NBSIR 75-816

NAT'L INST. OF STAND , &, TECH mr

PUBLICATIONS

PERFORMANCE CHARACTERISTICS OF A LIQUID HELIUM PUMP

p. R. Ludtke

Cryogenics Division

Institute for Basic Standards

National Bureau of Standards

Boulder, Colorado 80302

July 1975

Final Report

Prepared for:

Air Force Aero Propulsion Laboratory

Air Force Systems Command

Wright-Patterson Air Force Base, Ohio 45433

U.S. DEPARTMENT OF COMMERCE, Rogers C. B. Morton, Secretary

John K. Tabor, Under Secretary

Dr. Betsy Ancke r-Johnson , Assistant Secretary for Science and Technology

^C>0 NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Acting Director

1175
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4. TITLE (-aodrJSMbtmf^ ^ ., „„-< .,.^-... w-... . ——

) 'performance characteristics of a liquid helium '] PUMP ,

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18. SUPPLEMENTARY NOTES

19. KEY WORDS (Continue on reverse side If necessary and Identify by block number)

Cavitation; cryogenic motors; induction motors; liquid helium pumps; motor

efficiency; net positive suction head; pump efficiency; pump performance.

V.

: 20. AJ^5^^f'^'-T (Continue on reverse side If necessary and Identify by block number'* Part A of this report presents performance data for a simple, preinduced,

single stage, centrifugal liquid helium pump powered by a close-coupled submersible

cryogenic induction motor. Data on pump efficiency and the motor efficiency are

given, and the effects of decreasing the pump leakage loss and removing the preinducer were also investigated.

Part B describes a study of the cavitation characteristics of the pump in

liquid helium I. The net positive suction head at cavitation of the pump was extensively investigated, and found to be near zero; and effects of removing the

inducer, changing the pump inlet geometry, and varying the helium temperature were also determined.

OD 1 jan'^3 1473 EDITION OF 1 NO^SfiS IS OBSOLETE „ , j X Unclassifxed

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Table of Contents

Page

Terminology iv

Introduction 1 A. Pump Performance and Efficiency 2 A. 1 Measurement of the Electric Motor Efficiency 2 A. 1.1 Description of the Motor Efficiency Test Apparatus 2 A. 1.2 Motor Efficiency Test Procedure 2 A. 1.3 Motor Efficiency Test Results 7 A. 2 Pump Performance Tests 7 A. 2.1 Pump Performance Test Apparatus 15

A. 2. 2 Pump Performance Test Procedure 15

A. 2. 3 Pump Performance Results 21

A. 2.4 Pump Efficiency Results 22

A. 3 Conclusion from Pump Performance Tests 31

B. Pump Cavitation Study 31

B. l Cavitation Study 31

B.2 Cavitation Test Procedure 34

B.3 Cavitation Test Results 34

B.4 Analysis of the Cavitation Data 45

B.5 Conclusions from the Cavitation Tests 49

References 50

List of Figures and Tables

Figure 1. Motor efficiency test apparatus 3 Figure 2. Photo, motor efficiency test apparatus 4 Figure 3. Photo, motor efficiency test apparatus , 5 Figure 4. Photo, motor yoke assembly 6 Figure 5. Induction motor load characteristics 8 Figure 6. Induction motor efficiency in liquid helium, light loading 9 Figure 7. Induction motor efficiency in liquid helium, medium loading 10

Figure 8. Induction motor efficiency in liquid helium, heavy loading 11

Figure 9. Induction motor efficiency in liquid nitrogen, light loading 12

Figure 10. Induction motor efficiency, 8 and 11 percent slip 13

Figure 11. Assembly drawing of the helium pump and the original

inlet geometry 14

Figure 12. Photo, pump and motor assembly 16

Figure 13. Photo, pump interior 17 Figure 14. Photo, pump inducer, impeller, and vaneless diffuser 18

Figure 15. Centrifugal impeller for helium pump 19

Figure 16. Helium pump test loop and Instrumentation 20

Figure 17. Pump H-Q performance in liquid helium at 4.1 K 23

Figure 18. Pump performance, 6000 rpm, 11 percent slip, coolant holes open .... 24

Figure 19. Pump performance, 6000 rpm, 8 percent slip, coolant holes plugged ... 25

Figure 20. Pump performance, 5000 rpm, 8 percent slip, coolant holes plugged ... 26

Figure 21. Pump performance, 7000 rpm, 8 percent slip, coolant holes plugged ... 27

Figure 22. H-Q data normalized to the design speed of 6000 rpm 29

Figure 23. Flow coefficient control data 33

Figure 24. Helium pump with streamline entry 35

Figure 25. Cavitation data, 6000 rpm, streamline entry 36

Figure 26. Cavitation data, 7000 rpm, streamline entry ... 37

Figure 27. High-entry-loss cover plate 39

Figure 28. Cavitation data, 4000 rpm, high-entry-loss cover plate . . 40

Figure 29. Cavitation data, 6000 rpm, high-entry- loss cover plate ... 41

Figure 30. Cavitation data, 7000 rpm, high-entry-loss cover plate 42

Figure 31. Axial flow inducer 43

Figure 32. Cavitation data, no inducer, streamline entry , 44

Figure 33. Cavitation data, no inducer, hexagon and hemispherical nuts 46

Figure 34. Cavitation data, at temperatures of 2.35 and 4.7 K 47

Table 1. Liquid helium pump performance data 30

Table 2. Flow coefficient - entry velocity head values 48

iii
Terminology

Capital Letters

H

LL

N

Nsf

NPSH

NPSHc

Ig

2g

T

V

o

o2g

V,

_2

2g

W

sh

= area of the pump inlet at elevation Zo = total dynamic head

= helium liquid level above Z = motor or pump speed

speed of the stator field in the

electric motor

= net positive suction head

= net positive suction head at cavitation

= static pressure at the pump

inlet, elevation = Z

o = static pressure at the pressure

tap in the helium bath located

at elevation

= gauge pressure from the capillary

tubing terminating at Z^ = static pressure at the pump discharge pressure tap located at

elevation

= gauge pressure from the capillary

tubing terminating at the pump

discharge, elevation Z^

= pump capacity

= induction motor slip = Nsf - N Nsf ^

= temperature, kelvin

= average fluid velocity at the

inducer entrance, elevation Z

= fluid velocity head at the pump

entrance

= average fluid velocity at the

static pressure tap,

= average fluid velocity within the

pump discharge tube at the static pressure tap V^, elevation Z^

= fluid velocity head within the

pump discharge tube at the static pressure tap elevation Z^

= fluid power

= pump shaft power

= pump reference datum plane (at

inducer leading edge) = elevation above the reference

plane Zo

= ptrnp elevation head = 7 cm

Lower Case Letters

g = local acceleration of gravity

= pump entry head losses

h = absolute static pressure head at the pump inlet (elevation = Z^)

h = head generated by the ullage

pressure

h = head due to the vapor pressure of

the fluid at the pump entrance, at

elevation Zo m = mass rate of flow through the pump

u^ = inducer tip speed

v^ = specific volume of the vapor

Vq = specific volume of the liquid

Greek S3rmbols

ri^ = pump hydraulic efficiency

T]^ = total efficiency of the pump and

motor

ri = efficiency of the electric motor m

r)^ = pump volumetric efficiency

rip = gross efficiency of the pump

p = fluid density

$ = inducer flow coefficient, V /u

iv
PERFORMANCE CHARACTERISTICS OF A LIQUID HELIUM PUMP*

P. R. Ludtke

Cryogenics Division

Institute for Basic Standards

National Bureau of Standards

Boulder, Colorado 80302

Abstract

Part A of this report presents performance data for a simple,

preinduced, single stage, centrifugal liquid helium pump powered by a close-coupled submersible cryogenic induction motor. Data on pump

efficiency and the motor efficiency are given, and the effects of

decreasing the pump leakage loss and removing the pre-inducer were also investigated.

