HOT FLOW TESTING OF MULTIPLE NOZZLEEXHAUST EDUCTOR SYSTEMS
Daniel Roy Welch
NAVAL POSTGRADUATE SCHOOL
Monterey, California
THESISHOT FLOW TESTING OF MULTIPLE
NOZZLE EXHAUST EDUCTOR SYSTEMS
by
Daniel Roy Welch
September 1978
Thesis Advisor: P. F. Pucci
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Hot Flow Testing of Multiple NozzleExhaust Eductor Systems
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Hot Flow ModelMultiple Nozzle Exhaust Eductor Systems
20 ABSTRACT (Contlnuo an rorormo lido II nmcmmmmrr mnd Idmntity »y block numbor)
Hot flow model tests of multiple nozzle exhaust eductorsystems were conducted to evaluate effects of exhausttemperature on eductor performance. A one-dimensionalanalysis of a simple eductor system based on conservationof momentum for an incompressible gas was used in determiningthe non-dimensional parameters governing the flow phenomenon.Eductor performance is defined in terms of these parameters.
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(20. ABSTRACT Continued)
An experimental correlation of these parameters which waspreviously developed and used to correlate cold flow datawas found to be effective in correlating both cold andhot flow data for eductor systems. Temperature data wasobtained for the mixing stack wall and the exhaust flowat the mixing stack exit plane.
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Hot Flow Testing of MultipleNozzle Exhaust Eductor Systems
by
Daniel Roy WelchLieutenant, United States Navy
B.S .NavArch. , United States Naval Academy, 1971
Submitted in partial fulfillment of therequirements for the degrees of
MASTER OF SCIENCE IN MECHANICAL ENGINEERING
and
MECHANICAL ENGINEER
from the
NAVAL POSTGRADUATE SCHOOL
September 1978
Q.I
ABSTRACT
Hot flow model tests of multiple nozzle exhaust eductor
systems were conducted to evaluate effects of exhaust tem-
perature on eductor performance. A one-dimensional analysis
of a simple eductor system based on conservation of momentum
for an incompressible gas was used in determining the non-
dimensional parameters governing the flow phenomenon.
Eductor performance is defined in terms of these parameters.
An experimental correlation of these parameters which was
previously developed and used to correlate cold flow data
was found to be effective in correlating both cold and hot
flow data for eductor systems. Temperature data was obtained
for the mixing stack wall and the exhaust flow at the
mixing stack exit plane.
TABLE OF CONTENTS
I. INTRODUCTION 15
II. THEORY AND ANALYSIS 18
A. MODELING TECHNIQUE 18
B. ONE-DIMENSIONAL ANALYSIS OF A SIMPLE EDUCTOR —
I
9
C. NON-DIMENSIONAL SOLUTION OF SIMPLEEDUCTOR ANALYSIS 25
D. CORRELATION OF EXPERIMENTAL DATA 29
III. EXPERIMENTAL APPARATUS 31
A. COMBUSTION GAS GENERATOR 3i
B. EDUCTOR AIR METERING BOX 32
C. INSTRUMENTATION 33
D. EDUCTOR SYSTEM 36
1. Mixing Stack 36
2. Eductor Nozzles 3 6
3. Standoff Ratio 37
IV. EXPERIMENTAL METHOD 3 8
V. DISCUSSION OF EXPERIMENTAL RESULTS 40
VI. CONCLUSIONS 44
VII. RECOMMENDATIONS 45
VIII. FIGURES 46
IX. TABLES 95
APPENDIX A: Combustion Gas Generator Operation 118
APPENDIX B: Determination of the Exponent in theNon-Dimensional Pumping Coefficient 124
APPENDIX C: Uncertainty Analysis 125
BIBLIOGRAPHY — 129
INITIAL DISTRIBUTION LIST 130
LIST OF FIGURES
Figure Description Page
1 Schematic Diagram of SimpleExhaust Gas Eductor 46
2 Simple Single Nozzle Eductor System — 47
3 Schematic Diagram of CombustionGas Generator 48
4 Combustion Gas Generator 49
5 Schematic Diagram of Gas GeneratorFuel System 50
6 Gas Generator Fuel Supply System 51
7 Eductor Air Metering Box 52
8 Eductor Air Metering Box Arrangement - 53
9 Interior of Air Metering Box ShowingUptake Stack and Primary Nozzles 54
10 Interior of Air Metering Box ShowingMixing Stack and Primary Nozzles 55
11 Schematic Diagram of PressureMeasurement System 56
12 Manometer Board 57
13 Main Control Panel, DigitalPyrometers, Manometer ManifoldValves 58
14 Schematic Diagram of TemperatureMeasurement System 59
15 Entrance Nozzle Calibration Curve 60
16 Air Metering Box End Plate andMixing Stack Collar 61
17 Dimensional Diagram of PrimaryFlow Nozzles ° 2
Figure Description Page
18 Dimensional Diagram of PrimaryFlow Nozzle Plate 63
19 Primary Flow Nozzle Plate (Back View) — 64
20 Primary Flow Nozzle Plate (Front View) -- ^5
21 Illustrative Plot of the ExperimentalData Correlation in Equation 14 66
22 Primary Flow Nozzle TemperatureProfiles 67
23 Uptake Stack Temperature Profile 68
24 Comparison of Cold Flow PerformancePlots for L/D =3.0 69
25 Comparison of Cold Flow PerformancePlots for L/D =2.5 70
26 Composite Performance Plot for allTemperatures, L/D = 3.0 "71
27 Composite Performance Plot for allTemperatures, L/D = 2.5 72
28 Performance Plot, L/D = 3.0, Cold Flow - 73
29 Performance Plot, L/D = 3.0,TUPT = 550*F 74
30 Performance Plot, L/D = 3.0,TUPT = 650 °F 75
31 Performance Plot, L/D = 3.0,TUPT = 750 °F 76
32 Performance Plot, L/D = 3.0,TUPT = 850 °F 77
33 Performance Plot, L/D =2.5, Cold Flow - 78
34 Performance Plot, L/D = 2.5,TUPT = 550 'F 79
35 Performance Plot, L/D = 2.5,TUPT = 650 °F 80
Figure Description Page
36 Performance Plot, L/D = 2.5,TUPT = 750 F 81
37 Performance Plot, L/D = 2.5,TUPT = 850 °F 82
38 Mixing Stack Pressure Distribution,L/D =3.0 83
39 Mixing Stack Pressure Distribution,L/D = 2.5 84
4 Mixing Stack Wall TemperatureDistribution, L/D = 3.0 85
41 Mixing Stack Wall TemperatureDistribution, L/D =2.5 86
42 Comparison of Mixing Stack WallTemperature Distributions 87
4 3 . Mixing Stack Exit Plane TemperatureProfile, L/D =3.0 88
44 Mixing Stack Exit Plane TemperatureProfile, L/D =2.5 89
45 Schematic Diagram of Compressor Layout - 90
46 Cooling Tower Switches and CoolingWater Valve 91
47 Carrier Air Compressor, ButterflySuction Damper and Cooling Water Valve - 92
48 Auxiliary Oil Pump and Switch 93
49 Main Air Supply Globe Valve 94
LIST OF TABLES
Table Description Page
I. Summary of Results 95
II. Entrance Transition NozzleCalibration Data 96
III. Primary Nozzle TemperatureProfile Data 97
IV. Uptake Stack Temperature Profile Data — 98
V. Performance Data for L/D = 3.0 "VI. Performance Data for L/D = 2.5 104
VII. Mixing Stack Exit Plane TemperatureProfile L/D =3.0 113
VIII. Mixing Stack Exit Plane TemperatureProfile L/D =2.5 115
IX. Uncertainties in Measured' ValuesFrom Table VI 117
10
NOMENCLATURE
ENGLISH LETTER SYMBOLS
A - Area, in
C - Sonic velocity, ft/sec
D - Diameter, in
f - Friction factor
F - Functional denotation
Ff
- Wall skin-friction force, lbf
g - Proportionality factor in Newton's Second Law,
gc= 32.174 lbm-ft/lbf-sec 2
h - Enthalpy, Btu/lbm
k - Ratio of specific heats
L - Length, in
P - Pressure, in H_0
P_, B - Atmospheric pressure, in Hga
R - Gas constant for air, 53.34 ft-lbf/lbm-°R
S - Standoff distance, in
T - Temperature, °F, °R
U - Velocity, ft/sec
W, m - Mass flow rate, lbm/sec
x - Axial distance from mixing stack entrance, in
Dimensionless Groupings
A* - Secondary flow area to primary flow area ratio
K - Kinetic energy correction factor
11
K - Momentum correction factor at the mixingm i • .stack exit
K - Momentum correction factor at the primaryp nozzle exit
M - Mach number
AP* - Pressure coefficient
Re - Reynolds number
T* - Secondary flow absolute temperature to primaryflow absolute temperature ratio
W* - Secondary mass flow rate to primary mass flowrate ratio
