NUREG/CR-5558

EGG-2600

Generic Issue 87:

Flexible W~edge Gate' Valve

Test Program

Phase II Results and Analysis

Prepared by

R. Steele, Jr., K. G. DeWall, J. C. Watkins

Idaho National Engineering Laboratory

EG&G Idaho, Inc.

Prepared for

U.S. Nuclear Regulatory Commission

C
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NUREG/CR-5558

EGG-2600

RI, RM, IS

Generic Issue 87:

Flexible Wedge Gate Valve

Test Program

Phase II Results and Analysis

Manuscript Completed: July 1990

Date Published: January 1991

Prepared by

R. Steele, Jr., K G. DeWall, J. C. Watkins

G. H. Weidenhamer, 0. 0. Rothberg, NRC Program Managers

Idaho National Engineering Laboratory Managed by the U.S. Department of Energy

EG&G Idaho, Inc.

Idaho Falls, ID 83415

Prepared for

Division of Engineering

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555

NRC FINs A6857, B5529

Under DOE Contract No. DE-AC07-761D01570
ABSTRACT

Qualification and flow isolation tests were conducted to analyze the ability of

selected boiling water reactor process valves to perform their containment isolation

functions at high energy pipe break conditions and other more normal flow

conditions. Numerous parameters were measured to assess valve and motor-operator

performance at various valve loadings and to assess industry practices for predicting

valve and motor operator requirements. The valves tested were representative of

those used in reactor water cleanup systems in boiling water reactors and those used

in boiling water reactor high-pressure coolant injection (HPCI) steam lines. These

tests will provide further information for the U.S. Nuclear Regulatory Commission

Generic Issue-87, 'Failure of the HPCI Steam Line Without Isolation," and Generic

Letter 89-10, "Safety-Related Motor Operated Valve Testing and Surveillance.'

Physical inspection of the valves indicated that these valves, operating at design

basis conditions, were very near their physical fragility limits. The excessive bearing

pressure between the disc and the body guide materials resulted in yielding, spalling,

and gouging of the surfaces. In some of the designs, the guide clearances were large

enough to allow the disc to tilt during closure, resulting In significant damage to the

sealing surfaces. Review of the data indicated that all six valves required more force

during closing against high pressure flow loads than would be calculated using the

standard variables in the industry's motor-operator sizing equations. The test results

also indicated that the stem force equation used by industry for sizing motor

operators is incomplete.

A6857-Equipment Operability

B5529-Fallure of HPCI Steam Line Without Isolation (GI-87)

Hii
EXECUTIVE SUMMARY

The U.S. Nuclear Regulatory Commission (NRC) is sponsoring valve and

motor-operator functionality research at the Idaho National Engineering Laboratory

(INEL). Among the objectives of this research program is a task to determine what

factors affect the performance of motor-operated gate valves and to determine how

weli Industry's analytic tools predict that performance. These tasks support the

NRC's effort regarding Generic Issue (GI)-87 'Failure of [High-Pressure Coolant

Injection] HPCI Steamline Without Isolation.' GI-87 covers three boiling water

reactor (BWR) process lines: the HPCI turbine steam supply line, the reactor core

Isolation cooling (RCIC) turbine steam supply line, and the reactor water cleanup

(RWCU) process line. AU three of the BWR process lines communicate with the

primary system, pass through containment, and normally have open isolation valves.

The concern with the isolation valves is whether they win close in the event of a pipe

break outside of the containment. A high energy steam or hot water release in the

auxiliary building could result in common cause failure of other components

necessary to mitigate the accident.

This research program also supports the implementation of Generic Letter (GL)

89-10 "Safety-related Motor Operated Valve Testing and Surveillance," which is

applicable to all light water reactor (LWR) safety-related motor-operated valves

(MOVs) as well as selected position changeable MOVs in safety-related systems.

One of the major parts of the research program was two full-scale flexible wedge

gate valve qualification and flow interruption test programs, Phase I and IL The

Phase II program was performed in the late summer and early fall of 1989 at the

Kraftwerk Union (KWU) facilities near Frankfurt in the Federal Republic of

Germany. Six valves were tested, three 6-In. isolation valves typical of those used in

RWCU applications and three 10in. valves typical of those used in the HPCI

applications. One of the 6-in. valves was also tested at RCIC test conditions. In all,

seventeen flow interruption tests were performed, seven of these at design basis

conditions.

Two RWCU valves were tested during the earlier Phase I Test Program. As a

result of that work, we expected that the valves would require more stem force to

close than industry normally would have predicted. Therefore, for the Phase II

Program, we set the motor-operator control switches at higher-than-normal torque

values to ensure valve closure and then determine the strengths or weaknesses of a

given valve design from the recorded data.

The test results clearly showed that for the 01-87 concerns, all valves that were

subjected to design basis flow interruption tests required more torque and

subsequently more stem force to close than would be predicted using the standard

industry motor-operator sizing equation for disc load calculations with a common

coefficient of friction. The highest loads recorded were the result of internal valve

damage caused from the high differential pressure loads across the valve disc as it

attempted to isolate flow.

iv
The high loads encountered during the test series raise the concern that some

valves Installed in nuclear power plants may not have large enough motor operators

to ensure closure In the event of a design basis accident. A pipe break in one of the

GI-87 systems (outside containment) Is a low probability but high consequence event.

The study into the phenomena affecting the stem loads in a motor-operated gate

valve continues. However, the results to date Indicate that the phenomena taking

place Inside the gate valve are much more complex than previously thought. The

actual disc factor is much higher than previously believed, but this higher disc factor

can be moderated for some valve applications once the self-closing force balance on

the valve disc is understood.

V
ACKNOWLEDGEMENTS

The efforts of many people and organizations were required to successfully conduct

the Generic Issue-87 test programs. For the Phase II portion of the research

reported herein, we wish to thank Dr. U. Simon and W. Schoeder from Kraftwerk

Union Karlstein; H. Knuedler and AS J. Schmidt from the Bechtel Kraftwerk Union

Alliance; and S. B. Wilchek, T. Branum, and L Thompson from General Physics.

Dr. 0. H. Weldenhamer and Dr. Roy Woods from the U.S. Nuclear Regulatory

Commission assisted both in the planning and in the actual performance of the test

program. We would also like to thank P. Schwieder from the Idaho National

Engineering Laboratory for his help in selecting our diagnostic equipment options

and his programming efforts, which decreased the data reduction time, and M. J.

Russell for his insights and assistance in the interpretation of data.

vi
CONTENTS

ABSTRACT ............................. iii

EXECUTIVE SUMMARY .Iv

ACKNOWLEDGEMENTS .vi

ACRONYMS AND ABBREVIATIONS .xi

1. INTRODUCTION. 1

1.1 Background ........... 2

1.2 Motor-Operator Sizing. 3

2. APPROACH TO TESTING.. 8

3. TESTING .. 15

4. TEST RESULTS AND INTERPRETATIONS .. 20

4.1 Generic Issue 87 .20

4.2 Flexwedge Gate Valve Stem Force Requirements .20

4.3 Parametric Studies .36

4.4 Other Concerns .49

4.4.1 Concern 1: Performing a Design Basis Test .54

4.4.2 Concern 2: Motor Operator Dynamics (Rate of Loading) .54

4.4.3 Concern 3: Critical Test Measurements .59

4.4.4 Concern 4: Motor Performance Characteristics .59

5. CONCLUSIONS .67

6. REFERENCES .68

FIGURES

1. Typical motor-operated valve showing components important to calculating stem force. Two of

the test valves were modified by installing a load cell in the valve stem. 5

2. Valve assembly showing Instrumented spool pieces .10

3. Phase II valve assembly and local piping instrumentation .11

vii
4. Simplified KWU 6-in. valve test loop showing configuration, lengths, and piping sizes .... ........ 12

5. Simplified KWU 10-in. valve test loop showing configuration, lengths, and pipe sizes .... ........ 13

6. Valve stem force history from Phase II testing shows nonlinear performance of the valve,

indicating valve damage ............. .......................................... 17

7. Valve 1 in Phase 11, 6-in. RWCU system valve, same as Valve A in Phase 1, nonhardfaced guides,

refurbished, and tested with the opposite end as the inlet at line break flow, comparing actual versus

predicted stem force ................ .......................................... 21

8. Valve 2 in Phase II, 6-in. RWCU system valve, same as Valve B in Phase I except nonhardfaced

disc guides, line break flow, comparing actual versus predicted stem force .................. 22

9. Valve 3 in Phase 11, 6-in. RWCU system valve, nonhardfaced guide surfaces, line break flow,

comparing actual versus predicted stem force ....................................... 23

10. Valve 4 in Phase II, 10-in. HPCI system valve, nonhardfaced guide surfaces, line break flow,

comparing actual versus predicted stem force ....................................... 24

11. Valve 5 in Phase II, 10-in. HPCI system valve, nonhardfaced guide surfaces, line break flow,

comparing actual versus predicted stem force ....................................... 25

12. Valve 6 in Phase II, 10-in. HPCI system valve, nonhardfaced guide surfaces, line break flow,

comparing actual versus predicted stem force ....................................... 26

13. Valve 2 in Phase 11, 6-in. RCIC system valve test, closing on line break steam flow, comparing

actual versus predicted stem force ................................................ 27

14. Gate valve cross section showing possible valve tippage during flow ......................... 29

15. Typical valve damage observed .................................................... 30

16. Typical valve damage observed .................................................... 31

17. Typical valve damage observed .................................................... 32

1& Typical valve damage observed .................................................... 33

19. Typical valve damage observed .................................................... 34

20. Valve 3 construction details ....................................................... 35

21. Valve 2 closing at full flow. Effect of subcooling on break flow isolation ..................... 37

22. Valve 2 closing at full flow. Effect of subcooling at 1000 psig on the disc factor .... ........... 38

23. Valve 2 closing at full flow. Effect of pressure at 1001F subcooling on the disc factor .... ....... 40

24. Valve 2 opening at full flow. Effect of pressure at 100l F subcooling on the disc factor .... ...... 40

viii
25. Valve 2 closing at full flow. Effect of pressure at 100OF subcooling on stem force/AP ............ 41

26. Valve 2 opening at full flow. Effect of pressure at 100F subcooling on stem force/AP ........... 41

27. Valve 3 closing at full flow. Effect of pressure at 10OF subcooling on stem force/AP .... ........ 42

2& Valve 3 opening as full flow. Effect of pressure at 10F subcooling on stem force/AP .... ....... 42

29. Valve 5 closing at full flow. Effect of pressure at saturation on stem force/AP ..... ............ 43

30. Valve 5 opening at full flow. Effect of pressure at saturation on stem force/AP ..... ........... 43

31. Valve 2 closing at full flow. Effect of subcooling at 1000 psig on stem force/AP .... ........... 45

32. Valve 2 opening at full flow. Effect of subcooling at 1000 psig on stem force/AP ..... .......... 45

33. Valve 2 closing at full flow. Stem force at 1000 psig, cold water ........................... 46

34. Valve disc cross section showing pressure loads ........................................ 47

35. Valve 2 closing at full flow. Effect of pressure at 100F subcooling on the bonnet-to-under disc

pressure ratio . ............................................................... 48

36. Valve 2, closing at full flow. Effect of subcooling at 1000 psig on the bonnet-to-under disc

pressure ratio . ............................................................... 48

37. Valve 2 closing at full flow. Effect of pressure at 100°F subcooling on the under

disc-to-upstream pressure ratio .................... ............................. 50 SO

38. Valve 2 closing at full flow. Effect of suboooling at 1000 psig on the under disc-to-upstream

pressure ratio . ............................................................... 0

39. Valve 3 closing at full flow. Effect of pressure at 10°F subcooling on the bonnet-to-under disc

pressure ratio . .............................................. 5.............. 1

