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

sSTANDOFFDISTANCE

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PRIMARYNOZZLES

^_V

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UPTAKE

A

Mu

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WWUPTAKE MACH NUMBER

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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

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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 -

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63

FIGURE 19. Primary Flow Nozzle Plate (Back View)

64

Js FIGURE 20. Primary Flow Nozzle Plate (Front View)

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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

NAVSEA Code 0331Naval Ship Systems CommandWashington, D. C. 20362

7. Mr. Olin M. Pearcy 1

NSRDC Code 2833Naval Ship Research and Development CenterAnnapolis, Maryland 21402

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

NAVSEC Code 6136Naval Ship Engineering CenterWashington, D. C. 20362

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

3 2768 001 95199 9

DUDLEY KNOX LIBRARY