Part B describes a study of the cavitation characteristics of the

pump in liquid helium I. The net positive suction head at cavitation

of the pump was extensively investigated, and found to be near zero; the effects of removing the inducer, changing the pump inlet geometry,

and varying the helium temperature were also determined.

Key words: Cavitation; cryogenic motors; induction motors; liquid

helium pumps; motor efficiency; net positive suction head; pump

efficiency; pump performance.

Introduction

In an earlier report, the design, assembly, and preliminary testing of a small submersible centrifugal helium pump were described [1], That work showed that the pump performance

in liquid nitrogen and in normally boiling, supercritical, and even superfluid helium was consistent with standard scaling laws for predicting pump performance, though the developed

head was less than that specified by design. In the limited time available for testing, only

the combined efficiency of the pump and motor was measured, but not the efficiency of the

pump itself; also, some preliminary evidence of cavitation in boiling liquid was given.

This report presents further studies on the same pump. It is divided into two parts,

(A) evaluation of pump efficiency, and improvements in pump performance suggested by both the

earlier work and the current laboratory tests, and (B) an attempt to study the cavitation

characteristics using correlative techniques which have been developed in our laboratory [2]

for other cryogenic fluids.

A number of tests were conducted with the small liquid helium pump. The purposes of

the tests were: Part A

1) to compare the previous pump performance data with the new test data taken after

making improvements to the pump, and

2) to plug the motor coolant holes in the pump and determine if this improves the

pump performance characteristics, and

3) to measure the induction motor efficiency at various operating conditions in liquid

helium and liquid nitrogen, and

4) to calculate the pump efficiency at different operating conditions using the motor

efficiency and total efficiency values, and

Part B

1) to conduct a thorough investigation of the cavitation characteristics of the

helium pump, and

2) to use the cavitation data obtained to check two different cavitation models

(generated from cavitation data) and determine if the computer models can be

used to predict cavitation characteristics for different fluids (He) and pumps of

different geometries.

*Studies reported as Part (A) of this report were funded by the Wright-Patterson Air Force

Base, and studies reported as Part (B) were funded by the Defense Advanced Research Projects

Agency.
A. Pump Performance and Efficiency

A.l Measurement of the Electric Motor Efficiency

The helium pump and the induction motor are rigidly connected together, and the total

efficiency of both units can be easily determined by measuring net electrical power Into

the motor and fluid power out of the pump. It was decided that in our case the most

straightforward method of determining the pump efficiency is to measure the efficiency of

the motor separately as a function of shaft speed and input power and frequency, and then

calculate the efficiency of the pump using the motor efficiency and the measured total

efficiency.

The motor originally powered a small circulating fan in liquid oxygen vessels to inhibit thermal stratification. The motor is an 8-pole Induction tj^pe, designed for 3-phase

400 Hz input power, with a maximum power rating of 35 watts. At 400 Hz the stator field

rotation speed is 6000 rpmt. The circulating fan was removed and replaced by an agitator

in order to measure the efficiency of the motor.

A. 1.1. Description of the Motor Efficiency Test Apparatus

A special experimental apparatus was designed and built to measure the motor efficiency

in liquid helium. A cross-section of the apparatus is shown in figure 1, and photos of the

apparatus are shown in figures 2, 3, and 4. The induction motor is clamped into a movable

yoke which is suspended by a stainless steel torque rod from the top plate. The motor and

yoke assembly is thus free to rotate ± 15 degrees from a neutral position. There is a stationary baffle plate below the yoke and motor assembly that serves to center the motor

and inhibit swirling of the helium during a test. Three additional baffle plates were attached to the stationary rods around the periphery of the motor. This was necessary because liquid helium swirling around the motor could exert drag on the motor body which would

tend to decrease the torque reading from the motor. A ball bearing centers the motor shaft

in the baffle plate. Agitators of three different sizes were used to load the motor in

order to measure the efficiency under different loading conditions.

A thick disc of high density polystyrene foam prevents circulation of the cold helium

gas on the underside of the top plate. The liquid helium fill line and the vent gas line

pass through stainless steel standoff tubes attached to the top plate. This allows the

brass top plate to remain near ambient temperature during tests. The 3-phase power and

motor speed sensor wires are attached to the motor yoke, extending up through the foam and

out through an ambient temperature rubber seal on the top plate.

The entire assembly is placed inside of, and is supported by a 15 cm diameter glass

dewar. This dewar is then placed inside a larger glass dewar filled with liquid nitrogen

for thermal shielding. The motor and yoke assembly is supported by the ambient temperature bearing in the top plate. The bottom bearing carries no thrust, and any frictional

drag from this bearing introduces no error in the torque measurements, since it merely

exerts additional drag on the motor.

A torque transducer is firmly attached to the assembly above the brass top plate.

Flexible helical connectors were used on each side of the transducer to compensate for any

misalignment. The transducer senses the torque on the motor body which is equal and

opposite to the torque on the motor shaft, and a calibrated meter provides a visual indication of the torque value.

A 3-phase power supply having variable voltage (0-150 V line-neutral) and variable

frequency (50-3600 Hz) was used to power the motor. A polyphase wattmeter was used to measure the electrical power into the motor; a small correction was necessary for power

consumed by the wattmeter. A counter was used to sense the motor speed. Values for the

motor speed and torque were then used to calculate the power output of the motor.

A.l. 2. Motor Efficiency Test Procedure

Each of the glass dewars was purged with either nitrogen or helium gas. A thorough

purge of the inner vessel was necessary to insure that all moisture was removed from the

t In order to present the pump performance data concisely and in the terminology widely

accepted in the pump field, the pump and motor speed is given in units of revolutions per

minute (rpm) in the text. The SI units of radians /second (rad/s) are given on all the

pump performance graphs, along with the conventional units of revolutions per minute.

2
Anchor Pin

Ball Bearings

Support Rods

(Three)

Retainer Clip

Torque Tube

Speed Sensor

3* Induction Motor

Torque Transducer (0—720 centimeter— grams)

Helical Coupling

Ball Bearing

Shield on bottom side.

Brass Top Plate

Polystyrene Insulation

LN2 Shielding

Liquid Helium

Glass Dewars

Stainless Steel Tubing

Spur Gear on Motor Shaft

Motor Yoke Assembly

Ball Bearing

with Plastic Cage

Stationary Baffle Plate

Agitator

Figure I, Motor Efficiency Test Apparatus

3
Figure 2. Photo, motor efficiency test apparatus.

4
Figure 3. Photo, motor efficiency te st apparatus.

5
Figure 4. Photo, motor yoke assembly.

6
motor bearings. The shielding vessel was then filled with liquid nitrogen and the inner

vessel subsequently filled with liquid helium. The tests were conducted with the helium

liquid level > 6 cm above the yoke assembly. The motor was run at a given 3-phase input

frequency of 300, AOO or 500 Hz with line-neutral voltage varying from 20 to 130 volts;

efficiency data was taken at each 10 volt increment.

The electrical power into the motor, the motor speed, the torque, and the voltage were

recorded for each voltage setting. The net power into the motor, the power out of the

motor, and the motor slip were then calculated. Tests were conducted using three different

size agitators for varied loading. The zero reading on the torque transducer was checked

before and after each individual test, since the wires leading from the yoke assembly to the foam could be a source of error. However, the total movement of the yoke and motor

assembly for the maximum 720 centimeter-grams (10 inch-ounces) torque reading is only 0.6A

degrees. Since the torque meter returned to zero after each test and the full scale rotation of the motor assembly is so small, the error introduced by the wires is thought to be

negligible. The overall accuracy of the measurements is estimated to be ± 3 percent.