40. Valve 5 closing at full flow. Effect of pressure at saturation on the bonnet-to-under disc pressure

ratio ...................................................................... 52

41. Valve 6 closing at full flow. Effect of pressure at saturation on the bonnet-to-under disc pressure

ratio ...................................................................... 53

42. Damage to the seat and guide rail of Valve 4 because of closure loads ....................... 55

43. Valve B stem force measured at pipe break flow ....................................... 57

44. Valve B stem force measured at no flow ............................................. 57

45. Phase IT, 10-in. HPCI system valve, stem position and stem force at torque switch trip,

comparison for a loaded and unloaded valve test .................................... 58

ix
46. Limitorque calibration of motor operator output torque as a function of torque switch position

for an SMB4-0-25 operator .............. ........................... 61

47. Comparison of stem force versus motor operator torque for various stem nut frictions ..... ...... 62

48. Torque measurements during a 4-in. valve design basis flow interruption test .................. 63

49. Force measurements during a 4-in. valve design basis flow interruption test ................... 63

50. Stem force for Valve 5 closing at normal BWR HPCI operating pressure and temperature .... ... 64

51. Stem force for Valve 5 opening to 30% open and reclosing ............................... 64

52. Motor current for Valve 5, closing at normal BWR HPCI operating pressure and temperature .... 65

53. Motor current for Valve 5, opening to 30% open and reclosing ........... ................. 65

54. Comparison of Limitorque's motor performance curves and observed performance during motor

trip ......................................... 66

TABLES

1. Limitorque operator nuclear ratings .................................................. 4

2. Valve and operator information ..................................................... 9

3. Flow interruption test target temperatures and pressures .................................. 18

4. Valve qualification test pressures .................................................... 19

x

.§,
ACRONYMS AND ABBREVIATIONS

ANSI

ASME

BWR

CEGB

DAS

EG&G Idaho

EPRI

GI

GL

HPCI

IBM

IE

INEL

KWU

LVDT

LWR

MOV

NOP/NOT

NRC

PC

RCIC

RWCU

American National Standards Institute

American Society of Mechanical Engineers

boiling water reactor

Central Electric Generating Board

data acquisition system

EG&G Idaho, Inc.

Electrical Power Research Institute

generic Issue

Generic Letter

high-pressure coolant Injection

International Business Machines Corporation

Office of Inspection and Enforcement

Idaho National Engineering Laboratory

Kraftwerk Union

linear variable displacement transformer

light water reactor

motor-operated valve

normal operating pressure and temperature

U.S. Nuclear Regulatory Commission

personal computer

reactor core isolation cooling

reactor water cleanup

xi
GENERIC ISSUE 87,

FLEXIBLE WEDGE GATE VALVE TEST PROGRAM,

PHASE 11 RESULTS AND ANALYSIS

1. INTRODUCTION

The Idaho National Engineering Laboratory

(INEL), under the sponsorship of the U.S. Nuclear

Regulatory Commission (NRC), is performing

research' to resolve specific generic issues and

develop and Improve Industry mechanical

equipment qualification and operating and

maintenance consensus standards. TIhis overall

research effort includes a program that tested the

operability (opening and dosing) of six full-scale

motor-operated gate valves typical of the

containment isolation valves installed in boiling

water reactor (BWR) reactor water cleanup

(RWCU) process lines and high-pressure coolant

injection (HPCI) turbine steam supply lines. The

valves were qualified and parametrically tested

above, at, and below the pressures, temperatures,

and flow conditions of a worst-case downstream

pipe break in the RWCU and HPCI turbine supply

lines outside of containment. One of the RWCU

valves also was tested with steam to provide

insights for the reactor core isolation cooling

(RCIC) turbine steam supply line isolation valves.

The purpose of the test program was to provide

technical input for the NRC effort regarding

Generic Issue (GI)-87, Failure of the HPCI Steam

Une Without Isolation. GI-87 also applies to the

RCIC and RWCU isolation valves. All three of

the 01-87 BWR process lines communicate with

the primary system, pass through containment, and

normally have open containment isolation valves.

The concern with these containment isolation

valves is whether they will close in the event of a

pipe break outside of containment. A high energy

steam or water release in the auxiliary building

a. Work supported by the US. Nuclear Regulatory

Commission, Division of Engineering, Office of

Nuclear Regulatory Research, under U.S.

Department of Energy Contract No. DE-AC07-

761D01570.

could result In common cause failure of other

components necessary to mitigate the accident.

The test program also provides Information

applicable to the implementation of Generic Letter

(EL) 89-10, -Safety-Related Motor Operated Valve

Tlesting and Surveillance," for all light water

reactor (LWR) safety-related motor-operated

valves (MOVs) and selected position-changeable

MOVs in safey-related systems.

The analyses performed to date on the measured

data obtained during the 01-87 Wive lest

Program and the conclusions derived from the

analyses are discussed in this report A complete

analysis of the data required by the program

objectives will follow in a later report We are

issuing this report now because the findings to date

contain Information that may be beneficial to the

industry when responding to current regulatory

recommendations. Additionally, the Phase II

testing did not answer all of the test objectives

because we could not locate a dc motor in time for

the test program. That work and the open

questions regarding stem factor and rate of loading

will be performed on the INEL valve load

simulator.

For those individuals or organizations who wish

to do their own analysis of the data, the measured

data are available from the NRC Public Document

Room. The Phase I program is reported in BWR

Reactor Water Cleanup System Flexible Wedge Gate

Isolation Valve Quahjficawon and High Enedr Flow

Internuption Test [EG&G Idaho, Inc. (EG&G

Idaho), 1989] (NUREGICR-5406). The Phase II

Test Program actual measured data are reported in

Geneic Issue-87 Flexible Wedge Gate Valve Eht

Program Phase I1 Data Report, (EG&G Idaho,

1990). The Phase II data are also available in

International Business Machines Corporation

(IBM) personal computer (PC)-compatible format.

There is also a video tape documenting the post

1
test disassembly and Inspection of the test valves.

Both the video and the magnetic data can be

obtained through the INEL lTechnology Minsfer

Office at (208) 526-6042.

1.1 Background

Iao sets of experiments have been conducted

for the G1-87 research program. In Phase I, two

full-scale RWCU valves typical of those in

operating plants were tested with high energy

water. The results of the Phase I program, as

previously stated, were reported In

NUREGICR-5406. These results challenged the

validity of many gate valve design rules. The

two-valve sample was small and one valve

sustained damage as a result of the high flow

loading. Because of this, sofne experts in the

Industry did not accept the results as having

general applicability. In the Phase II program, the

NRC increased the valve sample size to six valves

representative of those Installed In both the HPCI

and RWCU Systems.

Phase II test program objectives were

reevaluated and modified to include the lessons

learned in Phase I and to include tests on the

RWCU valves that would be applicable to RCIC

valves. The resulting objectives were as follows:

* Determine the valve stem force required to

close typical RWCU, RCIC, and HPCI

system Isolation valves at typical operating

conditions and under blowdown conditions

. Compare valve closing loads to opening

loads at various system conditions

* Evaluate valve closure force components,

such as disc friction, packdng drag, stem

rejection load, and fluid dynamic effects

* Measure the effects of temperature, pressure,

and valve design on valve closing and

opening loads

* Evaluate the terms and variables In the

present standard valve and motor-operator

sizing equations

* Provide detailed information to assist in the

NRC effort regarding GI-87

* Correlate the data to develop a methodology

for In-situ motor-operated valve testing,

supporting the implementation of GL 89-10.

The results of the testing may also contribute to

specific guidelines being developed to Improve

valve qualification and in-plant test standards such

as American National Standards Institute

(ANSI)/American Society of Mechanical Engineers

(ASME) standards B16.41 (1983), OM-8, and

OM-lO.

Surveys of utility installations performed before

the Phase I program determined that in the BWR

systems of interest, the flexwedge gate valve with a

Limitorquea motor operator was the predominant

configuration. lb avoid duplication, we reviewed

previous applicable industry test programs. The

review included the Electrical Power Research

Institute (EPRI) power-operated relief valve/block

valve testing at Duke Power in 1980, the Central

Electric Generating Board (CEGB), United

Kingdom; gate valve testing performed at

Kraftwerk Union (KWU), Federal Republic of

Germany; and the KWU testing performed for

their own plant designs. The EPRI and CEGB

work were go-no-go tests; however, the results

showed that both flexwedge gate valves and

parallel disc designs did have problems. The KWU

work on their own gate valve designs resulted in

many gate valves being replaced with globe valves.

Where they could not replace the gates, they

developed a structurally stiff design with rather

close Internal clearances. The U.S. has valves with

close and large clearances, but none of our designs

are as structurally stiff as the KWU designs. The

KWU testing indicated that we might expect

trouble based on their findings.

a. Mention of specific products and/or

manufacturers in this document implies neither

endorsement or preference nor disapproval by the

U.S. Government, any of its agencies, or EG&G

Idaho, Inc., of the use of a specific product for any

purpose.

2
The previous test programs did not answer our

01-87 questions, however. It must be noted here

that after reviewing all of our test results, the most

flexible U.S. design performed better than our

moderately stiff designs. Therefore, a valve may

perform correctly without being stiff The valve

design needs to be tested with the worst-case

tolerances to ensure operability.

The NRC Office of Inspection and Enforcement

(IE) Bulletin 85-03, 'Motor Operated Valve

Common Mode Failures During Plant Transients

Due to Improper Switch Settings,' and OL 89-10,

require the utilities to develop and Implement a

program to ensure that the switch settings on

selected safety-related MOVs are set and

maintained correctly to accommodate the

maximum differential pressure expected on these

valves during both normal and abnormal events

within the design basis. The 01-87 valves are a

subset of this larger class of safety-related valves.

Industry has helped to meet these criteria by

developing new MOV diagnostic test equipment

and methods for in-situ valve testing. Prior to the

GI-87 Valve Test Program, the motor-operator

control switches settings were based primarily on

standard industry practices and analysis. Very little

design basis testing had been conducted in or

outside of the plants. Utilities typically verified

the analytically-determined MOV output torque or

stem force through valve seating or backseating

loads with little or no valve hydraulic loading. The

GI-87 Phase I test results cast some doubts on this

industry practice, primarily on whether the true

design basis load can be determined analytically.

1.2 Motor-Operator Sizing

The gate valve is a high recovery positive shutoff

valve and is typically used in systems where

minimal pressure drop is desired when the valve is

open. The design is ideally suited for isolation

purposes and usually is not used for throttling

flow. When the disc is in the seat, the upstream

pressure load on the disc assists in sealing. This

feature is less Important in the flexwedge than in

the parallel disc gate.

There are a number of calculations made to

determine the correct operator size for a given

valve application. The term sizing means that the

output power available from a given motor

operator is determined by its size. lypically, the

larger the operator the higher the allowable

output. Table 1 lists the typical maximum output

in torque and stem force for the various SMB

Model Umitorque operators. This table should

only be used to understand what sizing means.

Requalification of motor operators and other

limitations on valve hardware can affect the use of

a specific operator size.

The four most important calculations that are

made in sizing a motor operator for a specific

valve application are (a) total stem force necessary

to operate the valve at its design basis load, (b) the

operator torque necessary to produce that force,

(c) the operator gear ratio (including the stem nut

thread necessary to produce the needed valve

stroke time), and (d) the size and speed of the

electric motor necessary to produce the needed

operator torque for that gear ratio. lb be

conservative, other calculations are made, such as

degraded voltage concerns, which do not need to

be explained for a basic understanding of

motor-operator sizing. TWo of these calculations,

the gear ratio and the electric motor sizing, appear

to be well understood and the results are

repeatable In application. The required stem force

and the operator torque to produce that force

appear to be the areas that have not been

conservatively predicted In the past for some

classes of valves.

