

nents were computed by

[V

((7)0, o, z, a) = $2 mania) (3.22)

i = l
and the phase averaged rms calculations were computed by
1
N 2

(i/xr, o, z, a) = 117%? Z [Ui(iT+ta)- (U)(r,o,z)]2 (3.23)

i = 1
where T is the time for one full revolution of the fan and ta is the phase lag for a given
angle of 0° < 0t < 360°.

The phase averaged velocity magnitude was also of interest and was calculated by

 

_ 2 _ 2 _
(M) = J<U,> +<Ue> +<u,>. (3.24)
The phase averaged axial vorticity was also found from the velocity data. It was

calculated by

46

 

A second

equation 3

and a for“;

Where Ah is

(3.25)

 

8 U a U
((1)2)(7', 9, Z) = %(—('%—62— (36») .

A second order finite difference scheme was used to find the partial derivative terms in

equation 3.25. A central difference was used to calculate the interior points:

 

Ba) ai+ 1 " ai— 1
— = 3.26
ah ,- 2Ah ( )

and a forward difference was used to calculate the boundary points:

aa _ ‘3ai+4“i+1"ai+2

5’,‘ 2Ah

 

(3.27)

where Ah is the distance between data points.

47

 

 

A
A
_
—
l
=
t-
V
9
-
-
3'
I-
-
’v
—
\

 

AFRri

6C9}.

Fig“ re 3
that

 

 

l
l
|
g l
.é |
‘5’
= l
E".
° l
I‘ ' l
5,00 l — 5.5 in-lb l
(steady-state tare value '
l used for integral measurements) l
4.50 l l
4.00 l ,._ . ,- __ . -=.__1.___!_ ._._ .J.__._...- - ”fl :. . -__ 2 .. . ‘__._ . . .___ .l __.L-__L_l__l
0 20 40 60 80 100 l20 140 160 180

Time (min)

Figure 3.1 Torque measurements acquired over a period of time with the fan
removed from the shaft to measure a tare value

Radiator

l l ‘ llll II Il I ll . Ii l'~E3‘tl~HtII
lll lllilMIllflI, NIH in l. ‘Illlt’lmllllllll

l l

 

Radiator
support harm
a l

I
I
I
4-I

VIII/IIIIIII

     
   

 

 

   

  
   
  

   
   
 

     

II.

  
 
 

   

   

    

III

 

IIIIIIIIII TIIII

AFRD Upper \ /
Receiver

(

Figure 3.2 Schematic representation of the pressure rise across the fan blades. Note
that the quantity, APfan, was actually calculated from two separate pressure

measurements‘ (Preceiver " Patm) ' (Pradiator ' Patm)'

48

 

interr
“in

Figure 3.3 ..

“time 3.4 A :

 

 

 

 

 

 

 

 

 

 

 

 

 

lst frame
”I , , ,
interrogation ’ ’ ’ ’ ’
window ” ” 7 i 7
2nd frame
particle image
(double frame)

Figure 3.3 A schematic representation of the cross-correlation technique using in
PIV vector processing

., p Interrogation
y, ._ A Window

Neighbors
(50% overlap)

 

Figure 3.4 A schematic representation of the interrogation window overlap used in
PIV vector processing

49

 

Figllrej

 

 

 

 

 

Shroud

-0.2 / Fan Blade k

 

.5
h
llllrlllll

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

N O r
: Region 2 Region 1
0.2 :-
_ F an Drive Shaft
0.4 5
i— l l 1 L1 1 l l l l l l l l l 1 l l l l 1 1 l J l l 1 l l l l l l
0 0.2 0.4 0.6 * 0.8 1 1.2

1'

Figure 3.5 Schematic of PIV measurement regions relative to the experimental
configuration

50

 

 

Flgu

5.5,7 7 7- 77, 7 7 7 7-7 7 7 77-7 7 7777777

5* Phase Trigger (n=1) Phase Trigger (n=2)

9
9‘

Blade Position (n=1) Blade Position (n=2)

9’
“229‘. .4:

output volts

-‘ N
m noggin

 

 

 

 

 

0 0.5 1 1.5 2 2.5 3 3.5
time(seconds)

Figure 3.6 TTL voltage signals from the optical encoders for PIV phase triggering
taken at a shaft speed of 20 rpm

 

 

 

   

a) Side View b) Head-on View

Figure 3.7 Illustration of the double-x hot-wire probe.

51

Blower Inlet

   
  

Flow Conditioning

   

Traversing Mechanism

Hot-wire Probe

Figure 3.8 Hot-wire calibration facility (Top and side panel removed to show detail).

Horizontal Rotation Axis

I
SW mama 19391:».

Hot-wire Probe Holder

Figure 3.9 Front-view of the calibration facility. 'IWo axes of rotation allow the hot-
wire probe to rotated in the vertical and horizontal plane.

52

4.5

 

     

 

 

 

 

4 -
3.5
3 .
g” 2.5 .
§
O
> 2 -
1.5
l
1 .. d
0-5 + Voltage Signal
— Polynomial Fit
0: 135 2 2.5 3 3.5 :
Sample x 10

Figure 3.10 A representative pressure signal from a hot-wire calibration. The high
standard deviation was reduced by fitting the curve with a 15th order polynomial
function.

 

r‘ 7
“Hi-l.

Figure 3.11 A representative image showing a fluctuating tuft. This tuft was used to
find the mean flow direction at a given radial location in the wake of the fan.

53

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

, C 331%.In-_ a
g C
§ axis
fill: =.e.=.
l

 

 

 

Figure 3.12 Schematic of hot-wire probe positioning device.

10‘
2f; .,7. -37. . W ,7m

1.2l

Number of Samples

o .o .o .o
N A 0" 00
44 T' 4 ‘T fl

15:2

 

-10 0 10 20 l 30 '40

-30 -20
Angle from Probe Axis (degrees)

Figure 3.13 Histogram of flow angles measured by an improperly aligned hot-wire
probe.

54

 

Number of Samples

 

0.5 _ , .
‘ ll I‘ I
01* fl ‘ M“__. ..
—40 -30 -20 -lO 0 10 20 30 40

Angle from Probe Axis (degrees)

Figure 3.14 Histogram of flow angles measured by a properly aligned hot-wire
probe.

Predicted Velocity (m/s)

 

Angle (degrees)

Figure 3.15 Intersection of the predicted velocities from the two wires in a single x-
array for use in the processing algorithm to find the measured velocity and angle.

55

4.0 Results and Discussion - Integral Measurements

The results of the integral measurements are presented and discussed in this sec-
tion. Integral quantities are important in assessing the overall performance of the cooling
fan with and without the aerodynamic shroud. First, the baseline measurements will be
presented along with a discussion of the characteristic features of the performance curve
in Section 4.1. This is followed by a discussion of the change in performance due to the

aerodynamic shroud in Section 4.2.

4.1 Baseline Measurements (Unpowered Shroud Condition)

Baseline performance and efficiency curves without a powered aerodynamic
shroud (x = 0) were acquired to establish a benchmark from which to compare the x :t 0
results. For all shroud conditions the integral performance data are presented as
w = limb, x) , which represents, in non-dimensional form, the pressure rise across the fan
as a function of volume flow rate and shroud power. The corresponding efficiency data
are presented as nsys = nsys(¢, x) where n was defined in equation 3.8. The functional
dependence of these two integral quantities was discussed in Section 3.2.2.

Each performance test followed the procedure outlined in Section 3.2.3. Three
separate baseline data sets were taken to demonstrate the repeatability of the measure-
ments. Plots of the three separate characteristic curves are presented in Figure 4.1 and the
efficiency curves are shown in Figure 4.2. Although the majority of the powered shroud
data presented in this thesis used a fan speed of 950 rpm, a baseline test at 700 rpm is
included in the plots to demonstrate that these data scale correctly with fan speed when
presented in non-dimensional form. The data agree well between separate tests confirm-

ing the repeatability of the measurements. It is noted that these data were obtained with

56

the radiator upstream of the fan inlet, and that the pressure coefficient represents the non-
dimensional pressure rise across the fan blades only, as defined by equation 3.3. These
data do not reflect the pressure rise across the complete fan and radiator system. Conse-
quently, the plots do not extend to (l) = 0. A plot of the pressure rise across the complete
fan and radiator system as a function of flow rate is presented in Figure 4.3 for reference.
Data were acquired at a total pressure rise less than zero to mimic a less restrictive
upstream element.

Referring to the baseline performance data in Figure 4.1, a particularly notable fea-
ture is the prominent stall region of this fan. Fan stall is a condition in which air flow sep-
arates from the suction surface of the fan blade. A “stalled” condition typically develops
at a high pressure rise (relative to the maximum pressure rise at a zero flow rate condition)
caused by an increased system resistance. On the characteristic curve, fan stall is identi-
fied as a plateau or drop in the pressure rise across the fan blades despite a lower flow rate.
It is usually followed by a second pressure increase until the point of no delivery. (No
delivery is also referred to as the “full shut-off” condition.) The above description of fan
stall is illustrated by the annotated baseline performance curve in Figure 4.4. As shown in
the figure, the onset of stall for this fan is approximately 4) = 0.256. The stall point is fol-
lowed by a relatively large “stalling dip” until ¢ = 0.132 , when the pressure coefficient
begins to increase again. Although the fan has regained the ability to produce a static

pressure rise, it operates in an extremely inefficient manner (see Figure 4.2).

4.2 Effect of Aerodynamic Shroud (Powered Shroud Conditions)
The shroud off (x = 0) condition is plotted along with selected powered shroud

conditions (x > 0) in Figure 4.5. A clear effect of increasing shroud power (increasing x)

57

on the performance of the fan is the movement of the stall point to a lower flow rate. Gen-
erally, as the shroud power is increased, this effect is increased, although a unique func-
tion cannot be derived. At the highest shroud power (x = 22.0 x 10_3 ), the stall point is
moved to d) = 0.186 , a 27% decrease in the stall point flow coefficient as compared to the
x = 0 condition. Even at the lowest shroud power investigated (x = 1.13 x 10—3 ), the
stall point flow coefficient is decreased by 11%. For flow rates higher than the stall point,
there is no measurable change between the baseline condition and the powered shroud
conditions. A clear degradation of performance is seen at the low flow rates where the
x > 0 characteristic curves cross the x = 0 curve. In this region, the pressure coefficients
are much lower for the powered shroud conditions compared to the baseline curve. The
combined effect of the aforementioned changes in the baseline performance data create
the appearance that the aerodynamic shroud systematically “shifts” the defining character-
istics of the baseline performance curve to lower flow rates as the x value is increased.
By delaying the stall point to a lower flow coefficient, the effective operating
range of the fan is increased. This point is further clarified by considering the fan’s per-
formance curve along with a system load curve. A system load curve is a parabola repre-
senting the pressure drop across the restrictive elements upstream or downstream of the
fan as a function of volume flow rate through those elements. When plotted on the same
axes, the intersection of the two curves provides the operating point of the fan. Two sys-
tem load curves are presented along with a baseline condition and the x = 8.64x10_3
powered shroud condition in Figure 4.6. Given the load curve denoted as “System load
curve 1” in the figure, one can see that the operating point is still within the desirable oper-

ating range of the fan. If the system resistance was increased, for example by the addition

58

of an air conditioning condenser to a standard radiator, thereby increasing the system
resistance represented by “System load curve 2”, the operating point would now lie in the
stalled region of the baseline, x = 0 , condition; however, the fan would not yet be stalled
in the x = 8.64x10_3 shroud condition. Hence, the advantage using an aerodynamic
shroud is clear.