A. 1.3. Motor Efficiency Test Results

Efficiency tests were conducted in both liquid nitrogen and liquid helium. The

efficiency data for liquid helium are shown in figures 5, 6, 7, and 8. Figure 5 is a plot

of corrected power input to the motor vs motor slip for an input frequency of 400 Hz using

three different agitators on the motor, giving light, medium, and heavy loading. Similar

data was obtained for input frequencies of 300 and 500 Hz. If the data are plotted showing

motor efficiency versus induction motor slip as in figures 6, 7, and 8, it becomes immediately

apparent that the motor operates at maximum efficiency when experiencing slip in the range

of 5-10 percent. Also, the peak efficiency of the motor is slightly dependent upon input

frequency, and the magnitude of the efficiency is dependent on the load placed upon the

motor. With a 3-phase frequency of 400 Hz (the motor design input frequency) the motor has

optimum efficiency in the slip range of 5-10 percent. The efficiency of the motor with 8 percent slip and 400 Hz input is 58, 69, and 74 percent for the light, medium, and heavy

motor loads, respectively. This is higher than had been estimated (without data) in

reference [1].

For most applications where small electric motors might run in liquid helium, it is

desirable to operate the motors at or near maximum efficiency in order to boil away a minimum amount of liquid helium. Thus, electric motor efficiencies in the range of 60 to

75 percent are encouraging for applications using electric motors to power small helium

pumps to circulate liquid helium in superconducting power transmission cables and other

superconducting devices. Even higher efficiencies may be possible with additional innovation.

Efficiency data for similar tests in liquid nitrogen are shown in figure 9. The

efficiencies were not as high as in the liquid helium tests, possibly because of the

different operating characteristics of the motor at 4 and 76 K. Minimal data were taken

at 76 K. . The efficiency data plotted in figures 5, 6, 7, and 8 do not permit a quick determination of the electric motor efficiency for given operating conditions, because the

efficiency depends on the motor load which is not quantitatively specified in these

figures. However, if the motor efficiency is plotted vs power input/ 3-phase input frequency for a given motor slip, it is possible to quickly determine the motor efficiency

for a particular set of operating conditions. Curves are plotted for eight percent slip

and eleven percent slip in figure 10. Eight and eleven percent slip are plotted since the

motor was run with these two values for slip in the two sets of performance tests with the

helium pump. These data were then used to extract the pump efficiency from data on the

combined efficiency of the motor and pump.

A. 2. Pump Performance Tests

The rather simple helium pump used in these tests was designed by an independent

engineering company. The pump bolts directly to the motor as shown in figure 11. The

pump has an axial flow inducer, a centrifugal type impeller, and a vaneless diffuser. The

pump was specifically designed for liquid helium and has a design point of 15.2 m helium at 2.52 X 10~ m /s (50 ft helium (3 4 gpm) at 6000 rpm.

The pump and motor have been tested previously; the pump performance data are reported

in reference [1]. The previous pump performance data fell significantly below the design

specifications for the pump.

7


:^udOJ3(i 'ADI\13I3Ijd3 HOiOW

10
:^u9Djad 'A3N3I3IdJ3 aOiOW

11

75

I I I I

I

I I II

I

I I I I

I

I I I I

I

I I I I

I

I

70

u

s-

(U Ou

>- 65

o; 60

o

8 Percent

11 Percent

55

o 8 Percent Motor Slip (4. IK)

A n Percent Motor Slip (4. IK)

50

.001

LI i_L

.005 .010 .015 .020 .025 .030

POWER INPUT / 3-PHASE INPUT FREQUENCY , Watts/Hertz

Figure 10. Induction motor efficiency, 8 and 11 percent slip.

13
Stainless Steel Tubing

Helical Compression

Spring

6-32 NC Self -locking Nut

Low - temperature

Bearings with Gloss impregnated, PTFE Retainer

Submersible 3-Phase

Induction Motor,

400 Hz, 208 VAC L-L

120 VAC L-N

Power max 3 5 W

(95.25 mm)

Figure 11. Assembly drawing of the helium pump

and the original inlet geometry.

14
In view of this it was decided that, (1) the impeller position in the pump should be

checked, and adjusted if necessary for optimum performance, (2) the inlet geometry to the pump should be streamlined, and (3) comparative tests should be made to determine the

effect of the pump motor coolant passages on the performance characteristics of the pump.

The impeller position within the pump was checked. The base of the impeller (i.e.,

^1 P u ... u ^ J- ^, J X. J N u J +0.203 mm /+0.008"\

the flat surface where the vanes attach to the disc shaped body) had a „ < r -0.102 mm (-0.00A ) variation from the plane of the vaneless diffuser; the variation resulted mostly because the

back shoulder of the impeller was not parallel with the impeller base. For optimum performance, the impeller base should be in the same plane as the vaneless diffuser plane. The

back shoulder of the impeller was re-machined perpendicular to the axis, reducing the impeller

. . ^ +0.051 mm <+0.002" \ ^ ^ ^ , ^ ^ , ^.^^ variation to _ „ \ ^ nnncu{ with respect to the plane of the vaneless diffuser. There -0.013 mm 1-0.0005 ) ^ was also a slight machining error in the cover plate of the pump. An entirely new pump cover plate was machined with a much improved suction bell inlet. The original acorn shaped hexagonal shaft nut was also replaced with an ogive shaped nut to improve the streamline flow

into the pump. The photos of the motor and pump in figures 12, 13, and 14 show the new pump cover plate with the improved suction bell and the streamlined ogive nut on the inducer. A

drawing of the impeller is shown in figure 15. After making the above improvements, the pump was again tested for H-Q performance characteristics in liquid helium.

A. 2.1. Pump Performance Test Apparatus

The test apparatus used to determine the performance characteristics of the pump is

shown in figure 16. The apparatus is essentially the same as used in the previous work fl],

except for the turbine flowmeter which was added for the cavitation measurements. The

entire pump assembly and flow loop are placed in a dewar filled with liquid helium. The

liquid helium level is always maintained above the level of the flow control valve. The

helium flow path is as shown: entering the bottom of the pump, discharging vertically through

the venturi flowmeter and on through the turbine type flowmeter, then through the discharge

control valve and on down through a conical diffuser to the helium bath in the dewar. The

discharge control valve is used to manually adjust the rate of flow through the pump. The

venturi meter and the turbine meter both measure the flow rate of the pump. The flow rate

through the venturi is determined from the AP measurement across the meter and a calibration factor. The turbine type flowmeter is calibrated using the venturi meter.

The pressure lines from the test apparatus are stainless steel capillary tubing

(1.57 mm CD X 0.2 mm wall). The capillary tubes were thermally connected every 20 cm to maintain approximately the same thermal gradient within each tube, thus minimizing the liquid

level (head) differences within the various capillary tubes. This procedure was found to be

necessary to minimize random drifts and fluctuations in the venturi AP readings. The pressure

lines exit to ambient temperature diaphragm type AP gauges.

The capacitance type liquid level gauge used during the earlier tests was replaced with

a superconducting type liquid level gauge with improved sensitivity and resolution. The

improved liquid level sensor was necessary because we planned to run cavitation tests at

4.7 K, where the density ratio of liquid to vapor is only 4.17, making a capacitance type

liquid level indicator very marginal (see part B)

A barostat was used to control the ullage pressure of the helium dewar, thus con- trolling the temperature of the liquid helium bath. The remainder of the instrumentation ~

the 3-phase power supply, the polyphase wattmeter, the pump and motor speed counter were

described previously in section A. 1.1.

A. 2. 2. Pump Performance Test Procedure

Previous to conducting a test, the helium dewar was evacuated and subsequently

pressurized to near atmospheric pressure using helium gas. This cycle was repeated three

times in order to minimize condensible gas impurities and moisture within the experimental vessel. The nitrogen shielding dewar would then be filled, and subsequently the

experimental dewar would be filled with liquid helium to 40 cm above the level of the

flow control valve.

At the start of a test, the barostat would be adjusted, the predetermined 3-phase

input frequency would be set so the motor would operate with the desired slip, and the

motor or pvnnp speed would be adjusted to the desired speed by adjusting the input voltage.

