ORNL-2376.txt

 

 

 

   

    

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3 4456 D3L1LaY &

AEII RESEARCH A

 

CENTRAL RESE A

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DOCUMENT COLLECTION

MENT REPORT c-e. “Reacrors-steinf 7

Features of Aircraft Reactors

  
   
   
   
  

ART FUEL PUMP AND XENON REMOVAL
SYSTEM DEVELOPMENT, TEST EVALUATION,
AND AIRCRAFT APPLICATION

  
   
   
 

G. Samuels
M. E. Lackey

5 3 ‘ fl A
BECLASSIFIED
CLASSIFICATION CHANGED To: * x % (,S i SR, E./

e o e e i dem - % -

     
 

By Ayt “iou[ TY OF:_ _Q'Q.\__C- ____: — e .._._._(L"...,B

Lot o e
By: ..

fil@d..(_x.?b_m_é- el

  

 

OAK RIDGE NATIONAL I.ABORATORY
OPERATED BY

 UNION CARBIDE NUCLEAR COMPANY

Division of Union Carbide Corporation

vee

Y POST OFFICE BOX X * OAK RIDGE, TENNESSEE

 

 
 

 

 

 

 

WL';EGAL‘_‘NOTI_CE

Th!s reporr was prepared as -an account of Government sponsored work Ne lrher
nor the Commlsslon, nor: any’ person octihg on. behalf of the Commlsslon. . a5

A, Makes ony warrunly or représentation,- express or lmplled wufh respec,

 

completeness, or usefulness of the’ informotion contained ‘in’ H-ns ‘repo 't, oF rhut fhe Use
.uny mformohon -upporotus, rnefhod .oF process dlsclosed in ‘this rep)orf moy: . not |nfr| .
prwcnely owned rights; or A
B. Assumes any hol:nlmes wn'h respect fo the use of, or for dumages resulf
any mformohon, opperufus method or process disclosed in this reporf. '

As used |n the obove, person oclmg on beholf of the Commlsslon

or' distributes, or prowdes access to, any ‘information pursuant f6. hls employmenf or confruci
wuth the Commlss:on. ) : Coe s

 

 

 

 

    
   
 
 
 
 

 

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LT 1R 1 8

 

ORNL-2376 e
C~84%, AIRCRAFT REACTORS.

This documlent consists of 186 pages.
) _‘Copy':7éof 205 copies. - Series A.

Contract No. W-7405-eng-26

;-_A:'Lfcre.ft Reactor Engineering Division

ART FUEL PUMP AND XENON REMOVAL SYSTEM DEVELOPMENT
| TEST EVALUATION, AND AIRCRAFT APPLICATION e

G. Samuels
. M. E. Lackey

| DATE ISSUED
—DEC 194957

OAK RIDGE NATIONAL LABORATORY
Opera‘ted. by

UNION CARBIDE NUCLEAR COMPANY
A Division of Union Carbide and Carbon Corporatlon o
Post Office Box X - C ‘
Oak Ridge, Tennessee ‘

 
  
 
  
 

 

/I//WIT//I/fl//////i//fii/?/?]/fl/fll/?”ifl/?il??filfifi

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|

   

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1. S.
2. H.
3. D.
L. F.
5. E.
6. A.
7. C.
8. G.
9. E.
10. R.
11. D.
12. C.
13. R.
14, W.
15. B.
16. W.
17. F.
18. C.
19. L.
20. A.
21. J.
22. D.
23. W.
2L. B.
25. W.
26. A,
27. E.
28. C.
A.

H.

A

G.

W

M.

78-80.

81-82.
83.
8k,
85.

-1i-

 INTERNAL DISTRIBUTION

 

E. Beall 35.
W. Bertini’ 36.
S. Billington 37.
F. Blankenship 38.
P. Blizard 39.
L. Boch 4o.
J. Borkowski h1.
E. Boyd: 2.
J. Breeding b3,
B. Briggs Wi,
W. Cardwell L5,
E. Center (K-25) L6.
A. Charpie L.
G. Cobb L48.
Y. Cotton k9.~
‘B. Cottrell 50.
L. Culler 51.
W. Cunningham 52,

“B. Emlet (K-25) 53.
P. Fraas 5k,
H. Frye, Jr. 55.
E. Ferguson - 56.
T. Furgerson 5T.
‘L. Greenstreet 58.
R. Grimes - 59.
G. Grindell. 60.
Guth 61.
S. Harrill 62.
S. Householder 63 -6k,
W. Hoffman '

. Hollaender - 65-Th.
W. Keilholtz B T5.
H. :Jordan T6-7T.
T. Kelley :

EKTERNALIDISTRIBUTION

 

Alr Force Ballistic Missile Div151on

AFPR,
AFPR,
AFPR,
AFPR,

Boeing, Seattle

Boeing, Wichita . _
Lurtiss-Wright, Clifton
Douglas, Long Beach

4

e =

(')I.?:IC.'):fl'UI‘.l:ILq'J>

EHPPHARGPEORTANEE DGR

ORNL-2376

C-8k - Reactors-Special.
Features of Aircraft Reactors
M-3679 (20th ed. Rev.)

E‘:I"UUZU:JN‘?N.L!UUSSUID—J:U{Zt‘“'dt\'l<fi:>fl?clm_‘h>l?d

Lackey

Lane
Livingston
Manly

Mann

Mann
MacPherson
Megrehblian
Morgan
Murray (Y-12)
Nelson

. Perry

"McNally
‘Robinson
amuels
Savage .
Savolainen
Schultheiss

. Shipley

Skinner
Snell
Swartout

. Taylor

Trauger
Weinberg

. Whitman . '
. Wigner (consultant)

Winters

ORNL - Y-12 ‘Technical Library,

Document Reference Section
Laboratory Records Department
Laboratory Records, ORNL R.C.
Central Research Library

a
86-88:

90- 91
92,

- 93.

- 9k,
95.

96.
97-110.
111.
112.
113-115.
116.

117.

118.
119.
120-125,
126.
127.
128.
129,

130.
131.
132,
133.
134,
135,
136.
137.
138.
- 139-142.
143,

14k,
145,
146,
147,
148,
1h9

- 150.
151..
152,
153.
15k,
155.
156.
157.
158,
159-162.
163.

7 - ‘ : e SADPISE
-1ife - o R

AFPR, Douglas, Santa Monica

. AFPR, Lockheed, Burbank

AFPR, Lockheed, Marietta '

AFFR, North American, Canogda Park

AFPR, North American; Downey

Air Force Special Weapons Center

Air Research and Development - Command (RDGN)
‘Air Research and Development Command. (RDTAPS)
Air Research and Developmeént :Command (RDZPSP)
Alr Technical Intelligence Center

Air Unlver31ty Library = .- ‘

-ANP Project Office, Convair, Fort Worth
Argonne National Laboratory

Armed Forces Special Weapons Project, Sandia _
Armed Forces Special Weapons Project, Washington.
Assistant Secretary of the Air Force, R&D - -~
Atomic Energy Commission, Washlngton

Bettis Plant (WAPD)

Bureau of Aeronautics

Bureau of Aeronautics General Representatlve
BAR, Aerojet-General, Azusa =

BAR, Convair, San Dlego

BAR, Glenn L. Martin, Baltimore

BAR, Goodyear Aircraft, Akron .

BAR, Grumman Aircraft, Bethpage

Bureau of Ships '

Bureau of Yards and Docks

Chicago Operations Office

Chicago Patent Group

Curtiss-Wright Corporation

General Electric Company (ANFD)

Hartford Area Office

Idaho Operations Office

Knolls Atomic Power Laboratory

Lockland Area Office

Los Alamos Scientific Laboratory-

Marguardt Aircraft Company '

National Advisory Committee for Aeronautics, Cleveland
National Advisory Committee for Aeronautics, Washlngton
Naval Air Development Center

Naval Air Material Center

Naval Air Turbine Test Station

Naval Research Laboratory

Nuclear Development Corporation of America ]

Office of Naval Research

0ffice of the Chief of Naval Operations (OP 361)
Patent Branch, Washington :

Pratt and Whltney Aircraft Division

San Francisco Operations Office

et vt . ) L o
Carae ot e T . -
s | | - ~iv=

16k4. Sandia Corporation

165. Sandia Corporation, Livermore

166. School of Aviation Medicine .

167. USAF Headquarters

168. USAF Project RAND

169. U. S. Naval Radiological Defense Laboratory
170-1T71. University of California Radiation Laboratory, Livermore
172-189. Wright Air Development Center (WCOSI-3)
190-204. Technical Information Service Extension, Oak Ridge

205. Division of Research and Development, AEC, ORO

 
 

..v..

CONTENTS

Abstract . e+ e * & & 8 e e s ._~-o o _..;.o' -o LI o e @ -'. ._. . . .

. Section I -

DESIGN CRITERTA. ,'..; e e eeve e e e e e e

- Precepts and Requlrements . e s e e e s e . ,';fl.g{ . « v

Tllustrations and Nomenclature. . . « « & s o o o o s o ow'e o 4

Circuit Description . . ® . * * .u * . - . » 7:0 » .. e . ‘n ‘e u o

Primary Fuel Circuit + « o e o o ¢ o o o0 o o o o o o o o

BypaSS Fuel Circuito . _» o s & o s+ 2 s e & . ‘® e e & & o @ _.

Dynamic Seals 0 Ld 'C' . . e * . . . . * * . * . . * .. . e -

. Section II‘

DESIGN'AND DEVELOPMENT OF ART FUEL PUMP AND XENON REMOVAL SYSTEM_
Xenon Removal System Design « « + + «-.+ . ..;“.‘..f‘. .-, . .

- Xenon Removal System FLOW. . o o« & o o o o o 0 s o o o o o

- Xenon Removal System Gas Removal e« o o o .. I
ART Fuel Pump Development . . . . ctle e e e e e e
Comblned.ART Fuel Pump and Xenon Rémoval System.Develqpment .

SYStem Pressure Fluctuations’ . ;f.'.-. D ee e e e e e

SyStem Pressure Level R ,ou-c- ¢ el e e . ¢« o o _-o) s o o o ¢

Bypass Flow Determination &nd Control. « o o o o & 4 ¢ v o |
Fuel Leakage Past Upper Slinger Seal . . o o o s s o o e
- Experiment Results of Combined System Tests.. . « « & « .o &« s &

~ ART Fuel Pump Cavitation Characteristics . . . « « . . . .

Section III'

DESIGN OF AN ATTITUDE-STABLE XENON REMOVAL SYSTEM. .« + 4 o o+ - .

CiI‘CU.lt DeSCI‘:Lp'thIl - s e e e e [ e s e e ’ ¢ . s w | . . o e

System Pressure Control. c s e el e e e e e e e e e

Prlmary Serlo s 6 e e e % e e .+ o + e ‘e s s e e s @ - ‘

Second.ary SW].I‘J.. . o . . s . . » ol, - . * & '! . * . .0 .o . . .
Dynamic SealS. o« « o ¢ ¢ o o e s s 0 e e 000 e e s e

" Development Tests .. e e e e e e e e e e e e e e e e e

SyStem AnalySiS * . -l r » . . L) ¢ . LI * e+ v . * ¢ * . '.‘

13
13

Y

19

. 19

2l

2
, 2k
. 45

ks

» 51

66

- IO
. 195
. 15

go:.

. Bo
. B3

. 83

91>
-vi-

Contents (cont'd.)

Appendix A
SUMMARY OF WORK LEADING TO THE DESIGn;QF PUMPS FOR THE CFRE,‘ N, (0}
| Appendix B |

CENTRIFUGE THEORY AND C.ALCUI.I.ATIONS . “- . . s e . ’i. | . e o.- . 0 . . e . . - :’l]_':s

Appendix C

-Appendix D

CALCULATIONS FOR THE ART FUEL PUMP IMPELIER REDESIGN . . . % o . . . . 1133
Appendix E
BYPASS FLOW CALCULATION AND EXPANSION TANK HEATING . . + o » « « . . . 1145
| ,_Appendiqu | |

RADIOACTIVITY IN THE ART FUEL PUMP OIL SYSTEM o o v « 4 o « + o o o o . 2I55

.Appendix.G

Xe™32 POISONING + « ¢ v v v v v v v v e vt v e e e v e e e . 2165
ACmOWIEDEWMS ‘. " = ‘ ° » . B » . . » » - . * - . "‘ . 3 * ., . L J . ‘ . * . a 9‘1«73

REFERENCES . o ¢ o .‘.' v .. ‘e c‘ t e 2 e vo .c LI T ol - s e e s . c. . -‘o 17)-"
Figure

1.1

1.2
1.3
1.4

1.5°

2‘.1

2.2
2.3

'e.h.

2.5

2.6

2.7:
2.8
2.9u
$2.10
2.11
2.12
2,13

2.1k

—Vll—

 

LIST OF. ILLUSTRATIONS

ART Bypass Fuel Circuit « o« o« o« oo o &+ o o

System .« .

' Title'

| Vertlcal Sectlon Through Reactor Assembly
'ART Fuel Pump
, - Schematic ART Fuel Clrcuit.‘, “ i e s e

W1th Model 32 Impeller ;_.;

" ART Primary Fuel CLrcuit v . « o o o o

Brass ART Fuel Pump Impeller e s é- 9 s @

 ART Fuel Pump

Experiment

,Experiment
ART Fuel Pump
Experiment
ART Fuel Punip
Experiment
ART Fuel Pump
Experiment
ART Fuel Pump

Experiment_

ART Fuel Pump

Experlment_

ART Fuel Pump
PExperiment

ART Fuel Pump

Experiment

.. ART Fuel Pump

- Experiment

. ART Fuel Pump
_ Experiment
. Corrections Applied to the ART Fuel Pump " Total

Discharge Pressure for Obtaining the Centrlfuge

Performance CharacterlstlcsI
,No.

ART Fuel Pumpf

l-T&bleel.....

Performance Characterlstlcs,
NO. 2 -— T&ble 2 l ‘- e e " ® - @,

Performance Characteristics
NO.3-—T8.bl€21. e s s e

Performance CharacteristiCS‘-'
No. b - Table 2.1 . .

Performance Characterlstlcs‘
Nos. 3, 5, and 6 - Table 2.1

Performance Characterlstlcs‘
NO T - Table 2 l - . * '; .o'

Performance Characteristlcs-
Performance CharacteriStics
Performance Characteristics
NO. 13 - T&ble 2.1 o.o - ..
Performance Characteristics.
NO. ll" - Table 2.1 e e 0.8

Performance :Characteristics

Statlc Discharge Pressure . . . & + .

 

Frow w2
.

FulI Scale Plastlc Model of the Xenon Removal

Page
10
11
12
1k
15

- 20

25
3%

133
EN
35

36

37

38

39

4o

b1

ko

43
Figure

2.15

2.16
2,17

2.18
2.19
2.20

2.21

2.22
2,23
2.2k

2.25

2.26

2.27

2.28 .

2.29

P

- S

 

-viii-

Tftlé

. ART Fuel Pump Volute Performance Characteristics .

Comblned Xenon Removal and Fuel Pump Hot Test- Loop

ART Fuel Pump Bypass Flow Characterlstlcs with
Model 32 Impeller at a 3-1/2 in. Depth of -
Liquid (Water) in the Expansion Tank . . . . .

Upper Slinger Seal with Radlal Vanes . « + « « .+

Upper Slinger Seal with an Axial. Step R

ART Fuel Pump Performance Characteristics with
Model 32 Impeller at a 1/2 in. Depth of
Liquid (Water). in the Expansion Tank. . . .

. ART Fuel Pump Performsnce Characterlstics Vith

Model 32 Impellef at a1l in. Depth of

Liquid (Water).in the Expansion Tank ... . . .
ART Fuel Pump Performance Charactéristics'with

Model 32 Impeller at a 2 in. Depth of

Liquid (Water) in the Expansion Tank c e e e

ART Fuel Pump Performance Characteristics with
Model 32 Impeller at a 3 in. Depth of _
Liquid (Water) in the Expansion Tank ... .

ART Fuel Pump Fluctuation Characteristics with
Model 32 Impeller at a 1/2 in. Depth of
Liquid (Water) in the Expansion Tank . . . . .

ART Fuel Pump Fluctuation Characterlstlcs W1th
Model 32 Impeller at a 1 in. Depth of _'
Liquid (Water) in the Expansion Tank - . . .

ART Fuel Pump Fluctuation Characterlstlcs with
Model 32 Impeller at a 2 in. Depth of
- Liquid. (Water) in the Expansion Tank . . . . .

ART Fuel Pump Fluctuation Characteristics with
Model 32 Impeller at a 3 in.: Depth of

Liquid (Water) in the Expansion Tank . . . . .

Calculated and Experimental ART Fuel Pump Bypass
Flow Characteristics . . . . « « &« o0 .o « .

Cavitation Characterlstlcs of ART Fuel Pump W1th_
Model 32 Impeller at a Pump Speed of 2400 rpm
and a 3 in. Depth of quuld (Water) in the Ex-

' pan31on Tank . '.°'. . e s 4 e o e s e s e s

s

Page
L
46

50

35

57
58
"55 o
&
6
62
63

6L

65

.68
Figure

2 .’3‘-0-

2.31

2.32

3.1
3.2

3.3+

3.k
3.5
3.6
3.7
3.8
3.9
3.10
3.11
3.12
3.13
3.14

3.15

To.

 

- ix;- .
Title

CaV1tat10n Characterlstlcs of ART Fuel Pump with

~ Model 32 Impeller at a Pump Speed of 2700 rpm
~and.’a 3 in. Depth of quuld (Water) in the Ex-
~pansion Tank . . . . . “ e e e e e e eo e

Cavitation Characteristics of ART Fuel Pump. W1th
Model 32 Impeller at a Pump Speed of 3000 rpm
and a 3 in. Depth of Liquid (Water} in the Ex-
pansion Tank . . o o« o o o o s o o .6 = o o

Minimum Expan51on Tank Gas Pressure Required to
Prevent CaV1tatlon in the ART Fuel Pump e o o

' Attltude-Stable Xenon Removal System e o e e e

Attltude—Stable Xenon Removal System Rotary
Assembly Test Model No. 15 e e e e e eTe

Schematic of Attltude-Stable Fuel Circuit: . .
Attitude-Stable System Pressure Control . . .. .

Attitude-Stable System Primary Bypass. swirl . .

' Attitude-stable System Secondary Bypass Serl .

 Attitude-Stable System Pump Shaft Shroud . + . .
_AttitudeAStable System Centrlfuge,Sllnger Seal..

. Attitude-Stable System Swirl Chamber.Energy'

Balance c s e s s s 5 e s e e-3 o o e e e v e

Attitude-Stable Xénon Removal System Test Mbdel

. NO. 7 . ' ‘.' I . * * . e o o o . o . o 7‘ * .o *

Attltude-Stable Xenon Removal System Test Model
No ll . . o 0 . @ - 9 ‘e . . .- o » O G o . * 9

Attltude-Stable Xenon Removal System Test Model
NO. 15 - - - . . . ° . . . . ® - ° o s e o .

Attltude-Stable System Maln Fuel Clrcult Charac-
teristlc ‘. - > - . . L ] o > o . o 'D T e o e - &

-Attltude-Stable System Bypass Clrcult Character-:'

lstic s & s & & e o 8 o -0 e & & e @ u"o e e »

fAttitude-Stable System Parallel Main Fuel and

BYP&SS FlOW ’Cl'l&raC'teriS'biCS e o o5 o @ ‘e e s

!

2
Fes

‘Page

69

Y

. 178

179
181

182

. 184

;Bs,
187
189,
159

100

101
101

. 1101
Figure

A.l
A.2

A.3

B.1

B.2

B.3

C.1

c.2

C.3 |

C.4
C.5

C.6

Dll
D.2.

D.3

D.ha

D.kb

-X -

 

“Title |

.Prellmlnary Model No. 1 of an Attltude-Stable

. Xenon Removal System . v « o v o e o0 .

Prellmlnary Model No. .2 of an Attltude-Stable
- Xenon Removal System e s s s e s ee e-e- e

Preliminary Model N6. 3 of an Attitude-Stable

. Xenon Removal System « « o ¢ o ¢« o o o« &

Correlatlon of Drag Coeff1c1ent and Reynolds

Number for Bubbles inWater . . « i « ¢

Effects of Bubble Radlus and Exit Veloc1ty on
Centrlfuge Effectiveness at 2700 rpm. W1th

‘a Cup Radius of 2-7/84An. v 4 o e o ,

Effects of Bubble Radius and Exit Velocity on
Centrifuge Effectiveness at '1350- rpm with

a Cup Radius of 2-7/8 in. . . . . . ..o o

| Head Characteristlcs of a Fluid Rotatlng as a

SOlld Body * L o . . . - » a . - a 0 . . .

Leakage Characterlstlcs of a Dynamlc Seal W1th

an -Axial Clearance of:0.050 in. . . . « &

Leakage Characteristics of a Dynamic Seal‘fiith

- an ‘Axial Clearance of O. 128 iNe v o o o &

Leakage Characteristics of a Dynamlc Seal with

an Axial Clearance of 0.248 in. . . . .

- Leakage Characteristics of a Dynamic Seal W1th

an Axial Clearance of O. 481 in. . . . . .

Zero Leakage Characteristics of =a Dynamic Seal

Pump Inlet Velocity Diagram . . + . +/¢ « o .

Redesigned ART Fuel Pump Blade Trace and Fluld

Passage Layout « ¢« ¢« o o ¢ ¢ o o o & 2 0 &

Veloc1ty Proflle Entering the ART Fuel Pump
Impe lle r ® - e ® . *® . . e . *® * . ‘o . . *

Plane Parallel to Fuel Pump Shaft Centerllne

Plane Perpendlcular to Fuel Pump Shaf't Centerllne

.Pagé

107

110

118
119

120

125 |

126

. 127

128

129
130

133
135

136
140
140
Figure

D:hg
D.5a
D.5b .
D.5¢c

D.5d

E.3

E.4

G.1

G.2

G.3 -

G.l

 

_,,,:' |

o
B~ L
i 2

-xi-

Title

'.Fluld Streamline along Fuel Pump Shroud .« + + + 4

Blade Trace e * s+ s+ s s 0o . o . e . c s 0 s . ¢ o e

Fluid P&SS&SG o> . . . . . . o 'o . .-- -. o @ . ° . . e

Distance from Tip Along Reference Radius . . . .

Blade B:Lanl: . a. . . o e . s e -"' 9’ oo! o ..._,. _o' o

Calculated ART Fuel Pump Bypass Flow Charac-
terlsthS c - e e . o . e o * e & ~®- @ 0"'. .. .