Figure 1, a cutaway drawing of a typical

motor-operated gate valve, shows the components

Important to this discussion. The necessary forces

currently defined by industry to close the valve and

isolate flow must overcome the resistance imposed

by three loads: (a) the disc frictional drag load

caused by the differential pressure across the disc

as the valve closes, (b) the stem rejection load

caused by static pressure on the stem, and (c) the

packing drag load. Industry has developed a set of

equations for use in sizing motor operators. The

first equation in this set predicts the total stem

force, as detailed below. Each manufacturer

modifies the variables in the equation slightly,

however, in the long run the application of the

equation is the same.

3
Table 1. Umitorque operator nuclear ratings

Model-Size

SMB-000

SMB-00

Ratio'

Range

12.5- 30.6

33.5-100.0

102.0-136.0

9.7- 22.0

23.0-109.0

114.0-183.9

SMB-O 11.8- 26.1

26.4- 96.2

102.6-150.8

158.3-247.0

Rated

Torque

9ft-lb)

90

90

90

250

250

190

500

500

500

340

850

850

780

625

1800

1800

1250

950

Rated

Stem force

(lb)

8,000

14,000

24,000

45,000

70,000

Maximum

Threaded

Stem Diameter

(in.)

1375

1.75

2.375

2.875

3.5

SMB-1 11.6- 25.7

27.2- 88.4

92.4-171.6

191.7-234.0

SMB-2 10.6- 25.6

26.2- 82.5

84.8-150.0

153.0-212.5

a. Unit overall gear ratio.

4
Electric

-Direct thrust

measurement

(load cell)

Yoke -

Valve

stem

A-A "-=t

Figure 1. Typical motor-operated valve showing components important to calculating stem force. Two of

the test valves were modified by installing a load cell In the valve stem.

5
Ft = #AAP + AsP + Fp

where

Ft = total stem

force

(1) operators depends on a stem factor calculation

given by

T = psFt

where

(2)

d = disc factor

Ad = disc area

AP = differential

pressure

A = stem cross-section area

P = stem pressure

Fp = packing drag load

(a constant).

I dynamic

component

I static

component

For wedge-type gate valves, the disc factor (ad)

normally used by the industry in Equation (1) is

0.3. Note that In this equation the stem rejection

load can be either positive or negative,

dependingon whether the valve is closing or

opening. This is because the stem rejection load

is always in a direction out of the valve body, this

load resists valve closure and assists in opening the

valve. The packing load is typically constant and

depends on the packing design, the gland nut

torque, and the direction of operation.

The equation is divided into two components,

which will be referred to in the analysis found later

in this report. The components are (a) the

dynamic component, which includes the disc load

due to differential pressure, and (b) the static

component, which is the sum of the stem rejection

and packing drag loads. The pressure values

(P and AP) used in the force equation are supplied

to the valve manufacturer by each individual utility.

Determining the motor-operator torque necessary

to produce that force is complicated by the fact

that motor operators control output torque, not

stem force. Thus, in determining the necessary

output torque one must consider the conversion of

operator torque to stem force. The torque-to-stem

force relationship normally used in sizing motor

T = operator torque

x5 = stem factor

Ft = total stem force [from Equation (1)1.

The stem factor used in Equation (2) is a

function of stem diameter, thread pitch and lead,

and the coefficient of friction between the operator

stem nut and the valve stem. As in Equation (1),

the only variable that cannot be measured in the

stem factor equation is the coefficient of friction.

Normally, it is assumed that only damage and

lubrication of the stem/stem nut threads can

significantly alter the stem coefficient of friction.

Umitorque personnel, in their diagnostic work,

have measured coefficients of friction from 0.10 to

0.20 in actual operation. Losses internal to the

motor operator, up to the capacity of the electric

motor, will typically be accounted for by the torque

spring/switch position. Losses in the stem factor

will not be accounted for by the motor operator.

The problem with the conversion of torqueto-stem force (stem factor) is not in conservatively

bounding it in the sizing calculation. It is that the

stem factor appears to change with stem load.

This will complicate utility efforts to comply with

regulatory recommendations to develop and

implement a program that will ensure that

safety-related motor-operator torque switch

settings (the switch that regulates the

motor-operator output torque) are chosen, set, and

maintained correctly to accommodate the

maximum differential pressure load expected on

the valves during both normal and abnormal events

within the design basis. Additionally, MOV

torque-to-force relationships can vary with age and

maintenance. Torque springs age, changing the

torque switch setting in comparison to output

torque. The stem factor can change for two

6
primary reasons: (a) lubrication quality at the valve

stem-to-motor-operator stem nut interface, and (b)

degradation of the threads of either component

The stem factor can also improve with wear

between the stem and stem nut threads.

The INEL believes the biggest motor-operator

sizing problem today is verifying the capability of

the operators in the plants. This problem is

compounded for gate valves and to some degree

for all rising stem motor-operated valves according

to NRC 01-87 valve test results. This indicates

that the variables used by Industry in the past for

determining valve opening and disc force

[Equation (1)] at high flow were not conservative,

and that the stem factor may vary with load,

making it very difficult for the utility to

diagnostically determine operator capability in

place without design basis testing. This is not

always possible In a plant.

Motor-operator sizing topics that apply to both

01-87 and OL 89-10 will be discussed in more

detail in later sections of this report.

7
2. APPROACH TO TESTING

Six full-scale, representative nuclear valve

assemblies were tested under various normal

operations and design basis pipe break conditions

for the RWCU, HPCI, and RCIC systems. Table

2 lists the test valves and motor operators used in

both Phase I and I test programs.

As shown in Table 2, the two valves used in

Phase I were reused in Phase IL Phase I Valve A

was refurbished and became Valve 1 in Phase IL

The valve's internal manufacturing tolerances

allowed the disc to tip downstream during the

Phase I testing, causing damage to the disc guide

surfaces. Gate valves are bidirectional; therefore,

the valve was turned around for Phase IL, reversing

flow through the valve to see how the internal

manufacturing tolerance stackup in the other

direction affected valve performance. Valve B

from Phase I was returned to its manufacturer for

a valve disc replacement and became Valve 2 in

Phase IL The valve disc or gate in Phase I was

equipped with hardfaced disc guides, representative

of valve assemblies built after 1970. For Phase II,

the hardfaced guide disc was replaced with a

normal material disc guide, representative of those

valves made before 1970. With the exception of

Valve B in Phase 1, all other valves in both phases

had normal carbon steel disc guide surfaces

representative of the largest majority of the

installed gate valves. All valves in both phases had

hardfaced sealing surfaces on both the body and

the discs. Valves 3, 4, and 5 were new valves

obtained through canceled nuclear plant surplus.

These three valves and the two valves used in

Phase I were returned either to their manufacturer

or to a nuclear valve service center for

refurbishment and/or inspection prior to their use

in Phase II testing. Valve 6 was manufactured new

for the Phase II Program. Prior to Phase It

testing, each valve body was assembled with

instrumented spool pieces and instrument taps as

shown in Figure 2.

Table 2 also identifies the 460-Vac, 3-phase,

60-Hz Limitorque electric motor operators used on

the valve assemblies. For Phase II, the same

SMB-0-25 was used on all three 6-in. valve

assemblies, and the same SMB-1-60 was used on

the three 10-in. valve assemblies. The

motor-operator stem nuts and helical reduction

gears were changed to accommodate the valve stem

thread pitch and lead and to establish, as close to

as possible, a 30-s stroke time for all six valves.

For valve stem force measurements, two of the

valves were instrumented with direct stem-mounted

load cells, as shown in Figure 1. The other valves

used a set of four load cells mounted between the

valve yoke and motor operator. During final

checkout of the motor operators, Limitorque ran

a special torque spring deflection versus operator

output torque calibration on their dynamometer.

By using a linear variable displacement transformer

(LVDT) and the Limitorque torque spring

deflection versus operator output torque

relationships, we were able to monitor apparent

motor-operator torque on-line during the test

program. Figure 3 shows the general location of

the valve response instrumentation used during the

Phase II test program.

The Phase II test program was performed in the

late summer and early fall of 1989 at the KWU

facilities near Frankfurt, West Germany. Two test

loops were used. Figure 4 shows the loop used for

the 6-in. valve tests; Figure 5 shows the loop used

for the 10-in. valve tests. Both test loops used a

22-MW off-fired boiler for heating and test media

propellant. The 6-in. valve test stand also had the

capability of being charged with gaseous nitrogen

for cold water high flow testing. Both test stands

were equipped with bypass lines around the loop

blowdown device. These lines were used to

establish normal flow through the valves for

normal-service functional testing. Maximum flow

was established on the 6-in. valve test loop through

a rupture disc and on the 10-in. valve loop through

quick opening valves.

The data collection objectives for Phase II testing

were driven by the lessons learned from the Phase

I program, where the performance of the valves

under higher loadings was unlike previous industry

claims for valve performance. Valve thermal

hydraulic inlet conditions appeared to influence

these performance problems (i.e., the

measurements made during the Phase I program

8
Table 2. Valve and operator Information

Test Hardware

Valve Manufacturer and Valve Identification

Valve Designation

Phase I Phase II Size aass Ap

Anchor/Darline

Velan'

A

B

1

2

Walworth 3

6-in.

6-1nL

6-In.

l0-in.

10-in.

10-in.

900 lb

900 lb

600 b

900 lb

900 lb

Flexwedge Gate

1

Anchor/Darling

Wi. Powell

Velan

4

5 U

6 600 lb

Motor-Operator Manufacturer Size Valve Used

Umitorque

LUmitorque

Limitorque

SMB-0-25

SMB-1-60

SMB-2-40

1, 2,3 and B

4, 5, and 6

A

a. Valves A and B from Phase I were refurbished for Phase II as Valves 1 and 2.

were not sufficient to allow a thorough analysis of

specific thermal hydraulic influences). The Phase

II measurements, Instruments, and data-collection

tools were optimized as thoroughly as practical to

characterize the test valves' and motor-operators'

performance at the various test conditions. Inputs

to these measurement and data collection schemes

included suggestions from experts at the INEL,

from the NRC staff at planning meetings, and from

industry experts at several review meetings.

To quantify the test loop thermal hydraulic

conditions and valve response, an average of 70

channels of information were measured and

recorded on a high-speed tape recorder. The tape

recorder was an Integral part of the INEL data

acquisition system (DAS). The DAS transducer

excitation voltage was measured at the transducer

and the system adjusted the internal voltage to

account for line losses. The outputs from the

transducers were then measured by the DAS and

committed to the tape.

The DAS was also configured to make

Incremental calculations of the stem force

throughout the closing and opening cycles using

the measured values for the parameters in

Equation (1) (differential pressure, disc area, etc.)

and a constant disc factor, either the common 0.3

or a more conservative 0.5. We performed our

calculations in this manner to see how well the

equation modeled the actual stem forces during the

entire cycle. A comparison of the calculated forces

to the measured forces shows where the deviations

9
Temperature Temperature

Dynamic pressure Static pressure Dynamic pressure

ard temperature

Flow

'Rotated 90 for clarity

Figure 2 Valve assembly showing instrumented spool pieces.
t :4D _| : 413 At

I

bd~~~~0

EJ El WE]

Figure 3. Phase II valve assembly and local piping instrumentation.

Legend

Accelerometer (triaxial measurements)

--- *-'m- Strain guages

@3 Pressure (static)

@2d Pressure (dynamic)

[] Temperature

I' 'AP

E}J Valve position

[] Valve current

E] Valve voltage

- 1 Torue spring deflection

E Stem force load cells

[EII Motor speed

[EI Measurements In pipe diameter

04056
Water tank Intermedtlato

tank

Normal flow lIne

From

HP

boiler

22m3 l 3-anid

20-In.

I."