15
16
17

49.91 mm

• 19.05 mm

6.6 mm

0.76 mm

.79 mm

Figure 15. Centrifugal impeller for helium pump.

19
Pu APv APp

Liquid Helium

Environment

Discharge

Control Valve

Superconducting

Liquid Level Sensor

Conical Diffuser

Inductive Speed

Sensor

Frequency

Counter

Polyphase

Wattmeter

3-Phase

Power Supply

Centrifugal Impeller

Motor

Speed

Motor

Power

^—Liquid Level (LL)

Extended Valve Stem

Support Tubes

iz:^^- Stainless Steel '''^ Capillary Tubing

I Turbine

Speed

Frequency

Counter

Turbine Type

Flowmeter

I Venturi Type

Flowmeter

20 Tooth

Steel Spur Gear

7 cm

|7Pi^'T^P Static

I 1/ Pressure Taps

r^P V Z

Axial Flow Inducer

Figure 16. Helium pump test loop and instrumentation.

20
When the desired operating conditions were achieved, the following data were recorded:

the ullage pressure, the pump AP, the venturi meter AP, the pump speed, the liquid level,

the line-neutral volts, the power into the motor, the 3-phase frequency input, and the

turbine meter signal. After recording the above data for a given flow rate, the flow

rate would be changed by manually adjusting the flow control valve, the pump speed would

be readjusted, and another set of data would be recorded. By adjusting the flow control

valve in increments between the closed and open positions, pump performance data were obtained for the full flow range. During the pump performance tests, the liquid helium

level was always maintained at least 4 cm above the level of the flow control valve.

A. 2. 3. Pump Performance Results

The total dynamic head (H) of the pump was calculated in the conventional manner; however, the measurement procedure should be carefully explained. The total dynamic head

is normally defined as

2 2

g Q

P 2g 2 o

If we assume that, in our particular case, the flow is isothermal and the fluid is of

negligible compressibility, reference [1]

2 P - P

1

J. dP ^ _2 lo

g Q

P Pg

and

2 2 P - P -

pg 2g 2 o

Referring

""

to figure 16, Zol= Z, and thus Z.2o21o - Z = Z„ - Z, . P is the static pressure at

the pump entrance and P^^ is the static pressure at the same elevation out in the helium bath.

Considering the pressures at points (1) and (2)

,

2 2 p, vr p V

_1 + 1 =h, + ^

pg 2g I pg 2g

If is considered negligible because of the low fluid velocity at point (1) , and the pump

entry losses are negligible, as with the streamline entry plate,

P P- _o _ _1 _o

pg pg " 2g '

and the total dynamic head may be expressed as follows, using variables pertaining to points (1) and (2)

2 P - P V

Next consider the type of instrumentation and methodology used for measuring ^2 ^^'^ As shown in figure 16, the pressure taps at points (1) and (2) are connected to a AP gauge

with stainless steel capillary tubing, and it is assumed that each tube is filled with

liquid helium to the same level as the bath liquid level. The relationship between the

static pressures at P^ and P2 and tfie gauge pressures P2 and P^^ are as follows.

P^ = Pj^g + pg(LL-Zj^) and P2 = PgCLL-Z^)

P2 - = ~ ^Ig pg (1^^-22 - + V' °^

P„ - P- P„ - Pt 1£ = J: 1 + (z _ z^) (4)

pg pg '21

21
Thus, using this method of pressure instrumentation, the pressure head calculated

using pressures read on the gauge is equal to the actual pressure head developed by the

pump between points (1) and (2), plus the elevation head (Z^ - Z^), and the total dynamic

head of the pump may be expressed in terms of the direct measured or calculable values,

thus „

Pg 2g

Pump performance data were taken at 4000, 5000, 6000, and 7000 rpm. The performance

data are plotted in figure 17, along with the performance data for 6000 and 7154 rpm taken

prior to making the pump improvements „ As shown in the figure, the improvements to the

pump had a significant effect on the head-capacity performance of the pump. At high flow

rates, there is a 30 percent increase in developed head at 6000 rpm. This percentage de- creases appreciably at the lower flow rates. However, the pump performance at 6000 rpm is

still significantly below the predicted design point plotted in the figure. The pump

doesn't even meet design point specifications when operated at 7000 rpm. We have no certain

explanation for this discrepancy, which was discussed also in reference [1]. Because in

this work we have determined that the pump efficiency is reasonable (about 45%, maximum) we

suspect that the failure to meet the stated design specifications is not due to any unusual

losses or leakage within the pump, but rather may be due to an analytical error in the

original design.

After obtaining one set of performance data with liquid helium, the three motor coolant

passages in the top of the pump were plugged in order to determine if these passages have

any effect on pump performance. These holes are 2.24 mm diameter and have a total area of

2 0.116 cm . However, the holes are not the only flow path between the pump and the motor;

there is an annular space between the rotating shaft and the pump housing. The shaft

2 clearance is 0.43 mm and the cross-section area of the annular flow path is 0.125 cm , just

slightly larger than the area of the plugged coolant holes. Thus, plugging the three coolant

holes eliminated 48 percent of the total area available for liquid helium flow up through

the motor.

The pump performance data taken after plugging the motor coolant holes is also plotted

in figure 17. The H-Q performance of the pump appears to be the same as previous with

the coolant holes open, and one can easily conclude that the coolant passages have no perceivable effect on the head-capacity performance of the pump.

During the later cavitation testing the axial flow inducer was removed from the pump.

A set of head-capacity data were taken with the inducer removed; these data are also shown

in figure 17. Removing the inducer has little effect on the head developed at high flow

-4 3 rates. Below 1 x 10 m /s the head decreases significantly as shown.

A. 2. 4. Pump Efficiency Results

Two different series of liquid helium tests were conducted; one series with the motor

coolant holes in the pump open and the motor operating at 11 percent slip, and a later

series with the motor coolant holes plugged and the motor running near peak efficiency at

8 percent slip. Since the coolant holes have no effect on pump performance, we have

essentially one series of tests with the motor running slightly off peak efficiency at 11

percent slip, and another series of tests with the motor running near peak efficiency at

8 percent slip. Complete performance data for the 6000 rpm tests were plotted from each

series. The fluid power, total efficiency, motor efficiency, pump efficiency, power input,

and the total dynamic head are plotted for each test. The data is presented in figures 18

and 19. Performance data for the pump and motor at 5000 and 7000 rpm are also shown in

figures 20 and 21.

The input power to the system is simply the polyphase wattmeter reading less the

necessary correction for the power consumed in the wattmeter measuring circuitry. The

power output or fluid power is calculated as follows:

= mgH

where W is watts, m is mass flow rate (kg/s), g is the local acceleration of gravity

2 (m/s ), and H is the total dynamic head (m) . The total efficiency is the ratio of the fluid

power from the pump to the power into the motor.

22
V No Inducer on Pump

A Motor Coolant Holes Plugged

o Motor Coolant Holes Open

Old [l] Performance Data

PUMP CAPACITY, m^/s

Figure 17. Pump H-Q performance in liquid helium

at 4. 1 K.
3.0x10

PUMP CAPACITY, m7s

Figure 18. Pump perforaance, 6000 rpm, 11 percent

slip, coolant holes open.

24
I I I I

I

I I I I

I

I

I I I Liquid Helium

628 rad/s (6000 rpm)

Coolant Holes Plugged

8 Percent Slip

I

I I I I

I

I I I I

I

I I I I

Pump Efficiency

(Percent)

3.0x10

PUMP CAPACITY, m7s

Figure 19. Pump performance, 6000 rpm, 8 percent

slip, coolant holes plugged.

25
3.0x10

PUMP CAPACITY, m7s

Figure 20. Pump performance, 5000 rpm, 8 percent

slip, coolant holes plugged.