The aerodynamic shroud may prove to be beneficial by extending the useful oper-
ating range of this test fan, however, this gain is at the expense of the fan’s efficiency as
seen in Figure 4.7. The measured efficiency accounts for the added power from the
shroud and is defined by equation 3.8. The plot shows that the efficiency is monotonically
decreased as the x value is increased. At the most energized shroud condition,
x = 22.0 x 10—3 , the efficiency is lowered from 49.4% to 41.1% at its peak. As expected,
this is the worst case. However, even at the lowest shroud power (x = 1.13 x 10_3 ), the
efficiency is decreased to 47.0% at its peak (cf. 49.4%).

It is difficult to compare the results of the present study to the previous study by
Morris (1997). In that experimental investigation, the test fan, boundary conditions, and
test conditions differed considerably from the present study. Two main differences are the
geometry of the shroud outlet (a sharp edged rectangular shape in the Morris study) and
the magnitude of the tip clearance. Specifically, the test fan used for the Morris study had
a tip clearance of 20.7% of blade span, compared to the 4.0% tip clearance for the current
test fan. Despite the differences, the present study confirmed that shroud power variable
x “can be used to fully characterize the contribution of the shroud” (Morris and Foss,

2001).

59

In the data presented in Figures 4.5 and 4.7, it necessary to note that the data points
shown only include the measurements acquired “in one direction.” That is, the plot only
includes the data acquired while the operating point was moved from a high flow rate to a
low flow rate (increasing pressure rise). It is displayed this way for clarity and to show the
distinct separation of the different characteristic curves. However, the data were also
acquired while moving the operating point from the low flow rate condition to the high
flow rate condition (decreasing pressure rise). In doing so, hysteresis effects were discov-
ered in the x > 0 data. The x = 1.13 x 10.3 characteristic curve is the most prominent
case of this and is shown in Figure 4.8 as an example. The significant hysteresis can been
seen in the stall area where the stall point occurs at a lower flow rate when loading the fan

from a high flow rate to a low flow rate. The two stall points are labeled as 4) nd

incr a
ode” on Figure 4.8 to denote the stall point when increasing and decreasing the flow rate.
This hysteresis was found to be repeatable and, although it appears in most of the x > 0
data, it is less apparent as the shroud power is increased.

A similar occurrence of this hysteresis can be found in airfoil (stall recovery) and
compressor research, and has been referred to as a “hysteresis loop.” Experimental data
show that the angle of attack of an airfoil that has stalled must be decreased considerably
beyond the original stall angle for flow reattachment to occur (Mittal and Saxena, 2000).
A hysteresis loop also occurs in compressor stall and effort has been directed at studying
the cause (see D’Andrea et al, 1996).

Delaying or preventing stall is also a topic of interest in compressor research. If a

low speed axial fan (e. g., an automotive type cooling fan) stalls, the consequences may be

lower efficiencies, a lower pressure rise, and noisy operation. The same consequences are

60

 

relevant in compressors. However, compressors may also exhibit excessive vibration, and
extreme degradation of performance due to the higher pressures and rotational speeds dur-
ing operation. Considering that compressors are often used in applications in which stable
performance is critical, stall is unacceptable. Thus, a considerable amount of effort is
directed at this subject. In specific instances, active flow control has been used as an
attempt in controlling the stall. One technique was successfully used by Day et al. (1993)
in which the general concept was similar to that of the aerodynamic shroud. Specifically,
it involved injecting air into the tip region of the compressor blades using 12 injection
valves spaced evenly around the circumference upstream of the inlet. Using this method,
the author was able to delay stall and extend the useful operating range by 6.0%. Many
researchers have also documented the role of tip clearance in delaying stall. Smith (1970)
showed that increasing the tip clearance between the blade tips and the casing of a low-
speed compressor caused the onset of stall to move to a higher flow rate.

Despite the similarities in the results, it is noted that compressor studies are only
partly relevant to the present study involving an axial flow fan due to the difference in
design and application. For instance, a compressor usually has a high solidity (chord
divided by blade spacing) leading to greater interaction between two consecutive blades.
The hub-to-tip ratio is much greater in compressors than axial flow fans as well. In terms

of operation, compressors typically produce a much higher pressure rise.

 

 

 

 

 

 

 

 

0.50
0'45 :i‘I. l- Trial 1 -950 rpm
__ A l4 Trial 2 - 950 rpm
0'40 g -. ‘0 Trial 3 - 700 rpm
A
0.35 b " .‘M"
9 H A Q
0.30 ~ ‘0 . ..P“"‘ ’1‘...-
»- ‘ .0 ‘H A 5
‘5 A"
0.25 ~ A“!
h .
0.20 m a
t. o
0.15 —
010 H i l .2 l . l r l i I i l .- 1 i L .
0.00 0.05 0.10 0.15 0.20 0.25 0.30 0.35 0.40 0.45

¢

Figure 4.] Baseline performance data for three independent tests

 

 

 

 

 

 

 

 

55
50 j .‘M‘v-‘DI-
45 — .‘I‘ A
_ 9 ‘A3
40 f “o... a.
35 — I ’
’3 r "
33 30 - ‘
._ x.
s: _
25 . . r .
a rial 1 - 950 rpm
20 f ‘1’ A TrialZ -950 rpm
_ 0 Trial 3 - 700 rpm
15 F ‘1‘
I
10 ~— -
5 ; .J
0 1' r l r L r l r .I i l r l a t i 1 m
0.00 0.05 0.10 0.15 0.20 0.25 0.30 0.35 0.40 0.45

(l)

Figure 4.2 Baseline efficiency data for three independent tests

62

 

 

wsystem

 

lTlTlTlIlT—llllllelll

 

 

llll l l l i llllllll
0.05 0.10 0.15 0.20 0.25 0.30 0.35 0.40

(l)

Figure 4.3 Total system pressure rise plotted as a function of volume flow rate in
dimensionless form

 

-0.6 i

0.00 0.45

 

0.50 .
0'45 : Stall Point
0.40

0.35
0.30

0.25

 

0.20

Desirable

0.15

 

l

l

l

l

Stalled Flow

1 l

L

l

1

Operating Range

1

l

l

l

i

l

I

 

 

0.05

0.10

0.15

0.20

0.

5

0.30

0.35

0.40

0.45

It

Figure 4.4 Baseline characteristic curve showing the stall point and separate regions
of the characteristic curve

63

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

0.50
x ‘ 10-3
0.45 -I- 0.0
—t- 1.13
0.40 - _ ._ 3.00
-O - 4.13
0'35 ..._ 7.30
a 0 30 ‘7" 8'64
' x - at- 10.1
‘~,. -0 - 12.2
0'25 "t —-e— 22.0
0.20 —- '
0.15 L
O 10 _ l L 1 l l l l l L l l l l l l L 1 l 1
0.00 0.05 0.10 0.15 0.20 0.25 0.30 0.35 0.40 0.45 0.50
It)
Figure 4.5 Performance data for selected shroud conditions
0.5
: System Load Curve 2 l
7 + x=0.0
0.4 —- + x=8.64
K + Load Curve 1
: \ -+- Load Curve 2
0.3 ._
9 E ‘ ‘
0.2 or
: System Load Curve 1
0.1 m
0.0 lb ‘ l g i l l L J l
0.0 0.3 0.4 0.5

 

Figure 4.6 Two performance curves, a baseline condition and a representative
powered shroud condition, plotted with a low and high system resistance curve to
demonstrate the advantage of delaying stall.

64

 

 

 

 

 

 

 

 

 

 

 

 

 

 

x‘10'3
+0.0
-t— 1.13
-0- 3.0
—0 - 4.13
°\° -o—7.30
v -+- 8.64
C -x- 10.1
-o - 12.2
—II— 22.0
5 .
O 7 l l i l l 1 IL 1 i l l l l i l l l l
0.00 0.05 0.10 0.15 0.20 0.25 0.30 0.35 0.40 0.45 0.50
4)
Figure 4.7 Efficiency data for selected shroud conditions
0.50
0'45 :— ¢decr=0.227 ¢incr=0.245
0.40 I—
035 *-
3 _
0.30 *—
O.25 ._
0.20 r—
0.15 F
0.10 — l l 1 L L l 1 L L l g l l l l l l
0.00 0.05 0.10 0.15 0.20 0.25 0.30 0.35 0.40 0.45
It

Figure 4.8 Performance data at x = 1.13 x 10_3 showing the stall point hysteresis

65

5.0 Detailed Velocity Measurements

Hot-wire and Particle Image Velocimetry (PIV) measurements were acquired to
analyze the detailed mechanisms responsible for the integral performance results for three
selected operating conditions. The procedure in which the data were acquired for each
measurement method was discussed in Sections 3.3.3 and 3.4.4. The point-wise, phase
averaged measurement of the three components of velocity at a fixed axial location down-
stream of the fan is provided by hot-wire anemometry and is presented in Section 5.1. The
PIV data provide a spatial view of the velocity field using both a time averaged and phase
averaged technique. Additional flow field interpretations are available from these data
and the results are presented in Section 5.2.

Three different shroud conditions were selected for the detailed velocity measure-
ments including 1 = 0, 8.64 x 10-3, and 22.0 x 10"3 representing a “shroud off,” “low
power,” and “high power” shroud condition. The focus conditions of the research were
the operating points denoted as ‘A’ through ‘D’ as labeled on Figure 5.1 with the corre-
sponding performance curves. Table 5.1 contains the significant details of each operating
point. The stall point for the corresponding shroud condition is listed in the table for refer-
ence. The specific points were chosen to compare both a stalled (off-design) condition (B)
and two powered shroud conditions (C, D) to a baseline design operating condition (A).
The hot-wire study, in particular, investigated the x = 0 and the x = 8.64 x 10_3 shroud
conditions including a stall condition. (Specifically, these are test conditions A, B, and C.)
The PIV study investigated all three x values but did not include a stall condition. (Spe-

cifically, test conditions A, C, and D were studied with PIV.)

66

Table 5.1 Test Conditions for Detailed Velocity Measurements

 

 

 

 

 

 

 

 

 

 

 

 

Test Measurement Stall Point
Condition Method x ‘l’ ‘l’ (p)
A HWA, PIV 0.00 0.27 0.35 0.256
B HWA 0.00 0.22 0.31 0.256
C HWA, PIV 0.0086 0.22 0.38 0.204
D PIV 0.0220 0.21 0.39 0.195

 

 

5.1 Hot-wire Measurements

Hot-wire data were acquired at a single distance of 35 mm (z/Rn.p = 0.098)
downstream from the trailing edge of the shroud. The hot-wire was traversed across the
exit of the fan recording data at 17 radial locations from r* = 1.181 to 1* = 0.731

(r* = r/ R ) in 10 mm increments (Ar/R".p = 0.028 ). The fan was operated at 950

up
rpm. A continuous time series was acquired for 60 seconds at rate of 11.4 kHz corre-
sponding to one sample every 05° of fan rotation. The experimental mesh shown in Fig-
ure 5.2 was produced from the data where each grid point represents a data point. A
schematic representation of the fan is included to illustrate the location and extent of the
hot-wire survey. The hub of the fan is at approximately r* = 0.368 , however, the loca-
tions of the measurement only extend inward to r* = 0.731 . This is due to a large region
of reverse flow. Hot-wire anemometry cannot discern the direction of flow, thus, the
extent of the data was limited to the measurable range. Note that the schematic represen-

tation of the fan is drawn at an arbitrary angular position to show the general position of

the fan and hub relative to the experimental mesh. The exact position of the fan blade re]-

67

ative to the measurement location was not known. Also note that the experimental mesh is
35 mm downstream from the shroud.