p* - Secondary flow density to primary flow densityratio
Greek Letter Symbols
2\i - Absolute viscosity, lbf-sec/ft
3Density, lbm/ft"
{ + =r A /Am 2 w mB - K + % A /Ar
Subscripts
- Section within secondary air plenum
1 - Section at primary nozzle exit
2 • - Section at mixing stack exit
B - Burner
m - Mixed flow or mixing stack
P - Primary
s - Secondary
u - Uptake
w - Mixing stack inside wall
12
Tabulated Values
DELPN, PN
FHZ
P*
PA, B
PA-PS, APS
PEH
PMIX, PMS
PNH
PU-PA
P*/T*
T*
TAMB
TMIX
TUPT
UM
UP
UU
WP
ws
WPA
WPF
W*
Pressure drop across entrance transitionnozzle, in H^O
Fuel flow meter reading, Hz
Pressure coefficient
Ambient pressure, in Hg
Pressure differential across secondaryflow nozzles, in H
2
Uptake static pressure, in H~0
Mixing stack static pressure, in H2
Static pressure upstream of entrancetransition nozzle, in Hg
Uptake static pressure, in H^O
Dimensionless pressure coefficient
Absolute temperature ratio, secondaryflow to primary flow
Ambient temperature, °F
Mixing stack wall temperature, °F
Uptake temperature, °F
Average velocity in mixing stack, ft/sec
Primary flow velocity at nozzle exit, ft/sec
Primary flow velocity in uptake, ft/sec
Primary mass flow rate, lbm/sec
Secondary mass flow rate, lbm/sec
Mass flow rate of primary air, lbm/sec
Mass flow rate of fuel, lbm/sec
Secondary mass flow rate to primary flowrate ratio
13
ACKNOWLEDGEMENT
Sincere thanks go to the author's advisor, Professor
Paul F. Pucci/ whose expertise and inspiration provided the
foundation and catalyst for this work. Great appreciation
is also expressed to Professor T. Sarpkaya who was always
available with words of wisdom and support. The assistance
and ingenuity provided by the personnel of the Department
of Mechanical Engineering Machine Shop, especially Mr.
George Bixler, is also sincerely appreciated.
Special thanks and grateful appreciation are due to my
wife, Sharon, for the never ending encouragement and
understanding she provided during the long days and late
nights devoted to this study.
14
I. INTRODUCTION
The gas turbine engine is steadily becoming more and
more attractive as a prime mover for various shipboard
applications. One of the unique features of the use of
gas turbine engines is its relatively hot and voluminous
exhaust. This presents problems such as overheating of
antennae and other equipment by exhaust plume impingement
and the creation of an undesirable infra-red signature of
the hot exhaust plume. An effective means of reducing the
exhaust gas temperature is to mix it with ambient air prior
to its discharge from the stack. Exhaust gas eductor sys-
tems presently in service have demonstrated their effec-
tiveness in facilitating such a mixing process.
The subject of this investigation is the application of
multiple nozzle eductor systems for cooling the exhaust gas
from gas turbine powered ships. This research is an exten-
sion of work reported by Lt. C. R. Ellin [1], Lt. C. M. Moss
[2], and Lt. J. P. Harrell [3]. Whereas this previous work
has been carried out with cold flow testing, this investi-
gation is concerned with testing using hot gas as the exhaust
or primary flow. The scope of the work reported here in-
cludes completion of and subsequent changes to the combustion
gas generator designed and built by Lcdr. P. D. Ross [4]
.
For the purpose of this investigation, the exhaust gas
eductor system, illustrated schematically in Figure 1, is
15
defined as the portion of the uptake which discharges the
exhaust gas through nozzles into a mixing stack. The purpose
of the eductor system is to induce a flow of cool ambient
air which is mixed with the hot exhaust gas in order to
lower the temperature of the exhaust stack and exhaust plume.
These gas eductors must meet three major requirements. They
must pump large amounts of secondary (cooling) air into the
mixing stack, they must adequately mix the hot high velocity
exhaust gas and the cool low velocity secondary air, and
they must not adversely affect the gas turbine's performance.
A one-dimensional flow analysis of a simple single-
nozzle eductor system, as a unit, facilitates determination
of the non-dimensional parameters which govern the flow
phenomenon. An experimental correlation of these non-
dimensional parameters has been developed and is used to
evaluate eductor performance.
The geometric parameters which influence the gas
eductor' s performance include the number and size of primary
nozzles, the length of the mixing stack, the ratio of the
primary nozzle flow area to the mixing stack area, and the
ratio of the length of the mixing stack to its diameter.
Numerous combinations of and variations in these parameters
have been investigated and reported in References [1]
through [3]
.
The intent of this investigation was to obtain data
using hot flow testing of gas eductor systems to establish
16
the effect of uptake gas temperature on the eductor '
s
performance. Correlation of hot flow data with previous
cold flow data allows a validation of the hot gas generator
and a validation of the use of cold flow models for hot
flow prototypes.
Two exhaust eductor models were tested. Both geometries
were tested previously using cold flow facilities. One
geometry was tested by Moss [2] and by Harrell [3] , each
at a different scale; the other was tested by Staehli and
Lemke [5] . All tests were made at the same flow parametric
values.
17
II. THEORY AND ANALYSIS
Evaluation of the effects of eductor geometry on
prototype eductor performance through experimentation with
models requires the following: assurance of similitude
(geometric, kinematic, and dynamic similarity) between
model and prototype; the identification- of the dimension-
less groupings pertinent to the flow phenomenon; and a
suitable means of data analysis and presentation. Dynamic
similarity was maintained by using Mach number similarity
to establish the model's primary flow rate. Determination
of the dimensionless groupings that govern tne flow was
accomplished through the analysis of a simple air eductor
system. Based on this analysis, an experimental correla-
tion of the non-dimensional parameters was developed and
used in presenting and evaluating experimental results.
A. MODELING TECHNIQUE
For the flow velocities considered, the primary flow
through the model eductor is turbulent (Reynolds number
5based on diameter of approximately 10 ) . Consequently,
turbulent momentum exchange is a predominant mechanism
over shear interaction, and the kinetic and internal
energy terms are more influential on the flow than are
viscous forces. Since Mach number can be shown to repre-
sent the square root of the ratio of kinetic energy of a
18
flow to its internal energy, it is a more significant
parameter than Reynolds number in describing the primary
flow through the uptake.
Similarity of Mach number was therefore used to model
the primary flow. Mach number is defined as the ratio of
flow velocity to sonic velocity in the medium considered.
Sonic velocity, represented by c, can be calculated using
the relation
c = (g kRT)0,5
^c
if the fluid is assumed to behave as a perfect gas.
Geometric similarity was achieved through the use of
a dimensional scale factor which is influenced by test
facility flow capabilities, primary flow velocities and
availability of modeling materials.