26
45

I

I

I I

>-o

40

LlI I—

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I

' ' ' ' M I I I

Liquid Helium

733 rad/s (7000 rpm)

Coolant holes pluggej

8 Percent SI ip

I

I I I I

I

I

I I Pump Efficiency

(Percent) 75

70

35

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Induction Motor Efficiency

(Percent) 65

30 oX

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Total Dynamic Head (meters)

Total Power Input (Watts)-

60 ^

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45

40

Fluid Power (Watts) — o XT -O 35

I » I I I I I I I I I I I I I t I 1 I f I I 1 I I I I I

I I I I 0.5 1.0 1.5 2.0 2.5 3.0 3.5x10

PUMP CAPACITY, mVs

Figure 21. Pump performance, 7000 rpm, 8 percent

slip, coolant holes plugged.

27
The motor efficiency was obtained from the test data described in section A. 1.3.

The pump efficiency was calculated using the total efficiency and motor efficiency values

as, \

I m

Tabulated values for the pump performance test are presented in table 1. The estimated

uncertainty of the pump performance data is ± 5 percent.

Comparing the performance parameters in figures 18 and 19, the total dynamic head for

the two tests are essentially the same, and the fluid power is essentially the same. The

power input to the motor is higher for the test run at 11 percent slip than for the test run at 8 percent slip. The motor efficiency and consequently the total efficiency is higher

when the motor is run at 8 percent slip. The pump efficiencies are the same of course,

since the same pump is running at the same speed.

The variation of total dynamic head, pump capacity, and pump shaft power with the speed

of a centrifugal pump follow definite relationships known as the pump affinity laws, which

may be derived using general principles of similitude. The affinity laws for a given cen- trifugal pump are expressed by the following equations:

^shl )

The subscript "1" in the above equations denotes known tests conditions, and ri^, ri^^, and ri^ are the volumetric, hydraulic, and total pump efficiencies, respectively. Determination of

the volumetric and hydraulic efficiencies of the pump are beyond the scope of this program;

it will be assumed that the efficiency ratios for the different test conditions are equal

to unity, as is often done. The affinity laws may be applied to a test pump to predict

parameters different from the known test parameters, or they may be used to normalize pump

-test data taken at different speeds to one given speed.

The affinity laws were used to normalize the 5000 and 7000 rpm test data to 6000 rpm

for comparison to the actual H-Q test data taken at 6000 rpm. The normalized data and the

6000 rpm test data are shown In figure 22; and one can readily conclude that the pump performance data is in excellent agreement with the pump affinity laws for speed variations

within this range.

The specific speed is another parameter which is often used to characterize impeller

pumps; it also results from the general principles of similitude and is defined as,

and is usually computed at the best efficiency point (b.e.p.) of the pump. The best

efficiency point of this pump is not precisely defined because of the lack of data points

near the peak of the curve, but considering all of the 6000 rpm (design speed) data, the

b.e.p. would be near the point indicated in figures 18 and 19. The H-Q values at this

"A 3 point are 9.7 m and 1.6 x 10 m /s, respectively. Values of N are dependent on the units

used; in U.S. engineering practice it Is customary to use the following units: N(rpm),

Q(gal/min), and H(ft) . The value for N using the customary English units, N = 6000 rpm,

Q = 2.54 gal/min, and H = 31.82 ft, is ^ ^

== 712 (English units)

Centrifugal pumps have values for (English units) in the range of 500 to 3500; thus,

this pump is of relatively low specific speed.

28
14

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1 1 1 1 M M 1

1 M 1 1 1 1 1 1 1

1 1 M

r

(J Ll

— n _

o

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628 rad/s (6000 rpm) data o ^

733 rad/s (7000 rpm) data A. ~

Normalized to 628 rad/s "^oc —

524 rad/s (5000 rpm) data °

Normalized to 628 rad/s

Helium Temperature =4. IK

1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 0 5 1.0 1.5 2.0 2.5x10

PUMP CAPACITY, m^/s

Figure 22. H-Q data normalized to the design speed

of 6000 rpm.

29
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30
A. 3. Conclusions From Pump Performance Tests

There are several conclusions resulting from the pump perfoirmance tests and the motor

efficiency tests. These conclusions are:

1) The use of a submerged cryogenic induction motor to power a simple preinduced, single

stage impeller pump is a reliable and efficient method of generating easily con- trollable liquid helium flow, and

2) the efficiency of this cryogenic motor is dependent upon slip and load, and ranges up

to a maximum of 74 percent in liquid helium, and 60 percent in liquid nitrogen,

3) the efficiency of the pump at design speed (6000 rpm) and at the best efficiency

point (b.e.p.) is 44 percent,

4) the pump affinity laws are applicable to the non-cavitating H-Q pump performance

curves,

5) using a streamline entry, and positioning the impeller for optimum performance,

increased the developed head of this pump up to 30 percent at 6000 rpm,

6) plugging the motor coolant passages had no perceivable effect on the H-Q performance of this pump,

7) removing the axial inducer caused a pronounced decrease in the developed head of

-4 3 the pump at flow rates below 1 x 10 m /s,

8) using reasonable precautions to remove all moisture and condensables by using a thorough helium purge procedure, the pimip and motor operation in liquid helium

has been routine and trouble-free in our experience.

B. Pump Cavitation Study

In our earlier work fl] some evidence was seen for the beginning of cavitation within

the pump when the liquid level dropped to an estimated few centimeters above the pump

inlet. The original estimate of the pump designer was that cavitation might develop at

liquid levels less than 25 centimeters. The conclusion stated at that time was "a simple

inducer-impeller pump can operate satisfactorily in helium I and helium II with little

liquid cover. This result may not be universal and needs further testing and evaluation."

An important objective of this work was to explore the cavitation performance of the

pump much more carefully, and to correlate it with cavitation performance of other pumps

in other cryogenic fluids using some thorough analytical procedures developed within our

- laboratory [2]. Toward this goal, a broad program was outlined, (a) utilizing improved

instrumentation, especially for measuring the helitim liquid level, (b) including a wider

range of pump speeds, limited in the previous study by the small frequency range of the

three-phase power supply, and (c) including a temperature range of about 2.5 to 4.7 K, to test sensitivity of the results to changes in the liquid-to-vapor density ratio.

B. l Cavitation Test Program

A cavitation test program was outlined that would provide sufficient data such that

the computer model capable of predicting cavitation parameters in hydrogen, could be

tested to see if the model could be extended to helium. The test program required cavitation data at'three different temperatures (2.8, 4.05, and 4.7 K were chosen), at three

different pump speeds (4000, 6000, and 7000 rpm), and at three different flow coefficients

= 0.1, $2 ~ 0.16, and ~ 0.22 were chosen). Constancy requirements were placed

on each of the cavitation parameters in order to satisfy the percent error requirements of

the cavitation model. These requirements were:

• Constancy of 1/10% on temperature

• Constancy of 1% of the flow coefficients

• Constancy of 1/2% on the pump speed.

With the test program firmly set, we began to check the operating tolerances necessary to meet the constancy requirements. In order to meet the helium bath temperature requirements, it was necessary to control the ullage pressure within the pressure tolerances

listed below:

31
Temperatures Pressure Tolerances

2.8 K ± 80 Pa (± 0.6 mm Hg)

4.05 K ± 333 Pa (± 2.5 mm Hg)

4.7 K ± 600 Pa (± 4.5 mm Hg)

It was assumed that the helium bath would be in equilibrium with the ullage pressure of

the dewar, and considering the circulation provided by the pump, this appears to be a reasonable assumption. We planned to control the ullage pressure with an NBS barostat

which does have the above control capability. Thus the temperature constancy requirements

could be met. The pump speed tolerances are given below; the counts are from the 20 tooth spur gear

on the motor shaft.

Pump Speed Counter Tolerances

4000 rpm ± 7 counts/second

6000 rpm ± 10 counts/second

7000 rpm ± 11 counts/second

During the earlier motor efficiency tests, it was not difficult to maintain a counter

tolerance of ± five counts/second; thus the pump speed constancy requirements could be

easily met.