ART Reactor Fuel and NaK Temperatures at a Constant
Ppmp Flow of 645 gpm o o v o ¢ o0 o o o « o o &

| The Effect of Reactor Power Level on the ART Ex-

- pansion Tank Liquid Level o o o o o e o sive o s

Mixed Mean Temperature of Fuel in the ART Expan51on
Tank as a Function of Reactor Power to Fuel and
Pump Speed at a Constant Fuel Flow of 645 gpm
per PUlpP o o o o o o o o 0 o o0 o0 o 0 s o a

Schematic of Flow Through Sparger and.Expan51on
T&I].k ... o"o ® e s e e e e e 0 & & & & o @ b & e

Xenon Solublllty in Fuel-30 « « & o v o o o o o v

Xenon Concentratlon at Equlllbrlum 1n Fuel 30
with a Power Generation of 60 MW v « & & o « « &

,“Xe135 Poisoning in Fuel-30 at 60 Mw Power and a

Helium Bleed of 1000 Liter/day « « + « « « « « .

|

Page

1ho

. 1h1
. Ah41
L AhT

41

147

148

. 149

150

. 169

a70
171

172
_‘-xii-
o

LIST OF TABIES

 

1.1 Design Requirements « « o o « o o o o o oo o o s o T
1.2 Design DBte « o o o o o o o o o 0 e s 0 0 oo o o u 9 .
2.1 : Conditidns and Resuits of water:PErformafice |

Tests of ART Fuel Pump « + o « o ¢ ¢ 4 ¢ o o o o o _29.
3.1l Model'lS - PErforfiance Evaluafion e s s e e e 193
3.2 | Séquéntial Develdfimént of Aftitfidé{Stgble Xenon

Removal System « ¢ ¢ o o o ¢ o 0 0 o o o o 0 o o gl

33 . Mode% TmffifErqumance Evaluation T:_i'g..'..; - (9.
3.4 | Modei”lfi - Dimensions . .. .1;-. . ;‘: .l. o s e e s 198
F.l Concentration and Activity df“Kryptofi, Xenon and

Daughter Products in Expansion Tank and 0il _ _
System for a Helium Flow Rate of 1000 liters/day. . 162

G.1l Xenon Concentration in Expansion Tank Gas J
VOl'(me 0700?_0'.0000-{1o-o:.oo.‘-oo.. I68’
ABSTRACT

This report summarizes the design and development of the fuel pump and
xenon removal system for the Aircraft Reactor Test. The system is comprised
of two fuel pumps-with an expansion tank mounted just above énd between them,
and a complex of impellers, seals, and stators arranged to be a part of the
removable fuel pump assemblies. The xenon removalisystem is designed to
bleed appfoximatély 1.5% of the main fuel flow rate'ihto the expansion _cham-
ber through the xenon removal equipmenf; The stripped fuel is returned to
the main system by way of centrifuges'mounted immediately in back of the fuel
pump impellers. The .centrifuges serve to filter out any entrained bubbles
and pro#ide in conjunction with seal impellers a suitable main fuel system
pressure level. | A |

_The initial design and development work included as an obJjective the
achievement of a system insensitive to attitude. That work is included in
this report as an aid to future designs.

This report inclfides the following:

:l. -Thé‘undérlying design precepts apd requirementé upon

which the development was based.

2. The functioning and principle of operation of the finalized
design. - |

3. The chronological evaluation of the finalized design and

the major design changes in the development. '

4. Discussions and calculations pertinent:to five design
criteria: | . |
a. pump impéller deSign'
b. system pressure control
c. swirl stability.
d. bubble removal in the centrifuge

e. seal requirements and performance,

H—
.
A

Section 1

'DESIGN CRITERTA
 

DESIGN CRITERIA

Precepts and Requirements "

The high absorption cross section of Xe_35

makes 1t extremely desirable
to strip xenon continuously from all or a fractlon of the fuel¥* 1n high power j
density reactors. Xenon can be stripped from liquid fuels by éxposing the-
fuel to a gas. interface where the xenon molecules can diffuse from the - liquid.
The rate of xenon extraction from fuel at a given temperature 1s a function
of the xenon concentration in the fuel the partial pressure of xenon in the .
contacting gas, and the rate of exposure. Thus, a xenon removal, or Y <R",
system might function as follows: . - |
l, A fraction of the fuel could be bled 1nto a proce351ng chamber.
2. A large interface in thls chamber could be prOV1ded by means
| of agitation, bubble entrainment or spraying._m A
3. The xenon-free fuel could be returned to the system in such
' a manner as to av01d foam or gas bubble carry-over.

To appreciate the X-R system as applied to the. Aircraft Reactor Test, it
must be clearly understood that the initial conception involved five basic pre-
cepts, namely ' o . _

1. From l to 2% of the. fuel should be bled continuously from
o the main fuel stream 4
| 2. The interface in. the expansion tank should be used for ‘the f
: removal of. xenon.
3. The pump shafts should supply any . energy required for the
operation of the system: H . o _
L, A centrifugal field should be employed to prevent gas carry-
, over into the main fuel stream to.avoid affecting reactivity.
5. The X-R systempshould opefate in stable equilibrium.with the

fuel pump and maintain the desired fuel system pressure .level.

  
 

 

¥ A Teactor 1nherently'will reach'Some& % brium xenon concentration
determined by the balance between xenon generation on the one hand
and xenon decay and "burn-up"” on the other. Any concentration below
this inherent equilibrium concentration implies xenon stripping.

Tl
 

These design precepts indicate the intimate relationship of the design
of the X-R system to the control of the system pressure, expansion tank liquid
level stability, and the pump de51gn with all of its seal and access1bility
problems. The multitude of interrelated design requirements mean that the
X-R system mst do a great deal ‘more than remove xenon. A summary of the de-
~sign requirements is given in Table 1. 1. ' - |

The interdependence of these design requirements is not immediately ap-
‘parent. After preliminary study, it became evident that the complexity of
the complete system demanded that the X-R system and the fuel pump be ‘developed
separately and subsequently tested as a unit. The development of the X-R sys-
tem and the .pump is discussed in Sections II and III.

The basic design data and phyS1cal pr0pert1es used in the des1gn of the
X-R system are given in Table l.2. Most of the dimensions were predetermined
by the pump configuration and space 11m1tations. Certain critical dimensions

were computed, The detailed calculations are given in the appendix.

I1llustrations and Nomeénclature

:Fig.'l,l gives an overall picture of the fuel pumps'and'the X-R system.'
Illustrated particularly are the. expansion chamber in relation to the rotary
elements, the 1nter-connecting parts, and the direction of fluid passage. The -
expansion tank in the fuel system acts as a processing tank for Xenon stripping.
All other required Punctions are. carried out in the rotary assembly, a cross
section of which is shown in Fig. 1.2 with the various parts labeléd..

Fig.'l;3 depictsnthe system as a’hydraulic circuit. The numbers are only
illustrativemof the pressure_interdependence between the X-R system and the
main fuel circuit. The Punctional description of the ART-XR system to follow
is presented'from'the point of view of a hydraulic circuit.

=3

.
S
oy
g
4
1.

10.

1l.

12.

‘13,

1k,

15.

  

Table 1.1 - De31gn‘fl>quirements

The pump should have the highest possible degree of reliability'which

_ necessitates large radlal and. axial clearances for mov1ng parts to

mlnlmlze the effects of thermal dlstortlon. B

s The pump shaft seal and bearing must. operate at temperatures below 300° F.

The pump should not constltute a hole in the shleld,

'The pump should be replaceable without removal of the reactor ShlEld.

The bearing and seal should be sufflclently shielded frcm radlatlon to

give a satisfactory iubricant 1ife.

Cooling prov151ons must be adequate to remove the heat generated by
rediation at full power and at the same tlme no areas contacted by llquld
fuel may drop below the fuel free21ng p01nt at zero power.

-The pump-mnst.supply theé requlred fuel system head and flow.

The pump must be free from cavitatlon within the requlred operatlng |

pressure level.

The pump must meet the reactor geometry demands.

Pump, casing, and expansion tank should: be rugged in construction, un- _
likely to be subgect to thermal distortion, and have a.mlnlmum of moving
parts. ' '
Operation should not be sensitive to variations in pump,speed or to
differences in speed between pumps, includiné one pump stopped.

The pump shaft power required should be minimized.

Fuel entry into the cool region around the impeller shaft below the seal
must be prevented.

The system should be ‘such as to drain naturally when in its normal position.

Provision must be 'made for removing xenon with a high degree of reli-
aebility and sufficiently thoroughly to provide a substential margin over

the concentration acceptable from the reactor control standpoint.

 
16.

17.

18.

19.

20.
el.

22.

25«

 

To. effect gas removal from the fuel & fraction should be recycled from
the main'cirCuit through a ¢hamber,with a helium atmosphere.

To 1ncrease the dlssolved gas removal effectlveness, the fuel in the

gas removal chamber should be agitated to increase the helium entraimment
thereby increasing the surface area and reducing the diffusion path length
in the fuel. _ |

The stripped fuel must be returned to the main fuel circuit free of gas
bubbles. o

To prevent bubbles from enterlng the core the return c1rcu1t must be
designed without leakageo

.Thejgystem pressfire must be maintained at a predetermined lével§

An expansion volfime must be provided for the fuel.

Clrculatlon.must be maintained through the expansion tank to avoid over-
heatlng from fission product decay heat.

Reactor off-gas and fuel vapor contamingtion of the pump lubrlcatlng oil

should be & minimum.

 
 

Table 1.2 - Design Data

Fuel No. 30

Density'(#/ft3)
Surface Tension (dynes/cm)

Viscosity (#/ft-hr)

Liquidus Temperature

Water
Density (#/f£3)
Surface Tension (dynes/cm)
Viscosity'(#/ftfhr)
Pump Shaft Spacing (in.)
Fuel Pump Impeller 0.D. (in.)
Fuel Pump Impeller I.D. (in.)

Total Fuel Volume Outside
Expansion Tank (ft3)

Change in Expansiqn-Tank

- Fuel Volume from Fill Temperature
to Operating Temperature (in.3)

Bleed Flow Rate per Pump (gpm)

Value

= 246.4-0.03227(°F)
530°C-157
630 -132
730 -115
1100°F-21.3
1300 . -12.8
1500 =8.5
968°F

- 62.3
68°F-72.8
68°F-2.42

21
2.7
3.5

8.78

568
12

 

Source

(2)
(2)

(2)

(2)

(3)
(1)

)

Calculated

Calculated
(1)
FUEL PUMP

     
  
   
 

CONTROL ROD

Na EXPANSION TANK
NaK

FUEL EXPANSION TANK

HEAT EXCHANGER
ASSEMBLY

-

Na-TO-NaK
HEAT EXCHANGER

 
 
  
 
 
 

4= Na INLET

REFLECTOR
ASSEMBLY

 
  

“

2

-

3

THERMOCOUPLE z
~-

W

2

w

TILE LAYER

af—

BERYLLIUM |
REFLECTOR-—

    
 
  

FUEL~TO-NaK
_HEAT EXCHANGER

. SPHERICAL B,C
TILE LAYER

 
 

FILLER PLATES

 
 

THERMOCQUPLE SLEEVE
FUEL DRAIN {HOT)

FIGURE II VERTICALZSECTIBRATHRC

 
PSS

ORNL—-LR—DWG 18291

   
 
    

PUMP SHAFT

UPPER SLINGER SEAL
SPARGING
CHAMBER

 
   
 
 
    

EXPANSION TANK

LOWER SLINGER SEAL

_ANTI-SWIRL VANES

   
  
    

CENTRIFUGE SLINGER
SEAL

  

SEAL PLATE

   

0.050 in.

_y_

  

SHRGUD

 

   
 

  
  
   
  
  
   
  
  
   

0.05Q in. CENTRIFUGE CENTRIFUGE
BAFFLE SHROUD SKIRT

CENTRIFUGE 0.198 in.

BLADE
FUEL PUMP
VOLUTE.

HUB
0.062 in.
FUEL PUMP

IMPELLER BLADE

SHROUD

FUEL PUMP
SLINGER SEAL

  
   

BYPASS
FEED HOLE

 
   

FIG. 1.2 ART FUEL PUMP WITH MODEL 32 IMPELLER

-11=-
ORNL-LR-DWG. 23922

He+ Xe'>°
*He

MINUS '
Heyl 2 psi ’
(l? B :LUPPER SLINGER SEAL

i

Ap~25psi } — ] |
A # _— = tf&a_‘v-mss PUMP

Ap~Tpsi - |
' "‘G)LOWER SLINGER SEAL
‘ =

i~ £ ary

L
ESa g8

 

 

 

 

 

 

 

 
 
 
   

 

 

 

 

 

CENTRIFUGE

 

| ~

=4 .
CENTRIFUGE

' SLINGER S
Ap~47 PSI (62.5 psi) SEAL FUEL PUMP

=

 

AP~6i psi

 

 

 

 

AT EXCHANGE

 

 

NOTE:

NUMBERS DENOTING

PRESSURES ARE
ONLY ILLUSTRATIVE

FIG. 1.3- SCHEMATIC ART FUEL CIRCUIT
-12-
iy
4

Circuit Descrlfi%lon -

‘The ART fuel circuit is most readily understood by referring to Figs. .
1.2 and 1.3. For clarity the circuit will be divided into zones and described
separately. | | | |
Zone 1. Primary Fuel Circuit _ '

| Circult wise ‘the main fuel loop consists: of two pumps operatlng in paral-
lel with a common discharge. Half of this C1rcuit is shown in Fig.:1. b,
fuel disch&rges‘intofthe core header and flowsrthrough the core and the heat
exchangerslto the pump suction. o | '

The core header.serves‘as a plenum and as{a,mixing chafiber-into which
processed fuel is discharged from the céntrifuge (shown dotted in Fig. 1.k4).
Since the centrifuge and the fuel pump have a common discharge, it is seen
that the fuel pumprdischarée pressure is regulated.by the'centrifuge discharge
pressure and that the fuel pump suction pressure is regulated by - system resis-
tance of the circult |

Zone 2. Bypass Fuel Clrcult

The bypass pump passes approxlmately 2% of the main fuel pump flow rate
into the sparging chamber where it is intimately mixed with helium to dllute
and remove from the fuel the Xenon—>° produced in the reactor core (see Fig.
1.5). The llquld-gas mlxture is jetted into the expan51on tank . where part of
the gas is separated from the fuel by allowing the mixture to form a relatively
quiet.pool from which the gas escapes to the off-gas system. The fuel in. the
expangion tank, with small amounts of entrained gas, feeds by gravity into the
centrifuge where the last of the gas is removed.and fed back to the sparging
chamber. The centrifuge diseharges into the discharge of “the fuel pump there-
by returning the fuel to the main circuit as well as mainfaining the discharge

pressure of the main pump.

-13-
. ABY N
R

ORNL-LR-DWG. 23923

 

 

 

 

  
 

 

 

 

 

 

|
|
/ |
I - l
| - - - 7
L | o
BY-PASS PUMP [ ] | _ .
T — — 1 P - , ‘
\i/ | L__J |—‘—::'_ —
SPARGER -
{ EXPANSION .
" CENTRIFUGE TANK
| (- —a—— '
| N A
| | -
lf |
|
| | CORE HEADER
x |
FUEL PUMP

 

HEAT EXCHANGER

 

 

 

 

FIG.1.4- ART PRIMARY FUEL CIRCUIT

“14-
 

 

BY-PASS PUMP.

ORNL.-L R-DWG. 23924

s

 

 

 

'_"-.” T

 

 

 

 

 

CENTRIFUGE -

 

N .
| FUEL PUMP

 

|~ HEAT EXGHANGER'

——  em—— e e— epm—

| _Expansion

T TANK

 

 

 

FIG. 1.5- ART BWF&SS FUEL CIRCUIT

-15-
LA A . g L .
P Ry Y
G

Zone 3. Dynamic Seals

There are three dynamic seals which are essentially centrlfugal pumps,
in the.bypass.fuel,circult,erranged to pump against the pressure of fuel
efitering the seel impellers fron fihat.is.normallytthe.discharge end of.the
blades. The interrelationship of . thesevseals is shown in Figs. 1.2 and 1. 3.

The. upper sllnger seal serves to prevent the fuel from leaking up the
punp shaft towards the bearlngs and to allow the helium to flow down the
pump shaft and into the sparging chamber., The lcwer\slinger seal serveq. ”
to prevent the fuel from leaking-from the sparging"chamfier'into thelcentri-v
fuge and to allow. the gas removed by the centrifuge to flow into the sparging
chamber. The .centrifuge slinger seal limlts the amount of recycle around the

top of the. centrifuge.

-16-’
Section II

DESIGN AND DEVELOPMENT OF -
ART FUEL PUMP AND XENON REMOVAL SYSTEM
 

i

’:j?:-:(_-l
v s (F240ts
.
o

  

o

 

$
YL
‘.Jy"-i {
i R
DESIGN AND DEVELOFMENT OF
ART FUEL PUMP AND XENON REMOVAL SYSTEM

A
> S,

Xenon Removal System DeSign

The X-R system described in this section is the finalized.ART design.
The system originally:enviSioned for application to the ART was one that
would be insenSitive to attitude. Before the development work on an atti-
tude-stable. system was completed,“it was. found that the. "north head" region
of the reactor would have to be revised because of a severe stress problem.
Another factor which contributed to the decision to radically modify the
original system was a 50% increase in the anticipated reactor fuel volume.

- The redesign of the 'north head" region and the increased fuel vélume
| necessitated a change in both the volume and shape of the fuel expanSion tank.
The shape of the tank was changed from cylindrical to ellipsiodial with the
pump barrels. passing through each end of the oval. The original requirement
for this system. to be insensitive to attitude was to be accomplished by spin-
‘ning the fuel volume in the expansion tank and us1ng the resulting centrifugal
force to stabilise'the'liquid-gas interface. 'Changing:the shape of the expan-
sion tank to.its final design made-.the problem of stabilizing t’he‘fuei‘gas |
interface much more complex and resulted in a decision'to eliminate the atti-
tude stability requirement for the sake of an expeditious deSign. The develop-
ment work completed on the original design is included in Section III of this
report to give a background for the final configuration and for. poss1ble future
application in an aircraft reactor with an attitude- stability requirement.

A Simple plastic model of the X-R system'was deSigned and built to test

its gas removal ability and.to calibrate the.bypass_flow through the centri-
. fuges. A cross'section,through this model is shown in Fig. 2.1.
Sibsimss
ey,

>

ORNL-LR.DVWG. 23925

FUEL PUMP
SHAFT

EXPANSION TANK
REGION

BY-PASS PUMP

CENTRIFUGE

FEED

SHAFT SHROUD

: BY-PASS FUEL HOLE
CENTRIFUGE .

FUEL PUMP DISCHARGE
REGION ' '

 

FUEL PUMP
SUCTION REGION

FIG. 2.1 FULL SCALE PLASTIC M%fi!{fflfi THE XENON REMOVAL SYSTEM

-20-
 

Xenon Removal System Flow

As -the - des1gn of the reactor progressed the estlmated fuel volume ine
creased. ThlS 1ncreased fuel volume necess1tated a larger expan31on tank and
a. larger bypass flow rate through the expanS1on tank and centrlfuge for the'
follOW1ng reasons. "'_' ' ,",“' T e '

1. The fuel expan51on tank volume 1ncreased in capac1ty in
""prOportlon to the total fuel system Volume T
‘é. To malntaln -the xenon concentratlon the same as for the
smaller fuel volume, the strlpplng rate 1ncreased 1n
'dlrect proportlon to the fuel volume. ' _
3. 'To keep the greater expan51on volume from overheatlng
| and . creatlng a serlous 21rcon1um snow problem, the
;quantlty of low temperature fuel rec1rculated through
the expansion tank 1ncreased 1n proportlon to the fuel
volume. ‘ ' '

‘The meterlng of the bypass flow taken from the maln fuel c1rcu1t had to
be done by properly sizing the hole up the center of the pump shaft and the
four radial holes formlng the bypass pump. A second factor affectlng the. by-
pass flow-was the dlfference between the pump suctlon pressure and the expan-
sion tank gas pressure. ThlS second factor could not be determlned 1n the
model as it is controlled by the comblned performance of the centrlfuge, fuel
pump and pump volute. “

.During the model tests the difference between the .pump suctlon pressure
and the expansion tank gas pressure was calculated from performance data ob-
tained from separate tests of the fuel pump and centrlfuges.

From Fig. 2.1 it can be seen that there are. two pr1nC1ple volumes to the
model. The volume between the two upper plates 31mulates the expans1on tank
and the volume between the two lower plates represents the high pressure
region at the dlscharge of . the fuel pump in the maln fuel c1rcu1t At the
lower end of the shaft is a bearlng and seal arrangement to separate the |
large hlgh pressure volume from the hole in the center. of the shaft. The
reason for this arrangement vas to control the Dbressure in the small volume '
at the lower end of the shaft at a value near that expected in the suctlon

region of the main fuel pump .

 
 

To calibrate the bypass flow fhrough the model a line containing a flow
meter was confiected.befween the high pressure discharge region and the low.
pressure region at the lower end of the shaft. A valve in the line was used
to control the flow until the pressure at the entrance to the shaft matched
that exbected at the suction of the main fuel pump. If the flow up the shaft
was greater than that desired in the X-R system, then the bypass pump holes
were reduced, and conversely if the flow was too small the holes were opened.
The reasofis for metering the flow with the bypass pump holes were that . changes .
in ths holes size were simple to make and also were highly effeqtive as the
pressure drop in this region is very large. }

The bypass flow rate which fiaS'measured by the flow meter was that bleed
from the discharge region to the'region below the seal while the measurement
desired was that of the flow up the shaft. In the actual test the leakage
around the seal was large (approximately 2 to 4 gpm), and amount to 15 to 30%
of the total bypass rate. As this leakage could only be estimated, the first .
calibration could,easily be in error by.1l0 to 15%. This bypass rate was later
determined.muchfmore accurately in the single pump Inconel hot loop test rig
to be described later.

Xenon Removal System Gas Removal

| The 1ncreased bypass flow necessitated by the 1ncreased reactor fuel
volume 1ncreased the velocity of the fuel passing out of the centrifuge to
such an extent that gas bubbles were carried into the main system. This in-
gassing was eliminated by increasing the flGW-area.available to the fuel.
This increase in flow area was aécomplished by increasing the number of holes
in the outer wall of centrlfuge from 8 to 33. |

A skirt connected to the lower end of the shaft shroud was added to the

system after it was originally bullt to prevent the incoming liquid from
falling to the bottom of the centrifuge cup and being thrown directly out-
ward toward‘ths exit holes\‘ Without the skirt the liquid which is thrown out-
ward carries entrained gas bubbles into the high velocity region of the exlt
holes where the smaller bubbles Wlll then be carrled on into the maln fuel

circuit.

 

 
 

Two otheruitems which wereaadded,to the modelgwere centrifuge'baffle
and blades. . As the direction and magnitude of the flow in the centrifuge
cups are not known, the point at which ‘the smallest gas bubbles are screened
out in the cup. is not known. However,-it is logical that any- direct passage
through the cup would enable the liquid to. carry more gas through the cup |
than if the veloc1ty'were more evenly distributed over the full height of the
cup. . The baffle wag added to prevent the incoming fluid from having a direct
passage to the exit holes.

- The centrifuge blades were added to decrease the slippage between the
incoming fluid to the_centrifuge cup and:the cup. Prevention of this slippage
improves the system in two ways. The effectiveness,of'the~centrifuge_depends
primarily upon the speed of. rotation of the fluid. It is obvious then that
any slippage between the fluid and the cup decreases the effectiveness of
the centrifuge. . I - _ L

.The_second»reason.for reducing this:slippage stems from_thelnecessity'
of having to balance the discharge pressure of the centrifuge against the .
pressure rise across the centrifuge slinger'seal " _The discharge pressure of
the centrifuge must be maintalned above that of the centrifuge slinger dis-
charge to prevent the fluid from bypassing the centrifuge and carrying gas
1nto the maln system., With the blades installed it is p0351ble to predict

the centrifuge discharge pressure very closely from the equation

H = _H;E
= 5%
where :
H = Head, ft 1lb per 1b
U = Tangential velocity, £t per sec A
g = gravitational constant 32,2 ft per sec2.1’

The presence of slippage in the cup'introduces an unknown coefficient to
this equation and also makes the head delivered by the centrifuge more depend-
ent on the flow through the centrifuge. .