B. ONE-DIMENSIONAL ANALYSIS OF A SIMPLE EDUCTOR
The theoretical analysis of an eductor may be approached
in two ways. One method attempts to analyze the details of
the mixing process of the primary and secondary flows
which takes place inside the mixing stack and thereby
determines the parameters that describe the flow. This
requires an interpretation of the mixing phenomenon, which,
when applied to multiple-nozzle systems, becomes extremely
complex. The second method, employed in this study, analyzes
the overall performance of the eductor system as a unit.
Since details of the mixing process are not considered in
this method, an analysis of the simple single-nozzle
eductor system shown in Figure 2 leads to a determination
of the dimensionless groupings governing the flow. The
one-dimensional analysis that follows is essentially that
of Ellin [1]
.
The driving or primary fluid/ flowing at a rate W and
at a velocity U , discharges into the entrance of the con-
stant area section of the mixing stack, inducing a secondary
flow rate of W at velocity U . The primary and secondary
flows are mixed and leave the mixing stack at a flow rate
of W and a bulk average velocity of U .
m r J m
The one-dimensional flow analysis of the simple eductor
system described depends on the simultaneous solution of
the equations of continuity, momentum, and energy with an
appropriate equation of state and specified boundary
conditions.
The following simplifying assumptions are made:
1. Both gas flows are treated as perfect gases with
constant specific heats.
2. Steady, incompressible flow throughout the eductor
and plenum exists.
3. The flow throughout the eductor is adiabatic. The
flow of secondary air from the plenum (at section 0)
to the entrance of the mixing stack (at section 1)
is isentropic. Irreversible adiabatic mixing occurs
20
between the primary and secondary flows in the
mixing stack (between sections 1 and 2)
.
4. The static pressure distributions across the
entrance and exit planes of the mixing stack (at
sections 1 and 2) are uniform.
5. At the mixing stack entrance (section 1), the
primary flow velocity U and temperature T areir ir
uniform across the primary stream, and the
secondary flow velocity U and temperature T are
uniform across the secondary stream; but U does
not equal U , and T does not equal T .
6. Incomplete mixing of the primary and secondary
flows in the mixing stack is accounted for by the
use of a non-dimensional momentum correction
factor. K , which relates the actual momentum ratem
to the rate based on the bulk-average velocity and
density and by the use of a non-dimensional kinetic
energy correction factor, K , which relates the
actual kinetic energy rate to the rate based on
the bulk-average velocity and density.
7. Potential energy differences due to elevation are
negligible.
8. Pressure changes P to P, and P^^ to Pa
are small
relative to the static pressure so that the gas
density is essentially dependent upon temperature
(and atmosperic pressure)
.
21
m
9. Wall friction in the mixing stack is accounted
for with the conventional pipe friction factor
term based on the bulk-average flow velocity U
and the mixing stack wall area A .3 w
The conservation of mass principle for steady state
flow yields
W = W + W (1)m p s
where
W = p U AP P P P
W = p U A (la)s s s s
W p U Am m m m
Substituting for W , the bulk-average velocity becomes
W + WU = — rr-E. db)m Pm A
m m
Now, from assumption 1
pm - A- (2)
m
where T is calculated as the bulk-average temperature form
the mixed flow. Applying assumptions 4 and 6, the momentum
22
equation for the flow in the mixing stack may be written
W U W U W UKp [-V\ + l-^\ + P
1A1 = V^, + P
2A2
+ Ffr
c 1 c 1 c 2
(3)
with A.. = A . The momentum correction factor K is intro-12 p
duced to account for a possible non-uniform velocity profile
across the primary nozzle exit. It is defined in a manner
similar to that of K and by assumption 5 is equal to unity
but is included here for completeness. The momentum correc-
tion factor for the mixing stack exit is defined by the
relation
Am
Km = vfV / U2
2p2«* (4>
m m n
The actual variable velocity and a weighted average density
at section 2 are used in the integrand. The wall skin-
friction force Ff
can be related to the mean velocity by
U2
P
F = f A [-5 S] (5)fr w 2 g
For turbulent flow, the friction factor may be calculated
from the Reynolds number as
P U D.-0.2 , „ m m m , c ,
f = 0.046 (Re ) , where Rem
= (6)
23
Applying the conservation of energy principle to the
steady flow in the mixing stack with assumption 7
U2
U2
U2
wn [hn + 7-E-J + w= [h<=
+ 9-4-1 = W [h + K *-£-]P P 2gcl s s gc 1
m m egc2
(7)
where K is the kinetic energy correction factor defined
by the relation
Am
K = i—y / U9
3p 9
dA (8)
m m
It may be demonstrated that for the purpose of evaluating
the mixed mean flow temperature T , the kinetic energym J
terms may be neglected to yield
W Wh = r^ h + =3^ h (9)m W p W sm c m
where T = F (h ) only from assumption 1.mmSimilarly, the energy equation applied to the flow of
secondary air between the plenum entrance and the mixing
stack entrance may be reduced to
K11- 2^P 2 gK
s ^c
24
The foregoing equations may be combined to yield the
vacuum produced by the eductor in the plenum chamber
2 2i r W
r,w„ a
a o gcAm tp A
ppp
Ag P- " " F1T]
W 2
IK. + >]} (IDAm pm m 2 Am -m m
where it is understood that A and p apply to the primary
flow at the entrance to the mixing stack (section 1) , A
and p apply to the secondary flow at this same section,9
and A_ and p_ apply to the mixed flow at the exit of them m
mixing stack (section 2) . P is atmospheric pressure anda
is equal to the pressure at the exit of the mixing stack
P2
» This equation also incorporates the assumption that
(P-)-, = (p_) n so that p may be taken as the density of
the secondary flow in the plenum.
C. NON-DIMENSIONAL SOLUTION OF SIMPLE EDUCTOR ANALYSIS
In order to provide the criteria of similarity of
flows with geometric similarity, the non-dimensional
parameters which govern the flow must be determined. One
means of determining these parameters is by normalizing
equation (11) which leads to the following terms:
25
p - pla. ^_0
p sAP* = =— a pressure coefficient which compares
pPa " P
2° the "pumped head" for thegc
P s
secondary flow to the "driving head"U 2
2of the primary flow.
gc
WW* = rr— a flow rate ratio, secondary-to-
Am
WP
TP
pP
AP
primary mass flow rate.
TT* = ~ an absolute temperature ratio,
secondary-to-primary
.
P sp* = — a flow density ratio. Note that
since P = P and the fluids ares p T
1perfect gases, p* = ^ = ^ .
s
AA* = ~ area ratio of secondary flow area
to primary flow area
A-S. area ratio of primary flow area to
mixing stack cross sectional area
26
Aw
Amarea ratio of wall friction area to
mixing stack cross sectional area
K momentum correction factor forP
primary flow
K momentum correction factor for mixedm
flow
wall friction factor
With these non-dimensional groupings, equation (11) may be
written as
ad* A A A= 2 ^{[K - -rE. 3] - W*(l + T*) *£ 6
T* A L L p A WJ Amm c m m
+ w*2 TM^ (1 -2^V>6 -% b]} Ula!p m
where
m
27
For a given eductor geometry, equation (11a) may be
expressed in the form
AP*= C, + C
9W* (T* +1) + (V W* T* (lib)
where
CA A
1 " 2 A^P " X* B)
m p m
A2
C2
= -2(^H) 3 (He)m
C3 = 2 S^ 1 " 2-A^V'B " ^B>
m p m
Equation (lib) may be expressed as a simple functional
relationship
AP* = F(W*,T*) ' (12)
A second means of determining the governing dimension-
less parameters is through a dimensional analysis of the
mixing process within the mixing stack. A presentation of
this method by Ellin [1] yields the same simple functional
relationship found in equation (12)
.