The inducer flow coefficient for the cavitation tests is defined as: V V

^ o o u^ 2TrrN where equals the average flow velocity at the axial inducer inlet, and u^ equals the tip

velocity of the inducer. Values of were obtained from the previous H-Q performance data.

In order to use the same inducer flow coefficient for all tests, it is necessary to make

plots of the flovraieter signal vs the inducer flow coefficient as shown in figure 23. One

of these plots is necessary for each temperature. These curves allow one to arbitrarily

choose values for each of the three flow coefficients near the low, mid, and high flow

ranges, and also to readily determine the flowmeter signal required for tests run at different speeds and temperatures.

The inducer flow coefficient constancy requirement was more severe. In obtaining the

H-Q performance data, the mass flow rate was measured with the venturi meter. However, the

dial type pressure gauge used to measure the venturi flowmeter AP does not have adequate

resolution so that one can read pressure differences resulting from one percent variations

qf the inducer flow coefficient. The minimum discernible value when reading the pressure

gauge is 25 Pa (0.1 inches water), and the venturi AP pressure excursion for a one percent

variation of the flow coefficient at 4000 rpm is ± 0.005 inches of water. Consequently

a flow measuring sensor with increased sensitivity was required in order to maintain the

inducer flow coefficient values to tolerance of one percent. We decided at this point to place a turbine flowmeter in series with and downstream of the venturi flowmeter in order to obtain a more sensitive flow signal. The flowmeter chosen has an output of 0-600 counts/

second for our flow range of 0-38 grams/second. The turbine flowmeter count tolerances for one percent variations of the inducer flow coefficients at 4000 rpm become:

Flow Coefficient (^) Turbine Flowmeter Count Tolerances

$ = 0.1 ±1.4 counts/second

$ = 0.16 ±2.3 counts/second

$ = 0.22 ±3.1 counts/second

When conducting the tests, we found it possible to easily maintain the flowmeter reading to

± 1 count/second. Thus adding the turbine flowmeter provided the increased sensitivity

necessary to determine and maintain the flow constancy requirements.

32
s/s:^uno^ *y3i3WM01d 3Niaiini

33
At this point we were satisfied that the cavitation parameter constancy requirements

could be attained. The other important measurement for cavitation studies is the liquid

level. The liquid level sensor used in the previous cavitation tests was a capacitance

type, which has marginal sensitivity at liquid helium temperatures. Using a capacitance

bridge with a resolution of 0.1 pf , the measurement uncertainty is ± 2 cm in still helium.

To improve this, a superconducting liquid level sensor was installed. The sensitivity

of the sensor is 326 mV/cm. The liquid level measurement uncertainty is dependent upon

calibration procedure, and is estimated to be ± 2 mm in quiescent liquid helium. Having

determined the flowmeter signal for the different flow coefficients and being certain the

cavitation parameter constancy requirements could be met, we proceeded to run the cavitation tests.

B.2 Cavitation Test Procedure

The cavitation test apparatus is shown in figure 16. The procedure for conducting the

cavitation tests was as follows: the dewar would be filled with liquid helium to about

30 cm above the pump inlet (Z ). The pump would be started, the barostat adjusted to maintain ullage pressure, the discharge valve would be adjusted to attain the flow rate

required for the desired and the pump speed would be attained by varying the voltage

to the motor. After adjusting the above test parameters, the pump AP, the liquid level,

the ullage pressure, the pump speed, the power input, and the turbine meter signal would

be periodically recorded as the liquid level decreased due to boil-off. The above parameters

would be recorded until the pump ran out of liquid or cavitation occurred.

B.3 Cavitation Test Results

Six different series of cavitation tests were conducted. Five of the series contained

cavitation tests run at 4000, 6000, and 7000 rpm, using the three different flow coefficients but all the tests were run at 4.05 K. The last series Included cavitation tests

with helium at 2.35 K and 4.7 K. The first series of tests was conducted with the low-entry-loss suction bell and

the streamlined inducer nut shown in figure 24. The cavitation data is shown in figures

25 and 26. The numbers on the left side of the data points indicate that several sets

of data were recorded while the liquid level sensor was indicating a constant liquid

level due to splashing. For cavitation testing, the liquid level measurement is most

important because the liquid level is representative of NPSH as explained later in the

report. We had difficulties during all the cavitation tests in obtaining an accurate liquid

level reading near the main body of the pump. This is undoubtedly due to the discharge

^helium splashing onto the top of the pump. The superconducting liquid level sensor is

composed of a small superconducting wire inside a 8 mm O.D. plastic tube which had 1 mm- diameter holes in line, spaced about 25 mm apart. After the first cavitation test, the

holes in the vicinity of the pump base were shrouded to prevent discharge liquid from

splashing onto the sensor at this level. The shrouding helped the splashing problem but we still had difficulty in obtaining accurate liquid level readings in the vicinity of

the pump, particularly at the higher pump speeds and the high flow rate ($^). Because of

the splashing difficulty it was necessary to stop the pump soon after the developed head

began to decrease, and obtain an accurate liquid level reading. These liquid level values

are indicated on the lower left side in the figures. However, a correction for the liquid

in the circulation loop which drains into the bath when the pump is stopped has to be made;

3 this amounts to 35 cm of liquid which increases the helium bath level about 4 mm. The first series of cavitation tests with the improved instrumentation revealed that

instead of cavitating with a positive liquid level, the pump performed very well until it

appeared to abruptly cavitate or merely run out of liquid. During a typical test, the

pump AP, the pump flow rate, and the pump speed would remain constant until the liquid

level reached -4 mm, then suddenly the pump AP would decrease rapidly and the flow rate

would tend to follow. The pump speed would also tend to increase slightly, but this could

be easily maintained by adjusting the input voltage. We would try to maintain the constant

flow rate (flow coefficient $) by opening the flow control valve, but very soon it became

impossible to maintain the flow rate and the pump would be stopped. The liquid level after

stopping the pump would be near zero. Thus in the first series of tests, the pump did not

cavitate with a positive NPSH; the pump either loses suction and begins ingesting vapor or

it cavitates abruptly, close to zero NPSH, and in either case precludes the acquisition of

any meaningful cavitation data.

34
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After failing to observe cavitation in the usual sense, (i.e., a gradual degradation

of pump performance at a positive NPSH) , we decided to make a number of changes. We lowered

the liquid level sensor because the bottom of the sensor extended only slightly below the

bottom of the suction bell, giving some concern that we were not getting accurate liquid

level measurements in this region. In addition, we decided to shroud the holes in the liquid

level sensor near the pump base. We s-lso decided to use a different cover plate for the pump,

one having higher entrance losses. The NPSH of a pump operating in a closed vessel pumping

a boiling liquid is simply the static liquid level in the vessel above the pump reference

plane minus the head losses in the suction pipe (see section B.4). Thus if an entrance tube

with high head losses is used on a pump with minimal NPSH requirements, the pump should

cavitate at a higher liquid level. The cover plate shown in figure 27 was carefully designed

2 to give a total entry head loss of 1.08 Vjlg. However, the entry loss cannot be too high,

since we are also concerned about flash vaporization in the pump inlet. If the entry losses are sufficiently high to reduce h^ below h^, vapor could form upstream of the pump inlet and

void the tests. The new cover plate was installed, the liquid level sensor was lowered

7.5 cm and a second series of cavitation tests was run. These cavitation data are shown

in figures 28, 29, and 30.