" There is one advantage to having the sllppage in the cup. During'test'
with the model it became apparent that the’ discharge pressure of the centri-
fuge 1s a function of the liquid level in the expansion tank. 'As the liquid

 
 

level 1n the expans1on tank decreases the centrifuge pressure decreases, re-
sulting in a. like decrease in. pressure at the pump suction region° This
reduced pressure at the pump suction regults in a smaller flow up the pump
shaft and through the centrifuge. 'The decreased flow’ there causes a reduction
in the slippage’ which tends to increase the centrifuge discharge pressure. This
stabilizing ‘effect on the pressures ‘was not advantageous enough to overcome the
disadvantages stated above, '

The opening between the sparging chamber and the expansion tank was made
as large as possible° During tests made with this opening restrictive 1t was
found that the fluid was backed into the sparging chamber. When the fluid
level moved radially inmward from the tip of the lower slinger seal, it re-
stricted the gas flOW1ng upward around the shaft from the centrifuge to the -
expanS1on tank. With this normal path restricted, the gas was forced to
leave the centrifuge through the same opening 'ags that used by the incoming
fluid]which-reduced”the effectiveness of the centrifuge for larger'bjpass flovs.

ART Fuel Pump Development

‘The operability of a. face-type gas-seal sump pump pumping liquid metals
was experimentally demonstrated in the summer of 1951(5)In the fall of 1953,
G. F. Wislicenus designed an impeller (Fig. 2. 2) to produce the head and flow
reguired of the ART reactor using Flinek as the fuel. A volute was designed
within the space limitation of the reactor for this impeller in late 195k, ‘
Throughout this paper sections of the pump will be located by numerical sub-
scripts° The'position of these.sections and their numbers are shown in Fig.
2.2. The subscript (o) designates the "eye" of the passage just before the
impeller; (1) the impeller vane outermost inlet edge, and (2) the impeller
vane outlet edge or impeller rim.',A full scale model was built and water
performance test data were obtained in early_l955.' The model was installed
in a test”loop built ofu6 in. pipe with head and flow'measuring instrumentaq
tion and a throttling valve, The pump was driven by 1l5=hp variableuspeed
d=-c motor.. The first experiment was: performed by using a pumpasuction con- .
figuration that 51mulated the’ ART reactor design° The entrance region of the

- P
| 'g

 

2h-
 

, :QRNL—m- . 23924

 

 

 

 

 

 

SHROUD

 

 

 

m 1. '  ‘

 

 

 

 

BLADE _
SLINGER
SEAL

.

 

 

l’///////

 

 

FIG. 2.2- BRASS ART

 

 

 

oA P N G T RN
AN R
QLR RIS
RN
FUET™PUMP IMPELLER
- .

   

%é:fi$
pump was obscured by a flat plate located 1 in. below and parallel -to the

pump suction. A second test was performed with the plate removed and replaced
by an 8 in. pipe comnected directly to the pump suction. The_results of these
two experiments are shown in Figs. 2.3 and 2.4, The pump performance as indi-
cated by these two experiments was not affected by the simulated reactor en-
trance conditions. Since the flat plate was very. difficult to install, the

8 in. pipe suction was used for all subsequent testing.

The test conditions and results.from a series of experiments performed in
the test loop are given‘in Table 2.1. The'performance data from these experi-
ments are plotted in Figs. 2.3 through 2. 13h The performance data given in
Fig. 2.13 are representative of the best operation obtained during this series
of tests. The pump efficiencies are not considered accurate on an absolute
basis,'because the motor was not calibrated and the motor efficiencies were
obtained from the manufacturer's computed data. It isvestimated.that the
efficiency of-the pumps, exclusive of seal and beariné losses, is approximately
75% at the design point.

The impeller blade tip angle, 62, was increased to 26.5 deg from 22 deg in
order to produce the reactor design head and flow and a pump speed,approaching
‘the design speed; the results of this change are given in Fig. 2.5. Data ob-
tained by varying the impeller radisl clearance from 0.010 in, to_0.0hO“in.
indicated a'loss in pump performance of'approximately 8% at. low-flow hithhead
conditions and of approximately M% near the design point. These data are given
in Fig. 2 T

' A caVitation-like noise persisted throughout Experiments 1 through 12 at
- flows above 400 gpm and speeds in excess of 2000 rpm. The intensity of this
disturbance increased with increased flow or speed above the threshold values.
It was suspected that the noise'mightcbé;thééresultsfoffipoorlfluidflguidancefiat
the leading edges of the impeller uanes, An attempt was made to improve this
condition by changing the entrance angle of the blade to match that of the
fluid by machining the leading edges so that they lay on the surface of a cone
wnose apexlwent through the shaft centerline. The performance data from these

-26-

 
 

changes are given in Figs. 2.6 and 2.8. Modificatipns were;elso.made.to,the,
inlet radius of the suction 'eye and.to fihe.impeller nut to improve the entrance
conditions. These changes had no noticeable.effect upon the eheracter of the .
ndise;;-A new:impeller\blede,was designed to-give better entrance conditions
with no’ noticeable improvement.  The'performance data,of,this‘impeller blade. is
given in Fig. 2,10, | ,

It was determined that the noise was not prlmarlly due to cav1tat10n at the
blade leading edge but rather to a local condition that existed at the tongue,
of the volute. This was established by the use of-a'qarbon microphone probe..
A circuit was devised so that the microphone signal could be observed on a.
cathode«ray oscilloscope._‘With this arrangement, the region of maximum noise
was located near the volute tongue. The volute tongue was cut back to allow
more area for the. flow. This modification gave a decrease. in performance with
no improvement in the noise. 'The‘performance data for this modification is
given in Fig. 2.9. | | ‘

The need for addltlonal flow area in the volute is evidenced by the hy-
draulic unbalance of the-impellerfat the design point and .the occurrence of
the design speed maximum efficiency at approximafely 20% less than the design
- flow. A structural change in the design of the reactor north head region to
improve the stress condltions allowed the pump volute flow area to be increased;
however, this 1ncrease only compensated'for»the 1ncrease'1n-the fuel flow rate.
requlred for improved heat exchanger performance. To - obtain the additional
‘required flow area, the pump volute was redesigned W1thout diffusion cones.

Results of the water performance tests as glven in Fig. 2.11 with the re-
designed volute indicated excessive leakage losses between the impeller dis-
charge and the impeller suction. This is indicated by the sag in the perform-
ance curve at high-head low-flow conditions.~'The-éhroud-radial seal slingers
were extended to the impeller full diameter to increage their developed back
pressure. The resfilts of this modification are given in Fig. 2.12.' To further
reduce this leakage, the impeller'axiai clearance was reduced from 0.102 in.

to 0.053 in. The results of thisvfiodification are given in Fig. 2.13.

-27-
R
R

   

The différence in the pump total discharge head and.the average impeller
blade static discharge head as’a function of pump speed-and flow, as deter-
‘mined from Experiment No. 15, is shown in Fig. 2.1k. The data shown in -
Fig. 2.14 allows the -calculation of -the:volute efficiency (k). This effici-
ency, k, (Fig. 2.15) is defined as the fraction of the absolute impeller dis-
charge velocity head which.is converted to static head at the pump discharge..

It may be'séen from‘Fig,'2;13 that the design point lies in the region of
maximuh efficiency and‘that:the condition of.hYdraulid force balance on the
impeller occurs near the design point:'.Thg-nOise»present with the previous"
volute design was completely eliminated.  Based on the fierformance of the pump
indicated by Experiment:No;.l5, it was decided that‘thé redesigned volute and
the ofiginal impeller with a B, of 26.5 deg would be used in the reactor.

 
_68...

TABLE 2.1

CONDITIONS AND RESULTS OF WATER PERFORMANCE TESTS OF ART FUEL PUMP

 

Experlment

|Reactor Design

 

 

Impeller Impeller:: [Point at Time Suction . S
Number ' Design¥* Design Point of Tegt** Conditions Remarks and Results
1 Five vanes, Y70 gpm 606 gpm | Suction box, simu- Impeller design point met at
Blade Tip Angle, b5 £t head 50 ft head|-lating reactor; 1/2 in.| approx. 2650 rpm.- Reactor
(B ), 22 deg 2500 rpm . ~2300 rpm radius on suction eye design met at approx. 3050
2 ' e rpm. Pump very noisy.
2 Same as 1 Same as 1 Same as 1 | Straight 8-in. pipe No appreciable change in -
S L ' with four antiswirl performance as compared to
vanes; 1/2 in. radius Exp. 1.
on suction eye .
3 Five vanes, 606 gpm Same as"1l | Same as 2 Reactor design p01nt ‘met at | :
6235 26;5 deg .50 £t head ' approx. 2800 rpm; approx. a
S - 2850 rpm . 10% increase in efficiency.
o with respect to Exps- 1 with
| the ‘peak efficiency shifted
| towards the hlgher flows
Pump very noisy '
b Five vanés, Same as 3 Same as 1 | Same as 2 -Reactor design met at
B, - 22 deg : | | approx. 3100 rpm. Pump
leading edges wvery noisy. -
jcut back on an ' R
80 degcone
, angle L 1. _ L L ,
5 lseme as 3 Same as 3 | 'Same as 1 | Same as 2 Impeller radial clearance

 

 

 

 

 

was increased to 0.025 in.

from 0.010 in. No appreci-
able change in pump. perform~ {
ance., Pump very.noisy -

 

 

 
‘Table 2.1.(cohtinued)

 

 

Reactor Design

Suction

 

. -0t~

 

 

 

 

Experiment Impeller - Impeller Point at Time P
Number Design¥* Design Point of Tesgt** Conditions Remarks and Results
6 Same as 3 Seme as 3 Seme as 1 | Same as 2 Impeller radial clearance
- . . T was increased to 0.0L0 in.
~from 0.025 in. No appreci-
able change in pump perform-
ance at design point; 8% loss
in head at flows below 300
gpm. Pump_very noisy.-
7 |8ix Vanes, Same as 3 - 620 gpm Same as 2 | Impeller design point (Exp.
B, - 22 deg ’ 35 £t head | ' 1) met at approx. 2600 rpm.
- |1€ading edges cut A~ 2800 rpm Reactor design point (Exp. 1)
~ |back on a 56 deg ' ‘met at approx. 3050 rpm. Re-
cone angle ! actor design point (Exp. ')
o met approx. 2850 rpm. Pump
very noisy.
8 Same as T Same as 3 Same as T Same as 2 except Simiiar:to Expts. 1 and 7.
‘ ' ' ' ' ‘ 1/2 in. radius - S '
increased to 1 in.
on suction eye.
9 Same as 3  Same as 3. Same as 7 |Same as 8 Volute tongue cut back
" ' ’ 1 ,3/8 in. Performance simi-
lar to Exp. 3.
10 - |Same as 3 Same as 3| - Same as 7 |Same as 8 Volute tongue cut back

3/8 in. modified impeller
nut; performance similar.

to Exp. 3.

 

 

 

 

 

B h =’"i °
"TE"

Table 2.1 (continued)

 

 

Number

Experiment| ~

Tipeller
Design* =

-Impeller""

Reactor Design

Point at Tlme
of - TestX¥k:-

Suctlon

“‘“Condltlonsr

 

11

12

. 13

14

157

 

Same as 3

Five vanes;¥*

e2 - 26.5 deg

Same as 3

Same as 3

. |Same as 3

 

 

Design Point

Same as 3

620 gom
25 ft head
2800 rpm

Same as 3 .

Same as 3

Same as 3

 

Same as T

‘Same as 7

645 gpm
37 £t head

A 2800 rpm

Same as 13

Same as 13 -

 

. ‘Same as 8

- Same as 8

. Same as 8

- Same as 8

Same as 8

flow. and speed

.ameter,

 

Remarks and Results

Volute tongfie cut back 30
deg; performance showed a

' 'deCrease'in_head and flow.

No change in noise.

Volute tongue back 30 deg;

j:Performance similar to Exp.
11

New volute with increased
flow area. Impeller hy-
draulically balanced near

reactor design point. Head

is decreased at constant
. Pump very
quiet.

Extended sllnger on lower
shroud to full 1mpelle* di-
Increased hea® at
constant flow and speed.

'-.Eff1c1ency peak at de51gn
- p01nt

Impeller axial shroud clear-
ance reduced to 0.053 in.
from 0.103 in. Increased head
at constant flow and speed. |
Increased efficiency peaks at
design point.

 

o :‘E{F{: .

b
Sl

'

¥  The impeller blades were redesigned to provide entrance angles more nearly aligned with the fuel.
*% Due to the time lag between the impeller design and the 1mpeller testlng, a different impelleii ée51gn

point and reactor design point resulted.

 
- ze..

100

90

Head (.i:'ra__lh)

' omw‘ao_mr

 

WAL E Dol

 

wew Pump Efficiency (%) .-

 

 

e Pressure Balance Line

Reactor Deslgn Point

 

 

 

 

 

A
2700 mpn 300¢ rpm Pump Speed .

 

 

 

 

 

 

 

 

600 700 800

 

 

 

 

 

 

CAPACITY (gpm)
FIG. 2.3 ART FUEL PUMP PERFORMANCE CHARACTERISTICS EXPERIMENT NO, 1 - TABLE 2.1
ORNL-LR-DWG, 23924

 

   

 

  

 

 

 

 

 

 

 

 

 

 

  

 

 

 

 

 

 

 

 

 

 

 

£
_ : olutel Pressure Ballance Line
/ \ <--- Pump Efficiency| (%)
y ) 75
/’
- 70
" Reactor 7
Design
Point / \
P 4
/0 \
L/ ' 3300 rpm Pump Speed
1
[
Ciq." .
. L. 60
i ”~
A
NN\
3000 rpm
2700 rpm
© 2250 rpm '
2000 rpm
00 rpm
100 - 200 300 L0oo 500 600 700 800 900 1000
’ ’ CAPACITY (gpm) ‘ '
FIG, 2., ART FUEL PUMP PERFORMANCE CHARACTERISTICS EXPERIMENT NO.-A:;?‘x"Il“ifiEllfiliE 2.1

 

 
ORNL-LR-DWG. 23929

 

 

  
  
  

_ I .
Pressure Balance Line ---- Pump Efficiency (%)

 

 

 

 

 

3000 rpm Pump Speed

 

\ 2700 rpm

pm

 

1500 rpm Tm

L1000 Tpm
400 500

 

 

 

 

 

 

 

 

 

 

 

 

 

500 700 800 900 1000
CAPACITY (gpm) '

' *T.f'n':..‘s’fi
FIG, 2,5 ART FUEL PUMP PERFORMANCE O ARACTERISTICS EXPERIMENT NO. 3 -~ TABLE 2,1 SRR
-ge-

ORMNL-LR-DWG. 239!'

 

90.

80

  
 

Head (£2.1D)

 

=== Pump Efficiency (%)

 

 

 

Plressure Balahce Line

_—Reactor Design Point

 

 

 

70
60

 

504b
’-
3000 rpm Pump Speed

 

 

 

 

 

 

 

 

 

 

 

 

 

 

700 rpm
g 500 rpn
30 LOO 500 600 T00 800 200 1000
CAPACITY (gpm)

 

0 100 200

FIG, 2,6 ART FUEL PUMP PERFORMANCE CHARACTERISIICS EXPERIMENT NO, 4 - TABjE:Zulosswes
90

 

 

 

 

 

ft 1b

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

AR
ORNL-LR-DWG. 23931
80
kl\‘--\ ] .
70 = ::_: -
N Experiment No. 3
. < :
Hxperiment Ng. 5 <+~
- \‘\
60 \\
I NN
Experiment No.b\
50 5 \. F,Reactor Degign Point

F \\
® ko - \
O
=

30 > z
20 \
10
T00 200 300 LG 500 %00 700 800 960 1000
CAPACITY (gpm . )
S o | S
FIG, 2,7 ART FUEL PUMP PERFORMANCE CHARACTERISTICS AT 2700 rpm EXPERIMENT NOS. 3, 5, AND 6 - TABLE 2,1 o
-Le-

ORNL-L R-DWG. 23952‘

 

90

 

80

=== Pump E

fficiency (%

 

70

  

€ Pressure Balance Line

 

 

Head (EI%E)

 

tor Design Ppint

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

3000 rpm Pump Spepd
pm
® 1000 pm
300 Loo 500 600 700 800 900 1000
CAPACITY (gpm)

 

FIG. 2.8

ART FUEL PUMP PERFORMANCE CHARAGCTERISTICS EXPERIMENT NO, 7 - TABLE 2.1 RRNRRERY

 
 

 

. QRNL-LR-D¥G, 23933
8

---- Pump Efficiency (%)
7

 

 
  

 

Velute Pressure Balance Line

 

-88-

.Head (EEI%E)

 

actor Design

/

N\

Point

 

10

\ 2700

DO rpm

rpm Pump Spe

ad

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

0 100 200 300 400 500 600 700 800 900 1000
| CAPACITY (gpm) _
FIG. 2.9 ART FUEL PUMP PERFORMANCE CHARACTERISTICS EXPERIMENT NO. 11 - TABLE 2,1  -uan’E%
 

90

 

80

---- Pump Efficiency (%)

 

70

 

Head (%}E)

10

- Volute| Pressure Balance Line

 

 

/7 ‘?\Kflemtor Design Point

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

200

 

 

 

7 ~ - )/ETJU rpmPump $peed
Ve A
s _ ~
e 7
< s 24p0
7 rpm
N ,/ o~
” ’/'
70>\ ~
~® 2000 rpm
6§
300 L0O 500 300 700 800

CAPACITY (gpm)
FIG, 2.10 ART FUEL PUMP PERFORMANCE CHARACTERISTICS EXPERTMENT NO; e~ TABLE 2,1
-ov-

Head (%F]'E)

70

50

/)

ORNL-LR-DWG. 23935

 

 

~~== Pump Ej

Priciency (%

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

40
r'\ rpm Pump Speed
30
20 P~
\ | . / 2D00 rpm

) ‘\7\. 150 o

.\ /<Volute Pressure Halance Line

1000 rpm
rpm
0 100 200 300- 400 500 600 700 800 900 - 1000
CAPACITY (gpm)

FIG. 2.11 ART FUEL PUMP PERFORMANCE CHARACTERISTICS EXPERIMENT NO. 13 - TABLE 2.1

s
Head (231%;12)‘

- 17-

80

70

50

L0

30

20

10

QO

a/‘\\ / /
\\ - T~
e

\

ORNL-L R-DWG. 239!!

 

 

"em== Pump Efficiency (%)

 

64

 

[
/

~.

65
/
[

 

/

|

I

\
\

  

 

7

Z

N

L

 

rpm Pump Spepd

 

>

~
~

/
s
-

!
|

\

\

.S
7

 

100 rpm

 

 

 

 

 

1]

e
>+

 

~e .~
{

Volute

!
[
\
N\
\
g
\J\\\t
/7

 

® 1500

 

 

 

 

 

 

FIG. 2.12° ART FUEL PUMP PERFORMANCE Ci ARACTERISTICS EXPERIMENT NO. 1l - TABLE 2.1

" CAPACITY (gpm)

00 rpm
=% 55 rpm
100 200 300 400 500 700

 

 

 

1000
-z~

Head (£, - 1%

8o

70

50

30

20

10

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

IRy

 

 

CAPACITY (gpm)

 

 

\\\\
5 ~~w= Pump Efficiency (%)
| 7 %
‘\ II I’ \/i _
o /
\ / .’ ! ~..®
TN / / 7'\ . '
\,\'\ / . / / \
/ 1/ .
~ . [ 7’\.\ / / L. 80
< I’ / T~ /// 81\\\.-
\ / / / ' TN
\L\l l/ / / } ll 75
- es 70
T~ o/ I | Po%/ 7 N *
| , / 3000 rpm Pump $peed
| [/ / || /
\L\! // / > N // /
i O 7 D2
NI { i /| N
| , \ \ \ \ / // P /™ 2700 rpm
“\ \‘ SN \ / N\ - 7
\K \\-.-:' \/./ a4 2400| Tpm
‘ — —
_ \ ™~ —— {__.—-
2000 rpm
~ oluteé Pressure Balance‘ L;Ee
> — U I'pm
100 20 50 5 70 8 90 10

 

. FIG. 2,13  ART FUEL PUMP PERFORMANCE CHARACTERISTICS EXPERIMENT NO, 15 - TABLE 2,1 pumussii™®
It
Pump Total Discharge Pressure Minus Impeller Blade Static Discharge Pressure (—-.I#)

-
o

Y

ORNL-LR-DWG. 23938

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

100

m8 uoo _.r8 moo 80 ...8 moo woo
: CAPACITY (gpm)

FIG. 2,14 .oowwmo,fiozm APPLIED TO THE ART FUEL PUMP TOTAL DISCHARGE
PRESSURE FOR OBTAINING THE CENTRIFUGE STATIC DISCHARGE PRESSURE

-43-
 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

FIG., 2.15

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Els
g
8
qm..% / &
= .
8 E
- E
fl‘ / o
| | g &
S \ o o,
m \\ o um.
° TN B g g"
\ . o H B
/ M (ud3) Lousto}srd
: N0 wrgrixey 38 moTj dumg
\
\: o
\/// / M
\ "~ AN
N &
L / S
\. / / A,%.
N N &
\/ ‘ / g
. B
\ %// NS
.// o / . m
/\/ N m
/ _
/ B
\;///// _ N .mm
S g
,/;/////////, <
/ m.
/ o
E
,//, £
S
B
"
3 2 '3 R ‘S A & Q

(g ur £ouetoryyg eynfop dumd Temg) ¥

-l

ART FUEL PUMP VOLUTE PERFORMANCE CHARACTERISTICS
Combined.ART Fuel Pump and Xenon Removal System Development B

The £inal development work on the X-R system was’ done with a gingle pump
Inconel hot. 100p test rig (see Fig. 2.16). . This’ final work was originally
planned to ‘be -done- W1th a twin pump test loop very similar to the ‘north head
region of the ART. Construction difficulties with the twin pump loop delayed
its completion and the develoPment work was shifted to the single: pump loop.“
This looP combined the’ Latest X-R system tested with the plastic model and the
final pump design determined from .the pump. water tests.“ All development work

done With the Single pump rig used water as the working fluid.

System Pressure Fluctuations

At high liquid levels in the expansion tank (above 3 in.) it was found
that the pressure throughout the main: system was subject to rather Large and
irregular fluctuations. These fluctuations decreased as the expanSion tank -
level was decreased from 3 in. to about 3/h in. where the fluctuations again’
began to increase.‘ This increase at low levels was expected.' The fluctuations
at low. expanSion tank levels were observed in the plastic model and attributed
to a fluctuating liquid-gas interface in the centrifuge.. The flow into the
centrifuge is by graVity from the expansion tank and any decrease in the level
in the expanSion tank will decrease The flow to the centrifuge._ As the main '
fuel - circuit is filled with an essentially non-compressable fluid it is necces- |
sary’ that the flow to and from this system be equal, thus, . any. change in the
flow to the centrifuge will result in a change in the pressure of the main
system and therefore a change in the flow from the main. system back to the
expanSion tank° It can then be.seen_that as the liguid level in the expansion
tank decreases, the flow‘to and'through'the'Centrifuge-decreases‘and the main
system pressure‘decreases.- As shown. previously the main system pressure is
controlled by the discharge pressure of the centrifuge and any decrease in main
system pressure results from a decrease in the centrifuge discharge pressure.:
In the case_beingidiscussed the centrifuge-is running at constant speed and has

- a Tixed outside diameter. Therefore any decrease in discharge pressure must

Lo : .