Two geometric dimensionless quantities were added to
this investigation. The distance, S, from the primary flow
nozzle exit to the mixing stack entrance and the distance,
28
x, from the entrance to the mixing stack, normalized with
respect to the mixing stack diameter, D, were also defined
as non-dimensional quantities. The two additional quan-
tities are listed below:
xD
SD
ratio of the axial distance from the
mixing stack entrance to the diameter
of the mixing stack.
standoff; the ratio of the axial dis-
tance between the primary nozzle exit
plane and the mixing stack entrance
to the diameter of the mixing stack.
D. CORRELATION OF EXPERIMENTAL DATA
The previous experiments by Ellin [1] , Moss [2] , and
Harrell [3] were done in facilities which did not have the
capability for varying the primary flow temperature. Thus
T* / the ratio of the absolute secondary to primary flow
temperatures was determined by the rise in temperature of
the primary air in the blower supply and was near unity
(approximately .85). A means of presenting the experimental
data for a given geometric configuration in a form which
results in a pseudo-independence of the dimensionless
groupings P* and W* upon T* was developed. From reference
[1] a satisfactory correlation of P*, T* and W* for all
29
temperatures and flow rates is
AP*/T* = F(W*T*°'44
) (13)
The details of the determination of 0.44 as the correlating
exponent are presented in Appendix [B] . A plot of AP*/T*
44as a function of W*T* * from the experimental data yields
the eductor's pumping characteristic curve. Variations in
geometry will change the appearance of the pumping charac-
teristic curve and facilitate a direct one to one comparison
of pumping ability between various models and prototypes.
44For ease of discussion, W*T* will henceforth be referred
to as the pumping coefficient.
30
III. EXPERIMENTAL APPARATUS
Hot primary gas is supplied to the nozzle and mixing
stack system by the combustion gas generator and associated
ducting illustrated in Figures 3 and 4. The eductor system
being tested is mounted in a secondary air plenum which
facilitates the accurate measurement of the secondary air
flow through the use of ASME long radius flow nozzles mounted
on the secondary air plenum.
A. COMBUSTION GAS GENERATOR
The input air to the combustion gas generator is supplied
by a Carrier Model 18P350 centrifugal air compressor. The
compressor is located in an adjacent building and the input
air is piped underground to an eight-inch inside diameter
(ID) horizontal pipe with a butterfly-type shutoff valve
and a globe-type bypass valve. All air demands for this
testing can be met with the butterfly valve closed and the
globe valve open as necessary.
The input air travels through an entrance transition
piece that mates the eight inch ID compressor discharge
piping with the four inch ID system piping. This nozzle is
used to measure the primary air flow.
A portion of the input air travels straight through the
piping to the exhaust stack while the remainder passes
through the U-bend section to the combustion section. The
31
combustion section includes the burner can and igniter
assembly from a Boeing model 502-6A gas turbine engine.
Certain fuel system components from this engine were also
utilized. The fuel system is shown schematically in Figure 5
and pictured in Figure 6
.
After the air is heated in the combustion section, it
is mixed with the cooler air after both pass through the
turbine nozzle box containing the bypass air mixer. By
controlling the relative amount of air passing through the
burner and the amount of fuel to the burner, the exhaust
stack temperature can be controlled. The procedure for
system light-off and operation is included in Appendix A.
The hot gas then passes up the exhaust stack to the
primary nozzles and the eductor system. A flow straightening
section was added to the uptake stack to de-swirl the hot
gas after it leaves the turbine nozzle box.
B. EDUCTOR AIR METERING BOX
Secondary air flow is measured with a large metering
box designed to enclose the entire eductor assembly and act
as an air plenum. A set of standard ASME long radius flow
nozzles of varying cross-sectional areas were chosen to be
mounted in the metering box away from the eductor.
The metering box was designed with interchangeable stack
seal plates to enable variation of both exhaust and mixing
stack sizes up to 1 foot in diameter. The seal plates also
have a limited range of vertical movement to facilitate
exhaust and mixing stack alignment. The entire box was
designed to be movable along an angle iron track parallel
to the gas generator stack longitudinal centerline. This
enables variation of the mixing stack to nozzle separation
distance without adjustment and realignment of the mixing
stack. The mixing stack end plate was also designed to
be movable to allow centering for various mixing stack
lengths. An access door was added for eductor adjustment.
The metering box general arrangement is pictured in Figure 7
and a dimensional layout is given in Figure 8.
Appendix D of Reference [1] outlines the design and
construction of the ASME long radius secondary air flow
nozzles. Flexibility is provided this secondary air flow
measuring system by utilizing three different flow nozzle
sizes: four of four inch throat diameter, three of two
inch throat diameter and three of one and a half inch
throat diameter, various combinations of which produce a
wide variety of secondary cross sectional flow areas.
Mounted inside the air metering box are supports for
the uptake stack and mixing stack. The interior of the air
metering box is pictured in Figures 9 and 10.
C. INSTRUMENTATION
The performance of an eductor is calculated from pressure
and temperature data taken at various points in the system.
Necessary measurements include the primary mass flow rate
(air and fuel) , the secondary mass flow rate, the uptake
33
stack Mach number, and the mixing stack temperature and
pressure profiles.
Pressure measurements are made with one of several
manometers. Available are a 20 inch mercury upright
manometer, a 20 inch water upright manometer and a two inch
inclined water manometer. A manifold system allows selec-
tion of the instrument of proper range. Atmospheric pressure
(PA) is measured with a mercury barometer. A schematic of
the pressure measurement system is shown in Figure 11.
The manometer board and manifold system are pictured in
Figures 12 and 13.
Temperature measurements are made with either copper-
constantan or chromel-alumel thermocouples wired to Newport
model 267A digital pyrometers. The pyrometers are capable
of monitoring 18 inputs each through barrel-type selector
switches. Secondary air or ambient air temperature (TAHB)
was measured with a mercury-glass thermometer. A schematic
of the temperature measurement system is shown in Figure 14.
Fuel flow measurement is made with a Cox Instrument model
V4 0-A vortex flowmeter coupled to an Anadex Instruments
model CPM-603 frequency counter.
The calculation of the primary air mass flow rate
requires the measurement of the inlet absolute pressure to
the transition nozzle (PNH) , the pressure drop across this
nozzle (DELPN) , and the inlet air temperature. The calibra-
tion of this nozzle for the measurement of mass flow rate
34
as a function of these two pressure readings was previously
accomplished by Ross and details of this calibration can
be found in Reference [4] . The calibration curve is shown
in Figure 15.
The calculation of the secondary air mass flow rate
requires the measurement of the ambient pressure and tem-
perature, and the pressure drop across the secondary air
nozzles (PA-PS) . The secondary air plenum is equipped with
pressure taps mounted both in the rear section containing
the air metering nozzles and in the front section containing
the eductor under test. No measurable difference was de-
tected between the two taps so the pressure tap nearest the
eductor was used in the data runs.
The uptake stack Mach number calculation necessitates
the measurement of the uptake temperature and pressure as
well as the primary mass flow rate. The uptake temperature
(TUPT) is measured with a chrome1-alumel thermocouple inserted
through the primary nozzle plate at the centerline and pro-
truding approximately two inches into the stack. Uptake
pressure (PEH) is measured through a four-point averaging
pressure tap located approximately seven and one half inches
(one diameter) upstream of the eductor nozzle entrance.
The mixing stack was constructed with pressure taps
every one-half diameter down the length of the stack. The
mixing stack pressure distribution (PMIX) is easily measured.
Chromel-alumel thermocouples were welded every one-half
35
diameter to the outside of the mixing stack to facilitate
measurement of the temperature distribution.
Temperature profiles at the exit plane of the primary
nozzles and the mixing stack are obtained using a chromel-
alumel thermocouple mounted on an adjustable traversing
mechanism shown in Figure 16.
D. EDUCTOR SYSTEM
The eductor system includes the eductor nozzles and the
mixing stack. Figure 1 shows the general eductor system
arrangement.