Design pressure: 160 bar (2321 psi) Condensation

pool

Figure 4. Simplified KWU 6-in. valve test loop showing configuration, lengths, and piping sizes.
Water tank

Design pressure: 116 bar (1682 psl)

*0041

Figure 5. Simplified KWU 10-in, valve test loop showing configuration, lengths, and pipe sizes.
start to take place and allows us to look at the

fluid conditions and other parameters and

determine i influences other than the terms in

Equation (1) affect the stem forces during valve

operation.

At various times during the test programs,

parallel diagnostic measurements were made by

Bechtel-KWU Alliance, General Physics, Liberty

Technology, Limitorque, Movats, Westinghouse,

and Wyle Laboratories. This allowed the

diagnostic vendors to compare measured loadings,

similar to those that might be developed during

in-plant testing, to design basis loadings, which

typically cannot be developed in the plant.

14
3. TESTING

Six qualification tests and seventeen flow

interruption tests were performed. Seven of the

flow Interruption tests were design basis tests; the

other ten were parametric studies. The parametric

studies were an attempt to understand the effects

of pressure, temperature, and fluid properties on

valve performance. The fluid conditions and valve

operating responses provided information

concerning valve and motor-operator performance

at various valve loadings.

The two lists below summarize the qualification

and flow interruption test procedures for each

valve. The basic test procedure for each valve

through the design basis flow interruption test was

nearly the same, with a few exceptions (listed

later).

The following list is an overview of the test

procedure performed In accordance with the test

plan for each valve through the design basis flow

interruption test, which was always performed first

on each valve assembly.

* The valve was installed in the test loop

* The motor operator was installed and limit

control switches were set

* Instrumentation was installed

. The torque switch was set

* A baseline opening and closing test without

pressure was performed

* Cold leakage test (Annex A of ANSUASME

B16.41)

* Cold cyclic test (Annex B of ANSI/ASME

B16.41)

* Hot cyclic test (Annex C of ANSI/ASME

B16.41)

* Flow interruption test at the normal operating

pressure and temperature (NOPINOT) for the

valve's representative system (Annex G of

ANSI/ASME B16.41). The details of this step

are provided below.

Exceptions to this basic procedure are as follows:

* Because of operator limitations at the very

high torque switch settings used, some of the

under-voltage tests recommended by the

standard in hot and cold cycling were not

performed.

* The torque switch settings in some cases were

increased after hot cycling to ensure valve

closure. (Our objective was to close the valve

and determine from the measurement how

much stem force was actually required to

isolate flow and seat the valve. In one case

(Valve 1), we underestimated the closing force;

we obtained flow isolation but did not fully

seat the valve.)

The flow interruption test consisted of the

following numbered steps. The numbers reference

the headers of the Phase II plots used in this

report:

* Bring the test loop and valve up to NOP/NOT

(Step 7)

* Close the valve at NOP/NOT without flow

(Step 13)

* Depressurize the downstream side of the valve

(Step 15)

* Open the valve against NOP/NOT upstream

pressure (with the downstream valves closed)

(Step 17)

* Establish normal flow through the test valve by

opening the valve in the normal flow line (see

Figures 4 and 5)

* Close and reopen the test valve at normal

system flow (Step 18)

* Close the valve in the normal flow line

15
* Establish line break flow through the test

valve and close the test valve (Step 25)

* Reopen the test valve 30% open and reclose

at maximum flow (Step 26)

* Depressurize the test loop and open and close

the test valve while still at temperature

without system pressure (Step 30).

Following the flow interruption tests, we reviewed

the quick look plots, which served two purposes:

* To determine If the valve had been damaged

during the test. Valve stem force plots are

very good indicators of valves sustaining

damage during the high flow tests (see

Figure 6).

* To determine if the thermal hydraulic

conditions at the valve inlet met the test

objectives.

If the valve was damaged by the loading (as

determined from a review of the stem force plot),

we stopped testing and Installed the next valve. If

the valve was not severely damaged and the

thermal hydraulic conditions were met, we then

subjected the valve to parametric studies (varying

the pressure and/or temperature). If the thermal

hydraulic conditions were not met, we repeated the

test.

Table 3 provides the test matrix for the flow

interruption test sequences performed on each

valve. The temperatures and pressures listed for

each sequence Include the target pressures and

temperatures for a nearly closed valve. Facility

capability limitations necessitated that some tests

be started with a slightly higher pressure than the

target pressure. In some cases, the valve was

started from less than 100% open in the maximum

flow closures so that the target pressure would be

achieved as closely as possible when the valve was

nearly closed.

Table 4 provides the maximum working pressures

for the valve classes tested in this program and the

test pressures for the leakage and cyclic tests. Test

pressures were not always maximums for the valve

class because of facility limitations; however, they

did bound the qualification pressure for the

intended service.

16
-

C)

o

UE

a) -J 4-.

U)

5000

0

-5000

-10000

-15000

-20000

-25000

-30000

-35000

-40000

-45000

0 5 10 15 20 25 30 35

Time (s) S211 KGD-0490-02

Figure 6. Valve stem force history from Phase II testing shows nonlinear performance of the valve, indicating valve damage.
Table 3. Flow Interruption test target temperatures and pressures

Valve

No.

Test

No.

Target

Pressure

fgsig)

Actual

Pressure

(_DsiI)

Target

Temperature

("El

Actual

Temperature

Eo Media

6-ins Valve Tests

1 1

2

2

2

2

2

2

1

2

3

6A

6B

6C

1000

1000

1000

1000

600

1000

1400

1000

1200

1400

900

950

1040

750

600

1000

1300

920

1100

1300

530

530

545

<100

300

430

480

530

550

580

520

520

550

<100

450

470

520

520

550

570

Hot water

Hot water

Steam

Cold water

Hot water

Hot water

Hot water

Hot water

Hot water

Hot water

3

3

3

1

5

7

lOin. Valve Tests

4 1

S

5

1A

lB

1000

1000

1400

1000

1400

1200

750

800

1040

'990

1400

1100

545

545

590

545

590

570

510

520

550

580

590

550

Steam

Steam

Steam

Steam

Steam

Steam

6

6

6

1A

lB

1C

18
Table 4. Valve qualification test pressures

Maximum ANSI working pressure by

pressure class for these alOlys

Valve No. Class

1,2,4,5 900 lb

900 lb

Temperature

< 100

600

<100

600

Maximum Worldng

Pressure

(pSMi)

2250

1815

1500

1210

3,6 * 600 lb

600 lb

I Cold Lzeakae Test Annex A

Valve No.

Temperature

(*F)

Test Pressure

(Mfig)

1

2

3

4

5

6

<100

<100

<100

<100

<100

<100

2200

2200

500

2200

2200

1500

Cold Cyclic Test Annex B

Valve No.

Temperature

(FE)

Test Pressure

-- (Mg)

1

2

3

4

5

6

<100

<100

<100

<100

<100

<100

1650

1650

1200

1600

1600

1200

Hot Cyclic Test Annex C

Temperature

- * -F)

Test Pressure

Valve No. -(Mfig)

1

2

3

4

S

6

600

600

550

610

610

550

1650

1650

1200

1600

1600

1200

19
4. TEST RESULTS AND INTERPRETATIONS

4.1 Generic Issue 87

01-87 addresses whether containment isolation

valves will close in the event of a pipe break

outside of containment and downstream of the

isolation valves. (The question also applies to the

RCIC steamline and the RWCU supply line.)

Test results provided positive evidence that, given

enough stem force, the tlexwedge gate valve can

close at pipe break flows. Test results also showed

that the stem force necessary to close each of the

test valves was considerably higher than the force

that would normally be predicted using the

standard industry motor-operator sizing equation.

Thus, the potential exists for undersized operators

on these containment isolation valves.

Because of our Phase I test experience, where we

tested two 6-in. flexwedge gate valves at RWCU

linebreak flows, we expected higher stem force

requirements for closing the Phase 11 valves.

Therefore, we sized the operators and set the

torque switches to produce stem forces high

enough to close the valves under any credible

loading based on our previous experience. Our

plan was to close the valves and determine from

the measurements how much stem force was

actually required to isolate flow and seat the

valves.

4.2 Flexwedge Gate Valve Stem

Force Requirements

Figures 7 through 9 show the stem force histories

for the three 6-in. valves closing against high

energy hot water that were representative of line

break flow in the RWCU supply line. Figures 10

through 12 show the stem force histories for the

three 10-in. valves closing against high energy

steam flow representative of line break flow in the

HPCI turbine steam supply line, and Figure 13

shows the stem force history for one of the 6-in.

valves closing against high energy steam flow

representative of linebreak flow in the RCIC

turbine steam supply line.

The stem force histories In these figures were

compared with the forces that were calculated

using Equation (1) with a constant disc factor at

the commonly used industry 0.3 and at a more

conservative 0.5. This was done, as previously

explained, with the DAS using real time values for

the variables in the equation. The actual port size

for each valve (as provided by the manufacturer)

was used in the area calculations instead of other

mean diameters, which are sometimes used by the

industry. The ratio of the actual port size-to-pipe

diameter varies with the class of valve and

manufacturer, resulting in a variation in the degree

of conservatism. All conservatisms were removed

from the calculations so we could compare the

actual stem force histories with the calculated stem

force histories based on Equation (1).

The stem force histories are plotted against time,

representing the stroke time of the valve for that

closing. Due to facility limitations, some tests

were not started with the valve in the fully open

position. During the closing stroke of the valve,

the stem is In compression; thus, the negative

convention of the stem force history. The more

negative the stem force history the greater the

required stem force.

The test results indicate there are generally two

types of valve stem force responses, those bounded

by the calculations and those that are not. Those

that are bounded we call predictable and those that

are not we call nonpredictable. Figures 7 through

13 show several different shapes for the stem force

histories. As we stated, the stem force calculations

are made through the DAS and the same real time

forces are acting on the calculations as on the

valve disc. The only difference between the stem

force calculations and the measured stem force

history is that the calculations have an assumed

disc factor [Equation (1)] and they are not

influenced by fluid properties. Fluid property

effects will be discussed later in the report;

however, the differences for the RWCU 10F

subcooled hot water and the HPCI-RCIC steam

conditions are minor. As we examine Figures 7

through 13, the Major difference in the

relationship between actual and calculated stem

forces is due to the pressure distribution through

the valve and the disc guiding and sealing surface

drag, which are both lumped in to the disc factor

term in Equation (1). None of the actual stem

20
20000

10000

-0~ a.

L-0

o

E

a) -10000

e)

-20000

-30000

0 5 10 15 20 25 30 35

Time (s) S211 KGD-0490-04

Figure 7. Valve I in Phase II, 6-in. RWCU system valve, same as Valve A in Phase L nonhardfaced guides, refurbished, and tested with the

opposite end as the inlet at line break flow, comparing actual versus predicted stem force.
20000 V AV .. JJJI1I. I -JrLLJ~I.~'JI.U ULrlJI-U

10000

0~~

0 -

*+ -10000

C/)

-20000 Actual

---------- Calculated (,t = 0.3)

Calculated (,u = 0.5)

-30000 I I I I I I . I I 1 I I I I , , I I I a I L ,, I 

0 5 10 15 20 25 30 35

Time (s) S211 KGD-0490-06

Figure 8 Valve 2 in Phase II, 6-in. RWCU system valve, same as Valve B in Phase I except nonhardfaced disc guides, line break flow, comparing

actual versus predicted stem force.
20000

10000

30

a)

C.)