-
 

 

s B
UNCLASSIFIED
PHOTG 27742

 

 

 

7 o l; = B! -
3 . . " - o
i g i T il = 3 aL o
: il r = A - — - e 2 -
" - ia b = - i . -
- - §"F ~ i R . . V }
¥ g b Ly . - - 2 3 - 2 =t
- L o -3 e - = i
L) - I 1 5 . = B
[ i Sy S -
A : 5 : o :
i g e - T " .

FIG. 2.16- COMBINED XENON REMOVAL AND FUEL PUMP HOT TEST
 

result from a decrease in the suction pressure of the centrifuge or increase
in the effective inSide diameter. .In the case of the low expanSion tank level
the effective inSide diameter of the liquid—gas interface moves radially out-
ward decreaSing the pressure rise through the centrifuge._ This change in liquid-
gas interface was observed in some of the earlier tests with plastic models._ For
very low liquid levels this condition becomes. unstable with the centrifuge alter-
nately trying to becomemfirst empty and then as-the,outflow decreases to become'
full.‘ ;.. : - : - o . A S :i . .
The cause of ‘the fluctuations at high expanSion tank liquid level was not
nearly 50 obvious. | During the development work each of the various flow paths
were investigated and the difficulty Mas found to be caused by a fluctuation_
in the pressure at the inSide diameter of the centrifuge The cause of this
fluctuation was traced to the lower slinger seal . At very high liquid levels"‘
the centrifuge will run in a flooded condition and it is: probable that slugs.'
of liquid as well as gas Wlll flow upward between ‘the shroud and shaft to the
inlet of the lower slinger seal ‘This gas and liquid flow through the slinger
impeller would cause it to .surge with the surges being transmitted back to the
centrifuge as irregular periods of high and low pressure. These surges would
then be transmitted through the centrifuge to the main system. )

~To alleViate these surges the vanes were removed from the lower slinger
seal By remov1ng these ‘vanes the head of the lower slinger seal is reduced

and any flow surges in this region Wlll be dampened.

System Pressure Level

A difficulty encountered in the. early testing of the combined system was
an excessive system pressure level which results in 8 high stress condition |
being imposed on the plate. which separates the expanSion tank and ‘the fuél
pump discharge.- Since this pressure level is fixed by the centrifuge pressure
‘rise, attempts were made to reduce the centrifuge pressure rise by two different
approaches._ First, the inside diameter of the centrifuge cup was increased and
- second, the exit-area_fromflthe centrifuge was decreased., There were several ad-

vantages to the first'method; The change in the'existing centrifuges were very

“h7-
 

simple to-make; and also it was possiblento calculate ver& closely the new.in-e
side diameter of the centrifuge cup to get the desired change in the main system
pressure level. Another advantage is that the Cross sectional area of the en-.
trance to- the cup is 1ncreased and will 1ead to a more stable system pressure
at lower expan51on tank liquid levels.‘ One disadvantage is that the radial
depth of the 1iqu1d in, the centrifuge 1s decreased. This thinner 1ayer may
enable gas bubbles to pass into the main system under operating conditions
where they would have been stopped with the thicker layer. In the case of the
small exit=area from the'centrifuge the major disadvantage was the difficulty
of -making the change. _Any decrease in the exit‘area byudecreasing the nunmber
of exit holes or the diameter of these holes'will increase the-velocity through
the cup and thus reduce the degassing ability of the centrifuges.' To overcome
this difficulty the holes would have to be tapered with the hole area at the
inner wall of.the cup remaining the same and with the ares at the exit in the
outer diameter being reduced. Not only'would the construction of these holes
be difficult but their size would have to be determined by trial and error.
One advantage of this system is that the liquid level or liquid gas_interfacef
. in the centrifuge‘does’not change. .The .second and’ more important advantage
was never confirmed by test. This second advantage would show up only in the
case where one of-the two reactor pumps_had stopped while the other was running.
‘For this-one.pump out operation it is very desirable that the centrifuge have a
head flow characteristic such that the head decreases very rapidly with any in-
crease in through flow. All of'theuclearances around the centrifuges are large
and the leakage is controlled by a d&namic seal. When one pump stops the seal-
ig also lost so.that the. leakage becomes very large unless the pressure drop
across this clearance is decreased.. In. the case where the exit area is small,r
the‘pressurezdrop,across the_exit holes of thepoperatinglpump isclarge and will
increase rapidly with the increased flow resulting from 1eakage around the in< .
operative pump. In the case of the enlarged inside centrifuge diameter the ‘
head of the centrifuge must also drop off rapidly as’ the through flow 1ncrease,

but in this case, the decrease in head must be accomplished by the centrifuge,

 

 

 
the centrifuge becoming partially empty and'increasing the effective inside

 

diameter even more.  This increase in the diameter of the gas-liquid inter-

face leads to a breakdown in the effectiveness of the centrifuge.

Bypass Flow Determination and Control

One of the most important purposes of the combined system tests was to
determine and control the-bypass flow. As. previously noted the plastic model
served only. to give an approximate rate because of leakage of . the lower seal.

In the combined single pump--system the bypass flow passage was 1dentical to

that for the ART. A disadvantage of this‘system is the lack of a method of
directly measuring this bypass flow. The procedure used was to plug the hole

in the center of the pump shaft through which the bypass flow normally passes
and substitute an external line between the pump suction region and the expan-
sion tank. This external line contained a valve and flow meter to regulate

and measure the flow. During these tests runs the pump speed and main system
flow were held constant'while the flow through the external line was varied.

The expan31on tank gas pressure and the pump suction pressure were measured.
This" method yields a -pseudo head-flow curve for the centrifuge and is used as

a calibration for test runs where the actual flow up the pump shaft cannot,be
determined -directly. The use of the pump suction”pressure rather than the
centrifuge discharge pressure for calibration purposes 1is arbritrary. The use
of the suction pressure does simplify the immediate reduction of the data, and

- more important the pressure differences between the pump suction and the ex-
pansion tank is one of the factors controlling}thehbypass flow and is of more
interest than the pressure difference.between the centrifuge:discharge and the
expansion tank. -After this pseudo head-flow curve was established for a series
of operating points, the plug was removed from the pump shaft and the external
line closed. The operating point used in the above calibration runs were then
repeated and again the pump suction pressure and the expansion tank gas pressure
were determined. The data for the runs with the pump shaft plugged were plotted
as shown in Fig. 2.17.. The'difference between the pump suction pressure and the
‘expan51on tank gas pressure for the pump shaft open runs were then checked against
the identical calibration run and the bypass flow up the shaft was read from the
curve,

-.)4-9"'
ORNL-L R-DWG. 23940

 

Bypasrs Flow through the Expansion Tank - Bhaft Pluggefl .
' (LO0% = 17.6 gpm) ,

-og-

FUEL PUMP SUCTION PRESSUNLE MINUS EXPANSION TANK GAS PRESSURE

1z

10

~2

 

 

. Expansion [Tank Liquid

Level - Shaf

L Open

 

 

/

 

100%

[ Pump Speed)

 

5

A

RN Ty S A, fll

 

~\

A .

O
~N

« | 2400 rpm

 

 

&

Y
[y

\

 

 

)

LR

b0 rpm

 

 

 

 

 

 

 

 

 

 

 

 

 

 

3300 rpr

 

 

 

 

 

. FIG, 2,17 ART FUEL PUMP BYPASS FLOW CHARACTERISTICS WITH MOOEL 32 IMPELLER
AT A 3-1/2 IN, TEPTH OF LIQUID (WATER) IN THE EXPANSION TANK
 

The flow through the external line or up the pump shaft is not the total
flow through the centrifuge;"Part of the centrifuge flow recirculates over
the centrifuge slinger seal and back 1nto the cup. Thisfrecirculation flow"
is not ‘accurately known but s1nce it is not a function of the manner in which
- ‘the, flow gets ‘from the pump suction to the eXpansion tank, but instead is only
a function of the flow through the" centrifuge, it does not enter into the cali-

bration. =

Fuel Leakage Past Upper Slinger'Seal.

One item 1nherent in the X-R system design was the cause of much concern.
This was the poss1bility of fuel passing over the upper -slinger seal and being
forced or drawn up into the narrow annulus between the pump shaft and the bar-
rier shield. plug If any ‘fuel from the system were to. rise into this region
it could freeze because of the cooling of the lubricant circulating in this
region. 'The fuel on freezing becomes hard and could do damage to the shaft 5
and shield plug can ‘and could. pos31bly freeze the shaft to the stationary
shield plug can. This would necessitate a change of the pump rotary element.
If the reactor had been at power this task would be unpleasant.

- In the early test 1n the Inconel hot. loop, the upper slinger seal design
was the same ag used in. the plastic model. This design had twelve radial vanes
about 3/16 in. high for. the full radial distance. (see Fig. 2 18) In these
early tests 1t was found that the gas pressure in the expansion tank was. greater
than that in the oil catch basin region -of the rotary element. In many cases,
this. pressure difference was as large as three to four feet of. working fluid-
head. TFor the first de51gn this pressure difference was found to depend upon
the pump-speed,,the gas flow down the pump-shaft,annulus, and the 1iqu1d level
in the-expansionftankc“vThereaWas*a‘noticeable?"break“pointfi.in‘this.pressure
difference as the expansion tank liquid level passed 2 in. This is aiso~the~
“level at which the liquid. will uncover the ports between the sparging chamber
and the expans1on tank. As the liquid level was reduced from a. ‘high. level to

‘about 2 in., there was no significant change in the pressure difference. At

 
 

approximately 2 in. the pressure differerce dropped sharply and was then again
fairly. independent of the liquid level as it was reduced further. -The level
‘at which the break point occurred was not fixed but was-dependent on whether ..
the liquid level was ‘increasing or,deereasing. The.leéel.for'this break point
‘'was higher for the:case when the liquid level was.increasing-than»when the
level was decreasing. It was also noted even at the highest liquid levels:
that as the rafie-of.gas flow down from the eatch basin was increased the pres-
sure difference was decreased and for extremely high £low, several times the
normal'SOO-liters~per day, the pressure difference would reverse with. the.

. cateh basin gas pressure -being greater-than the gas'pressure in .the expansibn
tank. It was also found that this undesirable pressure difference increased
as the pump speed 1ncreased.» -

Undéer: normal operating conditlons & gas pressure reversal in this region
Wlll not cause any damage. The potential head of the slinger is greater than
the pressure difference 80 that no liquid will be drawn over the vanes. The
real danger of this condition is for;the‘case'where one or both pumps stdp
suddenly‘asxinwthewcase;of a power failure. If the slowing down rate of the
pump is so .rapid that the inflowing gas camnnot build up the pressure in the
lower pressure -region at the same rate the dynamic head of the seals is deQ
creasing due to.the decreasing speed, then the liquid will be drawn up into
- the annulus around the pump . shaft and probably require a replacement for the
'rotary assembly. ' '

All of the above .observations show that the upper slinger seal is.acting
as a gas pump_ The actual gas pressure buildup in the expan51on,tank is. far
greater than a pnmp'of-the size and type used should be able‘to'n;rmallj pro-
duce using gas as the working fluid. It appears that the pressure buildup is
due.to.thefreCireulatiOn of 1iqnid'through-and'over-the:_vaneso The upper
slinger seal acts' the same-as an-open‘face eentrifugal"bump’eperating under no
flow conditions. Due to the rather large axial clearance over“theztop of the
seal and recirculation is large,\ This recirculation appears as a flow outward

through the vanes and an inwardfflow through the clearance above the vanes.

 
-

The outward flow carries gas bubbles to the outer tip where it then escapes into
the sparging region and then into:the expansion tank._ As this process continues
the pressure difference Wlll increase and the length of Vane which is covered
by liquid will increase as-the liquid is drawn radially inward to produce a.
large head to offset this rising pressure difference. As this radial dimen31on
'of the 11qu1d increages, the centrifugal effect of the spinning 11qu1d begins
to screen or centrifuge the gas inward toward ‘The pump shaft. These counter-
acting effects thus 1imit the magnitude of the pressure difference the upper
‘;slinger seal can produce. This also explains the observed dependence of the
_ pressure difference on the 1iquid level, speed and gas flow. As the gas flow

is 1ncreased it will reach such a flow that the upper slinger seal cannot en-
train and pump -the gas as fast as 1t enters.. When this flow. is reached the
pressure difference will reverse.. In. the case of the effect of the liquid
level the observed hysteresis becomes clear. Before the upper slinger seal
can act as a gas pump it must become primed. When the liquid 1eve1 is in-
'creasing the upper slinger seal will "throw". the liquid avay from its tip |
until the level has reached a. point higher than the upper slinger seal or un-
til the 1iqu1d which obstructs the ex1t ports from the sparging chamber causes
the pressure to build up -in the region of the tip of the upper slinger seal.
When the expansion tank is filled to about 2-1/2 in., the liquid 1eve1 is
above the . upper slinger seal and 1t is primed when the pump is. started To
break this prime the 1eve1 must be reduced until the exit ports from the
sparging.chamber are open and the pressure:at_the outer_radius of the upper
slinger seal is reduced. This als0'accounts'for the sudden ghange in pressure
when.these points are‘reached."The ability'of the'upper slinger_seal to en-~
train and pump gas is obviously a function of the speed. |

In trying to overcome th1s problem, the basic ided was to reduce the

large recirculation over the upper slinger seal which with the large axial
clearance required is impOSSible to completely stop.‘ The first attempt was
»to replace the vaned upper slinger seal with one whose top surface was smooth
 The axial clearance was held the same as for the vaned upper slinger seal.

This reduced the unfavorable pressure difference but did not reverse it.

S
"u! .

-53-
 

Severa150ther methods were tried with some improvement but were not completely
Satisfactory. All of these changes were attempts to reduce the outward flow .
potential at the rotating gsurface. ' '

' The de51gnwwh1ch finally gave a pressure difference in-the desired direc-. .
‘tion, i.e., the catch basin’ pressure greater than the pressure in the expansion
tank, is shown in Fig. 2 19 The top surface. of the upper slinger seal was cut
back for a radial distance of 3/h in. This configuration gave a catch basin
pressure higher than that in the expansion tank by about 10 to 14 in. at design
'eondltlons and high llquid levels in' the expansion tank. This incréased clear-
ance reduces the.potential.head of the upper slinger seal, but as it has far
| more head capacity than required to protect the rotary element when operating
normally there is no deleterious effect in making this change. The outSide..
diameter .of the upper slinger seal could not be decreased to obtain the fabor-
able pressure'gradient.' It was found that a liquid seal must be maintained be-
Atweenfthe face of the upper slinger seal and thé¢ sparging chamber. With this .
Pavorable pressure gradient,.the»fuel.cannot enter the shaft annuluSJduring
the slowing down ‘of the pump.  Also once the pump has stopped, the gas flow : .
to the expan51on tank through the shaft annulus malntains the favorable pres-
sure gradient. ' ' ‘

Experimental Results :

' The test results of the final design configuration are shown in Figs.
2.20 through‘2,28. Figs. 2.20 through 2.23 give the-head flow characteristics
of the sYStem as a-funetion of speed. The value of the difference hetween the
pump suction?and expangion tank pressure is also shown. The small variatiens

in the head~flow curves frem'ene liquid level to another are probably not sig-
nificant and are within the*aceuracy~of the data. The values of the difference
‘between' the pump suction pressure (PS)'anduthe expahSion-tank-gaS'pressure'(PHe)

- are definitely a function of liguid level at lower levels, especially at 1 in.

and below for reasons given previously. -~ S S .

5o
 

. UNGL ASSIFIED
. ORNL-LR-DWG. 23941

3}“ 3“.‘ -
“ 76 xl_e V_ANES

 

 

 

 

 

 

 

 

| "FIG. 2.18
'UPPER SLINGER SEAL WITH RADIAL VANES

 

 

 

 

 

 

 

| ) FIG. 2.19 . o
UPPER SLINGER SEAL WITH AN AXIAL STEP

3

 
   
, —

The reason for the scatter in the data can be seen by examining Figs..
2.2h.through 2.27. -These figures show the pressure fluctuations throughdut
the -test loop as a function;of speed, flow, and liquid level. . While these

fluctuations are not large enough to cause any damage to the reactor, they
did makejanalysis.of the data difficult. The reason these relatively small
fluctuations_wére so important can be seen by studying Fig:.2717. If the
_ valueof-v-PS "PHe on the figure are shifted.by 0.2 ft, the error in.deter-

mining the actual bypass .flow at 645 gpm amounts to about .5%. At lower
liquid levels the fluctuations become larger and the experimental values .of
the bypass flow erratic. For this reason, the bypass flows for the lower
: liquidqlefels are calculated (see Appendix E). |

Calculated .and experimentally determined flow up the pump shaft or

that amount which bypasses the main fuel pump is shown in Fig. 2.28. This
is not the total reéirculation through the expansion- tank. In the test loop
in which the data was obtained there was é leakage .of approximately 0.6 gpm
from a point near the pump suction to the expansion tank. This leakage
should be added to the bypass flow to get the actual total flow th;ough the
expansion fiank{ In the réactor there is also a leakage very similar to the
one in the tést loop. This leakage is around the island where it paSses
through the floor of the expansion tank. The amount of this leakage is
estimated to be about 4 gpm or 2 gpm per pump and like the leakage in the
test loop will depend upon the operating pump speed and flow. ‘

' The total flow through the centrifuge cups is equal to the sum of thé
bypass flow, leakage and recirculation over the top éentrifuge'seal vanes.
The leakage over the vanes is large, about 6 to 8 gpm at design conditions.
This large leakége_does not effect the degassing ability of the cup. The
reason for this is that when the flow hasiincreased to the point where in-
gassing occurs the heé@ of the centrifuge has dropped to the point where it:
is less than the head of the vanes. Under this cpndition it is probable that
the centrifuge is not completely full and that the inside diameter of the
vanes receive gas only. The vanes then act as a slinger and there is no flow
through them in either direction, the location of the liquid-gas interface
being determined by the head required to match that of the centrifuge.-

-56-
- Ls—

(LL1b

HEAD

70}

50

Lo

20

ORNLM

 

 

 

 

 

 

N

 

b S 0L S

 

10—

 

 

 

 

 

 

 

 

 

 

 

 

 

160-rpm /“"/ 7 /f / / /.r\//x//////\x.{llBlLl }
1800;\\/‘://///'//\,.\<//{?/>;\///,>( 12
1,4“)0 rpm )./ /__\'\ /\./. >
300" o0 900_1;9‘:/-;'-—{.- -;..,___.; </ S _
e M-zoa | 300 — 130 500 %00 730 o - -

CAPACITY (gpm)

FIG., 2,20 ART FUEL PUMP PERFORMANCE CHARACTERISTICS WITH MODEL 32 IMPELLER

AT A 1/2_IN, DEPT OF LIQUID (WATER) IN THE EXPANSION TANK
- -8§-
ft 1b
b )

HEAD (

 

ORNL-“AI!

 

 

3300 rpm Pump Speed '

 

 

 

700 s %/ // A
Fuel Pump tton a;:ei:u—>//\//\<// / /7\//4/{ /

R
8/

500 rpm \/ [
5 /7
~X.

1200 1

/ 4..-.-
00
300 0 ron .~~~ ;
rpm___'_-.rép-—--.._ . . . .
100 200 300 1,00 500 600 700 800 900
CAPACITY (gpm)
ART FUEL PUMP. PERFORMANCE CHARACTERISTICS WITH MODEL 32 IMPELLER
AT A 1-IN, DEPTH OF LIQUID (WATER) IN THE EXPANSION TANK = - _
ST

 

 

 

l_l

 

 

 

 

 

— 1000

 

 

 

 

 

 

 

FIG, 2.21
| -65-

70

50

ft 1b
1b

HEAD (—— ) .

4O

30

:"20

19

<o

 

 

 

 

 

Fuel Pump Suct
Minus Expang
Pressure (f£4)

ion Pressure

ion' Tank Gas -

 

240

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

T -2 ____ A /
/ “.“"/‘-. ’
_ 300 rpn, | / '
T
200 1o ., 7
150 rpm zedl [T
100 200 300 110 500 400 700 800 900 1000
CAPECITY (gpm)

FIG, 2,22- PRT FUSL -PUMF PERFORMANCE CHARACTERISTICS WITH MODEL 3%2 THMPSLIER S
AT A 2-IN, DEPDI OF LIYUID (WATER) IN HiE EXPANSION TANK RS

1100
 

 

 

 

 

ft 1b
1b

AD (

: N ' ‘ - o
3000 rpm /, / / / / l\% 1 : |
, // 7 ./ 4\//// /// / /’ 7 ‘:'i-—F‘uel Puxfip Section Pressure MimL Expansion

 

-09-

TN
//%";(f"‘?é |

 

2100 rpm |4/

A
21,0 rpn: . //
| ///7%/{/ LY

 

L TR
LY

 

T200 rgm/.L_____- ’/

 

 

/ ~.

"-;--o

 

 

 

 

 

 

 

 

 

 

 

 

300 400 500 600 700 . 800 500 1000 1160
CAPACITY (gpm)

FIG, 2.23 . ART FUEL PUMP PERFORMANCE CHARACTERISTICS WITH MODEL 32 IMPELLER :
AT A 3-IN, DEPTH OF LIQUID (WATER) IN THE EXPANSION TANK &Egfifi;

 
ORNL-LR-DWG. !!'l.

 

 

 

 

)

3040 rpm Pump 3 . .
Pump Suction Pressure Fluctuations

\.@
S
/

 

1b

£t 1b

 

HEAD (.

-19-

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

2700 cpm I/ /
<
// 4
2]4301'?1“ . /
1/2>
X — N
A N AN
2100 3 -
mPs L
1800 npm ~Z ™\
Y~ \\ /
15001‘pm-.\ | ‘ /
| T~ /32 1/16
=~ .
350 0500 &0 ——750 8O0 %0 —mte —T

CAPACITY (gpm)
' FIG. 2.2l ART FUEL PUMP FLUCTUATION CHARACTERISTICS WITH MODEL- 32 IMPELLER -

 

AT A 1/2-IN, DEPTH OF LIQUID (WATER) IN TE EXPANSION TANK R - @@W
£ 1b
1b

-29-
ZAD (

H

70

60

50

40

onne .

 

 

 

 

jel Pump Suc

Fluctuations .

 

\ .
N\
N\

N

 

tion Pressure

N/

TN /'\'./
) ' | /8 1t
N/ /58 ek
S

N 7
ISL NG

£ Oy

 

 

 

1/32 \'\/ g
i

 

 

 

 

 

 

 

 

 

 

 

 

 

 

1209 //
1/16 -
\'\_-.‘ \(‘/
"--"
300 600 rpm | ' '
TPMmey,_ | ,
0 100 200 300 Loo 500 600 700 800 - 900 1000 1100
- CAPACITY (gpm)
. f » '
FI3, 2,25 RT FUEL PUMP FLUCTUATION.CHARACTERISTICS WITH MODEL 32 TMPSLLER ’ ‘
AT A 1-IN, DEPTH OF LIOUID (WATER) IN THE EXPANSION TANK —
PR
-€9-

ORNL-LR- . 23948

 

 

 

70 ; ; T ‘ - - 3300 rpm Pump Speed

50

 

 

: . ' - 07671~ S , \i-mel Pump Suction| Pressure Fljictuations
50 ‘ - / Y _ ; ) \ ) .