1. The Mixing Stack
The mixing stack was construe Led of 7.5 inch OD,
0.188 inch wall thickness stell pipe. Two lengths were
tested. First a 3 diameter long (21.366 inch) stack was
tested then a 2.5 diameter long (17.805 inch) stack was
machined from the long stack and tested.
The mixing stack is supported inside the secondary
air plenum by means of an adjustable saddle and held in
place by an adjustable metal band. The stack is also
supported by the adjustable collar at the plenum wall.
This collar can be seen in Figure 16. The adjustable saddle
and collar allows alignment of the mixing stack with the
primary nozzles.
2. Eductor Nozzles
The eductor nozzles investigated consisted of two
different four-nozzle geometries previously tested. The
36
first geometry was a mixing stack to total nozzle area
ratio of 3:1 and the second was a mixing stack to total
nozzle area ratio of 2.5:1. The nozzle elements were machined
from steel tubing and welded to a circular nozzle plate
which was bolted onto the exhaust stack. The nozzles are
shown schematically in Figures 17 and 18 and are pictured
in Figures 19 and 20.
3. Standoff Ratio (S/D)
Both geometries investigated were tested at an
S/D value of 0.5. Previous testing [2] has shown this to
be approximately the optimum standoff ratio.
37
IV. EXPERIMENTAL METHOD
The pumping coefficient, W*T* 0,44 , provides the basis
for the analysis of parameter variation effects on eductor
pumping. Figure 21 graphically illustrates the eductor
pumping characteristic curve defined by the experimental
data correlation of equation (13) . Design of the experi-
mental apparatus facilitates determination of the dimension-
less parameters in the experimental correlation with the
exception of the secondary flow rate at the operating point.
For the operational eductor system, little or no restriction
of the secondary flow is present. Modeling of this opera-
ting point precludes the use of restrictive flow measuring
devices, such as ASME flow nozzles used in model tests.
The technique of determining the pumping coefficient at the
operating point, then, is first to establish the pumping
characteristics of the eductor system. This is accomplished
by varying the secondary air flow rate from zero to its
maximum measurable value, using the ASME flow nozzles mounted
in the secondary air plenum and recording the temperatures
and pressures required to calculate the corresponding
dimensionless parameters. The "open to the environment"
condition is then simulated by removal of the end plates
on the secondary air plenum. Extrapolation of the charac-
44teristic curve to its intersection with the W*T* axis
locates the pumping coefficient for the operating point of
the eductor system.
38
The mixing stack pressure and temperature distributions
were obtained from a series of pressure taps and thermo-
couples at half diameter distances along the mixing stack.
These pressures and temperatures were recorded at the "open
to the environment" condition and then plotted versus the
ratio of tap location (X) to mixing stack diameter (D) for
each geometry tested.
A measure of the degree of mixing of the primary and
secondary flows was obtained by plotting the mixing stack
exit plane temperature profile at the "open to the environ-
ment" condition. Two temperature profile traverses were
made. A greater degree of mixing of the flows will result
in a flatter temperature profile.
39
V. DISCUSSION OF EXPERIMENTAL RESULTS
The intent of this investigation, as discussed earlier,
was to conduct hot flow tests of exhaust eductor systems
in an attempt to meet three primary objectives. The first
objective was to test and verify the proper operation of
the hot gas generator. The second was to validate the use
of the correlating parameter (W*T* ) . The third objective
was to obtain temperature data on the mixing stack wall
and of the exhaust gas at the mixing stack exit plane.
Initial testing of the hot gas generator was concerned
with ensuring that a sufficient range in uptake temperatures
could be obtained while maintaining the desired Mach number.
Uptake temperatures from 550°F to about 900°F were easily
obtained. The lower limit exists due to the requirement
for a minimum fuel pressure to the fuel nozzle. An attempt
to lower the uptake temperature below 550 °F necessitates
too low a fuel flow rate to achieve proper fuel atomization
and smoking or loss of ignition occurs. The upper limit
exists because an attempt to go to higher uptake temperatures
requires burner temperatures above the 1500 °F maximum. Lower
uptake temperatures are obtainable at higher Mach numbers
,
as are higher uptake temperatures at lower Mach numbers.
Primary nozzle exit plane temperature profiles taken by
Ross [4] indicated that the exhaust was swirling up the
exhaust stack. A flow straightener consisting of two wire
40
screens placed two inches apart was installed in the uptake
stack one foot from the nozzle box. The temperature profiles
shown in Figure 22 are basically consistent from one nozzle
to another and are considerably flatter, indicating that the
flow straightener is effective in taking the swirl out of
the exhaust flow.
A temperature profile across the uptake stack at the
mid-length point was taken and is presented in Figure 23.
The temperatures taken at this point are normalized with a
reference temperature taken on the stack centerline lj inches
upstream of the primary nozzles. The temperature profile is
essentially flat, with the maximum temperature deviation
less than 2%. The average value of this curve is approxi-
mately one. The reference position was therefore used to
measure the uptake temperature, since it is essentially
equal to the average mid-length temperature.
Verification of the experimental setup was made by
duplicating previous cold flow results using the hot rig
under cold flow conditions. Figure 24 shows the results
of the cold flow test done by Moss [2] and the results
obtained with this setup for an identical geometry but
different scale. The pumping coefficients at the open to
the environment condition (P*/T* = 0) differ by only 1.5%.
Figure 25 gives a similar comparison for data taken by
Staehli and Lemke [5] for a different identical geometry
and different scale. Again the difference at P*/T* = is
about 1.5%.
41
Pumping performance data was taken at a range of uptake
temperatures from 150°F (cold flow) to 850°F. Figures 26
and 27 show that although the performance data are contained
within a narrow band, a temperature related trend is present.
The cold flow data is at the upper edge of the band with data
at increased temperature fanning out below it. This result
is not predicted by the one-dimensional theory discussed
earlier and is an area of possible future study. Possible
causes include temperature effects on either primary or
secondary air flow measurement. For example, a leak in the
air metering box that is accentuated with temperature would
lead to an underestimation of W which would lower W* ands
44in turn lower W*T** , shifting the performance plot.
Figures 28 through 37 give the pumping performance plots
at each condition individually. The pumping coefficients
at the open to the environment condition are all within 8%
of one another. The pumping coefficients of the Am/A= 2.5,
L/D =2.5 geometry (to be called the "2.5" geometry) are
about 30% lower than those of the Am/A= 3.0, L/D =3.0
geometry (to be called the "3.0" geometry). This agrees
with data obtained by Moss [2] and Staehli and Lemke [5]
.
Temperature and pressure data was acquired every half
diameter down the length of the mixing stack. The pressure
distributions presented in Figures 38 and 39 show a rise in
pressure down the stack indicating that the degree of mixing
of the primary and secondary flows increases down the stack.
Previous cold flow data gives similar results.
42
The mixing stack axial temperature distributions are
given in Figures 40 and 41. Figure 42 shows the temperature
distributions at TUPT = 850°F and A /A =3.0 and L/D =3.0
compared to the same uptake temperature and A /A =2.5,nr p '
L/D =2.5. The smaller nozzles (A^/A = 3.0) cool the stack
more effectively than the larger ones, as predicted by
the greater pumping coefficient achieved by this geometry.
The maximum stack temperature for the 3.0 geometry is 368°F
compared to 4 28 °F for the 2.5 geometry at the same uptake
temperature (TUPT = 850 °F)
.
Temperature profiles taken at the exit plane of the
mixing stack and presented in Figures 43 and 44 show that
the peak exhaust temperature for the 3.0 geometry is 11%
lower than the 2.5 geometry. Again, this is as expected
based on the larger pumping coefficient. These measurements
indicate that the peak temperatures occur about one-half inch
to the right of the stack centerline. A calculation of the
misalignment angle required to produce a one-half inch offset
gave an angle of about l-~ degrees. A check of stack align-
ment revealed a small alignment error which would corrobor-
ate the offset. It is not felt that this offset had any
noticeable effect on any other data taken.