The presentation of the hot-wire investigation is divided into four sub-sections.
First, a description of the uncertainties involved in the measurement technique is pre-
sented in Section 5.1.1. This is followed by the presentation of the phase averaged veloc-
ity results for the three velocity components in Section 5.1.2. The phase averaged
fluctuation intensity (RMS) values for each velocity component are also presented to ana-
lyze the unsteadiness of the flow field. These data are presented in Section 5.1.3. Finally,
contour plots of the axial vorticity for each shroud condition are presented in Section

5.1.4.

5.1.1 Hot-wire Measurement Uncertainties

The major cause of uncertainty in the hot-wire velocity data was due to the dis-
agreement between the pre-and post-calibrations. Several data sets were discarded during
the experimental program because of this issue. As described in Section 3.4.3, the agree-
ment between the calibrations was checked by processing one calibration data set with the
other calibration coefficients (e.g., process the pre-calibration data files with the coeffi-
cients found from the post-calibration). Once this initial check verified that the calibra-
tions were sufficiently close, the complete data set was processed separately with both
calibrations. From this information, a sample-by-sample comparison was performed to
determine the true error from the calibration “drift”. After removing erroneous velocity
data, a percentage difference between each pair of data points from the two separate data

sets was calculated.

68

Figures 5.3 through 5.5 contain histograms of the percentage difference for the
three velocity components and velocity magnitude for each shroud condition. Although
the histograms presented are from one radial location, the results are representative of the
complete data set. It was desired to have error less than i2% for all velocity data. How-
ever, extreme difficulty was experienced in achieving this window of error for all velocity
components since the signal from all four wires of the double-X hot-wire probe must
remain consistent over the course of the experiment.

The error histograms for condition ‘A’ are presented in Figure 5.3. The radial
component agrees very well as the histogram is centered about the zero axis. The tangen-
tial and axial components both have a mean error of 3%. However, the radial component
is also much larger in magnitude than the other two components. This is reflected in the
error histogram for the velocity magnitude which is centered at approximately 1%.

Conditions ‘B’ and ‘C’ both exhibit the same trends in the velocity component
errors as shown by Figures 5.4 and 5.5. The histograms of the radial components are cen-
tered very close to zero and are within i2% . The highest error was observed in the stalled
condition ‘B’ in which the histogram extends over 110% for the axial component. How-
ever, because of the high radial velocity and low axial velocity, the mean agreement in the
velocity magnitude is still only 2.2%. Determining which of the two calibrations would
be more appropriate to use in processing the experimental data is difficult if not impossi-
ble. Therefore, the two data sets were averaged on sample-by-sample basis. It is these
data that are presented in the subsequent section.

The phase averaged velocities were also tested for convergence of the mean and

RMS quantities. Several data points were tested and the same trend among them was

69

observed. Representative plots of the data converging about the mean are presented in
Figure 5.6 for a single data point from condition ‘A’. The plot shown includes the conver-

gence of the mean and RMS of the three velocity components.

5.1.2 Phase Averaged Velocity Data

Phase averaged velocity data are presented as contour plots for the first 90° of fan
rotation. The flow is assumed axisymmetric and more detail can be observed in this trun-
cated view compared to the complete data set. The mean, RMS, and magnitude of veloc-
ity are scaled by the fan blade tip speed U“.p and are defined as (17*), (13*), and (IV*I) ,
respectively.

Figures 5.7 through 5.9 show the mean radial velocity contour plots for conditions
‘A’, ‘B’, and ‘C’. It is immediately evident that conditions ‘A’ and ‘C’ are qualitatively
similar. A distinct blade passing is demonstrated by an area of high radial velocity fol-
lowed by a much lower radial velocity. These high velocity regions are spaced 40° apart
corresponding to the spacing of the nine fan blades. The radial location of the maximum
radial velocity is centered around r* = 1.08 in both conditions, however, the azimuthal
location in condition ‘C’ is 5° behind condition ‘A’. Condition ‘C’ also exhibits a higher
radial velocity, although the velocity variation is. much greater. It is assumed that the
annular jet stays attached to the surface of the shroud which curves outward in the radial
direction. This shroud flow may entrain the main flow towards the outer radial regions
accounting for the larger radial component. In condition ‘B,’ the effects of fan stall are
immediately apparent. A blade passing is not readily distinguishable since the radial flow
field is no longer tightly coupled to the fan blades. The radial velocity is also much lower

and nearly uniform in the outer region. From these data, it is quite apparent that the aero-

70

dynamic shroud changes the nature of stall and structure of the radial velocity field in the
wake of the fan.

The mean tangential phase averaged velocities are presented in Figures 5.10
through 5.12. Once again, the periodic blade passing is evident in conditions ‘A’ and ‘C’,
but not in the stalled condition ‘B’. The “cells” of high velocity related to the blade pass-
ing are centered near the tip region (r* = 1 ) in condition ‘A’. This point is moved closer
to r* = 1.1 in condition ‘C’, another indicator that the shroud flow entrains the main flow
radially outward. Unlike the radial component plots, the tangential velocity is much dif-
ferent between conditions ‘A’ and ‘C’. From these data, it is clear that the aerodynamic
shroud increases the tangential component of the flow field. It is also interesting to note
that the areas of the highest and lowest radial and tangential velocities in condition ‘C’
exist at the same nominal locations; see Figures 5.9 and 5.12.

Contour plots of the mean axial component of velocity are presented in Figures
5.13 through 5.15 for each shroud condition. It is noteworthy that the maximum axial
velocities occur well beyond the tip radius of the axial fan; note the peak values at
r* z 1.2 and the indication that the actual peak values may exist for an even larger radius
position. The azimuthal location of the peak coincides with the lowest radial velocity.
When the fan is stalled (condition ‘B’) the blade passing is evident in the axial component,
in contrast to the radial and tangential components where it is not visible. Condition ‘C’
exhibits larger axial velocity variations than condition ‘A’ between two passing blades.
As the fan blade passes a given azimuthal position, it appears that the magnitude of the
axial velocity is maintained and “swept” radially outward from the hub. It interesting that

the axial component is smaller than the radial component in this axial flow type fan. Pre-

71

sumably, this is due to the operating point at which the measurements were acquired. It is
well known that as the pressure rise increases and the flow rate decreases, the flow
becomes more radial. Venturi shrouds with a curved outlet are often used to increase the
usable operating range of the fan by accommodating this radial flow.

Contour plots of the velocity magnitude as calculated by equation 3.24 are pre-
sented in Figures 5.16 through 5.18. These plots are included to compare the contribution
of each velocity component to the overall magnitude. It is apparent that the flow is pre-
dominately in the radial direction for all shroud conditions. It is also apparent that veloci-
ties in condition ‘C’ are consistently larger in magnitude, although the flow rate in
condition ‘C’ is lower as reported in the integral measurements. It is important to note that
the shroud flow is incorporated in the hot-wire measured velocities, whereas, the shroud

flow in the integral measurements is subtracted from the total flow rate (see equation 3.2).

5.1.3 Phase Averaged Fluctuation Intensity (RMS) Data

The RMS levels of the velocity components for each shroud condition are pre-
sented in Figures 5.19 through 5.27 as contour plots. These data are an indicator of the
unsteadiness of the flow. It is well known that high velocity fluctuations contribute to
aerodynamic noise generation, and lower flow rates generally have higher fluctuation lev-
els (Fukano and Jang, 2003). The RMS levels for the stalled condition (condition ‘B’) are
slightly higher than condition ‘C’ and even higher than in condition ‘A’ according to the
radial component plots. This is expected because fan stall involves a high level of
unsteadiness and increased noise generation caused by the separated flow over the suction

surface of the fan blade.

72

The tangential and axial velocity fluctuations show much more uniform areas of
RMS levels. The highest levels of unsteadiness are observed near the blade tips in all the
components of velocity. The axial component, in particular, has the highest RMS levels
for each shroud condition. A blade passing is clearly shown in this component as lower
levels of RMS are observed in the inferred region between two consecutive blades fol-
lowed by the region of higher fluctuations. The higher RMS values are most likely from
the wake of the pressure side of the fan blade. In addition, the tangential component
velocities are generally the least unsteady of the three components as indicated by the
RMS plots. Quantitatively, this is expressed by averaging the discrete RMS values over
the domains for each component and each shroud condition. The values are presented in
Table 5.2.

Table 5.2 Spatially averaged RMS levels for each component and each test condition

 

 

 

 

 

 

Coiggfion < 0”) < ()9) < ()1)
A 0.0950 0.0688 0.1006
B 0.1264 0.0896 0.1298
C 0.1018 0.0832 0.1154

 

 

 

 

 

5.1.4 Phase Averaged Vorticity Data

The vortical flow field in the wake of the fan has a significant effect on the aerody-
namic performance and noise characteristics. Common vortex structures observed in axial
fan flows include the tip leakage vortex, the leading edge separation vortex, and the trail-
ing edge vortex. The phase averaged axial vorticity was calculated according to equation

3.25 in order to analyze one component of vorticity present in the wake of the cooling fan

73

for each shroud condition. The results were normalized by 252 following the method of
Morris (1997) and are presented as contour plots in Figures 5.28 through 5.30.

Regions of negative axial vorticity, which correspond to counter-clockwise rota-
tion, are evident in the periodic blade passing signal. These regions, which are understood
to be the signatures of the tip leakage vortex, exist before the region of high axial velocity
and after the regions of high radial and tangential velocity in condition ‘A’. Specifically,
the highest negative axial vorticity for the 90 degree plot shown is centered at 0 = 60°
(r* = 0.48 on the vertical axis) compared to 0 = 55° for the radial, 0 = 56° for the tan-
gential, and 0 = 65° for the axial component. These correlations are given in Table 5.3
for further reference. It is apparent that the negative vorticity is not directly associated
with any velocity component, but is nearest to the highest velocity magnitude correspond-

ing to the localized low pressure at the core of the vortex.

Table 5.3 Tabulated values of the locations of the peak values for various measured
variables for both shroud conditions

 

 

 

 

Test Centroid Centroid Peak Peak Peak Peak
Condition - ( (Dz) ((DZ) ((7,) (U9) (U2) ( | VI)
A 60.0 52.5 55.0 56.0 65.0 55.0

C 61.0 44.0 49.0 43.0 79.0 43.0

 

 

 

 

 

 

 

 

 

The aerodynamic shroud clearly changes the fan blade dynamics in terms of rota-
tional fluid in the wake of the fan. The negative vorticity appears to be weakened by the
aerodynamic shroud; however, the positive vorticity following this area is much larger in
magnitude. Consider the vorticity in the boundary layer (dominantly (39) on the surface
of the shroud. The, assumed to be impulsive, axial velocities associated with the blade

passing will give these 009 vorticity filaments a strong reorientation as {006(817z/r30)}

74

which is the equivalent of a “source term” for 0)Z > 0. It is suggested that this mechanism
is responsible for the a), > 0 concentrations following the (T32 < 0 of the tip vortex. This
argument is strengthened by comparing the results for conditions ‘A’ and ‘C’. (The (7)9
will be stronger for ‘C’ given the aerodynamic shroud Coanda jet). By the argument pre-
sented, one would expect the tip vortex to be weakened by the “gap filling” Coanda jet in

agreement with the observations.