 

 

)

2700 rpm ‘\,//7

Lo +— /

N
N
3

 

g

  

 

S
N

HEAD (

~/

I

7N
~—
/A

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

0 §-68
3 2100 rpm :\'w / , \/
. ‘ .\. /7 .
: ‘ . — | / ~C 10.28° £t
20 : 1600 rpm_ . ' ,/ ~ 0.2L £t |
= 7 -
p— <l X
1500 rpm . o= b -
. . \ - /
10 000 5 0.02- S B, _
. . \  —— - 0. Oh \-
00 Me—s.
400 L. -
P -—
LSO rpm -. .nf'_ ~— ' ' | ‘
o 100 . 200 300 0o 500 600 700 _ 800 900 . 1000 1100

CAPACITY (gpm)
©  FIG, 2.26° ART.FUEL PUMP FLUCTUATION CHARACTERISTICS WIT{ MODEL 32 IMPELLER |
AT 2-IN. DEPTH OF LIUID (WATER) IN THE EXPANSTON TANK T
ORML-LR-DWG., 23949

 

 

330D rpm Pump Speed

 

><

3/8
: 5/16><'

 

WV = . |
: \§V—F\m1 Pump Suctiop Pressure Fluctuations

 ——

 

)

ft 1v

 

 

HEAD (

-v9-

/
SN

210D rp:/ €< AN ‘Y _ / | ‘\
5/15

 

2100 rpm

 

2700 rpn \/
/

N
N

/7
_ 1716 4
1500 rpm /

7
~
1800 krpm \//\/N\
/ /
1200 /\1/59/ > \ |

 

. = Y
N |

 

 

 

 

 

 

 

 

 

 

 

 

 

 

1/’ - TAe
| [ /16 "\
R ' |
S-.-‘\-.
300 %OTHR___
100 200 300 500 500 600 700 800 " 900 1000 1100

, | CAPACITY (gpm) ) ] _
FIG, 2,27 ART FUEL PUMP FLUCTUATION CHARACTERISTICS WI T MOTEL 32 IMPELLER
AT A 3-IN. DEPTH OF LIQUID (WATER) IN THE EXPANSION TANK

 
-g9-

ORNL-L R-DW¥G. 23%0

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

e Cplculated
= o =« = =" = = Experimental
o . .
55
2ho; " .-1"/26 Expahsion Tank
' X Liguid Lewvel
2100 rpm |
50 33 ™
v \ |
3/‘hw_..\ \\ \ \ o !{ 3300 rpm B
R 1 I
_‘.\' |- ‘\ /|
L A
AT
Ay 7] 2700 rpo
VoL
\ | ,./‘? 100 rm
e s T
"~ P100 rpm

 

 

 

 

 

 

 

 

 

 

 

 

 

 

FIG. 2.28 CALCULATED AND EXPERIMENTAL ART FUEL PUMP BYPASS F;L.OW CHARACTERISTICS
ART Fuel Pumpwgavitation Characteristics )
| Cavitation refers to conditions nhere,.because of a~local increase in

velocity as in regions of either boundary curvature or‘eddy.generation,

vaporization of the flowing liquid creates vapor filled cavities. These

cavities collapse . in regions of higher pressure.elseuhere in the system.

In order to form such vapor cavities,‘the pressure first has to drop to

the vapor.pressure of‘the working fluid; a condition that can be realized

as a local condition without a change in the average system.pressure. For

a pump oPerating at a constant-speed.a local pressure drop results from

separation and contraction of flow and deviation of streamlines from their

normal trajectory such as takes place in a turn or in passing,an obstruction

to the flow. The evils of the cavitation phenomenon are essentially three

in number: (1) reduction in efficiency due to constriction of flow and loss

of energy, (2) objectionable - if not structurally dangerous - vibration and

and noise; and (3) possible extensive pitting of houndary materials in the

zone of bubble collapse, due apparently to failure through fatigue after

countless stress reversals. '

The local pressure drop caused by the difference in the pressures on
the leading and trailing sides of theVimpeller blade is.the ma jor cause of
cavitation in a centrifugal pump. This local pressurefdrop is a function of -
the relative inlet velocity. (W), the nuiber of blades (Z), the thickness
of the blades (t) and the local absolute velocity (C ).

The.tegt loop from Which the original pump. development data was obtained
was not constructed in a manner to allow operation at an elevated temperature
or at a. suction pressure below atmospheric,' For this reason the determination
of the fuel pump cavitation characteristics was delayed until the construction
of a suitable loop._ The loop designed for the hot testing of the fuel pump
was used for obtaining the cavitation data. The Thoma cavitation parameter'
((f)(T)(B) was determined for the final 1mpeller design as a function of the
percent change in the pump maximum head at a constant flow and speed. The
parameter o is

-66- .
L egEAme s
. ' .y

where H is the pump total suction pressurenmeasured‘close to the. pump eye.in‘
order to avoid frictional losses, Fvap,ls the vapor pressure of the working
fluld -and H max is the pump-maximum totel head. -

The results of these’ experiments are shown in Figs 2.29 through-2.31. .
From these cavitation data and the data previously obtained on the:difference
between the expansion tank pressure and the pump suetion pressure, Fig. 2,32
was made on the assumptlon that the suction pressure requlred t0. prevent cavie

tation was independent of the liquid level in the expansion tank. Fig 2.32

e

afi&“

- gives the minimum expansion tank pressure required to prevent cavitation as &

function of the expansion tank liquid level and the pump speed:and flow.

-67-
-89~

Change in Maximum Head (%)

10

15

20

or I

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

% '——'?fl'-.—fll-“———-&—o—o—b 2 2 =‘i"""'—: —h—a— d—A—gb———h
;’*.— e . a )/— . h
. -1 / A .
n ‘ / 4
° 2 .
8/ ,
& / /1
& & | 600 gpm g = 0.61
= ® | 645 gpm. 0. = 0.825)
A 700 gpm oo = 1.40
0 0.5 0.6 0.7 0.8 0.9 1,0 1.1 1.2 1.3 1l LS 1.7 1.8

VALUE OF CAVITATION PARAMETER @& ,
FIG. 2.29 CAVITATION CHARACTERISTICS OF ART FUEL PUMP WITH MODEL 32 IMPELLER
AT A PUMP SPEED OF 2L0O rpm AND A 3-IN, DEPTH OF LISUID (WATER) IN THE EXPANSION TANK

1.6

 
-69—-

Change in Maximum Eead (%)

ORNL.L R-DWG. 23952

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

0 = +———F a4 4
° | L‘/
. A 4/
5 / i
w a o
10 | ‘
I. ® @ | 600 gpm O:':O.hé.
® |65 gpm gz = 0,51
4 | 700 gem g7 = 0,785
15
20 =
0, 0.2 0.3 0. 0.5 0.6 0.7 0.

 

 

 

 

 

 

 

VALUE OF CAVITATION PARAMETER O

FIG. 2,30 - CAVITATION CHARACTERISTICS OF ART FUEL PUMP WITH MODEL 32 IMPELLER -
AT A PUMP SPEED OF 2700 rpm AND A 3-IN. DEPTH GF LIQUID (WATER) IN THE EXPANSION TANK

it gt
2 ‘fl‘:ji'

 
-0c-

Change in Maximum Head (%)

10

15

20

onu“s

 

 

 

 

 

 

 

 

® | 600 gpm -, = 0.L412
@} 645 gom L . = 0.Lh3
A| 700 gmeNy = 0,583

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

0.1

0.2

0.3 0.4 0.5 0.5 0.7 o;
' ' VALUE OF CAVITATION PARAMETER O

FIG. 2,31 CAVITATION CHARACTERISTICS OF ART FUEL PUMP WITH MODEL 32 IMPELLER
AT A PUMP SPEED OF 3000 rpm AND A 3.IN, DEPTH OF LIQUID (WATER) IN THE EXPANSION TANK
_lé-

Expansion Tank Gas Pressure Minus V:)apor Pressure of Working Fluid
(£t abs

26

2k

20

18

16

1L

)2

10

ORNL-L R-DWG. 23954

 

700 gpm Kn. Expansion Liquid Level

 

 

 

3000 rpm
1 in,
e
, ‘ 1/2 in.
600 2 in. 1/2 1in, /\
gpm _
3 m.]\ 1 4n
1 in, 2 in,

: 3 in,

. 00
2 in 3 in, 700 gpm
r 700 gpm

 

 

 

] /]

600 /

 

645 gpm 6L5 gpm

 

 

T~7500 gpm

 

3000 rpm 2700 rpm- _ 2L00 rpm

FIG, 2,32 . MINIMUM EXPANSION TANK GAS PRESSURE REQUIRED TO PREVENT CAVITATION IN TYE ART FUEL PUMP
 

SECTION IIT

DESIGN OF AN ATTITUDE
STABLE XENON REMOVAL SYSTEM

 
 

 
 

DESIGN OF AN ATTITUDE-STABLE' XENON REMOVAL SYSTEM

In addition to the design precepts and requirements of a stationary up-
right xenon .removal system, one suitable for application to aircraft must meet
the additional requirement of attitude stability (9)_ An over- all picture of”
the original attitude stable X-R system is given 1n Fig. 8. 1. Illustrated
particularly are the swirl chamber in relation to the rotary elements, the
interconnecting parts, and the direction of fuel passage The swirl chamber
acts as an expansion tank in the fuel system and - as & processing tank for.
xenon stripping. All other’ required functions ‘are carried out in the-rotary
assembly,valcross section of which is-shown‘in'Fig.’3.2, which shows a full
scale test model of the rotary assembly. The nomenclature of parts as used
in the description is identified in Fig 3.2. o

Fig. 3.3 depicts the system as a. hydraulic circuit : The numbers are
only illustrative tonp01nt out the pressure interdependence between the X-R

system and the main fuel circuit.

Circuit Description

The bleed circuit is most readily understood by referring to. Figs 3.2

and 3.3. In the description below the circuit will be described in zones,

as follows:
Zone 1l: Centrifuge, core header region and nozzles
(pressure control) -

Zone 2: Nozzles, swirl chamber; and flow split stators '
(primary swirl)

Zone 3 Flow split stators, swirl pumps, and swirl chamber
(secondary swirl) . |

Zone l. .System Pressure Control
Circuit.wise, it is best to consider the centrifuge as a radial vane

centrifugal pump (see Fig. 3.4) discharging fuel to the core header region,

which serves as a mixing chamber into which processed fuel is discharged

 

 
60

ORNL-

TR
YA TG

 

5 PRt iy a0y

 

 

 

 

 

 

 

 

System

Fig. 3. Attitude-Stable Xenon Removal
ORNL-L R-DWG. 23955

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

  

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

? - ‘UPPER SILINGE{R SEAL ?
AN\ Y,
/ N _ /§ Y -
/ I > K / |
T 2
AT 1Y I / \
o 1 \§ TR
: LCENTRIFUGE
LW © / | | AN (e SLINGER
/ | ‘ | seac
SRR % CENTRIFUGE 2
9 (I %
{.I,llmll-'l-'l_s % o _ 2 CENJ(I-‘;‘II_FEUGE

FIGURE 3.2- ATTITUDE-STABLE: XENON REMOVAL SYSTEM

ROTARY ASSEMBLY
-77-
 

 

 

 

 

 

   

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

   
 

| rw B - [1
~ _— |
Dp~60psi I | :
. SLINGER
777
2 _
AL
, 7 .
; : wl
D p~30 psi Z 513
‘ 1 — &
‘ ' : | 2NN
XENON- REMOVAL SYSTEM | 7 & ;
: | SPLIT
BLEED. CIRCUIT < | g - , | 7 SPLIT
/A L - Z N SWIRL STATOR
* _ 2 ) x L CHAMBER
L 2
. W NOZZLES
8’ °© 4 CENTRIFUGE
4 ; 512« I SLINGER
L/ “ 2 a W SEAL
Ap~ 41 psi * > ‘ -
. Ap~ 43 psi —e | 40psi b . .
| C A CENTRIFUGE
N~ Y O
(” 7 - FUEL PUMP L

HEAT gxcHANGER

REACTOR CORE
-~ MAIN FUEL CIRCUIT

 

 

 

\ ‘

NUMBERS DENOTING PRESSURES ARE ONLY ILLUSTRATIVE.

FI1G. 3.3 SCHEMATIC OF ATTITUDE-STABLE FUEL CIRCUIT

e 78 -..~;_~_.,._,
E | : UNCLASSIFIED

ORNL-LR-DWG. 23956

 
  

 EXPANSION
TANK

 

 

 

 

 

- GORE HEADER
(PLENUM)

 

 

 

 

S EEEE G ISR IS TS AN TeE—" e e ST S—

FIG.3.4- ATTlTUDE'STAw,STEM PRESSURE CONTROL

-79-
and from which an equal'volume of fuel is bled for processing. Two nozzles
direct the‘bleed flow into the expansion - tank as'tangential jets to induce

‘a swirl, These nozzles represent the major resistance in the circuit.and,

| together with the centrifuge speed, determine the bleed flow rate and the
header region pressure. Since the header region is also the discharge region
for the main fuel pumps (shown dotted in Fig. 3.4) it is seen that the bleed
circuit in effect determines the main pump discharge pressure and thereby

the reactor system pressure. By making the pressure drop across..the nozzles
equel'to the mainfsystem fuel pressure drop, the main fuel pump suction pres-
snre can be maintained at swirl chamber pressure (controlled by the off-gas

system).

Zone 2. Primary Swirl

This part of the bleed circuit serves the primary purpose of agitating:
the fuel. The fuel in the swirl chamber is made to spin (reinforced by the
circuitsof Zone 3) by the tangentially directed discharge of.the fuel from. :
the nozzles (see Fig. 3.5) and careful contouring of the swirl chamber walls.
The resultinguhigh peripheral’velocities‘induce violent agitation; mixing, |
. and gas entrainment in the spinning fuel. The spin also serves to make the
fuel level insenstive to attitude. As seen in Fig. 3.5, ducting designed
into the .swirl chamber wall directs fuel into two protruding ports in the |
pump hou51ng. The - fuel scooped up by the ports is directed to the flow split
stators which divert such fractions of the fuel axially downward into the
'centrifuges as' is required to balance the centrifuge discharge’ (see Zone l)
~ Under normal operating conditions the centrifuge is simply flooded, the ex-'

cess being diverted upwards into Zone 3.

Zone 3. Secondary Swirl |

The upward directed excess of Zone 2 is passed to two auxiliary centri--
fugal "Serl" pumps mounted on the main fuel pump shafts. The discharge of
these pumps is directed by suitable porting and contouring (see Fig. 3. 6) to
reinforce- the spin in the swirl chamber. This additional energy is.not re-
quired until the level in the swirl chamber rises to about 40% full. : By

-80-
- 18-

. UNCLASSIFIED - .
ORNL-LR-DWG. 23957 -

 

 

 

 

FIG.3.5- ATTITUDE-STABLE SYSTEM PRIMARY BY-PASS SWIRL
-zg-

. UNCLASSIFIED
ORNL-LR-DWG. 23958

 

 

FIG. 3.6 - ATTITUDE - STABLE SYSTEM SECONDARY BY-PASS SWIRL
virtue of its elevated position the-SWirl pumps automatically cut.in when the

level in the swirl chamhéfrises sufficiently to prime the swirl impeller;
The swirl'impellér aisomserves to vefit the gas'liberated from fhe fuel

in the centrifugeé. A gas passage conneéting_the swirl impeller suction,side

with the centrifuge center is formed by.é shroud around the shaft as shown in
Fig. 3.7. | ‘

Dynamic Seals

| - To complete.the.circuit.descriptioh two important seals will be described
briéfly. Both seals are essentially radial vane centrifugal impellers arranged
to pump agalnst the pressure of fuel enterlng the impeller from what is normally
the dlscharge end of the blades. '

In the case of the upper slinger seal, its function ‘is to prevent any fuel
"leaking past the swirl pump from moving up the pump.shaft toward the bearings.
The upper slinger seal impeller radius exceeds the swirl 1mpeller radius in
order to prOV1de some safety margln and to permit 1nverted operatlon.»

The centrlfuge sllnger seal is an integral part of the centrifuge (see
Fig. 3.8) and serves to limit the amount of recycle around the top of the
centrifuge. The blade diameters must be such as to .develop nearly centrifuge
pressure but never to exceed it. A more detailed.discussion regarding design

and performancé is appendéd.

Test Development

 

The chronological stages of the development of the ART-XR system are
summarized diagrammatically in Table 3.2. Excerpts of a summary by A. P. Fraas
of the development of the first six models are abpended to this report. These
modeis first indicated the need'of_emfiloying a centrifuge and pointed out
several instability problems which led to the decision to build the swirl
chamber as a separate physical entity.. ' '

Denoted as Model No. 7 in Table 3.2 and 1llustrated in Fig. 3.10, such
a swirl chamber was tested independently of the pumps by supplying feed from
the plant sanitary water supply line. Tests were run at 2.6, 4.6, 8.9 psig

-

-83-
 

UNCL ASSIFIED
~ ORNL-LR-DWG. 23959

< 7

—

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

FIG.3.7- ATTITUDE -STABLE SYSTEM PUMP

SHAFT SHEOUD,
.fl‘-&‘r
| -Efi'gfi
'UNCL ASSIFIED
ORNL-L R-DWG 23960

   
 

el

   

 

 

 

 

 

 

 

S
RS

 

 

 

 

f ////////// N /////////

Fl1G. 38 ATTITUDE STABLE SYSTEM CENTRIFUGE
o SLINGER SEAL \ -
inlet pressure with various resistances in the discharge lines simulating
lift and. friction losses., The test data are summarized in Table 3.3.

In conducting the tests it was observed that the 1iquid 1evel in the
swirl chember sought 1ts own equillbrlum 1eve1 for any combination of inlet
and back pressure. Efforts to change liquid level independently.of pressures,.
by adding water directly to the swirllchamber, were'unsuccessful'-either.the
level reverted promptly to its original equilibrium level or a non-equilibrium
' condition resulted with the level continuously rising to overflowing.

The nature of this swirl instability was-subsequently analyzed.and is
schematically represented in Fig. 3.9 in terms of'an.energy balance. The

sequential dependence is as follows:

1. The amount of liquid retained by the centrifuge controls the

centrifuge delivery pressure.

2. The centrifuge delivery pressure acting on the nozzle deter-
mines the flow rate out of the centrifuge and into- the swirl
'chamber. N '

3. With the flow rate and pressure established the energy input
to the fluid is established '

L, For equilibrium conditions the input~energy,must equal the

losses. The losses can be broken down intos

,a.'-nozzle loss - a function of flow
b. swirl chamber discharge loss - a function of flow .
c. 1lift - fixed
d. swirl chamber agitation loss - a function of swirl chamber
| liquid level ' :.
_e. swirl chamber drag loss - a function of swirl chamber |
liguid level.

5. Since the swirl chamber also serves as the expansion volume for
the fuel systems, & rising fuel temperature will initially cause
a rising level in the swirl'chamber. . This changing level by way
‘the 44 and e dependence above, affects the system equilibrium in
a manner dependent on the initial equilibrium liquid level.
.

-86-
—18-

 

 

 

 

 

 

 

' UNCL ASSIFIED
ORNL-L R-DWG, 23961

 

 

 

 

 

 

 

  

 

 

  

 

   

 

 

 

 

 

 

 

 

 

r--—-—~~— -~~~ ~~—"~""~>"~—>"~>"™"™"™>"™"™""™>""""“""™"™"™"">""""" """+ [
| LIFT >
| jo
z
| I &lu
| T — , |2
| ) . SWIRL |GHAMBER : ccoom )%
’ | DELIVERED o ‘ I - FIXED 2 Jjo
L_{GENTRIFUGE ————w=< NOZZLE | | ST e A
| - WALL | . aciTATION o DISCHARGE
FRIGTION | ’AGITATION , ~_
T - ;Loss ) =
L0SS HIGH  LOW =
LEVEL LEVEL - @
___________ I ‘i
I
°/6 FULL - 1 9. FULL 1
| A . |
_- WALL AGITATION |
|- "HAaFRICTION N3, 1
N o A |
a \
o N\ I
o R |
S %% FULL .. |
I ' ‘ I
L — | EQUAL"Q" FOR EQUILIBRIUM}— — — — — ]

 

 

 

FIG. 3.9- ATTITUDE - STABLE SYSTEM SWIRL CHAMBER ENERGY BALANCE
)

From Fig. 3.9 it is seen that the agitation,losses_decrease
with rising level and .is the predominant loss at "low" liquid
levels. The drag losses, however, rise with 1ncre331ng level

and predominate at "high" liquid levels._

Instabillty.A. _
With the swirl chamber at an initially "low" liquid level, a
rising level (due to a rising fuel temperature);reduces'the _
swirl chamber loss and thereby inereases the flow out.ef the
swirl chamber. With the centrifuge cup initially full this.
does not, however, . increase the flow into the swirl chamber.
In effect, then, the swirl chamber reverts to its initial -
stable level, transferring the extra inventory to the space

7 -

above the centrifuge.. As born out by test experience, the

swirl chamber "refuses to act as an expansion chamber".

Tnstability B: |

With thé swirl chamber at an initially "high" liquid level, -
& rising level increases the swirl chamber loss and thereby
decreases the flow out of the swirl'chumber_andvinto-the.
centrifuge;cup. With ite supply reduced tne centrifuge 1iquid

‘_level'drops'thus reducing the energy supplied to ‘the swirl
chamber. With energy 1osses increasing {due to rising 1iquid
level from (a) the extra inventory due to temperature rise and
(b) the volume no longer retained by the centrifuge) and input

" energy decreasing, the system is more and more unbalanced, re-
sulting in a run awaylcondition ae bOrn out by test experience.'
The. swirl chamber may overflow and the centrifuge cups will lose
their primes.\'

At pressures above 7 psig suff1C1ent flow was obtalned to give good spin
and agitation. However, the tests indicated definite 1nstability tendencies

for certain combinations of liquid level, back pressure and inlet pressure.

-88-
    
  
 

ORML-LR-DWG. 23962

 

 

 

 

 

NOZZLE

 

 

 

     

 

 

S SN NSNS N

| |
SN SRS

 

AN

 

 

 

OOV NN NNNN

 

i

 

 

 

 

2z
&l

Ty

  
   

 

 

 

 

- / T T

 

 

 

 

b

[ ——
g

FIG. 3.10- ATTITUDE - STABLE XENON REMOVAL SYSTEM TEST MODEL NO.7
gx

L Ze e

.'; B -;14
IR

The 1nstab111ty tendency was more clearly apparent in Test Model No. (Table
3.2) when the swlrl chamber was placed’ in series with one centrifuge.
The analysis of the swirl chamber indicated a marginal energy balance be-_

o

tween the energy in the fluid issuing from the jets .and the fluid friction los=-
ses in the swirl chamber. At some minimum swirl chamber level a metastable
0perat1ng condition was attained. ;Any.small-increase"in frictional resistance
unbelanced the system to the point where the centrifuges lost all prime and
freelylpassed air. It was not possible'to reestablish stable Operating con-
ditions. In Test Model No. 9 every effort was made to reduce connecting line
friction losses. Furthermore, the swirl chamber was raised such that the
centrifuge could be supplied by gravity. However, even under these conditions,
the system could not be stabilized over the full range of liquid levels.

It was, therefore, decided to augment the energy supplied to the swirl
chamber by an;auxiliary impeller as shown in Model No. lO,.Table’3.2. The
centrifuge discharge was passed upward through the clearance space between the
centrlfuge cup and. the wall where it joined the auxiliary bypass fluid discharged
by the impeller mounted on. top of the centrifuge cup. .