After conducting the first cavitation test (3'^), it was quite evident that the helium

was still splashing on the liquid level sensor giving us an erroneous reading. Because

of this, during the next two tests at and the pump was shut off immediately after

a significant decrease in AP was noted. The liquid levels were -4.8 cm and -3.8 cm as indicated. The same procedure was used for and the 6000 rpm tests, and also for

the 7000 rpm tests. For the seven tests in this series, all of the liquid level readings

at the point where the pump head began to decrease abruptly were in the range of -3 to -5 cm, and it will be shown in section B.4 that the NPSHc ranges from -6.8 to -13.6 cm

for this series of tests. There is a possibility that the liquid level reached the bottom

of the plate (Z = -7.6 cm) letting the pump ingest vapor and lose suction, but we feel

our liquid level uncertainty is ± 2 mm when the motor is off, and that the pump did not

lose suction. Thus, once again there is no degradation in pump performance at positive

values of NPSH; it appears that the pump cavitates abruptly at negative values of NPSHc,

however, there is a remote possibilit}'^ that it loses suction due to turbulence and splashing in the dewar.

Since the high-entry-loss cover plate didn't have entry losses sufficient to raise

the liquid level above the pump when cavitation occurred, and we were reluctant to make

them higher because of the likelihood of vapor formation in the inlet, we decided to

abandon this inlet geometry and go back to the streamlined cover plate shown in figure 23.

We also removed the axial inducer from the pump and used a plain hub in its place. The

axial Inducer is shown in figure 31 and the hub would merely be an Inducer with the blades

removed. The pump datum base, which was initially set at the inducer inlet plane (Z^), was not shifted even though the inducer was removed. All the NPSH and liquid level values

in this report are referenced to Z as shown in figure 23. The liquid level sensor was

not moved.

A third series of cavitation tests was conducted, and once again there was splashing

on the liquid level sensor at high speeds and high flow rates. The cavitation data is

shown in figure 32. In all the tests, the liquid level after shutting off the pump was

about -5 mm, and if one subtracts another 4 mm for the liquid in the circulation loop, it

appears as if the pump began to abruptly lose head at a liquid level of -9 mm. However,

considering the turbulence in the vessel, there is a good possibility that the pump merely

broke suction and started ingesting vapor since the bottom of the suction bell is at a level of -14 mm. Results from this series of tests lead one to the same conclusion; the

pump either loses suction and starts ingesting vapor, or it cavitates abruptly at negative

values for NPSHc. This series of tests also indicates that the lack of an inducer has no significant effect on the cavitation characteristics of the pump.

We next decided to determine what effect the inducer nut might have on the cavitation

characteristics of the pump. "To do this a series of cavitation tests were run with a hemispherical nut on the inducer hub, and a second series was run with the original hexagonal

38
I

4.32 mm T

6 . 36 mm t — 6.38 -

17. 1mm

25.4 mm

26.1 mm 7.6 cm

-13.4mm

25.4 mm

+ z

-I,

- z

Square Corner

8.9 cm

ZqIs the arbitrary pump datum plane for all cavitation data.

Figure 27. High-entry-loss cover plate.

39
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43

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44
nut on the inducer hub. Again there was no inducer on the pump for these tests. The

results of the two series of tests are plotted together in figure 33. The liquid level

after shutting the pump off for most of the tests was about -10 mm, indicating that the

pump began to lose developed head when the liquid level was about -14 mm (taking into

account the liquid in the pump circulation loop) — the same level as the bottom of the

suction bell. There was no perceivable difference in the developed head or the cavitation

characteristics when comparing the two tests using the hemispherical nut and the hexagonal

nut; as the plots indicate, the tests are almost a 'carbon copy of each other. There was

no evidence of cavitation within the pump sufficient to affect developed head or flow

prior to the liquid level dropping to the same level as the bottom of the suction bell.

Data from this series of tests strongly indicate that the pump run out of liquid rather

than exhibit any type of cavitation.

Two more cavitation tests were conducted in order to determine if the temperature of

the helium had a significant effect on the cavitation characteristics of the pump. Since

theory [2] shows that cavitation is strongly dependent on the ratio v^/v£ or p^/p^, we

thought there might possibly be a difference in the cavitation characteristics of helium

at 2.35 K and helium at 4.7 K. The value of p„/p @ 2 . 35 K = 88 . 3 and p. /p @ 4.7 K = 4.17,

X. V X, V

a significant difference. These comparative tests were run but there was no significant

difference in the liquid levels when the pump lost developed head. The test data is shown

in figure 34. Once again, the liquid level after shut off averages - 10 mm. Considering

the liquid which drains from the circulation loop when the pump is stopped, it again

appears as if the pump broke suction and began ingesting vapor when the liquid level

reached the bottom of the suction bell. The data also indicates that the temperature of

the helium has no effect on the cavitation characteristics of the pump.

B.4 Analysis of the Cavitation Data

For the cavitation tests conducted with the streamlined suction bell, the NPSH is

calculated as follows. Using the terminology shown in figure 16, the net positive suction

head is defined as the absolute static pressure head at the pump inlet, plus the average

velocity head at the pump inlet, less the absolute vapor pressure head at the operating

temperature of the liquid at the pump inlet. Thus,

v2

NPSH = h + - h , (1) o 2g v

v2

and h + +h„ = LL + h

^ o 2g X/ u

and h = LL + h - — - h „

,

' o u 2g 1

so that NPSH = LL + h - h„ - h . U X. V

Now for an isothermal liquid in equilibrium with its vapor h^ = h^, and

NPSH = LL - hj^.

The long entry tube (figure 27) was designed such that the entry losses are equal to 1.08 V^/2g, thus - o NPSH = LL - 1.08

^ . (2) 2g

If the entry losses are assumed negligible, as for the streamlined entry,

NPSH = LL. (3)

Five different series of cavitation tests were conducted using the streamlined inlet

suction bell (figure 24). During these tests the developed head and flow rate began to abruptly decrease when the liquid helium reached the levels indicated below:

45
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47
Test series 1 Test series 3 Test series 4 & 5 Test series 6

- 1.0 cm

- 1.2 cm

- 1.4 cm

- 1.3 cm

The above data suggests that the NPSHc for the streamlined entry plate is about - 1.3 cm, the average of the above liquid level values. However, the bottom of the suction bell is

at an elevation of - 1.4 cm. Considering the turbulence in the dewar and the uncertainty

in the liquid level measurement, it is probably erroneous to speak of the NPSHc, because

the data strongly indicates that the pump simply loses suction and starts ingesting vapor

when the liquid level reaches the bottom of the suction bell.

Next, calculate the NPSHc for the cavitation tests at 6000 and 7000 rpm with the highentry-loss cover plate (figure 27) . As shown earlier the entry losses have to be considered

and they are approximately equal to the velocity head at the pump entry. The magnitude of

2 the term Y^/lg is a direct function of the pump speed and the flow coefficient. For our selected conditions, the entry velocity head values are given in the following table.

Table 2. Flow coefficient - entry velocity head values.

Pump Entry Velocity Head, V^llg (cm)

Pump Speed = 0.10 $2 = 0.16 = 0.22

4000 rpm 0.61 1.57 2.97

6000 rpm 1.38 3.54 6.7

7000 rpm 1.88 4.8 9.1

The NPSHc for the cavitation data taken at 6000 rpm are as follows. During the

$^ and $^ flow coefficient tests, the pump was stopped immediately after the head began

to decrease, then an accurate liquid level reading was taken. For the test at the

liquid level at the start of cavitation was - 5.3 cm. For the test at the liquid level was - 4.2 cm. The NPSHc calculated using eq. (2) is:

for -> NPSHc = - 5.3 cm - (1.08 x 1.38 cm) = -6.8 cm

for -> NPSHc = - 4.2 cm - (1.08 X 6.7 cm) = -11.4 cm Next consider the cavitation data taken at 7000 rpm. The corresponding liquid levels

at the start of pump cavitation are - 4.4 cm and - 3.8 cm, and

for cE'2 NPSHc -9.6 cm

for -> NPSHc = -13.6 cm

So the NPSHc for the pump using the high-entry-loss cover plate is negative, and in the

range of - 6.8 to - 13.6 cm. The liquid level at which the fluid entering the pump inducer could vaporize and

affect the pump performance may also be easily calculated. For the streamlined cover plate,

assuming the entry losses are negligible, „ V

h - h = LL - ^ . o V 2g

When h - h is £ zero, there is a possibility of helium vapor cavities forming at the pump

inlet, and ?he liquid level at which this could start occurring is equal to the inlet

velocity head of the pump, i.e., if h - h =0,

o V

LL = 2g

(4)

48
The velgcity head for the tests run at 6000 and 7000 rpm ranged from 1.6 to 9.1 cm, but there were no perceivable changes in pump performance when the liquid level was in

this range. Thus, if vapor cavities were occurring at the inlet, the cavities were not

sufficient to affect the performance of the pump.