L-.

o

UE

a)

+-j

(I)

0

-10000

-20000

-30000

0 5 10 15 20 25

Time (s)

30 35

S211 K(GD-0490-08

Figure 9. Valve 3 in Phase II, 6-in. RWCU system valve, nonhardfaced guide surfaces, line break flow, comparing actual versus predicted stem

force.
40000

30000

20000

Q 10000

-

0

U- -10000

E

wB -20000

en

-30000

-40000

-50000

-60000

0 5 .10 15 20 25 30 35

Time (s) S211 KGD-0490-09

'Figure 10. Valve 4 in Phase 11, 10-in. HPCI system valve, nonhardfaced guide surfaces, line break flow, comparing actual versus predicted stem

force.
40000 -7 vUlve -r. -JI4I4J v, LU. I /j J 1. 4rj- ri IOA

30000

20000

D 10000

U)

O 0

0

IL -10000 X

a) -20000

-30000 :-Instrument noise

-40000 -- :- i. Actual ----------Calculated (A = 0.3) -50000 Calculated (/.. = 0.5)

-60000

0 5 10 15 20 25 30 35

Time (s) S211 KGD-0490-10

Figure IL Valve 5 in Phase 11, 10-in. HPCI system valve, nonhardfaced guide surfaces, line break flow, comparing actual versus predicted stem

force.
40000 u, IU;L rvaiv ;5, LJ: t 5vvv p; 0., JTzJ I voaLur III

30000

20000

3 10000

a) 0

X- -10000 I

.~-20000

C/)

-30000

-40000 0 Actual

- ---------- Calculated (.t=0.3) -50000 - Calculated (A =0.5)

-60000

0 5 10 15 20 25 30 35

Time (s) S211 KGD-0490-11

Figure 12. Valve 6 in Phase II, 10-in. HPCI system valve, nonhardfaced guide surfaces, line break flow, comparing actual versus predicted stem

force.
20000

Because of facility capacity limitations, the test was started

with the valve 60% open.

10000

0 _ ......

0

UE

2 -10000

C,,

-20000 Actual

Calculated (Au = 0.3)

Calculated (/u 0.5)

-30000 .. . ... |.. l.. i-..|.. | .,

0 6 10 15 20 25 30 35

Time (s) S211 KGD-0490-07

FIgure 13. Valve 2 in Phase 11, 6-in. RCIC system valve test, closing on line break steam flow, comparing actual versus predicted stem force
force histories exhibit linear behavior, however, Valve 2 (Figures 8 and 13, tested with hot water

and steam, respectively) follows the general trend

of the calculations with a slowly increasing load.

Thus, the actual valve response is not modeled by Equation (1) calculated stem force histories, but it

can be bounded with a conservative disc factor.

This response we have called predictable.

The nonpredictable valve response is exhibited

in Figures 7, 10, 11, and to some degree, in Figure 12. The stem force history in Figure 7 departs radically from the calculations, and the stem force

load recovers at seating in Figures 10 and 11.

There is no recovery in Figure 7 because the valve

did not fully seat. We will be discussing this

anomaly later in the report; however, in the Phase

I test series we did seat the valve and it exhibited

the same characteristics as seen in Figures 10 and

11. You will also note commonality (the sawtooth

characteristic) in the three stem force histories.

Posttest inspection of these valves showed this

jagged response to be representative of internal

damage occurring to the valve guide and seating surface. The damage took three forms: plastic deformation of the guides, ripping and gouging of

the guide surfaces, and the disc machining off part of the body seat surface during the closing stroke.

The INEL believes all three of the damage mechanisms were allowed by the clearance between

the disc and valve body guide surfaces. The

clearances were great enough to allow the flow

forces to tip the disc downstream, as demonstrated

in Figure 14. Damage was observed when the

clearances were great enough to allow the nose of

the disc to contact the seat sealing surface. The

tipping disc also reduced the contact area on the

guide surface, concentrating the disc load on a

smaller surface area. The yield strength of the

material was exceeded, resulting in one of the

above damage mechanisms depending on the

relative roughness of the disc body guide surfaces.

The rougher surfaces ripped and gouged the

smoother surfaces. Some valves exhibited

combinations of both guide and sealing surface

damage. Those valves without antirotation

features on the stem also showed that stem torque reacting on the disc Influenced the damage. One

side of the disc was damaged more than the other.

Figures 15 through 19 are photographs of

representative damage.

These results indicate that for the valves we

have called nonpredictable, the disc factor

(coefficient of friction) is not a product of sliding friction or other internal valve loadings, but rather

the variable of additional stem force necessary to

overcome internal valve damage. The force

necessary to overcome internal damage is valvespecific depending on the internal clearances. lb

determine this additional force, one must test each

valve. The test would damage the valve, requiring refurbishment. Refurbishment will change the

internal clearances, negating the results of the test.

This Catch-22 may remain until we improve the

Internal design of these valves.

Valve 3 (Figure 9) falls into the predictable

category, but from a different response profile than

the others. All of the other valves are built with

rigid body guides (i.e., there is no flexing in the

disc body guide interface). Valve 3 has a relatively flexible, removable guide. It is anchored only at

the bottom of the body and at the top of the

bonnet; therefore, the guide is free to flex the

entire length of the body (Figure 20 shows the

construction details). Figure 9 shows a dip in the

force history at about the 15-s time line. This plot detail and the posttest inspection of the guide, disc, and body seat show the guide was plastically deformed by the large flow load acting on the disc, allowing the disc to make full seat contact during closing. The guide was not stiff and permitted the

disc to translate in the direction of flow. The disc

closed, bearing squarely on the hardfaced seat.

The sealing surfaces on the disc are also hardfaced.

This hardface-to-hardface contact resulted in the

lowest apparent disc factor (compared to

calculation) of all of the valves tested. All other

guide designs were rigid and resisted the flow

loads. Because of this flexibility, this guide design conformed to the load and the damage was not as

great as that of some other designs. The

disc-on-seat loading did cause some seat damage, however, and the valve leaked after the first

blowdown loading.

From the stem force history and posttest inspection, it appeared that Valve 6 (Figure 12) suffered a combination of both disc guide surface

damage and seat damage. These did not occur

simultaneously, however. During the period of

0-5 seconds, it appeared that damage was occurring

to the guide. Then, as more of the disc entered

the flow stream and more seat contact was made, the disc straightened out.

28
Flow

Electric

DIrect thrust

measurement

(load cell)

* Clearances resulting

in seat damage

* Clearances resulting in

guide damage

Valve

stemDisc

Section A-A

9-3007

Figure 14. Gate valve cross section showing possible valve tippage during How.
8

Figure 15. Typical valve damage observed.
Figure 16. Typical valve damage observed.
Figure 17. Typical valve damage observed.
w

Figure 18. Typical valve damage observed.
Figure 19. Typical valve damage observed.

I
FEigure 20. Valve 3 construction details.
At about the 10.5-s time line, the disc started to

shave the seat. At about the 13-s timeline, the disc

started to chatter across the seat, digging in and

releasing twice before final seating. Posttest

inspection also revealed the lower portion of the

welded-in body guides had plastically deformed

downstream. Valves 2 and 6 were manufactured by

the same company and were basically the same

design internally, but the stem force histories of

the two valves were different. Valve 2 fell into the

predictable category, while Valve 6 fell more into

the unpredictable category.

Valves 1 and 4 were also from the same

manufacturer. Both were in the unpredictable

category, but the damage mechanisms were

different. For Valve 1, the primary damage was to

the seat. For Valve 4, the guide surfaces showed

more damaged. The type of damage experienced

by Valve 1 (Valve A in Phase I tests, where the

flow through the valve was reversed from the

Phase 11 direction), was dependent on flow

direction. In Phase L, the damage was heavier on

the guide surfaces.

The dissimilarity in stem force histories between

valves of different sizes and valves of the same size

with reverse flow, and the fact that valve stem

force predictability appears to be detectable only

by testing will make it difficult to accurately

evaluate valve test data for other valves.

4.3 Parametric Studies

Valves 2 and 3 were both in the predictable

category, and, as seen from the stem force traces,

did not suffer extensive damage during the first

flow interruption test at NOP/NOI These valves

were then subjected to flow interruption tests at

higher and lower pressures and temperatures to

determine the influence of pressure and fluid

properties on valve stem force.

If all of the parametric studies could have

resulted in just one parameter being varied, then

the tests could have been compared to each other

to determine the effect of the parameter (e.g., fluid

properties). That was not the case, however.

Calculating the exact temperatures and pressures

in the test loop accumulators to provide the exact

conditions at the valve could not be verified. This

fact, along with other facility limitations (e.g.,

volume) resulted in tests that cannot be compared

without some type of normalization. An example

of this is shown In Figure 21, where Valve 2 was

subjected to four different fluid states at the same

initial target pressure of 1000 psig.

The individual stem force histories for the four

tests plotted in Figure 21 are very different in

shape and their comparisons to the calculated stem

force histories are very different. The calculated

stem force history with the 0.5 constant disc factor

bounds the steam test stem force history and it

marginally bounds the 10-F subcooled test;

however, it does not bound the other two

subcooled test stem force histories. Note also that

as the degree of subcooling is increased for each of

the tests, the valve stem force history at flow

initiation becomes increasingly more positive. In

fact, the cold water stem force history goes above

the zero stem force line at the initiation of flow,

representing tension instead of compression in the

stem. The stem force histories also appear to be

affected by fluid properties and possibly by forces

other than those identified in Equation (1). The

stem force histories in Figure 21 cannot be directly

compared because the actual test pressures vary

from the target pressure.

Our first attempt at normalizing the test results

was to use Equation (1) to solve for the disc factor

of each stem force history shown in Figure 21.

This comparison is shown in Figure 22. The plot

is read from right to left as the valve closes. As

time increases, the disc factor increases in the

negative convention and is plotted against stem

position. The zero stem position represents flow

isolation. At flow isolation, the area and the AP

terms in Equation (1) become constant. From 0 to

-10%, the stem travel involves seating and wedging,

and there are no terms in the equation to

represent these resistances. The calculation must

be conservative enough to bound these additional

loads. The figure indicates that the disc factor is

influenced by fluid properties, steam being the best

performer and cold water the worst. This is

contrary to the expectation that water should be a

better lubricant than steam.

Our next effort was to determine if the disc factor

was pressure- dependent. Figure 23 shows this

comparison for Valve 2 using the three parametric

tests where the fluid properties remained constant

and the pressure was varied. At flow isolation, the

36
20000

10000

B

a 0 0

UtL

E

!j -10000

-20000

-30000

Phase 11, 6-In. RWCU system valve, parametric study,

line break flow, comparing actual versus calculated thrust.

Vniv nf l4IPIP9gQ nsy.0ZFt0F Sub5ugoId.

Actunl 0

Coklcuated fs . 3

Calculated -j - 0.51

Phase 11 6-In. RWCU system valve, parametric study.

line break flow, comparing actual versus predicted thrust.

20000

10000

3

2 0 0

9L

E

2 -10000 In.

-20000

-30000

0 5 10 15 20 25 30 35 0 6 10 16 20

Time Ws)

25 30 36

lime (s) 5211 Efl0-04100.l 5211K N.X041041.

UJ

Phase 11. 6-In. RWCU system valve, design basis

lIne break flow, comparing actual versus predicted Ihrust.

Phase i1. 6-in. I1WCU system valve, parametric study. line break flow, comparing actual versus calculated thrust.

".v V.6dh4 I Tnsl 2 Siln 26 lnnn nel .I RA1P F It5I-.m u ... I vv..<. ..

20000

i0ooo

2 0 0

UE

!! .10000

Un

-20000

-30000

20000

10000

3

a

E

a,

U)

0

-10000

Actual

--______ Calculatod (it 0.3-

Calculatod /& - 0.51

-20000

-30000

0 6 10 15 20 25 30 36

lime (s) Sit 5 5I 0490461

Figure 21. Valve 2 closing at full flow. Effect of subcooling on break flow isolation.

0 5 10 15 20 25 30 35

Time (s) S2IICI O-041.SD
L0

4-'

C.)

C.)