Although the centrifuge flow rate could not be determined, this model was
able to remove bubbles from the plenum chamber below the centrifuge. Swirl
chamber operatlon was stable. However, no appreciable system pressure could
be developed. ‘ :

The performance of Model 10, Table 3.2 and the anslysis leading up to it
indicated that the system was. sufficiently'well understood. to make an attempt
at an improved de51gn appear promlsing. A two-pump integral plastic model was,
therefore, de51gned incorporating the following features:

1. Locating the SW1rl chamber above the centrifuge cup to permit priming.

2, Incorporating an auxiliary impeller wh1ch by superimposing a re-
C1rcu1at1ng Tlow supplied excess energy to the swirl-chamber.

!3. _De51gning a configuration such as to make passages between swirl
chamber and centrifuge assembly as short as possible to minimize

friction losses.

)
4. Introducing the centrifuge delivered flow-into the swirl.chamber
. through highly restrictivé nozzles to maintain the system pres--

surefand to provide .spin at low swirl levels.

The assembly as originally built is 1llustrated in Fig. 3 11.

The development of this unit to the final working model involved changes'
in the rotary assembly only These are illustrated in Table 3 2 Models 11 -
through 15. The problems encountered were primarily sealing problems to prevent
reciroulating short circuits. This development work is outlined in Table 3 2.
Finally, the entire assembly wags mounted on-a rig to permit inverted operation.
The performance is summarized in Table 3.k. The final critical dimensions for
Model 15, Table 3.2, are listed in Table 3.4, The final design'is illustrated
in Fig. 3.12. '

System Analysis (Refer to Fig. 3.3)

The manner in which the X-R system controls the main system pressure is
most easily seen from a study of the characteristic curves of the two systems.
Figs. 3.13 and 3.1k illustrate the characteristic curves for the main fuel
pump circuit and bypass flow circuit respectively. The ordinates.are—drawn to
the same scale but intentionally do not have their origin defined. The labeled
points have the following physical significance:
(a) fuel pump discharge pressure
(b) centrifuge discharge pressure
(c) centrifuge inlet pressure
(d) fuel pump suction pressure.
As previously described and illustrated in Fig. 3.4, both the centrifuge
and fuel pump discharge into a common header and therefore points (a) and (b)
must be.at the same pressure. Secondly, since the centrifuge inlet, because of .
its gas interface is_essentiallyrat swirl chamber helium pressnre, i.e., since

point (c) is at p = O (reference), all the other pressures are fixed relative to

(c) as shown in Fig. 3.15.

.-91"'
fic‘gi Ivlv,i -

The system pressure, particularly the pump suction pressure, is  therefore

controlled by the helium. pressure and the . bypass flow pressure drop. The bleed
flow Ap is primarily controlled by the nozzle resistance. This nozzle is made

highly restrictlve 1n order to keep the pump suction pressure high enough to

- prevent cavitation. Any pressure drop across the centrifuge cup holes reduces

the syétem pressure by an equivalent amount. Thus any condition that simulates .
an increased flow ;hrough'fhé céhtrifuge-cup holes effects a reduction in the - -

system pressure level.

.-92;

 
Hnas

Table 3.1 Model 15 - Performance Evaluation.

Test Condltlons
A. Speed range 700 rpm to 1800 rpm
'B. Storage volume range 20% full. to 80% full
'C. Attitude: 1) upright, 2) incllned, 3) inverted.

 

Performance Criterion  Test Condition . Evaluation
1. Expansion Volume Stability | A - Completely stable
: , : Completely stable
c ~ Completely stable
2. Agitation | A | Varies from fair to

excellent as speed’
" increases.

B , Best at low volumes -
generally good.
c Good.
3. Upper Slinger Seal Effectiveness A Complétely effective
B Completely effective
C Completely effective
L. System Pressure A Within 5% of theoretical
' - ‘ expected.
B o Drops with dropping vol-
"~ ume when less than 20%
- C ' . Drops off completely at
low volume.
5. Bubble Removal Effectiveness A © Good - best at lower
: o . "~ speeds.
B ‘ Good - best a high liguid
' -+ levels.
C o Fails at low volume and
~low speed.
6. Power Required A - Increases with speed.
B - Increases markedly with
o N : : high liquid level.
o e
7. One Pump Out , | .. Loss of system pressure;

poor agitation.

_93_
 

SEQUENTIAL DEVELOPMENT OF ATTITUDE-STABLE XENON REMOVAL SYSTEM

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

MODEL . o - o MAJOR CHANGES FROM
NO. REFERENCE FLQW PATTERN . PREVIOUS MODEL EVALUATION.
N o SLINGER
2 Fig. A.1l 1) Swirl insufficient to
STORAGE remove bubbles.
| | :
PUMP SWIRL 2) Agitation good
T - r TN 3) Slinger effective.
r—— ==~
i /T ]
T 4 | . * '
| L
2 Fig. A.l Jet added to assist swirl., Swirl stili_insufficient
to remove bubbles.
3 Flg. A.2 Centrifuge added above pump Insufficient 1ift avail-
impeller. able for getting fluid
into centrifuge.
4 Fig. A.2 Scoops added in swirl Effective up to 100 rmm
’ chamber to assist lifting only.
fluid into centrifuge.
5 Fig. A.3 1) Peripheral airfoil scoops | 1) Unstable operation.
added. 2) Passed air bubbles.
2) Expansion - and swirl '
chamber built as unit
chamber (baffle omitted)}.
3) Centrifuge discharge into
impeller discharge region
5 Fig. A.> Lebyrinth seal added between | No improvement,
. punp volute and swirl
chamber,

 

Decision: 1) To separate centrifuge and swirl

chambers.

2) Hew models not to include simulat.d

main circuit.

 

 

 
Table 3.2 (contd.)

 

MODEL

MAJOR CHANGES FROM

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

10. REFEREICE FLQW PATTERN - PREVIOUS MODEL EVALUATION
7 Fig. 3.10 Isolation of swirl | Test of swirl chamber
chamber from pump |with city water presswure
impeller. . . supply and variably
. o restricted discharge.
‘See Table 3.3.
SWIRL
CHAMBER
C——=
8 Fig. 3.10 __+__ Swirl chamber - 11) system unstable as
- | connected to single liquid level is
i I centrifuge unit. changed.
" )‘, 2) Centrifuge loses
prime.
[fifi' r_~1 _
(& 5
|
g rig. 3,10 = H Swirl chamber Marginal performance due
. o raised with respect| to insufficient spin
T B to centrifuge. energy .
. =] ) . R .
: ’(’ Flow reduced by
- _|plocking one nozzlel
_F _’_ " - * . .
-
t ‘
!
10 G Discharge through |1) Effective bubble

Fige 3.10

P

 

J |
jj]

 

 

 

 

 

 

 

)

 

 

 

 

 

 

 

 

 

connection opposite
impeller mounted
on centrifuge.

 

separation.
2) No system pressure.

%) Good spin except at
very high liquid level.

L) Centrifuge can lose
prime because of ‘
limited spin energy
and excessive fluid
friction losses.

 

 

Decision: .To build plastic model contaihing’twb centrifuge assemblies with integral swirl chamber between
shaeft centers. B ‘ ‘ '

-95-

 
Table 3.2 (contd.)

 

MODEL

MAJOR CHANGES FROM

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

i oV
NO. REFERENCE FLOW PA'I'I‘ERN PREVIOUS MODEL EVALUATION
11 Fig. 3.11 1) Two centrifuges Slinger seal ineffective.
: feeding common swirl '
chambers.
2) Auxiliary impeller
to reinforce swirl
“at high volume
levels.
12 Fig. 3.1 Slinger replaced by 1) Slinger impeller
' slinger impeller. effective.
2) Low pressure in main
systen.
]
A—
13 Fig. 3.2 | ' Impeller edded above 1) System pressure
' ~ h centrifuge. satisfactory.
?;/ f ' 2) Bubbles Ly-possed or
%% passed through
2 7 centrifuge.
YLllls 74
\‘\ E
/y_
1h ' Special test rig to Labyrinth seals in-
investigate sealing effective.
i . arrangements.
' ‘ gZzzz23
: [ 1 [
\ 1
15 rig. 3.1 Same as Model 13. Identical with Model 17| Satisfactory operation
3.2 but with rebuilt centri- with respect to:
3.12 fuge and carefully a) Dubble scparation

 

 

 

 

 

" |verting rig provided. -

machined inpeller. In-

 

 

b) Stability
c) System pressure
d) Attitude changes

 

 

 
Static Head

 

Table 3.3 Model 7--!Perfdrmance Evaluation

 

Spin Velocity Swirl *
Inlet  Back Measured 3/4 in. above Chamber Agitation Stability
Pressure Pressure scoop § ft/sec Flow Efficiency Evaluation Evaluation
£t £t _ cTs 'J_i'% —
6 0.52 5.25 0357 '1‘16;7 | fair stable
0.605 ' 5.78 .0357 -18.2> poor stable
unstable B - | | - unstable
10.6 0.52 3.38 .0k88 13.4 good ' stable
0.605 3.02 W73 13.7 good " stable
0.687 322 460 1h.1 fair stable
0.77 3.0k 0473 15.2 poor stable
0.855 5.25. OLT3 16.1 noné stable
unstable unstable
20.6 0.52 - 8.9 .067 10.8 -éicellent stable
0.687 8.181_ 067 . 11,6 . excellent . stable
0.935 - 9.62° 067 12.9 excellent stable
1.27 _9.62 067 1k.5 excellent stable
1.40 9.62 067 15.2 fair stable
unstable unstable

 

* Swirl chamber efficiency =

Total Head of Leaving Fluid
Total Head of Entering Fluid

 

_97-
 

Table 3 A
- . Model 15 - Dimensions
Bléed Flow Rate - .0297 ft3/sec

Swirl Chamber Nozzle Orifice - 1/16 in. x 5/8 in.
Expansion Volume required N 0.45 £43
 Expansion Tenk Dimensions. _ 0.D. 1k in.
| I.D. 5 in.

Height 5.75 in.

Centrifuge Cup Size B o 0.D. 5.75 in.
’ Lo I.D. 3.50 in.

Centriffige Holes ' : + © Number | 8
| | | | Diameter - 1/4 in,

Centrifuge Slinger Seal -~  radial vane I.D. 3.50 in.
| | ' " radial vane 0.D. 5-5/8 in.

 
—66—

 

ORNL-LR-DWG, 23!63

SECTION A-A = .. : .SECTION B-B

 

 

 

 

 

 

    

 

 

 

 

 

 

L

 

FIG. 3.11- ATTITUDE-STABLE XENON REMOVAL SYSTEM TEST MODEL NO. I -
iz
A B
S

  
  

SECTION B-B

 

 

 

S
%

=001~

N\

 

 

 

 

5o o -Et-r; -!;'!

 

 

 
UNCLASSIFIED
ORNL-LR-DWG. 23965

  

 

     

 

 

 

 

- PUMP DELIVERY PRESSURE | l o
| ~ SLIPPAGE
| 85 4r HoLeg
r5-———=———-- : | CENTRIFUGE
| DELIVERY
AP
AP : PRE SSURE
|
|
|
0 .44 Q 0 0268 Q
FIG. 3.13 | FIG. 3.14
ATTITUDE -STABLE SYSTEM MAIN ATTITUDE - STABLE SYSTEM

FUEL CIRCUIT CHARACTERISTIC BY-PASS CIRCUIT CHARACTERISTIC

 

 

 

FIG. 3.15

ATTITUDE -STABLE SYSTEM PARALLEL MAIN FUEL AND
BY-PASS FLOW CHARACTER!STIC |

-101-
Appendix A

SUMMARY OF WORK (PO JULY, 1954) LEADING TO THE
DESIGN OF PUMPS FOR THE CFRE
 

 
 

B SUMMARY OF WORK (TO JULY, 195k) LEADING TO THE
DESIGN OF PUMPS FOR THE CFRE |

(Refer to Models l through 6 Table 5 2)*

“The face-type gas-seal sump pump is attractive for use with high tempera-
ture liquids because, by inserting a heat dam, the seal can be operated at -
conventional temperatures with conventional materials, for example, hardened
steel'and-graphitar~in‘a petroleum ‘0il bath providing both lubrication and
cooling. . An experimental unit was prepared in the. summer of 1950 on the thesis
that the free surface.could be maintained in a centrifugal field and hence could
be made insensitive to -the vériations_in pump attitude and negative "g" loads to‘
be .expected in flight. A particular feature of the design was that it should:
act . to de-aerate the system -- a feature considered essential in a full-scale
aircraft type pump. . The principles of operation were demonstrated by operating
the pump with water in the fall ofkl950~ap@fapprokimatélyLSOersofqurther@testing
with sodium at temperatures up to 1000°F was carried out in fhe sumer of 19510.

. As the design:of the ART progressed it became evident that there is a
strong incentive to remove xenon, and that this might be done by agitating the
fuel in the -expansion tank so that it would entrain bubbles., This would serve
both to: increase the surface area by & large factor and to bring a high profiorf
tion of the fuel close to a surface within a short‘period.of'time.- Thus ‘1 or
24 of the fuel could be by-passed through the expansicn tank and the bulk of
the xenon stripped before returning it to the ‘main stream. . The arrangement had
the additionsl advantage that it would avoid overheating of the fuel in the
expansion tank as a result of fission product decay activity. While a separate
agitator could be provided, it was felt that the number of moving parts should
be minimized and that the pump impeller shaft could serve a double purpose.

The notion of utilizing & centrifugal field in back of the impeller to stabi-~

lize a free surface there had Worked-so#nicely in the pump.mefitioned'earlier

 

*A. P. Fraas, abstract taken fram "CFRE De51gn and Development Program.Handbook,
Prcject No. 3. . . P | , C .

. ~105-

 
 

that a modification of it to effect xenon removal seemed promising. A simple
model of work, glass, ‘and plastic was prepared to 1nvest1gate the possibilities.
A section through this model is shown in Fig. A.1l. The model was designed to
operate with the fluid in the expansion tank swirling under the impulse of fluid
accelerated and thrown off by the impeller. A coneiderably higher swirl velocity
vas .expected -in the annulus just above the impeller, and a diaphragm was inserted
to separate this region from the main part of the expan51on tank above. It was
hoped that the swirl veloc1ty‘1n this lower region would be high enough 50 that
the outer periphery_Would‘operate_free of bubbles. Thus, fluid leaking upward
through the labyrinth seal between the impeller and the. casing would pass into
the expansion tank where it would be violently agitated and mixed with»bubbles,
" This bypass flow from the high pressure region at the impeller discharge into
the low pressure region in the expansion tank would be returned to the impeller
inlet through holes in the periphery of the swirl chamber under the expansion
tank. It should be noted that the resistance of the main fuel system was simu-
lated in the model of Fig: A.l by a set of twelve l/h-in.-dia orifices in the ‘
lower of the two disks boundihg the annulus into which the.impeller discharged.
Thus the main stream flowed directly from the impeller discharge through these |
orifices back to the pump inlet while the bypass flow through the expansion
" tank followed a more devious route. A slinger was mounted on the impeller
shaft :just under the roof of the expansion tank to prevent fuel from entering
the clearance between the shaft and the heat dam between the expansion tank and
- the gas seal. |
When the model was built and tested, it was found that it performed
substantially as desired except that the swirl velocity in the expansion tank
was not adequate to'centrifoge.the bubbles out of the f£luid returning from
the expansion tank to the impeller,inlet. Tangential jets in'the floor of the
swirl chamber were drilled from the high pressure reglon at the pump discharge
to give a higher swirl velocity in an effort to get better bubble separation.
No improvement was obtained. Two features of the pump did perform surprisingly
well, however. The slinger was eminently effective in preventing the entrance
of fluid into the region between the slinger and the seal, and the agitation
and bubble entrainment of the fuel in the expansion tank gave good pramise of

- -106-

 
 

=01~

 

  

.
’ o . Omjvu

  
 
   
  

PLYWOOD

~SLINGER

 

GLASS POT

   

ORIFICES TO SIMULATE. EXTERNAL SYSTEM (2 EQUALLY
SPAGED. - '- ’ ] | |

—— . m— i — pm— e mmmaniss e — cv—"
S e e— e m—— e — A e — — — ——— —— o— g ——
- —— e — it e

  

RUBBER SEAL

 

FIG. A.1- PRELIMINARY MODEL NO.1 OF AN ATTITUDE - STABLE XENON REMOVAL SYSTEM
BB

removing the xenon. _

The results of this amalytical work clearly showed the need for a centri-
fuge operating with at least as high a.tip'speed as the imreller; On'this basis
the impeller was modified as indicated in Fig. A.2 to include a centrifuge cup
on its re#r face. Holes were drilled in the periphery of the cup so that fluid
would’discharge'into the swirl chamber. This arrangement was found to be much
superior to the first model, but it proved difficult to get fluid to flow into
the centrifuge cup at speeds above about 1200 rpm. Various scoops were tried,
including airfoil-shaped radial structs. The latter were hollow with inlet .
scoopé in their leading edges near the outer periphery so that ram pressure from
swirl invthe expansion tank vould force fluid through the strut and into the
centrifuge cup. The best of these scoops workedlfairly well up to about 1400 rpm,
and would give positive de-geration at speeds fram about 800 to 1400 rpm.

A careful re-exemination of the system led to the decision to make several
major changes. Since the first model had deteriorated considerably in the course
of the several modifications, it was decidedlthat a neW‘fiodel should be built as
shown in Fig. A.3. The baffle or diaphragm at the top of the centrifuge cfip was
cmitted and the fluid allowed to discharge from the centrifuge through 3/16-in.
holes into the regions behind the impeller vane tips. It was expected that
cmitting the diaphragm would increase the liquid level range in the expansion
tank for which it might be hoped that the pump could operate. The centrifuge
cup was made as long as possible in the space available to minimize the radial
through=flow:véloc¢ity. 'A new type of scoop was designed to fit in the bottom
of the expansion tank to lift the fiuid through the annulus between the tank
wall and an internal liner, and deliver it to the centrifuge. This arrange-
ment was found to be. satisfactory in every respect except that, at high speeds,
if the liquid level dropped so that the free surfacé of the expansion tank
vortex dropped below the top of the centrifuge cup, the fluid would be thrown
off the sheel intermittently and long streamers of air would extend down to and
into thevlabyrinth seal gt the top of the impeller and severe bubble entrain-

ment in the main stream would result.

-108-
 

' Separating the expansion tank from the pump impelier seemed;to bé a
pramising'arrangement if sufficient'agitatibn could be produced by jet action
to take care of both xenon removal apd pumping the bypass flow from the expan
81on tank to the centrlfuges. A c1rcular tank between the two putps as shown
in Fig. 3.10 was designed to accomplish both functions. Tangentlal Jets in the
bottam‘of the tank were de31gned to allow fluid to discharge ffom the'high‘
pressure region at the pump discharge into the expansion tank, which would run
at essentially’pump impeller inlet pressure. The large scale swirl induced in
the expansion tank could be made fo entrain gas bubbles through the use of
aspirator tubes. The ram.pressuré_on'tangential7scoops in the outer walls should
servé-to pump fluid through thé scoops to the centrifuge chamber. |

-109-
ORNL-LR-DWG, 23967

 
    
 

PLYWOOD

SLINGER '

   

CENTRIFUGE
CUP

 
 

 

   

 

LASS POT

   

ORIFICES TO SIMULATE EXTERNAL SYSTEM 12 EQUALLY
SPAGED. .

   

RUBBER SEAL

 
 

FIG. A.2- PRELIMINARY MODEL NO.2 OF AN ATTITUDE-STABLE XENON REMOVAL SYSTEM
ORML.LR-DWG. 239468

 

=LLL-

 

FIG. A.3- PRELIMINARY MODEL NO.3 OF AN ATTITUDE- STABLE XENON REMOVAL SYSTEM
Appendix B

CENTRIFUGE THEORY AND CALCULATIONS
 
 
 

CENTRIFUGE TEEORY AND CALCULATIONS

The centrifuges serre two purposes essential to the sys‘bem° First, they
remove the mixture of helium and enert fission gases from the fuel before it
-zreenters the main fuel system; and secondly, they serve as pumps to return the
fuel to the main system under pressure '

The abllity of the centrlfuges to remove the ges bubbles from the fuel’
depends primarily upon four factors: the speed of rotatlon of the centrifuges,
the radial dlstance to the p01nt of dlscharge, the velocity of radial flow
'through the centrifuges, and the size of the bubble to be removed. |

The relatlonshlp between these. factors may be shown as ‘follows:

The drag force on an obJect W1th motion relative to a fluid ‘is given

by the equatlon h o | :
| pve |
Fd = _cd_E-é—A
where -
F, = drag force (1b)
Cd = coefficient of drag

f = density of the fluid (llkft 3)

- V. = velocity of the obJect relatlve to the fluld

(ftcsec‘;) | _ _
g = gravitational acceleration (ft-seofe)_
A = area of object projected to the flow (fte).

The centrlpetal (flotatlon) force on a bubble in a rotatlng fluld (assuming

dens1ty of obJect is negligible) is given by the egquation

k3 w9x
Fc = 3 r f’
where S ‘
= centripetal force (Ib)
= bubble radius (ft)

. . 1
',{-Q"-r -(Sec )

= radial dlstancj from center of rotation to
center of bubble (fi).

-115-

angulargye

%

   

< = R
i
 

For the relative motion between the centrifuge cup and the bubble to be

zero, the two forces must be equal.

_»Cd_‘zg__j_ A= ll/sflrp-—*‘- and V = 3C

The value of C; is a function of Reynold's Number* and is as shown 1n Fig.

(10) a

B.l. Figs. B.2 and B.3 are plots of radial liquid velocity versus bubble
radius for..zero relative motion between the bubble and the centrifuge cup. The
radial distance to the bubble is assumed to be 2~ 7/8 in. in each case.

The above analys1s indicates several desirable des1gn features, namely, a
large diameter centrifuge to increase the centripetal force on the bubble and
also reduce the radial velocity, a tall cup to reduce the radial velocity and
a high speed to increase the centripetal force on a bubble. The diameter and
height of the cup are limited by shielding considerations and the speed is
limited by the main fuel pump design.

Another method of increasing the effectiveness of the centrifuge is to
increase the area of the holes that feed fuel back to the main system. This
reduces the radial velocity of the fluid at the periphery where the centrifuge
is most effective. This also reduces the tendency for the radial flow through

the holes to dlsrupt the.pressure field set up_byrthe rotating.liquid.

It has been calculated that any bubble smaller than approximately 0.003 in.
radius will passfout of the ART centrifuge if it reaches the exit holes when
operating at 2700 rpm on Fuel 30 (see Fig. B.2). It is probable that there
exists a point. at & smaller radius and lower radial velocity where bubbles
smaller than 0.003 in. radius will be screened. The- number of these small
bubbles which enter the system is a function of the surface tension of the
liquid. With clear water the number of bubbles entering the system was very
small,. By adding a small amount of soap, thus decreaSing the surface tension,
the system became very cloudy from the air bubbles pa551ng through the centri-
fuges. - All of the fuels presently being ‘considered have a surface tension ap-

proximately twice that of water.

 

 
The second function of the centfifuge is to déiiver‘the-sfripped fuel to
thé system under pressure. The preSSure delivered-by\thé centrifuge is a"
fuhétion of the outside diameter of the Cup_and either\the,diameter of the
.free surface when'thecgpis not.running full'Or'the inside diameter of the.
éentrifuge top when the cup is flooded. The pressure buildup.in a rotating
fluid with no flow is identicel to that of & radial vane pump. As the thorough
flow increases, there will be some'Slippage between the cup and the fluid;-giving
a pumping characteristic similar to that of a pump with backward swept vanes.
There will also be'a’pressure drop across the outlets frbfi the éentrifuge de-

pending on the through flow and hole sizes.