5.2 Particle Image Velocimetry Measurements

Time averaged and phase averaged PIV measurements were acquired in the r — 2
plane at a fixed azimuthal position for the two shroud conditions used in the hot-wire mea-
surements, x = 0 and 8.64 x 10_3 , and an additional “high power” shroud condition,
x = 22.0 x 10-3. All measurements were acquired with a fan speed of 950 rpm. Operat-
ing points slightly before the stall points were chosen corresponding to the points ‘A’, ‘C’,
and ‘D’ as illustrated in Figure 5.1. Therefore, two of the three test conditions (A and C)
can be related to the hot-wire results. Since a single PIV measurement region covered a
limited amount of radial distance, two regions were used to capture a larger area of the
velocity field as shown in Figure 3.5. The original experimental program included acquir-
ing PIV data near the fan hub to observe the hub recirculation region, however, an equip-
ment failure with the PIV laser did not allow the acquisition of these measurements.
Nonetheless, the PIV data presented is useful in analyzing the velocity field and provides a
validation of the hot-wire measurements.

The PIV data include the velocity magnitude and the axial and radial velocity com-

ponents, where the magnitude is defined as

75

lUrzl = ,lfifu‘if. (5.1)

It is important to note that the velocity magnitudes between the PIV and hot-wire data can-
not be directly compared due to the absence of the tangential component in the two-
dimensional PIV measurements. Consistent with the hot-wire survey results, the PIV data

were scaled by Rt.p and Utip, where U“.p = 35.37 m/s. The relevant non-dimensional

1

variables are defined as: r* = r/Rtip, z* = z/Rtip, (7*(r, z) = (7(r, z)/U“.p, and

*
lUrz

 

= loam/“p.
The presentation of the PIV results are sub—divided into four sections. The uncer—
tainties involved in the PIV measurements are described in Section 5.2.1. This is followed
by the time averaged PIV data presented in Section 5.2.2. A flow coefficient was calcu-
lated by the integration of the PIV velocity data and compared with the integral measure-
ments. These calculations are explained in detail and the results are presented in Section
5.2.3. Finally, phase averaged PIV data were acquired to investigate the blade to blade

velocity field variation and these data are presented in Section 5.2.4.

5.2.1 PIV Measurement Uncertainties

Errors in the PIV measurements were primarily a result of the low validation of
vectors in specific regions. The causal factors were glare, low quality seeding, or poor
laser light. The latter was most certainly the cause of a great number of invalidated vec-
tors. Specifically, when using a dual head laser, both heads must be working properly to
obtain good correlations between the two laser pulses. In the present study, one laser was
observed to be more intense and covered a wider planar area as compared to the other

laser. The low validation was observed in the corners of each PIV region where the laser

76

 

light intensity was lowest, while nearly 100% of the vectors in the middle region were val-
idated. A contour plot of the percent of validated vectors for condition ‘A’ is shown Fig-
ure 5.31 and displays this trend. This plot is representative of the other shroud conditions
that were tested.

Additional errors were introduced by using two separate PIV regions. Although
considerable care was taken to align and calibrate the image when re-positioning the cam-
era, the slightest mis-alignment would lead to a mismatched overlap region. Also, the two
regions were acquired for completely separate conditions. Since the test condition
involved establishing three separate variables (flow rate, pressure rise, and shroud power),
it became extremely difficult to obtain the exact flow condition as the corresponding mea-
surement for the other region. It is difficult to quantify these errors; they are noted to

explain discrepancies in the data.

5.2.2 Time Averaged PIV Data

The velocity magnitude lUrz“! for condition ‘A’ is presented in Figure 5.32. The
largest magnitude appears to be just downstream of the tip clearance region from r* = 1
to r* = 1.06. This high velocity region continues in a “path” travelling radially outward
and does not appear to follow the contour of the shroud outlet. These data suggest that the
shroud design may be improved if a larger outlet radius were utilized. The predominately
radial component observed in the hot-wire measurements is evident in the PIV data as
shown by the streamtraces superimposed over the velocity contours. This strong positive
radial component has been observed in other axial fan research near the suction side of
blades (Inoune and Kuroumaru, 1984). The authors note that this may be caused by cen-

trifugal effects acting on the blade’s boundary layer. (The Coriollis “force” will also pro-

77

 

vide positive radial components of the velocity vector). A reverse flow is observed
downstream of the fan closer to the hub, however, the velocity magnitude in this area is
close to zero. It appears that the strong radial flow near the outer regions entrains this
stagnant fluid radially outward. Glare from the laser reflecting off of the fan blade did not
allow measurement of the seeded flow directly downstream of the fan blade. However,
the measurement region was within 8 mm downstream of the fan blade and it shows posi-
tive axial velocity for r* 2 0.64. These data suggest that only the outer 36% of the fan is
contributing to the net volume flow rate. Tuft surveys conducted by Morris (1997) at a
high pressure condition in the previous aerodynamic shroud study showed reverse flow
from portions of the fan blade close to the hub but not across the complete span. Thus, at
test condition ‘A’, the flow may be separated at the inner portions of the fan blade causing
reverse flow, while the rest of the fan blade is providing axial flow.

The radial component of velocity for condition ‘A’ is presented in Figure 5.33.
The highest time averaged radial velocities appear to be at the radial extent of the shroud
downstream from the tip clearance region. These data agree well with the hot-wire mea-
surements in the location and magnitude of the maximum radial velocity. To analyze the
agreement between the two measurement techniques, a time averaged radial and axial
velocity was found for each radial location of the hot-wire survey. Also, the PIV data
were extracted at the downstream location of the hot-wire measurements creating a direct
comparison. These two sets of velocities were then plotted together to create a graphical
representation of the agreement. The radial velocities for condition ‘A’ are presented in
Figure 5.34. It is observed that the two measurement techniques agree well at the higher

r* values of each PIV region. From this plot, it is probable that the radial velocity profile

78

 

closely follows the hot-wire measured velocities and that the left half of Region I is under-
predicting the radial flow field. The inner radial region shows that the hot-wire overpre-
dicts the velocity as compared to the PIV. Nonetheless, the disagreement between the two
PIV regions (as discussed in Section 5.2.1) is quite apparent.

6

The axial velocity component for condition A’ is presented in Figure 5.35 and
shows that the highest axial component occurs just downstream of the blade tip. The hot-
wire data suggest that this region is in the inviscid core between the fan blades. Reverse
flow is evident from the lowest level contours where the highest negative axial velocity is
approximately (72* = —0.06 or about 2 m/s. The agreement between the hot-wire sur-
vey and the PIV measurements for the axial velocity field is presented in Figure 5.36. In
contrast to the radial velocity field, the two regions of PIV appear to match well at the
overlap region. The measured hot-wire axial velocities appear to have the same profile
across most of the area, however, the magnitude is higher by as much as 14%. A correla-
tion coefficient, K02 = W/ (72 (79 , was calculated for each hot-wire data location and
the resulting values are in range of 0.182 to 0.558. Since these velocity fluctuations are
correlated, it is suggested that the PIV data are biased because of particle migration from
the light sheet for the larger U 2 values.

Figure 5.37 presents the velocity magnitude for condition ‘C’. Qualitatively, con-
ditions ‘A’ and ‘C’ appear very similar signifying that the aerodynamic shroud is clearly
preventing stall. A strong radial component still exists and the direction of the flow has
not changed. However, the area of reverse flow has moved outward slightly to r* = 0.65 .

The aerodynamic shroud could be creating a larger area of reverse flow, or it is possible

that the lower flow rate has caused a larger portion of the fan blade to stall. Also, the mag-

79

nitudes are slightly lower in condition ‘C’ as compared to condition ‘A’. This is in con-
trast to the hot-wire results which show that the phase averaged magnitudes are higher in
condition ‘C’ despite the lower flow rate. The difference can be attributed to the absence
of the tangential component in the PIV data since the hot-wire data showed a large tangen-
tial component for condition ‘C’.

Figure 5.38 shows the axial and radial velocities for condition ‘C’. Again, the con—
tour plots are very similar between conditions ‘A’ and ‘C’. The radial velocity is still
strongest near the outer region of the shroud but this “peak” region has moved slightly
more downstream. Figure 5.39 shows that the hot-wire measurements are predicting a
much higher radial velocity (maximum of 16% higher) than the PIV measurements, how-
ever, the shape of the radial velocity profile appears to be very similar. The axial velocity
contour plot in Figure 5.40 is nearly identical to condition ‘C’ except that the area of pos-
itive axial velocity is concentrated more at the blade tip and the stagnant region closer to
the hub has become larger. When compared to the hot-wire measurements (Figure 5.41),
the velocity profile is very similar, and, again, the axial velocities are approximately 16%
higher as measured by the hot-wire. For this test condition, the calculated values for the
correlation coefficient K92 are in the range of 0.067 to 0.462 suggesting that the error is
due to the particle migration previously discussed.

The in-plane velocity magnitudes for condition ‘D’ are presented in Figure 5.42.
The contour plot follows the same trend as was observed in the previous two test condi-
tions. The velocity magnitude is considerably lower across the measurement area and the
area of reverse flow has moved outward even farther to r* = 0.69. Streamtraces show

that the flow is more radial than conditions ‘A’ and ‘C’. Also, the contours near the over-

80

lap between the two PIV measurement regions do not match as well as the other test con-
ditions. Since this is a relatively large shroud power, it is most likely that the error was
introduced when setting the experimental condition.

Figures 5.43 and 5.44 present the axial and radial velocities for condition ‘D’.
Although the flow rate is very close to condition ‘C’, the velocity components show a
much lower magnitude. Following the trend of condition ‘C’, the tangential component
may be an even larger percentage of the total velocity, however, it is not measured in this
two-dimensional view of the velocity field. The disagreement between the two regions is

very evident from the component contour plots as well.

5.2.3 Calculation of Flow Coefficient

The PIV results were used to calculate a flow rate by integrating across the fan exit
flow area. The integration is performed along a control surface extending from the outlet
of the shroud to the fan shaft (see Figure 5.45) and revolved around the circumference.

The integral is defined as

shr

Q = l (V.n)(2nr)dr (5,2)
0

where R S h r is the radius to the radial extent of the shroud and the dot product V f: repre-
sents the velocity vector normal to the control surface, or in this case the axial velocity
component. Due to the lack of data between the fan shaft and the inner PIV measurement
region, the velocity was assumed to be zero in this area. In actuality, this may not be true

and the validity of this assumption can be tested by the results. Axisymmetric flow is also

81

assumed in this calculation since the measurements take place at a singular azimuthal
position.