C),

0.0

-0.1

-0.2

-0.3

-0.4

-0.5

-0.6

-0.7

00

-15 -10 -5 0 6 10 15

Stem position (% of travel)

20

S268 KGD-0790-01

Figure 22. Valve 2 closing at full flow. Effect of subcooling at 1000 psig on the disc factor.
disc factor was not pressure-dependent, though

there was a small spread in the opposite direction

from what one might expect. Here the 1400-psig

test had a lower disc factor than the two lower

pressure tests.

Figure 24 again shows Valve 2 with the fluid

properties remaining constant and the pressure

being varied. In time, the plot reads from left to

right as the valve opens. The spread is not large at

isolation (zero stem position) but the 1400-psig

test has the lowest disc factor. This is not what

one would expect, nor would one expect the disc

factor to increase as the valve is opened further.

Comparing Figures 23 and 24, the magnitude of

the disk factor at isolation in the closing direction

is 0.3 to 0.35, whereas at opening the disc factor at

isolation is 0.5 to 0.55. Unless there are

mechanical effects involved, this is inconsistent

with Equation (1). The opening stem force

assisted by the stem rejection load should not

cause the disc factor to increase.

Our next effort was more basic. We took the

ratio of stem force divided by the differential

pressure for Valve 2, where the fluid properties

were constant and the pressure was varied in both

the opening and closing directions. Figure 25

represents this ratio for closing (read right to left)

and Figure 26 for opening (read left to right). At

flow isolation, the ratios are nearly the same, both

for opening and closing. Again, this should not be

the case if Equation (1) is valid because the stem

rejection load is always out of the valve. Valve 2

has a 1.75-in. diameter stem, which for the

1000-psig test represents a 2410-lb stem rejection

load, 20% of the stem force required to isolate

flow. If the stem rejection load calculation was

sufficient, we would subtract two times the stem

rejection load from the closing stem force to get

the opening stem force. Figures 25 and 26 indicate

that this is not the case.

We performed this same analysis using Valve 3.

Figures 27 and 28 show the closing and opening

stem forces divided by AP for the constant degree

of subcooling tests performed with this valve. The

conditions for these three tests were actual BWR

RWCU conditions of 10-F subcooled with the

pressure varied from 0.667 to 1.5 times the normal

operating pressure. Note the same kind of

pressure distribution seen with Valve 2 and the

same reverse order on opening. The 600-psig test

required more stem force per pound of AP than

the 1400-psig test. The closing traces are slightly

more muddled but show the same general

tendencies. This valve had a 1.5-in. diameter stem,

and for the 1000-psig test should have had a

1770-lb stem rejection load, or more than a 40%

swing between opening and closing. As with

Valve 2, the stem rejection load, which is a

significant percentage of a 6-in. valve opening or

closing stem force, did not show up in the actual

stem forces required to operate the valve. We

know that a stem rejection load exists because we

see it in the stem force load when the valve is

pressurized prior to initiating flow.

Next we looked at Valve 5, a 10-in. HPCI steam

valve. During the first high flow interruption4est,

there was some disc tipping and seat machining as

a result. On subsequent cycles, the stem force

histories looked much like those for other valves

that only sustained minimal damage. Inspection of

the valve internals showed that the disc machined

the body sealing surface from about halfway closed

to the mostly closed position, at which time there

was enough disc-to-seat contact to straighten the

disc. It appeared that after the initial machining of

the conical surface, the wider seatface guided the

disc similar to the performance of Valve 3, where

its flexible guide was plastically deformed and the

disc was guided by the downstream seat. This

analysis is subject to other interpretations, but all

of the evidence indicated that the valve, after the

initial closing, performed as well as any of the

other valves that were subjected to parametric

studies.

The other two 10-in. valves sustained too much

mechanical damage to be used for the thermal

hydraulic studies. These two opening stem force

histories do, however, support our analysis

conclusions for Valve 5. Figures 29 and 30 show

the closing and opening stem force divided by tP

ratios for Valve 5. The first closing cycle at 900

psi does show the effects of seat machining. The

others fall into a more normal pattern. The 10-in.

valve at flow isolation was very much like the

response of the 6-in. valves. The stem force

divided by the AP ratio was roughly 25 compared

to 9 for the steam test on Valve 2. This represents

a response ratio of 2.63 for the 6- to 10-in. valves,

while the port area ratio is 2.93. The 10-in. valve

opening stem force ratio is quite different in

comparison to the 6-in. valve opening stem force

ratios. The highest stem force per pound of AP is

not at seat lift-off, but comes later in the opening

stroke. At 10% open, the stem force to open

39
0.0

-0.1

-0.2 I-0

4-I.

CO

C,

U,

-0.3

-0.4

-0.6 _.

-- ------ -600 Psag -0.6 ____.---------_ 1000 psig -0.6 1400 psig

-0.7

-16 -10 -5 0 5 10 15 20

Stem position (% of travel) S268 KGD-0790-02

Figure 23. Valve 2 closing at full flow. Effect of pressure at 100(F subcooling on the disc factor.

1.0

0.9

0.8

Lto

0

co

C.

to

6

0.7

0.6

0.5

0.4

0.3

0.2

0.1

0.0

-15 -10 -5 0 5 10 15 20

Stem position (% of travel) 5268 KGD-0790-03

Figure 24. Valve 2 opening at full flow. Effect of pressure at 100F subcooling on the disc factor.

40
-3

C.

Ea) Iw

a)

CC4-

0)

a,

'6-

-o

.-._,

-6

-9

-12

-15

-18

-15 -10 -5 0 6 10 15 20

Stem position (% of travel) 5268 KGO-0790-04

Figure 25. Valve 2 closing at full flow. Effect of pressure at 100°F subcooling on stem force/AP.

to

'.0

a)

Co

CO

a)

L0.

.C

15

12

9

6

3

0

-15 -10 -6 0 5 10 15 20

Stem position (% of travel) S268 KGD-0790-05

Figure 26. Valve 2 opening at full flow. Effect of pressure at 100OF subcooling on stem force/AP.

a

41
-0

U1)

L..

U,

0.

-a

C

a)

Ia, to-

!to4-' UT)

IL..

0

-3

-6

-9

-12

-15

-15 -10 - 5 0 5 11

Stem position (% of travel)

10 15 20

S268 KGO-0790-0O

Figure 27. Valve 3 closing at full flow. Effect of pressure at 100F subcooling on stem force/AP.

01.

IU,

U)

0

a)

'a- a

'a.-

t4-.

L- 3-

15

12

9

3

0

-15 -10 -5 0 6 10

Stem position (% of travel)

16 20

S268 KGD-0790-07

Figure 28 Valve 3 opening as full flow. Effect of pressure at 10 0

F subcooling on stem force/AP.

42
v:-

co

0.

0

a)

a)

'a

4-0._

C

a,

L.

4-

'4-

:24.'

a,

Z-

-C

0

-56

-10

-15

-20

-25

-30

-35

-40

-15 -10 -6 0 5 10 16 20

Stem position (% of travel) S2M8 KGD-0790-08

g 2 .

Figure 29. Valve 5 closing at full flow. Effect of pressure at saturation on stein force/AP.

a)

L-.

.0

co

C',

-a

a,

(D

0.

'4-

%I4-

4IE

40

35

30

25

20

15

10

6

0

.5 5 5 5 I * S S S S S I * * ** SI S I * * S S

* --*-*-- 900 psig

* 1000 psig

-5--.-

I- *-*** -.----.

-.- 

,, -.- **45

,/ 2*** V.

-.4

--4.

I; --

S S S S S S S * I a sq I . . . a I . a . . S S S S I' * * ti 

-15 -10 -6 0 6 10

Stem position (% of travel)

15 t 20

5268 KGO-0790-O9

Figure 30. Valve 5 opening at full flow. Effect of pressure at saturation on stem force/AP.

43
equals the stem force to close at isolation. The

other 10-in. valves show this same phenomenon,

but the 10-in. valve response still does not show

the stem rejection load effects that we would

expect from Equation (1).

Next, we will look at our final parametric study

where the pressure, within the capability of the

facility, was constant and the fluid properties were

varied. This study again was performed on Valve

2, one of the 6-in. valves. Figures 31 and 32 show

the closing and opening stem force divided by AP

Within a given fluid, the opening and closing ratios

are close at flow isolation. However, the responses

to the fluids are very different. Steam has the

lowest stem force-per-psi ratio, cold water the

highest. The cold water opening test also shows

the same stem force response characteristics as the

10-in. valves, where the highest stem force is not at

unseating but later in the opening stroke.

The cold water closing stroke also has a unique

shape. Figure 33, the actual stem force history for

this closing, shows this response. The system is

pressurized at the zero line, and from zero to 2-s

the effects of a 1500-lb stem rejection load are

seen. At flow initiation, the 2-s time line, the stem

force takes a positive response. Because of facility

capacity limitations, the valve was 40% closed

prior to flow being initiated, so the disc was well

into the flow stream. The valve was commanded

to close at about the 3-s time line. We saw a small

reduction in the tension load here, but it remained

over the zero stem force line for the next 5 s of

valve traveL The magnitude of the stem force

history deviation from the calculated stem force

histories led us to our first real breakthrough in

understanding what was occurring inside of the

valve. Figure 34 shows the details of the areas

where pressure forces can act on the disc and stem

and where we drilled the three pressure

measurement ports into the valve bodies for

pressure distribution studies.

Review of the bonnet, upstream, under disc, and

downstream pressures, and the various internal

valve surfaces where these pressures can act has

started to provide some understanding of the valve

behavior outside of the loads accounted for in

Equation (1). Prior to flow initiation, a stem

rejection force as predicted by Equation (1) was

observed. However, after the initiation of flow, the

stem rejection component always decreased and, in

the case of cold water, reversed its direction.

Closer investigation of this phenomenon revealed

that before flow initiation, the pressures

throughout the valve were identical Thus, the

pressures acting on top of the disc that exerted a

downward stem force were identical to those

pressures acting on the bottom of the disc that

exerted an upward stem force. The upward and

downward stem forces did not balance, however,

because the areas on the top and the bottom of

the disc were not the same, differing by the area of

the stem. Thus, the classic stem rejection term in

Equation (1) is correct when the internal pressures

are identical

When flow started, however, pressure imbalances

developed within the valve, which affected the net

force balance on the disc. The figures on the

following pages present pressure ratios and provide

insight into the pressure loading forces and how

they affect valve performance. The first two

figures show the Valve 2 ratio of the bonnet

pressure to the under disc pressure versus valve

stem position. Figure 35 presents the effect of

pressure at a constant 100XF subcooling, whereas

Figure 36 presents the effect of subcooling at a

constant upstream pressure of 1000 psig. Notice

that changing the pressure at a constant degree of

subcooling had little effect on this pressure ratio,

which remained relatively constant In the 1.03 to

1.08 range. However, varying the degree of

subcooling had a very profound impact on the

pressure ratio; steam and the 10*F subcooling

liquid having the lowest ratio (1.02 to 1.04) and

cold water having the highest (1.10 to 1.25).

The area on top of the disc was less than the area

on the bottom of the disc, differing by the area of

the stem. Thus, a bonnet-to-under disc pressure

ratio in excess of 1.0 did not imply that a

downward or self closing force existed. Instead,

the pressure imbalance had to be large enough to

overcome the area differential, which is the classic

stem rejection term of Equation (1).

For Valve 2, the ratio of the area on the bottom

of the disc to the area on the top of the disc was

roughly 1.10. Although there were other pressures

acting on the disc that affected the net force

balance, a pressure ratio in the vicinity of 1.10

would have been necessary before a self closing

force would have been expected. A smaller

pressure ratio would have promoted a net outward

force, but this force would have been much less

than the classic stem rejection force.