 

-117-
-gi 1-

Drag Coefficient

Reynolds Number

FIG. B.1

UNCLASSIF IED
ORNL-LR-DWG. 23959

8 1
1

CORRELATION OF IRAG
COEFFICIENT AND REYNOLDS
NUMBER FOR BUBELES IN WATER

 
=slL-

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

‘Bxit Centrifuge Velocity (ft/sec)

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

FIG, B.2

 

 

 

BUBBLE RADIUS (in.)

o
_ o '

1/ \ F\lel-ll«h
-'-'-'--_.______; Fuel—_‘ifi

- Fuel-]

'\\ 4_—______—___________4—-——"""'— / Water

/’
—\ -__-__-—__—____‘
/
lé//
é

5,00 0.0 5.01 0,01 5,020 .02k 0,078 5.032

 

EFFECTS OF-BUBBLE RADIUS AND EXIT VELOCITY ON CENTRIFUGE EFFECTIVENESS AT 2700 rpm WITH A CUP RADIUS OF 2-7/8 IN,

 
-0z1-

Exit Centrifuge Velocity (ft/sec)

ORNL-L R-DWG.

23%1

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

- e —— ____—-__“-_'—'_‘—-u—_ Fuel
~ : Fuel-3)
Ea— ‘ Fuel-1l
'—-—._______________ ;
Water
‘7 ——__--_'_'—-—-_.__ _/
A /
A
/ /
74
/ g /
0.00k 0.008 0.012 0,016 0.020 0,02l 0,028 0.032
BUBBLE RADIUS (in.)
FIG, B,3 EFFECTS OF BUBBLE RADIUS AND EXIT VELOCITY ON CENTRIFUGE EFFECTIVENESS AT 1350 rpm WITH A CUP RADIUS OF 2-7/8 IN,
Appendix C

* DYNAMIC SEALS
DYNAMIC SEALS

In the X=-R system seals'are required in two critical locations, namely,
on the shaft to prevent‘fuel“léakage'up into the bearing housing and in the
region of the pump to prevent high pressure fuel from leaking around the
centrifuges. In the césg'of’the’uppér'Slinger séal, no leakage can be tole-
rated, while in the ‘centrifuge seal a small amount of leakage, properly directed,
does not seriously affect the system. The centrifuge seal is essential to the
system for two reasons: 1) to determine the pressure in the main fuel circuit,
and 2) to control the amount of leskage. "All fuel which passes through the.
centrifuge seal is recirculated through the centrifuge, thus increasing the
radial velocity and consequently ‘the size of the bUbblés-which'will escape to
the main fuel circuit. o - - =

A very similar sealing aerrangement was used for each location. The shaft
seal used was a slinger attached to- the shaft (see Fig. 1.2). The potential
delivery~pres$ure of this slinger was greater than the pressure of the fuel in
the region of the seal. As the theoretical pressure rise across the seal was
greater than its discharge pressure, a gas pocket exists around the shaft.

In the case of the centrifuge seals, radial vanes were attached to the top
of the centrifuge cups (see Fig. 3.8). For this arrangement to stop all leak-
age, the discharge pressure from the seals must exactly match the pressure from
the centrifuges. This condition would be impossible to maintain in operation.
In the actual system a radial vane impeller was designed to deliver a pressure
slightly less than that from the centrifuges. The leakage around the centri-
fuge cups will then be so directed that no fuel bypasses the centrifuges. As
the head potential of the seal and the centrifugecare very close, the leakagé
is small and the additional flow through the centrifuge is held to an accept-
“able amount. Figure C.1l shows the head delivered as a function of vane length
or cup diameter for various speeds. The net head delivered may be determined
by subtracting the pressure corresponding to the inéide diemeter from that

for the outside diameter.

_123-

 
 

Flgures C.2 through C.5 give the amount of leakage through the centri-
fuge seals as a function of speed and the difference in pressure delivered
by the centrifuges and seals for several axial clearances. These axial .
clearances are the distance from the top of the seal vanes and the stationary
plate above (see Fig. 3.8). | .

-As the speed increases, the difference in pressure delivered by the centrl-
‘fuges and the seal impellers will increase and the flow delivered by the centri-
fuges will also increase. In order- to reduce the bypass leakage caused'by the
larger pressure differentiasl, it is necessary that the holes in the centrifuge
cups be slightly restrictive. This restriction will also idd stability to the
system by enabling it to sense and counteract any variations in leskage.

. It was found experimentally that the ability of the system to screefi gas
bubbles depended upon the precision with which the centrifuge seal impellers
were'fiade. 'One unit in which the blades were glued in place passed bubbles
very badly. Upon replacing these seals with new ones in which the vanes were
milled into the top of the centrifuge cups, the system operated satisfactorily.

Figure C.G_gives head and speed characteristics of a seal at no flow

conditions.

-124. -

 
UNCLASSIFED
-L 239;.'2

e
7

FIG. C.1 HEAD CHARACTERISTICS OF A FLUTD e
ROTATING AS A SOLID BODY

2000
o
2250
8

2500
2750

—
‘0
- @
.
-
.8
o
—
1N
B

 

100
SEAL LEAKAGE X 100 (ft3/sec)

3.k

3.0

2.6

2.2

1.8

. 006

0.2

UNCLASSIFIED

ORNL-LR-DWG. 27973

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

SEAL SPEED (rpm)
FIG C.2 LEAKAGE CHARACTERISTICS

~126-

\ ‘ -S¢al Dimensions-- 5 1/k in. 0.D.,
' 2 7/8 in. I.D.,
®
o
\
e
\
* o
| o
ft (Seal Prlessure Drop)
.
b \
.\.__\.
\.‘\
.-_1;5;-.-=
1.15 |ft
® o
— ot @ o
1500 1700 1900 2100 2300

OF A DYNAMIC SEAL WITH AN AXTAL CLEARANCE OF 0,050 in,
SEAL LEAKAGE x 100 (ft.5/sec.)

ORNL-LR-QWG, 23974

UNCL ASSIFIED

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

.'SEALVS?EED (rpm)

AN AXJAL CLEARANCE OF 0.128 IN.

-127-

3J+\\\ ‘ , .
“ ) "\ . Seal Dimensjons: -5 l/h in. OD .
N\ | " 27/8 in. ID
\\\ - 1/2 in{ High
300 ‘ \ -
4.61 pt.
®
*
2.6 . -
o
2.2 \
'\ ’
\ v.
1.8 |
L.h 5
2.3 ft. (Sedl Pressure Prop)
.
10 ‘-:.-"'":--..._.__ ¢ .
. @ - 4
0.6 ——
B —o : - L <
0.2 .
- 1300 1500 1700 | 1900 . 2100 . 2300

" FIG. C.3 ‘IEAKAGE CHARACTERISTICS OF A DYNAMIC SEAL WITH
3.4

3.0

2.6

o
.
N

'—J
.
&

SEAL LFAKAGE X 100 (ft/sec)
|_l
™

1.0

0.6

0».2

UNCL ASSIF [ED

ORNL-LR-DWG. 23973

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Sehl Dimensiong-< 5 1/k in} 0.D.,
2 7/8 in} I.D.,
1/2 in} high
. .
\\\ L.61 ft
*
- N
Ao
° °
\. \ |
¢ . _ ~ ® .
\ \.
\
K\LB £t (Beal Pressure Drop)
5.““‘=-_
1\ O ——
° \.&ft ’
) —~———|_
.o -N-.-—._—
1300 1500 1700 1900 2100 2300

SEAL SPEED (rpm)
FIG C.L4 LEAK CHARACTERISTICS

OF A DYNAMIC SEAL WITH AN AXTAL CLEARANCE OF 0.2L48 in.

-128-

 

 
SEAT, LEAKAGE x 100 (£t /sec.)

UNCLASSIFIED

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

SEAL SPEED (rpm)

AN AXTAL CLEARANCE OF 0.481 IN.

-129-

.\
3.6 1 ‘ —
\ Seal Dimensjons: 5 1/4|in. OD
o e ] - | 27/8/in. m
. \\\\\E:6l rt) oo 1/2 i%,'High |
| ‘\\\\\: Sl -
2,4 }— N
\\ .
.6\
* o | |
® 2.3 £t. (Seal Pregsure Drop)
® i '
\
1.2 ,
. ‘ o | "
_ S ———
o.8j:\‘\‘
| T~
1.15 tf. .
o
0.k o T
1300 1500 1700 . 1900 2100 2300

FIG. C.5 LEAKAGE CHARACTERISTICS OF A DYNAMIC SEAL WITH

 
UNCL ASSIFIED
ORNL-LR-DWG. 23977

Seal Dimensions-- 5 1/l in. 0.D., 2 7/8 in. I.D., 1/2 in, high
Axial Clearances-- 0.957, 0.837,_0.71h, 0.481, 0.2h6, 0.128, 0.050 (ip.)

P~
9
>
i
+
4
p
5
|
3
w0

w—- Calculated
@ Experimental’

 

7 8 ©

SEAI SPEED (rpm)

FIG C.6 ZERO LEAKAGE CHARACTERISTICS
OF A DYNAMIC SEAL

-130-
| | B Appendix D
CALCULATIONS FOR THE ART FUEL PUMP
. IMPELIER REDESIGN -
DRy
g
R
CALCULATIONS FOR THE ART FUEL PUMP IMPELLER REDESIGN

Assuming . no prerotation, Fig. D.1 is the inlet velocity diagram for the

impeller at the outermost inlet radius from Wthh it can be seen that

 

2 2 '
W= U+ Ca (1)
where
. . . D, = Impelier blade diameter at
Elg. D.1 - Pump Inlet Velocity Diagram } 'the outermost. inlet edge -
‘N = Impeller speed, revolutions

per sec.

Referring to Fig. D.2, assume the fluid streamlines entering the blades..
curve in an axial plane with radii, R, which is a funetion of the distance from

the shaft centerline, r, according to the'fo;lowing reiationship
R = a +br - - (2)
where a and b are cohstants evaluated'as_followsf’

0.25 £t

when r, = 0.09637 Tt R, =
r, = 0.15ft Ry, = 0.08333 £t
b = -3.108'and a = 0.5495
or R = 0.5495 - 3.108 r - (2a)

The blade leading edges lie on the surface of a cone having a 28 deg half angle.
This surface is approximately perpendicular to the throughput component of the
inlet flow, Ca, therefore h |

A = %%E%%%Tfl;' A = area perpendieular (3)
& to throughput com-

ponent of flow -
As the fluid turns the corner from the inlet eye 1nto the 1mpeller blading, it

follows the law of constant moment of momentum, i. e.,‘
Ca = where K = a constant (L)

(ka)

or Ca =

 
and from the law of continuity

" Qp = | Caan . o (5)

Qp = Total flow into impeller
. (includes leakage)

 

r .
27K 2 T :
QT T sin 28 deg - j a + br ar , | (53)
rll . :
5, : | Ty - T ~ a '+ br, | -
Qp “sin 28 deg b 2 T a+br,

 

 

 

| Gy sin 20 deg 1 | j\ (6a)
or K = ——————————
2 r. -r : a + br
- 271 a 1
b + '—é In a + br
; | _ P S | o
Substituting in equation (6a) K = 2.405
- 2.405 -
and  Ca =g 5495 =3 o6z L - Tt
 28.86 .
r in

 

or Ca_” 6’59h - 3.1007

Fig. D.3 is a plot of Ca as a function of r which is the velocity profile enter-
ing the impeller blades. From Flg D.3 it can be seen that at r = 1.8 in. .y
= 28.86 ft/sec. ‘

U1‘.= nDlN = %26 # 45 = h2?39'ft/sec N assume to Ee'hB rpé
wi = Ca’ + ui = 28.86° + li2.39°
. - |
Wy = 2629,8
wi"fi 51.28 £t/sec
o a1Ca o, -128.86 _ ., -l .
?l' = tan 'U—l -= ‘tan m —-A! ta.n. 0.6808

{:faf\-‘ P‘n[‘"d'vq anp e
SIS
13‘—

  

&3
~GE1-

ORNL-LR-DWG. !J'WH

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

7 R
- 1.0 1.0 -
N,
l
2,875 &
R ‘
2.3619 = 2.75 2.79|
1.15658 L3R f Y
i
Fluid Passage
A 2 - 26.5
2,875 E / :
s |
o - 3
A
o 1.0
7 B -
g .
‘u <
n
. ., .
Blade Trace 4, - 26,5° ,E .
\ Five Bladed Impeller /
. . A1.0
At
0 0.5 1.0 1.5 2.0 2.5
' Inches Circumferentially A
Fluid Streamline glong Shroud
i | ] | B
FIG, D,2 REDESIGNED ART FUEL PUMP BLADE TRACE AND FLUID PASSAGE LAYOUT

 
AXTAL VELOCITY, Ca (ft/sec)

© ORNL.LR-DWG. 239%9

 

 

L0

 

30

 

 

 

 

N

 

 

 

 

 

 

 

 

 

 

 

1.0 - 1.2 1. 1.6
RADIUS, r (in.)

FIG D.3 VELOCITY PROFILE
ENTERING THE ART FUEL PUMP IMPELLER

=136~

 

1.8

]

ek

8 2""‘&’”‘

oy
d
 

34,25 deg ag3h.deg

Py =
‘Bmv = 24,1 deg AJQh'deg'
By = 19.32 deg ~19 deg - -

‘In order:forithe impeller blade to be aligned-with the fluid, the angle
formed by the 1ntersect10n of the blade mean line and the tangent.of its termi-
natlng circle whose center is common with the shaft centerline must equal the
fluid angle. For ease¢of fabrication the trace of the blade:mean line on the
hub-firbjected on a plane perpendicular to the shaft centerline was made. circu-

“lar, - Due to the fluid fiassage curvature the angle that the flow sees is not’
-the’same angle ‘as:.that seen in the plane perpendicular'to;the-shaft centerline.

A-correction is made.for this deviation with the following equation:

tan fi cos 7 = tanB' 7- S _ | _._. (N
where -
B = angle fluid sees
"y = angle between the tangent to the fluid passage
- "~ and the plane perpendicular to the shaft center-
line ,
. B! = angle seen in the plane perpendicular to shaft
centerline.

To allow for contraction through the blades a Bi of 20 deg was used for
the blade entrance angle on the hub. From Fig..D02‘7 was measured and found
to be 33.17 deg. | ‘

tan 20 deg ces_33tl7 deg = tanB{

B,' = 16.94 deg

From the experimental work, the blade. 52 was determlned as 26.5 deg° Nb cor~
rection is requlred for the fluid passage curvature at the discharge since for
ithls 1mpeller the tangent of ‘the passage is parallel to the plane perpendlcular
to the shaft centerllne. The radius of therhub'tfaee is calculated from the

e uation a2 a2 ‘ ' ‘

q o | R2 - R_i K o (8)
2(R, cosB,-R; cosBiAHEE

-137-

 
 

Substituting in equation (8),f3 is calculated to be 2.3619 in. The locus of
the centers of curvature of the blades was found from the law of c051nes to lie
on a 2.600 in. diameter c1rcle. The hub blade trace is shown in Fig. D.Z2.

To.align the fluid along the shroud with-the’blade, the streamline of the
fluid along the shroud is assumed to follow a portion of a two to one ellipse.
‘The relative entrance angle of this streamline is equal to Bl and the relative
exit angle equal to B2 The ordlnate of this partial ellipse represents the
arc length of the fluid passage at the shroud in a plane parallel to the pump
centerline, and the abscissa represents the circumferential .distance along the
shroud blade trace in a plarne perpendicular to the shaft centerline. The arc
length of the partial ellipse is the arc length of the blade trace in the plane
perpendicular to'the'shaft-centerline. The fluid passage arc length is obtained
from the pump layout as shown in Fig. D.2. The angles Bl and Bg are determined
by W and U. The tangents of Bl and 52 are the tangents of the ellipse at the
blade entrance and exit. With these values, “the portlon of the ellipse between
Bl and 62 is defined as is shown in Fig. D.2. |

The blade trace on the shroud is constructed from the partial ellipse .and
the shroud contour in a plane'parallel'to the shaft centerline. This construc-
tion is done by plottlng on-the plane perpendicular to the shaft. centerllne
concentric circles (see Fig. D.4b) whose radii are the distance from the shaft
centerline to a set of arbltrarlly chosen points. on the shroud contour (see
Fig. D. ha). o . ‘

~ Starting at the dlscharge end of the blade, measure a chord length (C) ob-

-tained from Fig. D.4c on an average circle as shown in Fig. D. 4, ‘and connect its
termination to the shaft center. The point of intersectlon of this radian and
the next smaller concentric circle lies on the shroud blade trace. This process
is continued as shown in Fig. D.4b until the smallesteconcentric circle is
reached. The points of intersection obtained determine the blade trace as
shown in Fig. D.2.. | |

‘The shape of the blade blank is determined from the blade trace and the
fluid passage, Figs. D.5a and D.5b. Plane eect;ons perpendicular.to the refer-
ence radius are taken through the blade as shown in fig. D.Sa, as sections A,

B, C. etc. These sections are construoted as shown in_Fig.~D.5o with the

 
distance "a" obtained from Fig. D.5a and the distance- "d" obtained from Fig..
D.5b. From the plane ‘sections shown in Flgs.-D 5a and D. 5c, sectlons -D, E,

¥, etc., parallel-to the reference perpendicular plane are constructed‘ The
center“Of curvature andfifadii of these sections. is-also'the center;of curva-..
ture and radll of the blade blank The blade mean: llne and blank is laid out
from these radii as shown in Fig:.D.5d. Since this 1mpeller 1s of welded con-
structlon, it was desirous to make the blades and fluld passage parts of sur-

faces of revolutlon.

-139-
“-{{ et -::'-‘ i.:(“.‘:‘b:
Erdye S b, UNCL ASSIFIED
R ORML-LR-BWG. 23980

 

 

 

   

 

 

 

 

Fig. D.4a Plane Paralle/ to Shaft Centerline

Fig.D.4b Plane Perpendicular to Shaft Centerline

 

 

 

 

 

 

 

9 fe
£8
e
@ 7
|
o 64
< e | ¢
Q 5 T . . :
S 4
@ E 3
v
o 3 4
Q.
'1-_7:: 21 /9, A7 ;"1‘”??1 s,
— RS e X 3 3
w | 1 BT

 

 

Circumferential Distance

Fig.0.4¢ Fluid Streamline Along Shroud

-]49_
=lyl=

Section Parallel to Reference
Perpendicular Plane

Center of Biade
Curvature

Distance From
Reference

Perpendicular A
Plane

 

 

 

 

F‘ig, D.5¢ Distance From Tip Along
Reference Radius

    
 

—a

 

Fig. D.5a Blade Troce

Referenpe
Radius

Reference

Perpendicuiar |

Plane

Qo

UNCL ASSIFED
ORNL-LR-DWG. 23981

    
  

Shroud
Mean Line

Leading Edge

 

 

Shaft

Fig. D.5b Fluid Passage

 

 

 

Fig. D.5d Blade Blank
Appendix E

BYPASS FLOW CALCULATIONS AND EXPANSION TANK HEATING
 

Lo ok g TS
,--:;3;:53;;;%x‘a-xi’sr?‘%-.:;j%efl;,
 

BYPASS FLOW CALCULATIONS AND EXPANSION TANK HEATING

In examining the experimental data it was found that the accuracy

required to closely determine the bypass flow rate directly was greater
than the accuracy of the data that could be obtained. zThis is espécially
true at the 1ower expansion tank 1iquid levels where the system (pressure)
fluctuatlons were quite large. :

 An equation was derived to determine the pressure losses in the bypass
circuit, that is, from the pump suction to the expension tank. This circuit
was then tested separately with water to determine the validlty of the de-
rived equatlon. In the water tests the shaft and spool piece were not rotat-
ing so that the test was only a check of the assumed loss coefficients. It -
was found that the predicted value was within 4% of the experimental value.
Also a cheéck of the experimentally determined value of bypass flow at high
expansion tank liguid levels agreed very well with the calculated values (see
Fig; 2. 28)'"'The total loss in this circuit was comprised of the following:
1) the loss entering the pump shaft, 2) the loss entering the four radial
shaft holes, '3) the loss enterlng the four holes in the spool piece, and
) the loss leaving the spool piece. The driving force or pressure differ-
entlal was taken as the difference between the pump ‘suction and the eXpan51on
tank pressure plus the head developed in the radlal ‘holes which were assumed
to be acting as radial impellers. '

The eqpatlon used is as fbllow5°

 

e .2
'43 = j"O..5(ll) ;g— + 5. 5(12) B2 +‘O,23(ll)';§— +l.ve ;
where: , : .
v, = velocity in the hole up shaft, ft/sep,
ve = velocity at exit of radial spool holes, ft/sec,
AP = _(PS -:Pfie) + QP'(pgad developed by radial holes);‘ff.

~145-
 

Bl
Substituting in the value of the areas

o - 19u (22,

where

Q = bypass ‘flow 1n gpm

Calculated bypass flow to the fuel expan51on tank through each pump shaft
is shown in Fig. E.l.

- The temperature of the fuel in the expansion tank will depend upon four
factors: 1) the fuel volume in the tank, 2) the total recirculation rate
through the tank, 3) the power density of the fuel in the tank, and 4) the
temperature of the fuel entering. The temperature of the fuel entering the
tank will be approximately that of the fuel leaving the heat exchangers and |
will depend on the manner in which the reactor will operate. Flgure E.2 shofis
the entfance, exit, and equatorial temperatures of the fuel in the,core.- These
temperatures ere based on an operating procedure where the system is started
from an isothermal temperature of 1200°F. As the power is increased, fhe cofe
equator temperature is held constant until the inlet NaK temperature drops to
1070°F; at higher operating power the NaK temperature is held at 1070°F while
the fuel temperature Increases. The power shown in Figs. E.2, E.3, and E.h
is that which is received by the fuel and not the total nuclear power. The
design power to fuel is about 55 Mw. ' '

Figure E.3 gives the expansion tank 11qu1d level as a function of the
power to the fuel. 1In this figure it is assumed that the reactor is filled
at 1200°F to a level of 1/2 to 3/4 in. The final level was based on the mean
fuel.tempefature and is.approximately 25°F less than that of the core equator
temperature at 55 Mw of power to the fuel. Thé élight rise in liquid level
from O to 20 Mw is due:to a shrinkage in theireactor voiume. It has been .

3

estimatedclf} )thé.,t the volume decreases by about 70 in.” from assembly tenipera—
ture to the design operating conditions. 'ifi'Fig. E.5 this was assumed to be
linear fiith pOWET | | ' |

The mixed mean fuel tempéiature‘in the ekfiansion tank is shown in
Fige. E.bt. In determining th;s temperature the initial f£ill level was assumed -

-146-

 

 
=Lyl-

ART Fuel Pump Bypess Flow (%) 1008 = 17.6 gpm

65

60

55

50 |

L5

   

 
  

3300 rpm
Pump .Speed

 

 

 

 

 

2400 rpm

 

700 gpm

 
 
  

 

 

 

 

 

   

550 gpm

. FIG, E.1  CALCULATED ART FUEL

PUMP BYPASS FLOW CHARACTERISTICS

 

 

3-1/2 in.

2 in..