The flow rates integrated from the PIV results are presented in Table 5.4 along
with the measured flow rate found from the integral measurements. Note that the mea-
sured flow coefficient is the total flow (measured by the flow meters) since the shroud
flow cannot be subtracted from the PIV data. The results show that the integrated flow
coefficient is consistently lower than the measured flow coefficients with a difference of
approximately 10.0% and 8.3% for conditions ‘A’ and ‘C’ respectively. Furthermore, the
10% difference for condition ‘A’ (no shroud flow) indicates that mass flow can only be
balanced to within 10% and the 8% difference for condition ‘C’ is most likely within the
uncertainty of the calculation. Despite this, the small difference indicates that a large per-
centage of the net flow rate is contained in the PIV measurement regions. Considering
condition ‘D’, the 25.3% difference is actually closer to 15% given the 10% difference in
the baseline condition. The large disagreement in this condition further proves that the
flow is not axisymmetric. This is a plausible explanation because the high pressures in the
shroud plenum will be more likely to create a non-uniform annular jet as a result of the
defamation of the upper Coanda jet surface.

Table 5.4 Values of the dimensionless flow rate found by integration of the PIV
results compared to the measured flow rate from the integral measurements

 

 

 

 

 

Test %
Condition W ¢m°asur°d ¢im°gm°d Difference
A 0.35 0.26 0.23 10.03
C 0.38 0.22 0.21 8.30
D 0.39 0.21 0.16 25.27

 

 

 

 

 

 

 

5.2.4 Phase Averaged PIV Measurements

82

Phase averaged data were acquired in the wake of the fan for conditions ‘A’ and
‘C. Eight phase planes, separated by 5° , were used to cover the complete area between
two fan blades starting with the laser aligned with the blade’s leading edge. Synchroniza-
tion of the PIV laser pulse with the fan blade was described in Section 3.3.3. For a given
phase position, 100 image pairs were acquired. The resulting contour plots are presented
in this section for both shroud conditions.

The results for condition ‘A’ are presented in Figures 5.46 through 5.53. Starting
with Figure 5.46 (at the leading edge), a small “cell” of high velocity is observed in the
wake and is centered at approximately r* = 1.1 , 2* = 0.08. This cell is marked by “or”
on the plot. It appears to originate at the blade tip just downstream of the fan blade. The
streamline directions show that this flow is strongly radial and that it follows the contour
of the shroud. A negative radial velocity is observed in the low velocity, reverse flow
region near the area closer to the hub (see B on the plot). At a point farther upstream, the
negative radial velocity is turned to the positive radial direction. The implication, for this
phase position, is that the flow is drawn inward until it is entrained outward by the high
velocity radial flow.

The progression of the high velocity cell is evident at the next phase position (5° ,
Figure 5.47) where it is now centered at r* = 1.15 , z* = 0.10 , see y. It still appears to be
originating from the blade tip. The negative radial velocity has decreased in the reverse
flow region as the flow is almost entirely in the negative axial direction. At 10° past the
leading edge (Figure 5.48), the high velocity cell (see 8) has moved even farther down-
stream and outward to the edge of the measurement area. The “trail” from which it

appears to originate is now moved farther out past the blade tip. This trend continues in

83

the next three phase positions: 15° , 20° , 25°. At 15° past the leading edge (Figure
5.49), the highest velocities extend directly downstream from the tip clearance region and
the flow has a larger axial component. As shown, the flow no longer follows the contour
of the shroud.

Figure 5.50 presents the contour plot at 20° past the leading edge. It is instructive
to note that the trailing edge of the fan blade occurs at approximately 22° past the leading
edge. It is observed that the highest velocity magnitudes occur at or near this phase posi-
tion. The location (downstream of the tip clearance region) of the “path” of high velocity
fluid suggests that there are strong centrifugal forces acting on the suction side of the
blade. This high velocity region (t) downstream of the tip region, is continued in the next
phase position (25° , Figure 5.51). At 30° past the leading edge (Figure 5.52) a new cell
of high velocity fluid (K) appears to form close to the blade tip. The direction of the flow
has also reverted back to the radial direction and a higher negative axial velocity exists
closer to the hub. Figure 5.53 shows that the progression of the velocity field through the
last phase position agrees well with the phase position at the leading edge.

The phase averaged PIV data for condition ‘A’ agree well with the general trends
of the phase averaged hot-wire data for condition ‘A’. Specifically, the PIV data show the
movement of the areas of high velocity magnitude radially outward as the fan blade passes
the azimuthal measurement location from leading edge to trailing edge. Referring to Fig-
ure 5.16, the hot-wire data show this general outward movement; however, the magni-
tudes cannot be compared directly because of the absence of the tangential component in
the PIV measurements. The PIV data also verifies the inferred regions from the hot-wire

data since the exact blade position was not known in those experiments. Namely, the

84

pressure side of the fan blade, near the leading edge (Figure 5.46), shows a strong radial
velocity followed by a change towards lower radial and higher axial velocity after the
trailing edge (Figure 5.51).

The phase averaged PIV results for condition ‘C’ are presented in Figures 5.54
through 5.61. Characteristic similarities and differences exist between the two test condi-
tions. For instance, the “cells” of high velocity that appeared in condition ‘A’ near the
leading edge are still apparent (see A on Figure 5.54). However, the location is not as far
downstream in condition ‘C’. This agrees with the 5° lag in the location of peak velocity
magnitude observed in the hot-wire data when comparing conditions ‘A’ and ‘C’ (Figures
5.16 and 5.18). A similarin observed when comparing the PIV conditions ‘A’ and ‘C’ is
that the maximum velocity magnitudes appear to be nearly equal. One would expect the
velocity to be lower in condition ‘C’ corresponding to a lower flow rate.

The phase positions of 5° and 10° past the leading edge show that the aerody-
namic shroud adds a greater radial component to the velocity field. Although this was
already observed in the time averaged data, the mechanism is more apparent in a phase
averaged plot. Specifically, Figure 5.55 shows that the flow is forced into an almost
entirely radial direction. This effect is even more prominent at 10° in Figure 5.56. More-
over, the flow near the comer of the shroud in the 10° and 15° phase positions is at a very
low velocity implying that either the wall jet created by the aerodynamic shroud does not
stay attached to the shroud, or that there is a secondary, rotational flow that exists at this
location. This results in a decreased area from which the flow exits the fan. This may par-
tially explain why the velocities in condition ‘C’ are comparable to those in condition ‘A’

despite the lower flow rate.

85

The effects from the stronger reverse flow imparted by the aerodynamic shroud
becomes apparent at 15° past the leading edge (Figure 5.57) where a small recirculation
region is observed at r* = 0.70 , 2* = 0.08 (see u on plot). It appears to be created as the
reverse flow interacts with the positive axial and radial flow in the same area. This recir-
culation is much larger at the 20° phase position (Figure 5.58) and is centered farther
downstream at r* = 0.77 , 2* = 0.16 (see v on plot). It does not appear to be connected to
the axial vorticity observed in the hot-wire derived vorticity data. Secondary flow features
such as this one lead to greater losses due to the enhanced dissipation of kinetic energy.

The next phase positions, 30° and 35° past the leading edge (Figures 5.60 and
5.61), show that the recirculation region is no longer visible. It is interesting that the flow
at these positions is more axial than the same phase positions as condition ‘A’. Thus, the
flow is more radial near the leading edge, and more axial in the area between the fan
blades as compared to condition ‘A’. This agrees well with the phase averaged hot-wire
data in which all velocity components have larger spatial variations in the powered shroud
condition as compared to the “shroud off” condition. This may also explain why a larger
positive vorticity was observed in the axial vorticity plots. Specifically, the change in
flow direction between two consecutive blades could cause this clockwise rotation to

develop.

86

 

0.50

 

 

 

_ -3
1‘10
0.45 + 0.0
-A- 8.64
0.40 -o- 22.0

 

1
3 0.35
0.30

0.25

 

0.20

 

 

OJSililililmlililililr
0.00 0.05 0.10 0.15 0.20 0.25 0.30 0.35 0.40 0.45 0.50

It

Figure 5.1 Selected shroud conditions labeled with the operating points used for the
detailed velocity measurements (see Table 5.1 for numerical values)

 

87

0.5

 

\ll\ll||lll||l|lllll|lllllllllllllllllm

“I“ I III II” I "III
‘m\\i\i‘\:,‘\‘\‘,‘lliii‘Ilel‘i‘illllllllIullInIlitlllllilimflfllllI/I/llllj‘rlvumflfl
\\‘\\‘\:‘\\I.ll\\‘\‘\‘i\\\ WM.(.InnnunuIiinInnunmm:u,,,,’l'lIIIi.,,/,',(;m0, m
\ .mI‘}.I\.I.IIi ll\l\\“”|.|\.n.ll\l‘ In. IIII IIIIIIIIuInu {II/II/.,,,"I/I.

.II I” ‘mu ‘.

I .. . .

,. . .V .

 
   

IIIi-Innm ‘ll‘Ill l in n”, .. .,

‘Il;“;".‘I,II5I II v Area of Plots
“ ' " / Figures 5.7 - 5.30

II.II
. .
IIII

I ..I. I .l l ;

”HI ruIIIIItIIgI‘ H

I III‘ ‘lllillll‘llnllltI‘I 1“ I l
.. . IIII, . .. \\\I
..,. .,,’ ”II..“ M Int. .. .III
.I II. II.... ...
” 1H“,mll‘l‘ll"IllllllllllllllfllIll”

.

I 1%,
( r

. 1.,

"I‘T'II
nun .
m “m“,
in
"l um
um

., .
/ .
/I 1. ”III, ”I .
a 1. Nu ‘l'IIIlIII‘mm I inuminnm
li/ll/I/llllullll/t'lllII/mml”llnlllIIllllllllllllllhlllllllul m
l/lllIm"“llltllllllllllllllllIl||l|l|||lllllllll\ll “
l"IImmmmuIIunimumumnnu

.ot
mu"
m\‘

\. -
“In
\“(IIuI

Figure 5.2 Mesh of hot-wire phase averaged data locations in r-0 plane at z = 35mm

(z/Rflp=0.098), Q shows direction of fan rotation

88

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

x10 x10
2.5 . . . 5
a) b)
:6 2 - § 4
Q.
g e
m 1.5 (:3 3
“-4 (H
O O
8 1 h“. 2
.D .0
E E
g 0.5 g 1.
A . O n
-10 -5 o. 5 10 -10 -5 0 5 10
Percent Difference Percent Difference
4 4
x10 x10
4 2 - - c
.. c) .. d)
(D Q.)
.— 3 'T 1.5
E" 2‘
(S! G!
U] U)
“5 2 “5 1
‘6 E3
.0 .D
E 1 S 05
Z Z
. 1 O . .
-10 -5 O 5 10 -10 -5 0 5 10
Percent Difference Percent Difference

Figure 5.3 Histograms illustrating the “sample by sample” difference between
velocities found using the pre and post calibration constants for Condition ‘A’: a)
mean radial component, b) mean tangential component, c) mean axial component,

and d) velocity magnitude

89

x10

 

N
m

8)

H
u—5 l): N

Number of Samples
9
L11

 

 

 

 

 

-5 o 5
Percent Difference

 

Number of Samples

 

 

 

-5 0 5
Percent Difference

Numb er of Samples

Numb er of Samples

 

 

 

 

 

 

 

 

 

 

3

2

1 .

O y 1

-10 -5 0 5 10
Percent Difference

x 10
4 -
d)

3 i

2 .

1 .