Next consider the high-entry-loss cover plate,

h - h =LL- -^T^+hn

o V 2g Jo or h - h =LL-2.08^ . o V 2g

Again, when h^ - h^ £ zero, vapor cavities could form at the inducer inlet. So, there is

a possibility of liquid vaporizing at the pump inlet when

LL = 2.08 — . (5) 2g

The cavitation data taken at 6000 rpm with the high-entry-loss cover plate are shown

in figure 29. During the $^ and flow coefficient tests, the pump was stopped immediately

after the head began to decrease, and an accurate liquid level reading was taken. For the

test, the liquid level at the start of cavitation was - 5.3 cm, and for the test, the

corresponding liquid level was - 4.2 cm. Now check these liquid levels with the liquid

levels required for vapor cavities to occur in the pump inlet:

LL, = 2.08 X 1.38 cm = 2.87 cm

LL. = 2.08 X 6.7 cm = 13.9 cm 0

3 Next consider the 7000 rpm data in figure 30. The liquid levels where vapor may

' begin to form in the inlet are

LL. = 2.08 x 4.8 cm = 10 cm

$2

LL. = 2.08 X 9.1 cm = 18.9 cm

$3

So the liquid levels where vapor cavities could form in the inlet is in the range of

3 to 19 cm for the high-entry-loss cover plate. However, the pump performance at these

liquid levels and at lower liquid levels is very steady with very little perturbation in

ptamp head or flow rate. Pump performance is very steady until the liquid level decreases

to the - 3 to - 5 cm level, where the pump head and flow rate decrease abruptly.

In comparing the predicted values of liquid level where cavitation should begin with

the actual experimental liquid levels where cavitation actually occurred, there is a significant difference. One can only conclude that if vapor cavities are forming at the

inlet, they are pvnnped without Impairing the performance of the pump, or the helium passes

through the inlet in a metastable state and no vapor cavities occur. B,5 Conclusions from the Cavitation Tests

Several significant conclusions resulted from the cavitation testing. 1) It was impossible to obtain any meaningful cavitation data from any of our tests which could be

used for correlation with liquid hydrogen data in a computer cavitation model. Additional

tests were not run at 2.8 and 4.7 K as originally planned because the pump would not

cavitate in an acceptable manner. This particular pump shows exceptionally uniform performance over the entire range of positive NPSH values, with no perceivable degradation

in developed head, flow rate, or speed until it either fully cavitates in an abrupt fashion

(near zero NPSH) or, more likely, just runs out of liquid and starts ingesting vapor. In

either case, there is no period during a test wherein there is a gradual degradation of

pump performance due to cavitation, during which time conventional cavitation data could be

taken.

49
2) There is no loss in pump performance even when analysis indicates that the static

inlet pressure is below the saturated vapor pressure at the operating temperature. This

indicates that the helium may be in a metastable state, or slight vapor ingestion may not

measurably degrade the pump performance; a possible explanation based upon a few millidegrees of temperature stratification seems not very plausible in this highly turbulent

bath. 3) Preliminary cavitation data from the earlier tests [1], conducted with marginal

instrumentation for that type of measurement, indicated that cavitation began at an NPSHc of about 9 cm. However, in these tests, with improved instrumentation, no such

limit was observed until the liquid level fell below the pump inlet. We suspect that

the earlier results were in error, due to an inaccurate liquid level measurement.

4) The temperature of the helium I bath appears to have no appreciable effect on the

cavitation characteristics of this pump. 5) The non-cavitating performance of the pump

at positive values of NPSH conforms with the conventional affinity laws for centrifugal

pumps. 6) Significant inlet geometry changes, including removal of the axial inducer,

addition of a turbulence-generating hexagonal nut at the inlet, and addition of a long

entry tube providing moderate entry losses, do not cause cavitation in this pump. 7) The

pump did not meet the manufacturer's design point specifications, most probably because

of an analytical error in the original design. However, it should be stressed that it is

possible to accurately design and predict pump performance for a liquid helium pump using

conventional pump design parameters and the known thermodynamic properties of helium.

Because of the relatively simple construction of the pump and its freedom from

cavitation, even when subjected to the variations described above, we believe that this

freedom from cavitation is probably characteristic of the operation of centrifugal pumps

in liquid helium.

References

[1] McConnell, P. M. , Liquid Helium Pumps, Nat. Bur. Stand. (U.S.) , Interagency

Report 73-316 (June 1973).

[2] Hord, J., Cavitation in Liquid Cryogens, Vol. IV, Combined Correlations for Venturi,

Hydrofoil, Ogives, and Pumps, NASA CR-2A48 (1974).

50
U.S. DEPT. OF COMM. BIBLIOGRAPHIC DATA

SHEET

1. in;ni,i( A i ioN OR Ki-:i'c)K'i' no. NBSIR 75-816

AFAPL-TR-75-40

2. Ciov 't Actc ssion

No.

3. i<< ( ipiciii s At cc.'%sion No.

4. n 1 AM) .^I'lrm i.i: Performance Lnaracteriscics oi a L,j.qui.u nexiuui rump

5. I\ibl l< .it ion Date- June 1975

6. Hcrlormin;; Organizai ion ' oij'

7. AT I IIOKIS) Paul R. Ludtke

8, Prrlorrmng (Jr;;aii. Ki-pori No.

9. P1:1U"(1R.M!NC. 0Kc;AN1/A I KXN- .\AM1-: ANO ADDRICSS NATIONAL BUREAU OF STANDARDS

DEPARTMENT OF COMMERCE

WASHINGTON, D.C. 20234

10. l^roicct/ Ta.'^k/Work Unii No. 2750154

1 1 . ( oni rai l (irant .No.

12. -'^pon •<or ing Ornaniz.uiun Namr atui ( oniplctc- AJdr<-,'.s (Street, City, State, ZIP)

Part (A) Wright-Patterson Air Force Base

Dayton, Ohio 45433

Part (B) Defense Advanced Research Projects Agency

1400 Wilson Blvd., Arlington, Virginia 22209

13. Type of Ki-pori &; \>rru,A

( .ovcrc iiFinal

14. .'~<pon.Norinf; .'\griu y Coiii15. Sljl>l^I.l-Mi:NTAI<Y NOTh.S

16. A1^STRA(!T (A 200-word or less factual summary of most si^idcant information. If document include.^ a .^significant bibliography or literature survey, mention it here.) Part A of this report presents performance data for a simple, preinduced, single

stage, centrifugal liquid helium pump powered by a close-coupled submersible cryogenic

induction motor. Data on pump efficiency and the motor efficiency are given, and the

effects of decreasing the pump leakage loss and removing the pre-inducer were also

investigated.

Part B describes a study of the cavitation characteristics of the pump in liquid

helium I. The net positive suction head at cavitation of the pump was extensively

investigated, and found to be near zero; the effects of removing the inducer, changing

the pump inlet geometry, and varying the helium temperature were determined.

17. KEY WORDS (six to twelve entries; alphabetical order; capitalize only the first letter of the first key word unless a proper name; separated by semicolons

)

Cavitation; cryogenic motors; induction motors; liquid helium pumps; motor efficiency;

net positive suction head; pump efficiency; pump performance.

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