Table 1 gives results of the findings for each geometry
tested.
43
VI. CONCLUSIONS
The experimental results were presented in Section V
and the resulting conclusions are summarized here.
A. The hot gas generator performs as desired. It was
verified that a wide range of uptake temperatures
could be easily obtained. The system proved to be
stable, repeatable and relatively simple to operate.
44B. The pumping coefficient (W*T* ) is an acceptable
parameter to measure a system's pumping performance.
The performance plots were contained within a narrow
region, which shows that the use of cold flow data
presented in this way accurately correlates to hot
flow tests. Cold flow pumping coefficients were
corroborated with hot flow data.
C. Mixing stack wall and exhaust temperature data
follow the trends predicted by cold flow testing.
The A /A =3.0, L/D =3.0 geometry cools the exhaus-m p
gas more effectively than the 2.5 geometry.
44
VII. RECOMMENDATIONS
In addition to meeting the objectives mentioned in
Section V, this research has also raised questions to be
solved by further research. Some recommendations for further
study are listed below.
A. A study should be done to determine the cause of
the slight temperature related spread in the
performance data.
B. Previous studies have used exhaust velocities as
a measure of the mixing and hence, as a measure
of the exhaust temperature profiles. Previously
obtained velocity profiles predict a slight tempera-
ture depression at the stack centerline that is not
present in the data obtained in this study. Exit
velocity profiles should be obtained, allowing a
direct comparison of velocity and temperature data.
C. The end plate of the air metering box should be
redesigned to allow greater freedom of mixing stack
alignment.
45
VIII. FIGURES
7K
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PRIMARYNOZZLES
^_V
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UPTAKE
A
Mu
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AMBIENT COOLING AIR(SECONDARY FLOW)
WWUPTAKE MACH NUMBER
UPTAKE PRESSURE
FIGURE 1.
TURBINE EXHAUST(PRIMARY FLOW)
Schematic Diagram of Simple ExhaustGas Eductor
46
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FIGURE 9. Interior of Air Metering Box Showing Uptake
Stack and Primary Nozzles
54
FIGURE 10. Interior of Air Metering Box Showing
Mixing Stack and Primary Nozzles
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VAp = 2 '5 Am/K = 3,0m p
A 1.251 1.154B 1.126 1.024C 1.770 1.770D 2.520 2.520E .250 .250F .125 .125G .500 .500
All dimensions in inches
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FIGURE 17. Dimensional Diagram of Primary Flow Nozzles
62
W 2 - 5 W 3 -
A 10.000 10.000
B 45° 45°
h 1.126 1.029
*2 1.251 1.154
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R4
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FIGURE 18. Dimension Diagram of Primary FlowNozzle Plate
63
FIGURE 19. Primary Flow Nozzle Plate (Back View)
64
Js FIGURE 20. Primary Flow Nozzle Plate (Front View)
65
w*r
FIGURE 21. Illustrative Plot of the Experimental Data
Correlation in Equation (14).
66
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IX. TABLES
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PNH + B APN /(PNH+8) -APN ma-
in. Hg) (in. H20) ( (in.Hg) -in.I^O) (lmb/sec)
55.46 13.05 26.902 1.91556.30 13.05 27.105 1.93355.40 12.85 26.681 1.89954.35 12.60 26.168 1.86253.26 12.55 25.853 1.85053.55 12.45 25.820 1.83552.25 12.15 25.195 1.78951.11 12.00 24.765 1.78651.05 11.90 24.647 1.74050.05 11.65 24.147 1.70348.15 11.15 23.170 1.63646.76 11.00 22.679 1.63246.05 10.65 22.145 1.56644.90 10.40 21.609 1.52443.66 9.80 20.684 1.49643.26 10.20 21.005 1.51140.16 9.45 19.481 1.39438.51 9.50 19.127 1.34236.91 8.60 17.816 1.28835.71 8.00 16.902 1.21835.01 7.58 16.290 1.17934.26 7.00 15.486 1.11234.01 6.60 14.982 1.11533.26 6.00 14.126 1.03032.51 5.05 12.813 0.93331.96 4.05 11.377 0.83431.56 3.30 10.205 0.75331.41 3.00 9.707 0.72530.91 2.05 7.707 0.58430.46 1.05 5.655 0.42430.29 0.60 4.263 0.330
TABLE II. Entrance Transition NozzleCalibration Data
96
DiametralPosition(inch)
.25
.50
.751 .001 .251 .501 .752 .00
3
.25
.50
.751 .001 .251 .501 .752 .00
4.251.504 75
1. 001. 251. 501. 752. 00
i 25a 501 75
1. 001. 251. 501. 752. 00
T (°F)
800808816818820820814810802
804810817820818820818816809
806815823826826818810806792
801810815820822822812808798
TUPT (°F)
852852854853855853853854855
854854854855855854855856855
856856856857858860858858858
855855856856856855854854855
Nozzle A
Nozzle B
Nozzle C
Nozzle D
TABLE III. Primary Nozzle Temperature Profile Data
DiameralPosition(inch)
i
.25
.50
1 .00
1 .50
2 .00
2,.50
3,.00
3.,50
4.,00
4.,50
5.,00
5. 50
6.,00
6.,50
7.,00
7. 25
7. 50
TUPT (°F)
773
816
826
840
849
852
856
860
848
850
854
855
852
849
842
832
835
819
TREF (°F)
851
851
853
855
857
856
856
863
852
850
854
858
858
859
857
848
852
848
TABLE IV. UPtake Stack Temperature Profile Data
98
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112a
DIAMETRALPOSITION(INCH)
• 25
• 50
• 75
1. 00
1. 25
1. 50
1. 75
2. 00
2. 25
2. 50
2.,75
3.,00
3.,25
3.,50
3,.75
4.,00
4,.25
4,.50
4,.75
5 .00
5 .25
5 .50
5 .75
6 .00
6 .25
6 .50
7 .00
7 .25
7 .50
TEXIT TUPT(°F) (°F)
393 845
397 845
400 847
416 846
420 846
432 847
441 847
449 849
452 850
463 850
468 849
475 849
482 851
496 852
497 853
497 851
501 851
501 851
500 853
498 853
494 850
492 852
484 854
480 855
471 856
462 855
458 856
452 856
440 856
418 858
TABLE VII. Mixing Stack Exit Plane TemperatureProfile L/D =3.0
113
DIAMETRALPOSITION(INCH)
• 25
• 50
• 75
1. 00
1. 25
1. 50
1. 75
2. 00
2. 25
2. 50
2. 75
3.,00
3.,25
3.,50
3.,75
4..00
4,.25
4,.50
4 .75
5 .00
5,.25
5 .50
5 .75
6 .00
6 .25
6 .50
6 .75
7 .00
TEXIT TUPT(°F) (°F)
385 848
395 848
403 852
412 852
419 851
426 851
437 851
443 852
450 853
455 « 853
460 855
464 855
471 855
480 857
483 855
491 852
498 851
500 851
491 852
486 852
480 851
476 851
472 851
468 853
• 458 852
452 852
446 853
440 851
435 853
TABLE VII. (Continued)
114
DIAMETRALPOSITION(INCH)
.25
.50
.75
1.00
1.25
1.50
1.75
2.00
2.25
2.50
2.75
3.00
3.25
3.50
3.75
4.00
4.25
4.50
4.75
5.00
5.25
5.50
5.75
6.00
6.25
6.50
6.75
7.00
TABLE VIII. Mixing Stack Exit Plane TemperatureProfile Data L/D = 2.5
TEXIT(°F)
TUPT(°F)
416 856
512 856
523 857
530 858
540 859
548 858
555 857
564 858
572 857
579 858
585 858
594 860
594 859
596 863
594 862
590 862
582 859
571 856
560 855
546 855
532 857
516 858
502 856
490 857
480 859
460 862
452 862
436 862
429 863
115
DIAMETRALPOSITION(INCH)
i.25
<.50
<.75
1,.00
1,.25
1,.50
1..75
2,.00
2,.25
2,.50
2.,75
3..00
3.,25
3.,50
3..75
4..00
4.,25
4..50
4..75
5.,00
5..25
5.,50
5.,75
6..00
6.,25
6.,50
6.,75
7..00
PEXIT(°F)
TUPT(°F)
476 857
483 857
497 858
506 857
519 858
532 859
548 859
561 859
569 859
577 860
581 860
588 859
592 860
593 861
593 861
588 861
580 861
570 861
558 861
551 861
537 860
523 860
509 860
500 862
494 862
480 863
468 862
456 862
446 862
TABLE VIII. (Continued)
116
Variable Value Uncertainty
Ts 525°R ± 1°R
T 1316°R ± 1°R
B,Pa 30.04 in Hg ± .01 in Hg
APN 6.05 in H2<D ± .;05 in H
2
PEH 8.50 in H2
. ± .05 in H2
APS .52 in H2
± .005 in H2
<D
FHZ 103 Hz ± 1 Hz
PNH 3.90 in Hg ± .02 in Hg
r 1.126 in ± .005
Values in this table are from L/D = 2.5,A /A = 2.5, TUPT = 850°F (Table VI)m p
TABLE IX. Uncertainties in Measured Values fromTable VI
117
APPENDIX A
OPERATION OF THE COMBUSTION GAS GENERATOR
A. COMPRESSOR LIGHT OFF
The primary air flow is supplied by the Carrier Model
18P350 centrifugal air compressor located in Building 248.