O , 1

-10 -5 0 5 10
Percent Difference

Figure 5.4 Histograms illustrating the “sample by sample” difference between
velocities found using the pre and post calibration constants for Condition ‘B’: a)
mean radial component, b) mean tangential component, c) mean axial component,

and d) velocity magnitude

9O

 

x10 X10
3 . . . fl - -

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

15
a) II)
31".: E
E” 2 . 3‘ 10 I
N (d
U] U]
‘H CH
O O
b b 5 J
. ,0 .
E’ a
Z Z
O 1 1 0 1 1 1
-l 0 -5 0 5 10 -l O -5 O 5 10
‘ Percent Difference 4 Percent Difference
x 10
2.5 m
d)
e a 2
CL CL.
8 E
g 8 1.5
‘H Q—i
O O
b b 1
.0 .0
E E
,2 .2 0.5 .
1 0 1 1
720 -1O 0 10 -10 -5 0 5 10
Percent Difference Percent Difference

Figure 5.5 Histograms illustrating the “sample by sample” difference between
velocities found using the pre and post calibration constants for Condition ‘C’: a)
mean radial component, b) mean tangential component, c) mean axial component,

and d) velocity magnitude

9]

 

 

 

 

 

 

 

 

 

 

   

 

 

 

‘3 5 — — 7——7 77 #7 u) 5-
% E I
I I
E E I I
e 9 I I
g OI g 0 W W l
i e
3: .—
.. D I
O I
°\°_5 ’13) . 4' ,, .. -.. .fi' °\°_ b).—._.._ _ ._ _l_—_.. .1.____-..._1 -l
200 4Olzlumber 888Iam pl e5800 1 000 200 40(Number Wgamplesaoo 1000
C 5 77 —-—--—- (I) 5. I «m .
CU I
I” E
z I
E E If g; l
9 .: l
g 0 g 0 W .
92 9 ;
a) ~93 .
3: a:
- o
2-519214 _e _______ __ a 4).. -_ . , I
200 4Olzlumber 89gample3800 1 000 200 40(Number tiggtamplesaoo 1000
C
m "*7 ’ ‘~ I — —‘. m 5 :‘ A:
o
2 E
E e
m
8 8 0
E g I
a’ a: l
a: ._
. I o
$-55 9)— . 1 g _ El o\°_ f) i .*_. 1 . -
200 40lzlumber 898ample5800 1000 200 4095lumber f(sifgamplesem 1000

Figure 5.6 Plots of convergence for a representative data point acquired in the hot-
wire measurements: a) mean radial component b) RMS of radial component c)
mean tangential component d) RMS of tangential component e) mean axial
component 1') RMS of axial component

92

 

1 _.
Level <Ur>
0.65
0.8 0.60
0.55
0.50
1. 0-6 - 0.45
0.40
~ 0.35
0'4 . 0.30
- 0.25
0.2 a 0.20
0 V 1

 

 

0 ‘ 02010.4 80.61" 0.8 1 1.2

Figure 5.7 Phase averaged mean radial velocity for condition ‘A’
(x=0.0, ¢=0.27)

1.2

  

Level <U:>
‘“; 10 0.65

0.8 9 0.60
8 0.55
7 0.50
“ 0-6 6 0.45
5 0.40
— 4 0.35
0'4 — 3 0.30
2 0.25
02 1 0.20
0

 

 

0 ' A 0.21 ' 0.4 l 0.0 I l 0.8 1 1.2

Figure 5.8 Phase averaged mean radial velocity for condition ‘B’
( 0 22)

x=0- ,¢= .

93

  

Level <U:>
10 0.65
0.60
0.55
0.50
0.45
0.40
0.35
0.30
0.25
0.20

—NUJJBMO\\IOO\O

 

 

01 ‘ 0.2 I l 0.4 I i 0.6 0.8 1.2

Figure 5.9 Phase averaged mean radial velocity for condition ‘C’
(x=0.008, ¢=0.22)

0.33
0.30
0.27
0.24
0.21
0.18
O. 15
0.12

 

 

 

Figure 5.10 Phase averaged mean tangential velocity for condition ‘A’
(x=0.0, ¢=0.27)

94

  

Level <U;>
-- s 0.33
0.30
0.27
0.24
0.21
0.18
0.15
0.12

HNw-bUIQfl

 

 

01" 0.21 l I 10.41” 0.6 A 0.8 1 1.2

Figure 5.11 Phase averaged mean tangential velocity for condition ‘B’
(x=0.0, ¢=0.22)

Level <U;>

-- 0.33
0.30
0.27
0.24
0.21
0.18
0.15
0.12

v—an-IiUICNNOO

   

 

 

0111I111111111‘171r
0 0.2 0.4 0.6 0.8 1 1.2

Figure 5.12 Phase averaged mean tangential velocity for condition ‘C’
(x=0.008, ¢=0.22)

95

  
    

Level <U;>
3 9 0.45
7 0.40
0.35
0.30
0.25
0.20
0.15
0.10
0.05

l

#NUJ-bUIOQOO

 

 

01 L 0.2 1 L 0.4 L i 0.6 0.8 1 1.2

Figure 5.13 Phase averaged mean axial velocity for condition ‘A’
(x=0.0, ¢=0.27)

1.2

Level <U;>

 

 

 

'7': 9 0.45
0-8 . 8 0.40
7 0.35
* 7 6 0.30
“ 0-6 . 5 0.25
4 0.20
3 0.15
0'47 2 0.10
1 0.05
0.27
0’“ .1..I.1....1...
0 0.2 0.4 0.6 0.8 1 1.2

r*

Figure 5.14 Phase averaged mean axial velocity for condition ‘B’
(x=0.0, ¢=0.22)

96

  

9

t-INUJhtJIONflOO

 

 

Level <U;>

0.45
0.40
0.35
0.30
0.25
0.20
0.15
0.10
0.05

Figure 5.15 Phase averaged mean axial velocity for condition ‘C’

(x=0.0os, ¢=0.22)

11
,1 10

     

0.4 '

I—NM-kUIQN

0.2

  

 

 

0 1111111111

1 I l l 1 A .. _
0 0.2 0.4 0.6 0.8 l 1.2
r*

Level < lVl >

0.80
0.74
0.68
0.62
0.56
0.50
0.44
0.38
0.32
0.26
0.20

Figure 5.16 Phase averaged velocity magnitude for condition ‘A’
7)

(x=0.0, ¢=0.2

97

r*

r*

0.6

0.4 5

0.2 '

 

0

() .

1 1

1....

0.2

[111

0.4

l I l l l
0.6 0.8
r*

  

 

Level < lV.l >

 

0.80
0.74
0.68
0.62
0.56
0.50
0.44
0.38
0.32
0.26
0.20

Figure 5.17 Phase averaged velocity magnitude for condition ‘B’
(x=0.0, ¢=0.22)

0.6 ;

0.4

0.2 _

 

L 14L J_14_

 

0.4

0.6 0.8
r*

 
    

l

 

    

fiLevel < lVl >

0.80
0.74
0.68
0.62
0.56
0.50
0.44
0.38
0.32
0.26
0.20

Figure 5.18 Phase averaged velocity magnitude for condition ‘C’

(x=0.008, ¢=0.22)

98

  

10

—Nw.t>u-O\\ioo\o

 

 

0‘ l 0.2 ' ' 10.4 l l 0.0 A 0.8 l 1.2

Level <U:>

0.35
0.31
0.27
0.23
0.19
0.16
0.12
0.08
0.04
0.00

Figure 5.19 Phase averaged RMS of radial velocity for condition ‘A’

(x=0.0, ¢=0.27)

  

 

 

W 10
i -1 9
8
7
6
5
4
3
2
1
0 .. . I i I. . . I . .. 1 ._ ’_
0 0.2 0.4 0.6 0.8 1 1.2
r*

Level <U:>

0.35
0.31
0.27
0.23
0.19
0.16
0.12
0.08
0.04
0.00

Figure 5.20 Phase averaged RMS of radial velocity for condition ‘B’

(x=0.0, ¢=0.22)

99

  

, Level <U:>
0.35

 

 

0‘ l [0.2 l l 0.4 l 0.6 l 0.8 l 1.2

Figure 5.21 Phase averaged RMS of radial velocity for condition ‘C’
(x=0.008, ¢=0.22)

1.2

Level <U;>

fl 6 0.25

 

 

 

0.8 0.20
0.15
0.10
*“ 0-6 , 0.05
, 0.00
0.4 i
0.2 .
O 1 1 1 1 a
0 0.2 0.4 0.6 0.8 1 1.2

Figure 5.22 Phase averaged RMS of tangential velocity for condition ‘A’
(x=0.0, ¢=0.27)

100

  
  

Level <U;>
6 0.25

L 5 0.20
4 0.15
3 0.10
2 0.05
l 0.00

 

 

0L L 0.2 L L 0.4 L L 0.6 L L 0.8 1 1.2

Figure 5.23 Phase averaged RMS of tangential velocity for condition ‘B’
(x=0.0, ¢=0.22)

1.2

Level <U;>
I :1 0.25
I : ,I 0.20
135 0.15
0.10
0.05
0.00

       
 
   

0.8

.. 0.6»

HNWAUIQ

0.4 L

0.2 ;

 

 

 

0 LL . I 141 I 1 1 L I
0 0.2 0.4 0.6 0.8 l 1.2

Figure 5.24 Phase averaged RMS of tangential velocity for condition ‘C’
(x=0.008, ¢=0.22)

101

0.8
u 0.6 1
0.4

0.2

 

 

 

I t l l I l

0.6
r*
Figure 5.25 Phase averaged RMS of axial velocity for condition ‘A’
(x=0.0, ¢=0.27)

 

  

 

 

1.2

HNW-fimO‘

 

 

Level <U;>

*1

0.25
0.20
0.15
0.10
0.05
0.00

0.25
0-8 0.20
0.15
0.10
*-
I‘ 0-6 0.05
0.00
0.4
0.2:
0711114pglglllll 1
0 0.2 0.4 0.6 0.8 1.2
r*

Figure 5.26 Phase averaged RMS of axial velocity for condition ‘B’
(x=0.0, ¢=0.22)

102

0.8
*1. 0.6
0.4 _

0.2 L

 

 

 

00L L L 0.2 L L 0.4 L 0.6 0.8 1 1.2

Figure 5.27 Phase averaged RMS of axial velocity for condition ‘C’
(x=0.008, ¢=0.22)

Level <0);>

' 0.60
0.40
0.20
0.00
-0.20
-0.40
-0.60

t—INw-FKIIQQ

 

 

 

L L 0.8 1 1.2

C III
0
N
9L
#.
0
ex

Figure 5.28 Phase averaged axial vorticity for condition ‘A’
(x=0.0, ¢=0.27)

103

0.8
.. 0.6»
0.4L

0.2 .

 

 

 

00L L L 0.2 L L 04 L L 0.6 L 0.8 1 1_ 1.2

Figure 5.29 Phase averaged axial vorticity for condition ‘B’
(x=0.0, ¢=0.22)

 

 

<0);>
0.60
0.40
0.20
0.00
-0.20
-0.40
~0.60

 

Figure 5.30 Phase averaged axial vorticity for condition ‘C’
(x=0.008, =0.22)

104

 

: x
‘0-1 7 Percent
' Va_li_dated

0 :2:...,.;::;'~-9LI'.:*_LL"L'-‘°5I'5:.I:3g._ ., .. -- ~ . iii-erii-zéL 90

. .. - , 8O
70
60
50
40
30
20
10

 

  

 

1
o ”“1";

'Ifinl
O L. 'L"- .L’

0.1

0.2

    

0.3

 

0.4

lllllljj

I
.—
_
_
h
p
1—
F—
r-
l—
-
I—
p
r—
I-
b
h.
l-
I—
b
-
r
b
h
F-
-
h
-
_.
1.
-
I-
_
_

 

p—A
N

0.9 1.1

.o
00
p—

0.7

.o
ox
"1
«-

Figure 5.31 Representative contour plot of percent validated vectors for condition
(A?