44
-3

a,_

04)

C',

a)

Ia)

CL

-a

4-

(1)

I-

-6

-9

-12

-15

-18 '--

-15 -10 -6 0 5 10 15 20

Stem position (% of travel) S268 KGD-0790-10

Figure 31. Valve 2 closing at full flow. Effect of subcooling at 1000 psig on stem force-AP.

co

.0

.a)

C')

a)

C

a)

4-.

!-C

:I2

15

12

9

6

3

0

-16 -10 -5 0 5 10

Stem position (% of travel)

15 - 20-

S268 KGD-0790-11

Figure 32. Valve 2 opening at full flow. Effect of subcooling at 10(0 psig on stem force/lP.

45
20000

10000

-0

C) 0

0

LL

EA) -10000

-20000

-30000

0 5 10 15 20 25 30 35

Time (s) S211 KGD-0490-18

Figure 33. Valve 2 closing at full flow. Stem force at 1000 psig, cold water.
pressure

0e3055

Figure 34. Valve disc cross section showing pressure loads.

47
1.2

o 1.1 _ 4-P

_

Lu

L.w

X~ 1_ _

D

I-.

0.9 _

0.8 - _ _ _

-20 0 20 40 60 80 100

Stem position (% of travel) 5268 KGO-0790-12

Figure 35. Valve 2 closing at full flow. Effect of pressure at 1Q00F subcooling on the bonnet-to-under

disc pressure ratio.

1.4

1.3

0

IU

C

0

a.

1.2

1.1

1.0

0.9

0.8

0.7

-20 0 20 40 60 80

Stem position (% of travel)

100

S268 KGO-0790-13

Figure 36. Valve 2, closing at full flow. Effect of subcooling at 1000 psig on the bonnet-to-under disc

pressure ratio.

48
The effect of subcooling was very evident In other

pressure comparisons throughout the valves. For

instance, the pressure under the disc relative to the

upstream pressure, as shown In Figure 37, displays

the effect of fluid subcooling. The difference was

the result of fluid acceleration through the valve

and the expansion potential of the fluid state. For

Instance, as cold water accelerated through the

valve, the pressure dropped. The fluid state had

virtually no expansion potential and was also

sufficiently removed from saturation so that

flashing did not occur. Thus, this fluid state had

the largest pressure drop In the throat region of

the valve relative to the upstream pressure. Steam,

on the other hand, could expand but had no

flashing potential. Thus, the pressure in the throat

region of the valve dropped relative to the

upstream pressure, but to a lesser extent than cold

water.

The 10 and 100-F subcooled states had the

potential to both flash and expand. Thus, the

pressure ratio did not drop to the extent observed

with the cold water and steam fluid states.

Furthermore, the 10F subcooled fluid was closer

to saturation than the 100-F subcooled fluid and

thus flashed easier. As a result, the drop in the

pressure ratio of the 10-F subcooled fluid was less

than the drop of the 100-F subcooled fluid.

Conversely, Figure 38 reveals that the effect of

pressure on the upstream-to-under disc pressure

ratio was relatively nonexistent. Although there

appeared to be a trend,' closer examination

revealed that all the pressure ratios were relatively

constant in the 0.80 to 0.84 range. Any observed

spread was more likely the result of slight

variations in the actual subcooling of the fluid

ratio than the result of a pressure effect. The

other pressures monitored on other valves during

the tests displayed similar trends relative to

subcooling and pressure. As such, they are not

presented.

Having observed several pressure trends in Valve

2, it was Interesting to see if these same trends

existed in the other valve pressure ratios.,

Remember, however, that parametric testing with

the other valves involved pressure variations only.

Only Valve 2 included a parametric study 'that

addressed subcooling.

Figure 39 presents the bonnet pressure-to-under

disc pressure ratio versus stem position for

}ave 3. The ratios are indeed very similar to

those observed for Valve 2 and again reveal no

effect of pressure. The other pressure ratios also

reveal trends similar to those observed in Valve 2.

As such, they will not be presented.

Figure 40 presents the bonnet pressure-to-under

disc pressure ratio for Valve S. Remember that

this was a 10-in. valve, whereas the previous valves

were 6-in. Although this pressure ratio exceeded

1.0, the magnitude was much less than that

observed with the smaller valves. The pressure

ratio for 10-in. Valve 6 (Figure 41) displayed

similar results. These larger valves had larger

internal clearances than the smaller valves. Such

differences may have affected the internal flow

path resistances, flows, and the resultant flow and

pressure distributions. As such, the resultant

performance of a valve may not have been truly

scaled to size. The pressure ratios discussed earlier

are virtually identical to the trends presented with

Valve 2. As such, they will not be presented.

We have not made all the physical valve

measurements necessary to create a complete

investigation. However, it is apparent from the

responses of the valves during the various

parametric studies that Equation (1) is incomplete

and that the missing terms have a first-order effect

on the observed responses of the valves.

Although not quantitatively defined to date, the

loads that will most likely be included in a more

complete stem force equation are disc drag,

packing drag, a net force balance on the disc (top

to bottom), and fluid properties effects.

4.4 Other Concerns

lbst results from the gate valve test program

have shown there were two classes of valves and

valve responses to the flow interruption test loads:

valves bounded by the disc load calculations (albeit

at a higher disc factor than previously used by

industry) and those damaged by the loading. The

amount of damage and the resulting loads were

vale-specific, depending on the stack up of

internal valve tolerances and clearances, which we

consider nonpredictable even with a large disc

factor to cover the unknowns.

49
1.1

1

0

4-'

E

.U

aa.

0.9

0.8

0.7

600 Psig ---- ----- 1000 psig

1400 psig 0.6

0.5

-20 0 20 40 60

Stem position (% of travel)

80 100

5208 KGD-0790-15

Figure 37. Valve 2 closing at full flow. Effect of pressure at 100°F subcooling on the under

disc-to-upstream pressure ratio.

1.1

1.0

0

E

0.

a-

-5-

0.9

0.8

0.7

0.6

0.5

0.4

0.3

-20 0 20 40 60 80

Stem position (% of travel)

100

5268 KGD-0790-14

Figure 38. Valve 2 closing at full flow.

pressure ratio.

Effect of subcooling at 1000 psig on the under disc-to-upstream

50
It

1.2

0

CU

L.

0~

C

C

0

.0

0L

1.1

1.o

1000 psig

----------- 1200 psig

1400 psig

tA

* 0.9

0.8 I-

-20 0 -20 40

Stem position (%

60

of travel)

80 100

S268 KGD-0790-16

Figure 39. Valve 3 dosing at full flow. Effect of pressure at 10°F subcooling on the bonnet-to-under disc pressure ratio.
1.2 I I I I a III * * *U II I I I I I I I I I I I

1.1 P 0

Ca._

w

CL.

z

0-

c8)

C

1.0

(A

w~

900 psig

----------- '-- 1000 psig

L I I I I t- - I I I I I I . a I I I a

0.9 _

0.8

-20 0 20 40 60 80 100

Stem position (% of travel) S268 KGD-0790-17

Figure 40. Valve 5 closing at full flow. Effect of pressure at saturation on the bonnet-to-under disc pressure ratio.
1.2 - - - - - - - - . . . . . .

I I I I I I I I I I

0

4 -P

w)

0~

*0

C

C

.0

.0 a1.1 F1.0

1000 psig

1200 psig

1400 psig

I I I I a I I . A I I I I I I Ig I . I I

0.9

0.8- -

- - -

-20 0 20 40 60 80 100

Stem position (% of travel) S268 KGD-0790-18

Figure 41. Valve 6 closing at full flow. Effect of pressure at saturation on the bonnet-to-under disc pressure ratio.
In addition to the GI-87-related test results

described to this point, we have identified other

concerns from the testing that may help the

industry as they attempt to understand the

concerns of GL 89-10 and GI-87.

4.4.1 Concern 1: Performing a Design Basis Test

Not all nuclear plant valves that open or close

communicate with large reservoirs such as the

primary or secondary system vessels. Some,

however, must be capable of opening against large

differential pressure loads or isolating potentially

high pressure flows. Valves that can be tested

in-situ at the worst-ase loads, including the design

basis pressure, temperature, and flow, are not a

part of this concern once they have been

successfully tested. These in-plant tests will

determine the margin available in the operator and

careful examination of the stem force, torque, and

current and voltage plots should indicate possible

damage from the loading.

Valves that must operate at potentially high loads

and cannot be tested in place at those loads are

the concern here. This includes valves that cannot

be tested with the full design basis pressure,

temperature, and flow loading over at least 60% of

the full closure stroke. We have found that closing

flexible wedge gate valves from less than this

amount will not identify valve designs that are

subject to nonpredictable performance. Thus, the

actual stem forces required for closure cannot be

detected.

Careful inspection of the valves after the GI-87

blowdown tests showed that valve damage started

at about the two-thirds open point on the seats

and guides. The buildup of material in the guides,

or the depth of the seat machining at final closure,

was influenced by this early damage. Early and

late damage affected the maximum stem force at

flow isolation. Measured stem forces at final

seating, the last 5-10% of the closure, were less

than at isolation in some cases.

Figure 10 contains the stem force history for

Valve 4 during the simulated full-scale pipe break

closure stroke. Note how the stem force history

became very jagged when the valve was

40% closed. This jagged shape typically indicated

damage to the valve during the flow isolation cycle,

resulting in higher closure stem forces than would

be expected from sliding friction. Also, note how

the stem force decreased in Figure 10 during the

last 15% of the closure. Figure 42 is a photograph

of the damage to the valve guide rail observed

during posttest inspection of Valve 4. Here we see

the damage to the guide rail that began at about

the 40% closed position and increased throughout

closure to flow isolation. It is unlikely that a

closure of only the last few percent would have

caused the same damage and revealed the high

closure forces necessary to obtain flow isolation at

the design basis loading.

4.4.2 Concern 2: Motor Operator Dynamics (Rate

of Loading)

The research program also found that

measurements of torque, stem force, and motor

performance are needed to completely characterize

MOV performance. The measurement of torque

or stem force alone will not identify problems in

the conversion of torque to stem force (such as

abnormal high stem factors) [Equation (2)]. At

normal valve loading, torque and stem force data

can be misleading because of motor-operator

dynamic response (rate of loading). The rate of

loading phenomenon is a highly visible issue in the

industry today and its basis is not understood

thoroughly. We believe that the effects of rate of

loading are seen only when a valve does not seat

fully enough to develop its potential seating

stiffness. The best description of the problem is

that the delivered stem force at torque switch trip

is greater when the valve stem is loaded at a faster

rate.

Limitorque informed the INEL that on rare

occasions they have also seen the rate of loading

effect on motor operators on their dynamometer

testing. The difference in delivered torque was

small. With the advent of in-situ diagnostic

testing in the plants, there have been a number

valves experiencing the rate of loading effect. This

is typically found when they compare the difference

in delivered stem force for a no load static test and

a differential pressure test. Industry has developed

many complicated analytical theories to explain the

phenomenon. From our experience, we believe the

most valid theory is the simplest one. If less stem

force is obtained in the differential pressure test

than in a no load static test, the valve application

is operating in the margin between fully seating

and some smaller degree of closure.

54
Figure 42. Damage to the seat and guide rail of Valve 4 because of closure loads.
We saw the effects of rate of loading twice in our

test programs: once on Valve B during Phase I

testing, and once on Valve 1 during Phase II

testing. Both times, the rate of loading effect was

the result of not setting the torque switch high

enough to fully seat the valve under the differential

pressure loading. In Figure 43, we see the stem

forces measured as Valve B closed against three

different pressures at pipe break flow in Phase I.

Note that at the same torque switch setting, the

stem force when the torque switch tripped in the

600-psig test with high flow is less than (18,100 lb)

the stem force when the torque switch tripped in

the no-flow static pressure test (19,900 lb)

(see Figure 44).