N

1/2 in, Expansion Tank Liquid Level

 

 

 
1700

1600

1500

11,00

TEMPERATURE  (°F)
=
W
8

=
no
o .
o

A

1100

1000

 

ORNL-LR-DWG. 29983

 

 

 

Fuel to

/ |

 

Hes?

1t Exchangers

Fuel af
Core Equatd

pd

 

 

'/

Fuel Fy

Heat Exclw/
] -

\

 

_—

/ .

 

NaK {

.o Heat Exchd

inger
g ~.

 

 

 

 

 

 

 

 

 

 

10

20

30 - Lo

"POWER TO FUEL (Mw

-148-

. FIG E-2 ART REACTOR FUEL AND NAK TEMPERATURES
AT A CONSTANT PUMP FLOW OF 6L5 gpm

50 60
'(in;)

EXPANSION TANK LILUID LEVEL

 

 

 

 

 

 

 

/ :

3/l in. Fill Le

2 in. Fill Level

 

 

 

 

 

 

 

 

 

 

 

 

 

Y 16~ 20 30 ) £o

POWER TO FUEL (Mw)

FIG E.3 THE EFFECT OF REACTOR POWER LEVEL
ON THE ART EXPANSION TANK LTIQUID IEVEL

-149-

60

——

Lo

ORM- -OWG. 23984~ %
-051-

Mixed Mean Fuel Temperature in Expansion Tank (°F)

1375

1350

1325

1300

1275

1250

1225 -

1200

1175

ORMNL-L R-DWG, 23985

 

2100 rpm pump Speed

NOO rpm

&
.

had

~d

»
e

s

gy

 

2700 TP T
%
3300 rpm T

60 Mwi?

A

2

 

 

 

50 Mw

 

/)

 

Lo Mw (Reéctor Power)

 

 

 
 
 

30 Mw

 

FIG. E.4 MIXED MEAN TEMPERATURE OF FUEL
IN THE ART EXPANSION TANK AS A FUNCTION
OF REACTOR POWER TO FUEL AND PUMP SPEED
AT A CONSTANT FUEL FLOW OF 6i5 gpm PER

 

 

 

 
s

to be 5/8 in. at 1200°F and the bypass flow was taken from the calculated
value shown in Fig. 2.26. The.leakage around the island through the floor
of the expansion tank was calculated to be 4 gpm at 2700 rpm and is e
function of the pump speed. The temperature of the fuel entering the expan-
sion tank was obtained from Fig. E. i.' The heating rate in the tank is the
sum of the after-heat rate (11 w/cmelh) -and . the fissioniheat release

(10 /cmifl'l These heatlng rates are for 60 Mw of nuclear power or 55 Mw
of power to the fuel. The healing rate was assumed to-be linear with power.
The hlgh degree of turbulence in the fuel expan31on tank as revealed by high
speed movies taken in this region indicate that the mixed mean temperature

“will probably be the true fuel'temperature,'

-151-
 
   

gg’éffi

z?&@@

T
~ Appendix F
. 'RADIOACTIVITY IN THE ART FUEL
PUMP OIL SYSTEM
 
 
 

RADIOACTIVITY IN THE ART FUEL,PUMPIOIL SYSTEM

This section treats the transfer of radiocactive gases from the expansion
tank region of the ART reactor to the fuel pump lubricating oil system. A
report by W. K. Stair(lS) covers tests on the back transfer of helium against
argon in the fuel pump to give some indication of . the back transfer. of xenon
and krypton agalnst helium. Two numerlcal values from this report were used
for the calculatlons of this sectlon, i.e., an attenuatlon of lOu between the
concentration of radioactive gases in the expansion tank region and the con-
centratlon in the 0il catch ba31n, and a leakage rate of aas from the oil catch
basin to the 011 reservoir of 0.8 1n.3/day. In determlning the leak rate of
0.8 in. 3/day, it was assumed that the 0il was completely saturated with argon
at the beginning of the test run and that all leakage across the seal was col-
lected in the oil reservoir. The value of 10 for the attenuation in concen-
tration between the fuel expansion tank and the pump seal is & minimum. The
helium concentration in the argon buffer gas used in the experiment reported
by Stair was 10'”, hence the actual attenuation in\concentration betfieen the
tank and the seal mayihave'been:greater.

Because of the shertflhalf—lives of the fission:product gases the attenua-
tion .in the amount of activity between the fuel ekpansion tank and the oil
regervoir is greater than the attenuation in gas concentration. Thus the rate
at which the fission proancfi gases enter:the.oil system is of significant im-

portance.,
Method of Calculation

In the ART as well as in the above mentioned test, 500 liters per day of
buffer gas flows down through the clearance around the pump shaft into the ex-
pan51on tank and 50 liters per day Wlll be bled directly from the oil catch
basin. The attenuation of 10 represents the back transfer of helium against
the 500 liters per day of argon. This back leakage was determined from analy-
sis of the 50 liters per day bled from the catch basin. |

 

~155-
 

The equation representing the numbér (N) of nuclei of any nuclide in the

catch basin as a function of time (t) may be written

dN i
=S - (A, + A )N
where
Sb = source or leak rate into basin -
' AP = reciprocai of average dwell time in the catchA basin
Ag = decay constant of radioactive nuclide.

Assuming an attenuation of lOLAL and a purge rate of 3000 in.3/day (50 liters
per day) the value of S“b and AP may be calculated as follows:

assume X concentration in expansion tank and
o 10“)+ X concentration in catch basin

10 in.3 = volume of gas in catch basin

= leakage of X into catch basin
T X leakage from catch basin

8, = 10 'l‘ x 3000 in.3/day 3.47 x 10 6 3/sec of

X concentration
< ¥ (_expansmn ta.nk)

 

Ap T BE,L00 x 10

S, = 3.47 x lO X Volume (expansion tank)
= 3.86 x 10 -8 N/sec
3000 3.47 x 1073 sec™t

and N may be ex;presséd as

- Sb lil-e'-(APJr Ad)t:]
vos . xp*')tdz : .

 

-156-
 

.. The leak rate into the oil system from the'eatch basin is O.8-in.3/day
or 9.26 x 10"6 in.3/sec. To convert this rate to'nuclide/sec the constant

. 9.26 x 10'6 must be multiplied by the nuclide/in._3 in the catch basin. The
volume of the catch basin is 10 in.3~and the leak rate is: :
. ‘ 9 56 x 10 6 3/sec x Ngcatch~b251n)‘=_9456-x 10-7‘Nu:izde

10 in.

Nuclide Concentrations in Expansion Tank
In the calculations for the ART the Tollowing assumptlons are made:

Helium flow rate down shaft = 500 standard 11ters/day

Gas volume in expansion tank 90 1n.3 |

‘Entire reactor fuel volume passes through xenon re- |
moval system every 190 sec where all fission

gases are removed.

. Sr = rate of formation* of a particular nuclide in fuel.at
- 60 Mw = 186 x 1016 X the yield of that nuclide per fission
- | | dt fuel = - (_AS + ?\d)N

- Assuming complete removal of the,gaseous nuclide in the course of its transit

through the fuel expansion tank, then

reciprocal of average dwell time in the reactor’

As
s T 1%0“527"103
5. = 186 x 107 16 x yield
at equilibrium | -
- 16

186 x 107> x yield

 

N, . =

fuel 527 %2073 4 A,
Se = _seuree of fission gases into expansion tank = ;Kstuel
_ . . T S S ;
- _ta‘rfli .
' — - A
3t AgNpyer = (Ag + M )Neary

 

¥ There are approximately 3.1 g

 
W

A purge rate of helium 1nto expansion tank at tank temp. and pressure
. fi, . _ . - - gas volume in tank _

= 1;96 X 10"2 sec™T
at equilibrium o

Astuel'

Ytk T A WA

After the fission gases are separated from the fuel, their decay products
are then treated as gases. This assumption is conservative but the actual be-

havior of these nuclides is not known.

| Nuclide Concentrations in Oil System
The following equations give the total number of each nuclide in the oil

system feeding each pump:

 

 

 

 

 

 

 

 

 

0. N
o l --e 1
R S (X7 + A5
‘ | SALt T : | LT
2) n° = ¢ {L--e °) )% + AL )
2 A”B Ao | u2+hg TR +2) (k-+XT
oAt =Xt T _
+cCyfe T e Z) M
- - ALt T T
(3) n.° = ¢,c (1-e °) > ) )5 + [ ’\2 e
| - ) T.
. l+'11' L M 7\,2. NCL
"Ry T TRy ) TAp Ay | TARFA)
At =Agt (T .\
+ C,Cp (e 7; < 2 5 F1 4
T e ALt - ALt
. Ao :INl N C e '3 Ao
Ry A | TApFag [ 7% T =7y (3 A

 
(v)

Il

Ca

+

-+ )

.+.

P -A’-l-t A NT .
C (1-e ) ll' + [l + 3 ] 3
B 7\)_|_ (A)_fi') ) : | (7(‘)1'_+ AP) .(A3+2P),

[i + A2 o+ A 3 » ]
R ) (?ufi?\) (l+)) )

 

 

 

(A2+A ) (23+)P) (Ay+2a)
T
1

Az Ap A N 1
()4+A ) (A3+A ) (At ag) (Al+APS

 

(e

37 A
c.C DR + |1+ 2
W TR | TR Ty o)

o o |
Na | "Al‘ Ao A1

 

. 3 _ ,
(X o~ A3)] (A Ag) [1 TRt A) TG A Tagr )

A3 As ; AL ° Hksl. | Ao

T (7\2-23)(71”) T G g

- oy
| Ag 32 1N |
(31 A3) (g 7 )] vy

 

-A,t -A _ T
2" _ y° A3 N

| (e ) ] T2 | Ay
CA TR A R g (R Ay [ TR )

 

 

A ) G [ e
-2t - zt - | T
€,Cy (e 1te 2 Ag e N |
T T ay) Tag A Tagr i)

-159-
where
N = fiotai number of any one nuclide_ -
( AP = .purée:constant for oil catch basin = 3.&7_x 1073 sec L
)J = decay constant for nuclide_undef'cohsideration
C, = leakage constant to .ca.tcl‘l basin = 3.86 x 10~ (n“:i;dej)a S,
Cp = leakage conetant to oil system = 9.26 x 10 7(EE§%%99)
t - time (sec)
superé. 0 =-oil eystem
T = expansion tank
subs . 1 ; parent nuclide (fission gas)
2 = first daugptef
3 = se‘condA daughter
4L = +third daughter

To calculate the dose rate near the oil reservoirs all of the radioaetive
nuclides are assumed to be concentrated as a poiht souree in-the oil reservoirs.
In the calculationé it is assumed that the 3/8 in. thick steel wall of the
reservoir will give no attenuation for gamma, energies above 0.03 Mev, an attenu-
ation of 10 in the energy range of O. 01 to O. 03 Mev and 75 for energies below
0. Ol Mev. In cases where the exact percentages of the gamma, energies are not
known, the maximum possible percentages are used. The data for yields, ener-
gies, and percent gammas are taken from Blomeke (16). The data for dosages

are taken from Rockwell(l7)
CalculatiqnmResults

The results of the calculations are given in Table F.l. Columns 6 and 7
give the B and y energy release rates in the fuel expansion tank. These are
equal-to 15.7.and 0.1 Kw, respectively. Column 9 tabulates the disintegration

per sec of each nuclide and Column 10 converts the B energy release of these

';;607
   

disintegrations to Mev/sec. Columms 9 and 10 refer to the activity and é%ergy
release of only one of the two oil systems. Column 11 gives the y dose at a
distance of 1 ft from either’of the two fuel punmp o0il reservoirs. This ¥ dose
rate is about 30 mr/hr at a distance of 1 ft ‘and 1s inversely proportional to
_the square of the distance from these reserv01rs. : -

The approx1mate volume of oil in each pump lubricant system is 30 gal
Iflthe total B emitters in the_oil_are assumed egually dispersed throughout
the system, the dose to the oil is founduto‘be about 20 rads/hr. This dose for
500 hr of operationvat-GO erafter-equilibrium is reached amounts to lOh rads,
which is insignificant compared to,the.lO8 rads which the oil is capahle of ab-
sorbing Without serious damage. _ - o

If one distinguishes between the fission gases and their daughter products,
it is found that the fission gases account for about T0 to 75. per cent of the
total activity in the Oii system but only 25 per cent .of the y dose calculated

in Column 11.

-161-
10

8 9

Table F.1
1

6

2

L

3

CONCENTRATION AND ACTIVITY OF Xr, Xe AND DAUGHTER PRODUCTS IN EXPANSION TANK AND OIL SYSTEM FOR A HELIUM FLOW RATE OF 1000 LITERS/DAY

 

1l

2/
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SPTIoMN

 

 

 

SECRET

31.6

3753

87.68
(0.1 Kw)

9795
(15,7 Kw)

 

TOTAL
~ Appendix G

Xe”” POISONING
 

 

 
oy

Xe™ > POISONING

This final section iS'COncerned with the xenon poisoning in the ART.
In the calculations which follow the removal of xel35 by neutron capture
is neglected. As the neutron capture rate of removal is only about 5% of
the total removal rate, the calculations will not be greatiy affected. Also;
the results for poisoning will be comservative in that it would be smaller
were the burnup rate considered.

The derivation for the equilibrium xenon concentration in the fuel can
best be followed by referring to Fig. G.l. The total fuel flow through the
expan51on tank including the leakage around the island is about 23 gpm or
1450 cmB/sec.- The helium flow rate is treated as a parameter and is varied
fram 1000 to 5000 liters/day at standard temperature and pressure. The
solubility(la) of xenon in fuel is shown in Fig. G.2. - If the value of the
sparging efficiency is. defined as N | ‘

(gg gol)ln (gm mol)
cm
7 = 2 (l)

(*‘3"m moly o - (solubili‘ty), P,

 

where P is the absolute pmessure of the xenon in the expansion tank. ' Then

T—ELB’

-ill-_—s—o-"- 14-.3 x 10 Px

. \

 

-

Z = 6.23x10°p +L . (2)

For equilibrium conditions, ¥ must be equal to the rate of formation of
xenon minus its decay before reaching the sparger and is the rate at which
xenon enters the sParger and thns the expan51on tank This term is the
same as ‘the source to the expansion tank as defined in Appendix F, and is
shown in the last column of Table G. lf The value of P&h can be determined

-165-

L
 

from the total xenon nuclides in-the expansion tank. The value for the total
nuclides p;esént are taken for tfie case where the free gas volume.in the ex-
pansion tank is 90 in.3,'and the method of calculation is again the. same.as
Appefidix Fo _ |

The nuclides of each xXenon which is presént in any appreciable amount
is shown in Table G.l. - A total pressuré of 2 atm and a temperature 1300°F
were used to determine the partial pressure of xenon.-;The'paftial pressures
at helium Plow rates of from 1000 to 5000 standard liters per dey are given.
in Table G.l. - Substituting the values in Table G.l in Eq. (2) gives the
xenon concentration in the fuel. Figure G.3 gives the total xenon concen-
tration in the fuel for helium flow retes of 1000 and 5000 standard liters
per day. Also shown in Fig. G.3 is. the Xe 155 concentration in the fuel.
This was obtained by teking 14% of the total xenon (the percent shown in
Table G.1 is 13.63). o |

The Xel35 poisoning(lg) in the reactor is the number of thermal neutrons
absorbed by the Xe™®? to those absorbed by the fuel or

LR )
For a fuel with 5 mole % of UFL the U235 concentratlon is about . _

8.2 x 1020 atoms/cm3. The. ratio of the xenon-to-fuel cross section G~ fifF
Por a Maxwellien distribution at ART thermal is h 9 X lO3 (20) As approx1-
mately 22% of the xenon cross section is scattering, the value of 0* Ar is
multiplied by 0.78. Therefore '

P o= 466 x 1000 e, | (4)

The results of Eq. (h) are shown in Fig; G.4 (upper curve). This equation
is only for the poisonlng effect on the thermal neutrons.- In the ABT about
4O% of the neutrons reach thermal energy. The lower curve of Fig. G.4 shows
the poisoning based on the total neutrop den51ty of the reactor and is

assumed to be 40% of that for the thermal neutrons. This is a gross simpli-

fication of the actual case, but in the present reactor in which the purge

-166-

 
rate is high compared to the burn_up rate the ass_u_mption shofild be fairly
accurate. | | - | | ' |

The effect of the p01soning may be more easily understood by comparison
to the. temperature coefficient, which for the ART is expected to be about
2x 10 5/ F. Thus, & poisoning of 1 x 10 =5 1s equivalent to a temperature
change of 50 F. ‘ ' ’

-167-

 
Table G.1l

XENON CONCENTRATION IN EXPANSION TANK GAS VOLUME

 

 

Nuclide in Expansion. Tank x-lo‘;7 for Helium Flow Rates of

Decay Constant Source to Tank

 

 

 

Nuclide | ‘Sec—.l) n ti(x)'(sx}day litzgcs)?d.ay .ngg?aay 1i tle*gg‘/)day 1 t_zg(s)?day Nu'c.:jlfid.e/sec‘ % 10-16
xel?l 0 28.2 1.1 9.40 7.05 5.6 5.5%"
132 o 43.8 21.9 4.6  11.0 - - 8.76 - 8.58
133" 3.49 x 10°° 1'.5'2' , 0.76 . 051 0.38 o.3§ 0.30 -
?155 : 1.52 x 10'6 62_;_8 .. _31.l+ | - 20.9 157.7 | 13.0 12.3 -
@f 134 0 ™2 27.1° 2h.7 18.6 . 11».8' o 1k.5
135% 7.#0 x 107 | 13.7 6.95 .h.67 3.51 | 2.81 2.78
135 2.11 x 1077 62.5 3.0 20.7 15.5 124 12,2
136 0 | 60.9 56.5 20.3 15.2 S 12.2 - -_ | 1.9
137 2.9Tx 1000 3.7 17.5 11,9 ; 9.05 _‘7.29' 1%
138 6.79 x 10‘1,’ 49.9 . 25.3 17.0 12.8 | 10.2 - -10.1‘_-'
139 1.69 x 10°2 7.63 4.97 3.68 2.93 2.43 o 2.78
1l+o‘ h.33 x 1072 1.95 1.48 . 1.20 | 1.01 0.87 . : 1.23
Totals 1—5_9.80 222,56 149.56 - 112.73 90.70 T 89.59
Percent Xelo? S 1k 13.9 13.85 13.78 13.68 13.63
P 0.00kOk . 0.00205 0.00137 0.00104 0.000834 |

 

.
ORNL-LR-DWG. 23986

X liters/day He | |
an X IiterS/day of He

   

Wam mOIS/sec of Xe

 

SPARGER
~ AND f
_emd. EXPANSION

1450 /sec of fuel| TA NK

(Z-Y)gm Mol/eae of X

3
1450 “M7sec of fuel

Z gm m’ol‘/sec of Xe

 

 

 

FIG. G.I- SCHEMATIC FLOW THROUGH SPARGER AND
- EXPANSION TANK

169
)

gm nol
cm3 atm

SOLUBTLITY X 108 (

ORNL-.LR-DWG. 23987

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

1100

1300

FIG G.2

- 1500
TEMPERATIRE  (OF)

XENON SOLUBILITY IN FUEL-30

-170-

1700
. - . ORNL-LR-DWG. 23988

100
FIG G.3 XENON CONCENTRATION
AT EQUILIBRIUM IN FUEL-30 WITH A
POWER GENERATION OF 60 MEGAWATTS
&
Q
™~
0
£
3
(11}
b
O
—~
4
_ 10 1
o
&
§ _ 000 liters/day
5 7 Helium Bleed
5 000 liters/day
: Helium Bleed
%
' < ' 000 liters/day
1 EFFICIENCY * % Helium Bleed

 

-171-
    

100
L
e
2
=
o
N
. -
o
Q‘ .
=3 10
— .
b
AT
a
L'H]
b
1 Neutrons
FIG G.L Xel3® POISONING
IN FUEL-30 AT 60 MSGAWATTS POWER
AND A HELIUM BLEED
OF 1000 liters,day tal Neutrons
1

10 EFFICIENCY (7 ) % .. 100
-172- |
&

“ .

VACKNOWLEDGEMENTS

A pumber of people and groups played 1mportant roles in the de31gn
-and development of the ART fuel pump and xerion removal system. The authors'
are 1ndebted to the follow1ng people: ' ' ' '
' W. Lowen contributed to the design of the attitude- stable
xenon removal system.

E. T. O'Rourke was responsible for the assembly and functlon-.
ing of all eqplpment 1nvolved in the plastic models of the
xenon removal system.

¢. D. Whitman and J.J.W. Simon followed the assembly and
construction and subsequently operated the iron fuel pump
and the combined Inconel fuel pump and Xenon removal system
test loops, respectively. ,

We G. Cobb and L. Wilson were responsible for the layout of
the fuel pump and xenon removal system. '

The authors also wish to thank the members of. the Puwer Plant Engi- -
neering and Experimental Englneering Groups of the Aircraft Reactor Engl-
neering Division for their contributlons and assistance in the’ de51gn and
development of the ART fuel pump and xXenon removal system

1734
1.

2.

10.

11.

12.

13,
1h.

. 15.
16.

17.
18.

19.

. ORNL CF-54-12-52,

 

REFERENCES |
J. L. Meem, The Xenon Problem in the ART, ORNL CF-54-5-1 (May 3, 195L4).

S. I. Cohen et al., A Physical PTOPerty Summery for ANP Fluoride Mixtures,
ORNIrElBO (Aug. 23, 1956)

AREDbANP De51gn MEeting No. 55-h ORNL CF-55 1-209 (Jan. 26, 1955)
ANP Progect Drawing E50-J4U7-4, Rev. 8 (Dec. 9, 1954).
A. P. Fraas,ACFRE-Design and Development Handbook, Project No. 3.

H. Rouse, Fundamental Aspects of CaV1tat10n, Proceedings of the National
Conference ~ Industrlal Hydraulics, p. 31, 30-7 (19h7)

D. Thoma, Bericht zur Weltkraftkonferenz, London, 192h Ze Ver. ‘deut.
Ing. 79, »- 329 (1935). . e ’

D. Thama, Verbalten einexr Krelselpumpe beim Betrieb im Hohlzog Bereich,
Z. Ver. deut. Ing. 81, p. 972 (1937). o

 

' W. Lowen and G. Samuels, ART Xenon Removal System, unpubllshed (FEb. Ts

1955).

F. H. Garner and D. Hammerton, Circulation In51de Gas Bubbles, Chem. Eng.
Sci. 3, No.. 1 (Feb. 9, 195&) _ N

Flow of Fluids Through Valves, Fittings and Pipe, Crane Company Techni- -
cal Paper No. 409 (May, 1942).

J« B. Baker and J. W. Michel, Pressure Changes With‘thE‘FlOW'Of‘WHter
Through Tees, K 998 (Mar. 16, 1953).

 

L. A. Mann, personal communication.
A:. M. Perry, personal communication.

W. K. Stair, Back Transfer of Gas in ART Pumgs, Experlmental Report Task
No. Th18, Exp. Ho. 1 (Aug. T, 1956). _

Jd» O Blomeke, Nuclear ITo@erties of U235 Fission Products,

T. Rockwell III, Reactor"Shieldihg Design Manual, Fig. 2.1, p, 19.

 

' ORNL Special Progress Report - Program 4400, ORNL-2249, p. 11 (Dec., 1956).

S. Glasstone end,M{ C. Edlund, The  Elements of Nuclear Reactor Theory,
Ps» 335, D. Van Nostrand Company, New York, 1952,

 

A+ M. Perry, personal communicatiohgu