This compressor's cooling system is piped into the cooling
tower system located behind the building. Figure 45 gives
a schematic of the compressor layout.
In preparation for compressor light off ensure that the
cooling water valve to the Sullivan compressor is closed,
then start the cooling tower pump and fan by pushing both
start buttons located on the south wall of Building 248
(see Figure 46) . The compressor can then be started by
completing the following steps.
1) Ensure that the compressor butterfly suction damper
in the airstream between the filter (on the roof)
and the compressor is closed (Figure 47)
.
2) Open the inlet water valve to the oil cooler
(Figure 47) wide enough to obtain an adequate flow
of cooling water.
3) Start the auxiliary oil pump by positioning the
on-off automatic switch (Figure 48) in the "hand"
position, thereby by-passing the auxiliary oil
pump out-in control.
4) When the oil pressure rises to at least 16 PSIG,
adequate pressure exists in the bearings and the
118
compressor may be started by pressing the start
pushbutton. The compressor will then come up to
speed, at which time the auxiliary oil pump switch
is turned to the "automatic" position.
5) Open the compressor butterfly suction damper.
Precautions ;
1) During the period in which the compressor is coming
up to speed, the operator should check for:
(a) oil pressure in the range 20 to 22 PSIG
(b) any undue noise in the motor, gear, or compressor
2) During operation, check the bearing thermometers
periodically to ensure the bearing temperatures
do not exceed 185°F.
B. GAS GENERATOR LIGHT OFF
After the supply air compressor is in operation, the
following is a recommended starting sequence.
1) Energize the main power panel and the thermocouple
and mass flowmeter readouts.
2) Calculate the required mass flow rate to achieve
the desired uptake Mach number, M . The formula
for this calculation (derived in Reference [4])
follows:
.0.5
M0502 (m + m^)TUPT
a f
u(
PU) + B
( 13TT7 ) + B
where
119
TUPT = uptake temperature (DEC R)
PU = uptake pressure (in. H~0)(PU can be assumed to Be 0.0 for thefirst iteration)
B = atmospheric pressure (in. Hg)
m = mass flow rate of air (lbm/sec)EL
mf
= mass flow rate of fuel (lbm/sec) .
(nu can be assumed to be .01 lbm/secfor this calculation)
3) Figure 15 gives the primary air mass flow rate versus
the pressure product. The pressure product comes
from the transition nozzle calibration and is
defined as:
( (PNH + B) * APN)*
where:
PNH = nozzle high pressure (IN. Hg)
B = atmospheric pressure (IN. Hg)
APN = pressure drop across nozzle (IN. H20)
From Figure 15 find the pressure product corresponding
to the required mass flow rate found in Step 2 above.
4) With the burner air valve 100% open and the bypass
air valve (see Figure 3) 50% closed, open the main
air supply globe valve (Figure 49) until the desired
pressure product is reached.
120
5) Adjust the bypass air valve until the U-bend
pressure difference (APu) is one inch of water.
6) Turn on the fuel supply pump and the high pressure
fuel pump.
7) Adjust the fuel control valve to obtain 150 PSIG
on the high pressure fuel gage (see Figure 5)
.
(Note: It is desired to obtain about 115 Hz on
the fuel flow meter, but this reading is not avail-
able until the fuel shutoff valve is opened. The
fuel pressure is therefore used as an initial
approximation of fuel flow rate.)
8) Energize the igniter plug and glow coil by depressing
the igniter switch. Hold this switch down for a
few seconds before opening the fuel shutoff.
9) Open the fuel shutoff valve by putting the emer-
gency shutoff switch in the "on" position. Watch
the fuel flowmeter; the reading should quickly
come to the 110-120 Hz range. Ignition should be
noted within 3-4 seconds. If ignition does not
occur quickly, turn off the emergency shutoff switch.
10) If ignition does not occur and the fuel flowmeter
indicated a flow outside the 110-120 Hz range,
adjust the fuel control valve to achieve a reading
in this range and repeat the procedure starting at
Step 7.
11) If ignition does not occur and the fuel flowmeter
indicated a flow in the 110-120 Hz range,
121
a) Check to ensure the U-bend pressure differential
(APu) is one inch of water. If not, adjust
the cooling air valve to achieve this.
b) If the U-bend pressure differential is one inch
of water, check the igniter. The igniter can
be checked by activating the igniter switch with
no fuel flow and watching for a 3-5 degree
increase in burner temperature.
12) When ignition does occur:
a) Deactivate the igniter.
b) Begin closing the bypass air valve immediately
while monitoring burner temperature (T_.) .
a
Continue closing the bypass air valve until TDa
stabilizes. (Do not allow burner temperature to
exceed 1500°F.)
C. TEMPERATURE ADJUSTMENT
The temperature adjustment is an iterative process
consisting of the following steps.
1) Adjust the fuel control valve to achieve approximately
the desired uptake temperature (while monitoring the
burner temperature)
.
2) Check the pressure product. Re-adjust the main
air supply globe valve to obtain the correct value.
3) Adjust the fuel control valve and/or the bypass
air valve (see Figure 3) to achieve the desired
temperature. Rough temperature control can be
122
achieved with the bypass air valve and fine control
with the fuel control valve.
4) Go to Step 2 and continue until the pressure product
and temperatures are satisfactory.
SYSTEM SHUT DOWN
1) Close the fuel shutoff valve.
2) Turn off the fuel supply pump and the high pressure
fuel pump.
3) Allow the system to cool for 5-10 minutes.
4) Close the compressor butterfly suction damper.
5) Turn off the compressor. Immediately after turning
off the compressor turn the auxiliary oil pump
switch to the "hand" position.
6) Allow the bearing temperatures to -reach 80 °F before
turning off the oil pump and the cooling tower pump
and fan.
123
APPENDIX B
DETERMINATION OF THE EXPONENT IN THENONDIMENSIONAL PUMPING COEFFICIENT
The method used to determine the value of the exponent
n in equation (13) is outlined below.
(1) Select a given geometry, assume reasonable values
for K , Km and f, and calculate C, , C2
and C- for use in
equation (lib)
.
(2) Set T* = 1.0, AP* = 0, and solve for W*max.
Equation (lib) plots as indicated in Figure 20; for AP* =
and T* = 1.0, the intersection of the curve with the w*T*
axis yields the value of W*max. Note that for each value
of T* < 1.0 (T* = T /T and T < T therefore T* < 1.0) as/ p s p
different curve will result.