105

 

l 12 0.55
- 11 0.50
10 0.45
0.40
0.35
0.30

B 8151

 

t—aNUJ-AUIONQOOO
O
N
O

0.4 _—

 

iijiIIIlII4_l

‘ l | I l l 1 l l L 1
0.6 0.7 0.8 0.2 l 1.1 1.2
I'

Figure 5.32 Time averaged PIV data: velocity magnitude in the r-z plane for
condition ‘A’ (x=0.0, ¢=0.27)

106

 

 
 
     

 

 

-01

Ur

0 0.50

0.45

0.40

0-1 0.35

*~ , 3 1::

0'2— 6 0.20

5 0.15

~ 4 0.10

0.3

7 3 0.05

2 0.00

0.4» 1 -0.05
.1..l.<. .1. 111 L. 1 1 .l. I I III 1 1 II
0.6 0.7 0.8 0.3 1 1.1 1.2

1‘

Figure 5.33 Time averaged PIV data: mean radial velocity for condition ‘A’. The
axial location of the hot-wire survey is indicated on the plot. The two ‘x’ marks
denote the radial range of the measurements.

 

 

 

 

 

0.60
0.50
0.40 L
*h T
[D 0.30 l—
0.20 0 PIV Region I
I PIV Region II
0_10 + HWA
0.00 _
-0 10 H .4] l | J l l L ‘_L l l l l l l l l l l l l l | l l
O 6 0 7 0 8 0.9 1 0 1 1 l 2 1 3

r*
Figure 5. 34 Comparison of the radial velocities measured by the hot-wire survey and
the two PIV regions for condition ‘A’

107

 

 

 

-0.1 p
:' U2
OI 0.40
0.35
0.30
0" _ 0.25
*~ : 2 3:2
0'2 5 0.10
4 0.05
3 0.00
0'3 2 -0.05
, 1 -0.10

0.4 r

 

IIII1I IiIIIIiIIIIIIIIIIIIIII
0.6 0.7 0.8 0.3 1 1.1 1.
r
Figure 5.35 Time averaged PIV data: mean axial velocity for condition ‘A’. The
axial location of the hot-wire survey is indicated on the plot. The two ‘x’ marks
denote the radial range of the measurements.

 

0.30

0.25 —

0.20 _

   
   

O PIV Region I
I PIV Region II

0.05 “‘ HWA

 

0.00 H

—0.05 _

 

 

_010 ' l_;l 1‘ l l | li;l;lil;l lnglgll l l
0.6 0.7 0.8 0.9 1.0 1.1 1.2 1.3

r*

 

Figure 5.36 Comparison of the axial velocities measured by the hot-wire survey and
the two PIV regions for condition ‘A’

108

2*

l T I T 7

 

0.1_

0.2 A

 

  

LeveI mg

12 0.55
11 0.50
10 0.45
0.40
0.35

u—aNUJ-FUIO'xfloob
O
N
C

Figure 5.37 Time averaged PIV data: velocity magnitude in the r-z plane for

condition ‘C’ (x=0.008, ¢=0.22)

109

U.
0.50
0.45
0.40
0.35
0.30
7 0.25
6 0.20
5 0.15
4 0.10
3

2

1

 

 

0.05
0.00
-0.05

0.4

IIITIII

 

lllll‘lllllllllli‘lL I_Lli'

0.6 0.7 0.8 032 1 1.1 1.2
I'

Figure 5.38 Time averaged PIV data: mean radial velocity for condition ‘C’. The
axial location of the hot-wire survey is indicated on the plot. The two ‘x’ marks
denote the radial range of the measurements.

 

  

 

 

0.60 IL
0.50 E
040 E 2
-x-
p“ 0.30?E
0.20 F '———“
‘ o PIV Regionl
010 L I PIV Region II
' if -o—HWA
0.00
‘— _l_l___gJ_l_J_l_.i_l__i_l‘l_l_l_L_l_l_l_l._L_l._l_l_l_
0.6 0.7 0.8 0.9 1.0 1.1 1.2 1.3

I-*
Figure 5.39 Comparison of the radial velocities measured by the hot-wire survey and
the two PIV regions for condition ‘C’

110

 

Level [7

11 0.40
10 0.35
0.30
0.25
0.20
0.15
0.10
0.05
0.00
-0.05
-0.10

000

 

t—‘NUJ-kUIONfl

 

9
.5
l l l I I

lllIlllllll'

I I l l l l l l I | l L l
0.6 0.7 0.8 032 l 1.1 1.2
1'

Figure 5.40 Time averaged PIV data: mean axial velocity for condition ‘C’. The
axial location of the hot-wire survey is indicated on the plot. The two ‘x’ marks
denote the radial range of the measurements.

 

0.35 l.
0.30 f
0.25 F
0.20 —
N I
ID 0.15

0.10

     
 

o PIV Region I
l PIV Region II
—o— HWA

0.05

 

0.00

-0.05

_0'10 Ll;l;l ll IglLIL‘L‘Q ll ll llI ‘1
0.6 0.7 0.8 0.9 1.0 1.1 1.2 1.3
r*
Figure 5.41 Comparison of the axial velocities measured by the hot-wire survey and
the two PIV regions for condition ‘C’

 

 

 

Ill

Level I U ml

12 0.55
11 0.50
10 0.45
0.40
0.35
0.30
0.25
0.20

 

 

p—anJ-hUIONQOOW

 

-lllllL"ll‘llll114l_l I [ll IIJ

0.6 0.7 0.8 0.32 l 1.1 1.2
1'

Figure 5.42 Time averaged PIV data: velocity magnitude in the r-z plane for
condition ‘D’ (x=0.022, ¢=0.21)

112

 

 

LII Ill-llll

l l l l l l J I I I I I
0.6 0.7 0.8 0.)? l 1.1 1.2
I'

Figure 5.43 Time averaged PIV data: mean radial velocity for condition ‘D’
(x=0.022, ¢=0.21)

 

' w 11 0.40
10 0.35
9 0.30
0.25
0.20
0.15
0.10
0.05
0.00
-0.05
-O.10

 

v-‘NUJ-hUIO‘flOO

 

0.6 ‘o.7‘ £08 10.9! 1 w1.1 11.2
r*

Figure 5.44 Time averaged PIV data: mean axial velocity for condition ‘D’
(x=0.022, ¢=0.21)

113

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

-0.6 _ —
-o.4 -
- Shroud —
: / \ _
-O.2 Fan Blade \ \
*N O; _____ _\ ..........
_ \
: \ Control Surface
0.2 _-
_ Fan Drive Shaft
0.4 _-
:IIIIIIIIIIIISILIIIIJIIIIIIIIIIIII
O 0.2 0.4 0.6 0.8 1 1.2
I'*

Figure 5.45 Fan and shroud configuration showing the control surface used for the
integration of the flow rate from the PIV velocity measurements

114

 

  

0 7 ‘ ‘ t

4 Level I UnI

_ - 3 ~1 7 0.60

0-1 _ k 6 0.50
*N _ 5 0.40
- 4 0.30

0-2 3 0.20

2 0.10

1 0.00

0.3 A

0.4 l

 

1141

L 0.8

I I l I I i
0.
,2

Figure 5.46 Phase averaged PIV data: velocity magnitude for condition ‘A’ at fan

1'11

1

 

I l I l I I
1.1

l
1.2

Iblade leading edge, note high velocity cell, a, and negative radial flow, B

III I

 

 

'1

1

l‘\
/
//

  

\

\N}

\.

 

  

Y

 

 

0 H .
I IUEI
0.60
0-1 _ 0.50
*N 0.40
0.30
0-2 _ 0.20
I 0.10
0.3 0.00
0.4 —
CI 1 , I J I I I I l I I I I I I I I 1 l 44%;; J
0.6 0.7 0.8 0.3 1 1.1 1.2
1'

Figure 5.47 Phase averaged PIV data: velocity magnitude for condition ‘A’ at 5
degrees past leading edge, note the high velocity cell, 7

115

 

0.1:

2*

0.2

 

0.3

0.4 -

 

H1111111111111 IIIIIIIIIII

I I I I I
0.6 0.7 0.8 0.3 1 1.1 1.2
1'

 

Figure 5.48 Phase averaged PIV data: velocity magnitude for condition ‘A’ at 10
degrees past leading edge, note high velocity cell, 8

 

 

 

 

0 .
Level IU'ZI
7 0.60
0.1 h 6 0.50
*N _ 5 0.40
V 4 0.30
0-2 3 0.20
_ 2 0.10
0.3 ~ 1 0.00
0.4 E
F I I I l I I I I l I I 1 I l I I I I l I I l 14 1 I l
0.6 0.7 0.8 0.3 1 1.1 1.2
I'

Figure 5.49 Phase averaged PIV data: velocity magnitude for condition ‘A’ at 15
degrees past leading edge

116

 

0.1:

2*

0.2 H

 

Figure 5.50 Phase averaged PIV data: velocity magnitude for condition ‘A’ at 20

‘11

L41 I 1
0.7

  

Level I Ur]

"NW-hUICh

0.60
0.50
0.40
0.30
0.20
0.10
0.00

degrees past leading edge, note the “path” of high velocity, 1

 

0.1 H

*

N
0.2

0.3

0.4 f

 

Figure 5.51 Phase averaged PIV data: velocity magnitude for condition ‘A’ at 25

0.6' '

 

r—nNUJ-hUIONN

degrees past leading edge, note the “path” of high velocity, 1

 

   

Level I U;I
I 7T: 7 060
01, 6 050
'R 5 040
I 4 030
02I 3 020
2 010
1 .
03? 000
044

 

06“

Figure 5.52 Phase averaged PIV data: velocity magnitude for condition ‘A’ at 30

degrees past leading edge, note high velocity cell, 1:

 

 

Level I UQI

"‘7 060

047 6 050

'g I i 5 o40

. 4 030

02- 3 020

2 010

03' 1 000
0.4;

 

0.61 i 1

Figure 5.53 Phase averaged PIV data: velocity magnitude for condition ‘A’ at 35

degrees past leading edge

118

  
  
 

Level I Ur;I

I I 7 0.60
0-1 ‘ 6 0.50
I a 5 0.40
N i ' 4 0.30
0.27 3 0.20
g 2 0.10
A 1 0.00
0.3?
0.4-

 

V I I l I I I l l I I I I I l i I I I I I I I I I l I I
0.6 0.7 0.8 032 1 1.1 1.2
1'

Figure 5.54 Phase averaged PIV data: velocity magnitude for condition ‘C’ at the
fan blade leading edge, note the high velocity cell, A

 

Level I U;I

I I C 7 0.60
0.1 _ r- ‘ 6 0.50
*N 5 0.40
I 4 0.30
0-2 . 3 0.20
2 0.10
I 1 0.00
0.3
0.4 -

 