(3) For the same geometric configuration and other
values assumed and calculated in step (1) , calculate AP*/T*
using equation (lib) with W*T*n for different values of T*
in each case varying W* from to W*max in equal increments
of W*max. For each new value of T* tried, vary n until the
resulting plots of AP*/T* vs W*T*n for T* < 1.0 come close
enough to the initial plot obtained in step (2) where T* = 1.0
that, for all practical purposes, all such plots can be
represented by a single curve.
(4) The value of n which most effectively collapses all
performance curves onto the T* = 1.0 case is n = 0.44.
124
APPENDIX C
UNCERTAINTY ANALYSIS
The experimentally determined pressure coefficient and
pumping coefficient are used in determining eductor operating
points which in turn provide the basis for comparison and
evaluation of eductor system performance. A determination
of the uncertainties in these coefficients was made using
the method described by Kline and McClintock [10]. Data
for the eductor configuration described in Table VI is
considered a representative case and is used to calculate
representative uncertainties in the pumping and pressure
coefficients.
For a single sample measurement the value of a specific
variable should be given in the format:
x = x ± 6x
where
x = mean value of the variable x
6x = estimated uncertainty in x.
Variations for the variables in the defining equations for
the two coefficients are listed in Table IX. Having des-
cribed the uncertainties in the basic variables of a
125
relationship, it is now necessary to determine how these
uncertainties propagate into the result. Consider the
relation where the result R is the product of a sequence
of terms.
R = Xl
X2
X3 ^
A reasonable prediction of the uncertainty in the result
R is obtained by using the Second Order Equation suggested
by Kline and McClintock [10]
.
o R r. « ^ . <3R p > Z /oR c> \-3-i
6r = [(air «*!> + (35- 5V + (^T 6x3
) ]
1 2 2
(b)
Evaluating the partial derivatives appearing in equation (b)
,
and normalizing by dividing through by result R yields the
simplified form of equation (b) which will be used in this
analysis.
XT3 a 6x 9 b 6x 9 c 6x 1/2
TT= [(-3E
Jl) + {~1TJl) + (~ST
2* J (c)
Determination of the uncertainty in the pressure
coefficient is facilitated by writing it as the product
of a series of terms,
£E± = (p^" 1(AP) (U
p)" 2
(T*)"1
(d)
126
where P represents the pressure difference (P - P rt )
.
a
Constants such as 2 g in the equation for the pressure
coefficient will be cancelled out when used in equation (c)
and are therefore not included in this analysis. Applying
equation (c) to the pumping coefficient in equation (d)
yields the following expression for its uncertainty:
AP*6 %r (-1) 6 PT* _
r,
v
1H18.2 , , (1) 6(AP)X2
AP* L K
p '+ ( AP '
m* S
(" 2) 6U
p 2 (-1) 6T* 21/2
+ ( oE
) + r l)
T * )
Z] (e)
P
Taking into account the respective equations defining the
individual variables, the terms of equation (e) are
expanded as follows:
P 6p2
<5P2
6T2
PEHt
6 PP]2 = [^PEH] 2 + ^ m\ 2
p = R T ' P„ PEH T^p P P P
UW 5U ~ 6W 6p 5A^
?-. I^l 2= t(^H) 2
+ (-E) 2- (^£)
21
P P A 'L U J L W '
v
p ' A^F p P p P P P
127
A^ = 47rr2
, h-£]2 = C
2 6r)
2
P ' L A J ^ r '
Wp
= .0310 + .0704((PNH+B) -APN)
*
5+ .00009GHz,
<5W2
Cl
6PNH 9 C o 6B o c o <SAPN C. 6FHz^ = ^ )+ ^4—
)
2+ C-Sj ) + (-^ )
2]
P p P . P P
C1
= .0310
C2
= .0704
C3
= .00009
!* rST* 2f\2
r
6TPl 2
P s p
Using the values of the variable and their respective
uncertainties listed in Table IX, the uncertainty in the
pressure coefficient is estimated to be
,A_P_*«<^>AP-
= .0187 = ± 1.9'
By a similar process, the uncertainty in the pumping
coefficient is estimated to be
6(W *T*'!^ ) = .0213 = ± 2.1% .
W*T**44
128
BIBLIOGRAPHY
1. Ellin, C. R. , Model Tests of Multiple Nozzle ExhaustEductor Systems For Gas Turbine Powered Ships ,
Engineers Thesis, Naval Postgraduate School, June 1977.
2. Moss, C. M. , Effects of Several Geometric Parameterson the Performance of a Multiple Nozzle Eductor System ,
MS Thesis, Naval Postgraduate School, September, 1977.
3. Harrell, J. P., Experimentally Determined Effects ofEductor Geometry on the Performance of Exhaust GasEductors for Gas Turbine Powered Ships , EngineersThesis, Naval Postgraduate School, September, 1977.
4. Ross, P. D., Combustion Gas Generator for Gas TurbineExhaust Systems Modeling , MS Thesis, Naval PostgraduateSchool, December 1977.
5. Staehli, C. P., and Lemke R. J., Performance of MultipleNozzle Eductor Systems with Several Geometric Configura^tions, MS Thesis, Naval Postgraduate School, September,1978.
6. Keenan, J. H. and Kaye, J., Gas Tables , John Wiley andSons, Inc. , 1963.
7. Pucci, P. F., Simple Ejector Design Parameters , Ph.D.Thesis, Stanford University, September 1954.
8. Boeing Airplane Company, Boeing Model 502-2E Gas TurbineEngine , February, 1953.
9. Carrier Corporation, Operating Instructions for CarrierModel 18P352 Air Compressor , October, 1955.
10. Kline, S. J. and McClintock, F. A., "DescribingUncertainties in Single Sample Experiments," MechanicalEngineering , p. 3-8, January, 1953.
129
INITIAL DISTRIBUTION LIST
No. Copies
1. Defense Documentation Center 2
Cameron StationAlexandria, Virginia 22314
2. Library, Code 0142 2
Naval Postgraduate SchoolMonterey, California 93940
3. Department Chairman, Code 69 2
Department of Mechanical EngineeringNaval Postgraduate SchoolMonterey, California 93940
4. Professor Paul F. Pucci (Code 69Pc) 10Department of Mechanical EngineeringNaval Postgraduate SchoolMonterey, CAlifornia 93940
5. LT Charles R. Ellin 1
13512 Westwind DriveSilver Spring, Maryland 20904
6. Mr. Charles Miller 1
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7. Mr. Olin M. Pearcy 1
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8. Mr. Mark Goldberg 1
NSRDC Code 2833Naval Ship Research and Development CenterAnnapolis, Maryland 21402
9. Mr. Eugene P. Wienert 1
Head, Combined Power and Gas Turbine BranchNaval Ship Engineering CenterPhiladelphia, Pennsylvania 19112
10. Mr. Donald N. McCallum 1
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No. Copies
11. Lt. J. P. Harrell, JR., USN 12004 Cloverleaf PlaceArdmore, Oklahoma 73401
12. Lt. C. M. Moss 1625 Midway RoadPowder Springs, Georgia 30073
13. LCDR P. D. Ross, JR., USN 1
673 Chestnut St.Waynesboro, Va. 22980
14. Lt. D. R. Welch 2
1036 Brestwick CommonsVirginia Beach, Virginia 23512
15. Lt. R. J. Lemke 1
2902 No. CheyenneTacoma, Washington 98407
16. Lt. Chris P. Staehli 1
Route 2 Box 648Burton Washington 98013
17. Professor R. Nunn, Code 69Nn 1
Department of Mechanical EngineeringNaval Postgraduate SchoolMonterey, Calif. 93940
131
2*7*1
Thesis 1781*7*1W3885 Welchc .l Hot flow testing of
multiple nozzle ex-
haust eductor systems.
thesW3885 ,
Hot flow testing of multiple nozze exha
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