‘ l I I I ‘ l I I I I l I ‘ ‘ l I l l I I

0.6 0.7 0.8 0.}? l 1.1 1.2
r

Figure 5.55 Phase averaged PIV data: velocity magnitude for condition ‘C’ at 5
degrees past the leading edge

119

0.1

~11-
N
0.2

0.3 _
0.4 f

0.61 I

0.1

2*

0.2 g

0.3 _

 

 

I..I1
0.7

1 I

0.8 A

I l I I I I l
0. 1
1‘:
Figure 5.56 Phase averaged PIV data: velocity magnitude for condition ‘C’ at 10
degrees past the leading edge

 

0.4 f

 

Ill

 

  
 

I l

 

0.6 l

I1IIII1I
0. 1
r2

Figure 5.57 Phase averaged PIV data: velocity magnitude for condition ‘C’ at 15

1.1 ‘

1.2

Level IUI;I
' 7 0.60
‘ 6 0.50
i 5 0.40
4 0.30
3 0.20
2 0.10
1 0.00

Level I U;I

I 7 0.60
J 6 0.50
' 5 0.40
4 0.30
3 0.20
2 0.10

1 0.00

degrees past the leading edge, note the recirculation region, 11

120

 

 

 

-0.1
0 a
lull
_ 7 0.60
0.1 6 0.50
IN V I 5 0.40
I " 4 0.30
0-2 3 0.20
2 0.10
1 0.00
0.3
0.4}
III III II

 

0.6

IJIII
0.7

0.8

I.1III
0.
,2

Figure 5.58 Phase averaged PIV data: velocity magnitude for condition ‘C’ at 20

degrees past the leading edge, note the large recirculation region, v

 

 

 

—0.1 k
0 I .
Level IUer
-_ 3 7 0.60
0-1 ‘ ‘ 6 0.50
*N 5 0.40
- 4 0.30
0-2 I 3 0.20
2 0.10
A 1 0.00
0.3 r
0.4 L

 

III

111

IlIII

 

0.6 J

0.7

0.8

IIIIII
0.
1*9

1

Figure 5.59 Phase averaged PIV data: velocity magnitude for condition ‘C’ at 25
degrees past the leading edge

121

 

 

 

0.17

*

N

0.2?

0.3»

0.4}
LIIIIl IIlI III I III
0.6 0.7 0.8 0.9 1 1.1

r*

Figure 5.60 Phase averaged PIV data: velocity magnitude for condition ‘C’ at 30
degrees past the leading edge

 

 

 

 

 

  
 

Level I Ur;I

0.60
0.50
0.40
0.30
0.20
0.10
0.00

—I
' I
I

t-‘NUJJXUIO‘N

-0.1
0 3 .
w .. Level IUnI
\r :1 7 0.60
0-1 r 6 0.50
*N I 5 0.40
4 0.30
0-2 3 0.20
2 0.10
F 1 0.00
0.3 7
0.4 :-
3 I I l I I I l I I I l I I I l I l I I I l
0.6 0.7 0.8 0.32 1 1.1 1.2
I‘

Figure 5.61 Phase averaged PIV data: velocity magnitude for condition ‘C’ at 35

degrees past the leading edge

122

6.0 Summary and Conclusions

An experimental investigation of a large scale prototype cooling fan utilizing a
unique aerodynamic shroud has been conducted. Integral quantities were measured first
to establish a baseline performance condition and to analyze the effect of the aerodynamic
shroud on the overall performance of the cooling fan. Detailed velocity measurements
were then acquired in the wake of the fan for selected operating conditions. The following
conclusions are supported by the results of these experimental investigations.

1. The aerodynamic shroud effectively delays the onset of stall. The integral mea-
surements clearly show that as the shroud power x is increased, the stall point flow coef-
ficient is decreased. In fact, for the highest shroud power tested, the stall point flow
coefficient is decreased by 27%. A hot-wire survey supports these data by showing that
the characteristics of the flow field in the wake of the fan for the powered shroud condi-
tion are qualitatively similar when compared to the unpowered shroud condition before
stall. The delay of stall can be beneficial because it extends the usable operating range of
the fan. Thus, a higher system resistance can be tolerated without sacrificing the undesir-
able effects of fan stall, such as increased noise generation and decreased efficiency.

2. The aerodynamic shroud does not increase the flow rate or efficiency of this
large scale cooling fan for flow coefficients larger than that for the onset of the unpowered
shroud stall condition. This result is in contrast to the previous aerodynamic shroud study
conducted by Morris (1997) involving an automotive cooling fan in which the flow rate
and efficiency were increased for high flow rates and decreased for low flow rates. The
magnitude of tip clearance and the nature of the flow field in this area is the most likely

cause for this new result. Specifically, the tip clearance area accounted for 31% of the

123

total effective flow area in the Morris (1997) study as compared to 5.6% in the present
study. Moreover, the baseline efficiency of the fan in the present study is much higher
(approximately 50% compared to 25%).

3. The dimensionless shroud power variable, x , can fully characterize the contri-
bution of the aerodynamic shroud. That is, only the combination of shroud pressure and
shroud flow rate is necessary to determine how the aerodynamic shroud will affect the
performance of the fan. This means that a small annular jet gap height (small volume flow
rate) can be used with a large shroud pressure; or a large jet gap height (large volume flow
rat) can be used with a smaller shroud pressure to achieve the same result. It should be
noted that this can only be stated for the limitations of the present study (a 1 mm to 5 mm
jet gap height). Thus, the combination of the shroud pressure and shroud flow rate allows
a reduction of variables so that the pressure coefficient is only dependent upon the flow
coefficient and the dimensionless shroud power. This result confirms the findings of Mor-
ris (1997) in that aerodynamic shroud study.

4. The aerodynamic shroud has a strong effect on the vortical flow field in the
wake of the fan. The hot-wire axial vorticity data suggest that the areas of negative vortic-
ity, that are assumed to be signatures of the tip leakage vortex, are weakened by the high
momentum annular jet of the aerodynamic shroud. In fact, the peak negative vorticity is
reduced by as much as 15% in the representative powered shroud condition. This is sig-
nificant because the tip leakage vortex plays an important role in the total losses of the
system. In particular, this secondary flow leads to increased viscous dissipation and can
interact with the main (through) flow in the blade tip area causing additional losses from

turbulent mixing.

124

5. The positive contribution of the aerodynamic shroud is strongly dependent on
the design of the fan and the nature of the specific loss mechanisms present in the fan
flow. Losses in a fan flow can occur from a number of sources including blade profile
loss, shroud and passage loss, tip clearance loss and secondary flow loss (Neal, 2002).
Lakshminarayana (1996) attributes 20-35% of total losses to tip clearance effects. How-
ever, Storer and Cumpsty (1994) suggest that a more realistic proportion attributed to the
tip clearance flow itself is 10%. Nonetheless, tip clearance loss varies widely according to
application. While the aerodynamic shroud may, in fact, decrease losses attributed to tip
leakage flow, it may also be increasing losses from other sources such as secondary flows.
The experimental results provide examples of this including the phase averaged PIV data,
which show that, for the powered shroud condition, a region of recirculating flow down-
stream of the fan was observed that was not seen in the baseline shroud condition. The
larger reverse flow region in the powered shroud condition is an example of the aerody-

namic shroud negatively affecting other aspects of the flow field.

125

REFERENCES

Baranski, B., (1997), “Designing the Engine Cooling Fan,” SAE #740691.

Bohl, D., (1996), “An Experimental Study of the Near Field Region of a Free Jet With
Passive Mixing Tabs,” MS. Thesis, Department of Mechanical Engineering, Michigan
State University, East Lansing, MI.

D’Andrea, R., Behnken, R., Murray, R., ( 1996), “Rotating Stall Control of an Axial Flow
Compressor using Pulsed Air Injection,” ASME J. of T urbomachinery, 119, 742-752.

Day, I.J., (1993), “Active Suppression of Rotating Stall and Surge in Axial Compressors,”
ASME J. of Turbomachinery, 115, 40-47.

Foss, J .F., Morris, S.C., Neal, DR, (2001), “Axial Fan Research for Automotive and
Building Ventilation Applications”, Tech Brief in ASME J. of Fluids Eng., Spring 2001 .

Fukano, T., and Jang, CM. (2003), “Tip Clearance Noise of Axial Flow Fans Operating at
Design and Off-design Condition,” J. of Sound and Vib., Corrected proof, page numbers
not assigned yet.

Inoue, M., Kumoumaru, M., (1984), “Three Dimensional Structure and Decay of Vortices
Behind an Axial Flow Rotating Blade Row,” Journal of Engineering for Power; 106.

Jang, C., Furukawa, M., and Inoue, M, (2001), “Noise Reduction by Controlling Tip Vor-
tex in a Propeller Fan,” JSME Int. Journal, 44(4), 748-755.

Keane, R.D., Adrian, R.J., (1990), “Optimization of particle image velocimeters. Part I:
Double Pulsed Systems,” Meas. Sci. Tech, 1, 1202-1215.

LaVision GmbH., (2002), FlowMaster Manual, Gottingen, Germany, p. 37.

Longhouse, RE, (1978), “Control of Tip-Vortex Noise of Axial Flow Fans by Rotating
Shrouds,” J. of Sound and Vibration, 58(2), 201-214.

Mellin, 1980, “Noise and Performance of Automotive Cooling Fan,” SAE paper No.
740691.

Mittal, S., Saxena, P., (2000), “Prediction and Hysteresis Associated with the Static Stall
of an Airfoil,” AIAA Journal, 38, (5), 933-935.

Morris, S.C., (1997), “Experimental Investigation of an Aerodynamic Shroud for Cooling

Fan Applications,” M.S. thesis, Department of Mechanical Engineering, Michigan State
University, East Lansing.

126

Morris, S.C., Foss, J.F., (2001), “An Aerodynamic Shroud for Automotive Cooling Fans,”
ASME J. Fluids Eng, 123, 287-292.

Morris, S.C., Neal, D.R., Foss, J .F., Cloud, GL., (2001), “A Moment-of-Momentum Flux
Mass Air Flow Measurement Device,” Meas. Sci. Tech, 12, 9-13.

Morris, S.C., Good, I.J., Foss, J .F., (1998), “Velocity Measurements in the Wake of an
Automotive Cooling Fan,” Exp. Thermal and Fluid Sci, 1 7, 100—106.

Morris, S.C., (2002), “The Velocity and Vorticity Fields of a Single Stream Shear Layer,”
Ph.D. Dissertation, Department of Mechanical Engineering, Michigan State University,
East Lansing, MI.

Neal, D.R., (2002), MS. thesis, Department of Mechanical Engineering, Michigan State
University, East Lansing.

Potter, M.C., Foss, J .F., (1982), Fluid Mechanics, Great Lakes Press, Okemos.

Smith, L.H., (1970), “Casing Boundary Layers in Mult-stage Compressors,” Proceedings
of the Symposium of F low Research on Blading, Elsevier Publishing Company.

Sorensen, D.N., (2001), “Minimizing the Trailing Edge Noise From Rotor-Only Axial
Fans Using Design Optimization,” J. of Sound and Vibration, 24 7(2), 305-323.

Wu, S.F., Su, 8., and Shah, H., (1998), “Noise Radiation From Engine Cooling Fans,”J. of
Sound and Vibration, 216(1), 107-132.

127

 

   

IIIIIIIIIIIIIIII111111511111