The valve closing at 1000 psig shows a higher

load before Isolation of flow than in the 600-psig

test. Just before this test, the valve stem was

lubricated and a slightly higher stem force

(18,600 lb) was obtained when the torque switch

tripped. However, the valve stem position and the

subsequent reopening of the valve indicated that

the valve was lightly seated and the measured stem

force was a greater reflection of closing load than

of seating load. During the closing at 1400-psig

inlet pressure (the design basis for this operator

sizing and the torque switch setting), the valve

marginally isolated flow but did not seat, and the

operator tripped on disc load. The stem force

when the torque switch tripped was lower (at

16,500 lb, a 17% reduction in the stem force at

torque switch trip and a 25% reduction in final

stem force) than with the lightly loaded case

(Figure 43).

Figure 45 shows the same rate of loading

comparison for Valve 6. 'No stem force traces

and two stem position traces are shown on the

same plot. We have shown only the last 4.5

seconds of these full valve closures for clarity. The

Step 30 trace is unloaded and Step 25 is the line

break differential pressure test. You can see that

in both of the stem force traces the rate of loading

is equal prior to torque switch trip and the stem

force at torque switch trip is the same. The stem

position traces show a slight variation In operator

momentum because of the loading but it is not

significant. The time base for each trace was

shifted slightly for clarity. This valve does not

show the rate of loading effect. The valve seat was

the resistance to motor operator momentum in

both cases prior to torque switch trip and the valve

achieved the same internal stiffness resulting in the

same stem force.

To understand the rate of loading issue, one must

visualize the two motion paths in the motor

operator

* The worm drives the worm gear, resulting in

valve stem motion

* The worm climbs the worm gear, compressing

the torque spring to torque switch trip.

The energy output from the electric motor takes

the path of least resistance, and the fraction shared

by each is determined by the relative resistance or

stiffness of each path. When a valve is tested

under low load conditions, the seat is the first

significant load the motor-operator encounters.

When this happens, the valve becomes a highly

resistive or stiff assembly and the first motion path

becomes essentially static. All the energy from the

motor is then transferred to compressing the

torque spring, virtually eliminating the losses

associated with the first motion path. The

assembly is very stiff, and the work input results in

high stem forces.

When a valve is tested under high differential

pressure conditions, which results in a high stem

load, the spring pack is being compressed

simultaneously with the rotation of the stem nut.

Thus, the motor output energy is split between the

two motion paths and their frictional losses. The

additional inefficiencies can result in lower stem

forces compared to the low load condition if the

torque switch is tripped prior to achieving the

same seating stiffness as in the unloaded case.

Therefore, the relationship of motor-operator

torque-to-valve stem force with both motion paths

active can only be ensured under design basis

loading.

If the loads during design basis closure can be

overcome to reach the same structural stiffness

prior to torque switch trip, there will only be one

motion path active when the torque switch trips,

and the resulting stem force will be primarily the

same as in the unloaded case. Once determined,

the no-load relationships can be baselined and

monitored for degradation. Valves that experience

the rate of loading phenomenon at torque switch

trip are not as fully seated as those that do not.

56
Valve S, Stop S, Tests 4, 2, and 3 (Una break flows)

°x CD -5i CD

K

.e

-t° -Valve B. 00 pslg. 430 IF

- Valve S. 1000 psl. 530 'F

15 --Valve S. 1400 pslg. 560 'F

Turt

-20'Tr

-25 ., I, I, I .

0 S 10e 1 20

Time (sec)

Figure 43. Valve B stem force measured at pipe break flow.

25 30 35

010012M

Valve 8, Test 2. Step 1, 1000 pslf, 530 'F (no flow)

10.0

_ 10.

E 0.0

0

-0 .0

- -10.0

2 -20.0

-30.0

0 5 10 15 20

Time (sec)

25 30 35

Figure 44. Valve B stem force measured at no flow.

57
VdI1V U. IxI Iat -t 4UIItUo l "I1 LVUsI anlu IU .11J

Loaded test,

-10000 . '* - step 25,

-10000 ; 5300F. 1000 psi

I'. ---------- Unloaded test,

step 30,

-20000 500°F, 0 psi 0.0

00

U..-30000 \\c

E ,.,,-Rata of loadin E

1-0 .. becomes equa 0)

-40000 -0.5 *5

Torque switch trip

-50000

-60000 * * I I * * I . I * * p .. -1.0

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5

Time (s) S298 KGO-1090-03

Figure 45. Phase II, 10-in. HPCI system valve, stem position and stem force at torque switch trip,

comparison for a loaded and unloaded valve test.

58
These valves are operating in the margin between

successful and unsuccessful closure. When

possible, increasing the torque switch setting

enough to fully seat the valve at design basis

loadings will reduce or eliminate the rate of

loading effect.

4A.3 Concern 3: Critical Test Measurements

Our experience has shown that torque, stem

force, and motor current and voltage should be

measured during all valve tests in addition to

motor operator switch position. As discussed

above, torque and stem force measurements will

provide the analyst with the stem factor (ratio of

torque divided by stem force) for the valve. The

motor current and voltage measurements allow the

analyst to determine motor margins.

If the utilities could subject a valve to an in-situ

design basis test every time testing is required by

maintenance activities or regulation, they may not

need to subject the valve to diagnostic testing at all

(valves would have to be checked for damage).

Routine design basis testing is not typically done in

the plant. Instead, a utility may ideally subject a

valve to prototypical or in-situ testing at design

basis loads and then use diagnostic measurements

to determine the force required to open or close it.

This establishes the baseline data for the valve.

Testing then could be done under static conditions

(no pressure or flow in the line) to determine

functional degradation.

Measuring both torque and stem force during the

baseline differential pressure and static tests

establishes the torque-to-stem force ratio (stem

factor) for the valve. This is particularly important

for valves that exhibit any rate of loading effects in

the differential pressure test. These baseline

values can then be compared to the calculated

values for a given motor operator and valve

assembly to determine the credibility of the

measured values and to identify possible problems

In the stem nut-to-valve stem conversion.

We recently were involved with design basis

testing of a 4-in. gate valve, which was not part of

the Phase I or II test program. The measurements

of torque and stem force did not correspond,

indicating a problem in the stem factor or in the

credibility of the measurements. The 4-in. valve

had a 1-In. diameter stem with 1/6 pitch and lead

thread configuration. Using the Umitorque

dynamometer-determined torque switch position

versus output torque calibration and the power

thread calculation equations, we constructed a

calculated stem force-versus-torque switch position

plot for various stem nut coefficients of friction

(stem factors). Figure 46 is based on the

Limitorque calibration of motor-operator output

torque as a function of torque switch position.

Using stem nut friction coefficients varying

from 0.1 to a worst case of 0.2, we constructed the

stem force versus torque curves shown in

Figure 47. These curves allow one to determine

what the stem/stem nut coefficient of friction is

(within the 0.1 to 0.2 friction range) using both the

actual torque and stem force measurements.

Figures 48 and 49 are the actual torque and

stem force measurements, indicating 144-ft-lb

torque and 9800-lb at torque switch trip. However,

Figure 47 indicates the Intersection of these two

points to be off-scale with regard to the 0.1 to

0.2 friction range. The rate of loading shown in

the torque and stem force plots and the overshoot

in the torque plot, indicate that the valve was

seated. Comparing the measured stem force with

the calculated stem force indicates that either the

stem nut had a very high coefficient of friction or

that the torque or stem force measurement was in

error. The torque measurement for a torque

switch setting of 2 appears to be correct. The

measured motor current was comparable to the

Umitorque motor-operator calibration for a torque

switch setting of 2.

A faulty stem force measurement without a

reliable torque measurement could have led the

analyst to erroneously recommend a higher torque

switch setting, possibly overstressing the valve

and/or operator components. (The sister valve to

this one required a torque switch setting of 2.5 to

obtain approximately the same stem force as the

test valve with a setting of 2.0.) If the coefficient

of friction had been excessive, we believe the

problem in the valve could not have been

discovered without both measurements.

4.4A Concern 4: Motor Performance

Characteristics

In anticipation of higher stem force demands,

Limitorque is in the process of requalifying some

of their operators for greater stem force capacities.

In some cases, utilities may need to set their

59
torque switches higher and it will be necessary to

determine if the valve structural capacity and the

electric motor capability of the motor operator can

take the higher loads. rpicalWy In the past, the

diagnostic motor current measurement has been

made for trending purposes. Our research has

shown that comparisons between the motor

manufacturer's motor performance curves and

measured current and voltage are very important for evaluating motor operating margins. Included

In this are the calculations that should be made for

degraded voltage concern and the length and size

of the power cables. A small amount of motor or

cable heat can have a large effect on a marginally sized motor's performance.

An example of the motor capacity problem is

shown in Figures 50 and 51. These plots show two

consecutive stem force traces for Valve 5, closing

to isolate full pipe break flow at normal

BWR HPCI operating pressure and temperature, then opening to30% open, and reclosing to Isolate

full flow again. The time axis indicates that the

two steps amount to about 38 s of highly loaded

operation for this 15-min duty cycle motor. A

look at the current traces (Figure 52) shows that

during the first closure, the current Is exceeding or

has exceeded the 20-A. maximum range of the

transducer at torque switch trip. A "normal trip

was experienced, however. In the reopening and

reclosing step (Figure,53), the current again exceeds the maximum range of the transducer at

the end of the valve stroke, but the motor is never

able to generate enough torque to reach torque switch trip, resulting in a stall as evidenced by the

current transducer remaining saturated for about

4 s before we manually stopped the motor.

Figure 54, which shows the manufacturer's

motor performance curves, gives some Information

about this behavior. The solid line Indicates the

expected performance of the motor, both motor

speed and current-versus-motor torque. During the test, we observed that the voltage at the motor

dropped from 460 V ac to as low as 380 V because

of the response of the power supply, the resistance

in the motor cables, and the stall or near stall

currents. The dashed line In the motor speed plot shows the expected motor performance, corrected

for the actual measured voltage. Actual motor

performance observed during this test is shown by the dotted line, indicating that the motor was

actually performing as anticipated at the measured

voltage.

Unfortunately, we designed our current

measuring system based on a Limitorque recommended 40-ft-lb motor. Limitorque later

notified us that a 60-ft-lb motor would be

necessary for Valve 5. If our measurement system could have handled the unexpected higher current

reading from the larger motor, we could have

checked the current demand after the design basis

test. We then could have determined whether we

were operating the motor so far out on the knee of

the motor torque speed curve that there was little

safety margin left in the motor, or that a voltage drop was driving the motor's poor performance.

In our case, both conditions occurred. The

important thing to remember, however, is that the

verification of the motor margins can prevent a

failure of the valve to function at the design basis

load condition. Valve, operator, and motor

momentum will typically carry a marginal motor

through a normal torque out at low or static loads.

60
160 I 5.vIu-vv-du

150

.,,140

4-

a 130

120

.. ° t g 12 0 / / 0 0 1I..

Go 100-/......,,,,..

90

80

1.0 1.5 2.0 2.5

Torque Switch Setting - S211 KGD-0490-73

Figure 46. Limitorque calibration of motor operator output torque as a function of torque switch position for an SMB--25 operator.
30000

25000 / _

o 20000 01/

0

UE

CD15000020

10000

5000 I , I I , , I I I I . @ * * *-

80.0 90.0 100.0 110.0 120.0 130.0 140.0 150.0 160.0 170.0 180.0

Operator Torque (ft lb) S211 KGD-0490-74

Figure 47. Comparison of stem force versus motor operator torque for various stem nut frictions.
4-,

'4-

ci-M

0

FH

-80

-90

-100

-110

-120

-130

-140

-150

-160

-170

-180 _ _ _ _ _

0 6 10 15 20 265 30

Time (s) S211 KGD-0490-71

Figure 48. Torque measurements during a 4-in. valve design basis flow interruption test.

8000

6000

4000

2000

0

a)

'-2000

0

-4000

-6000

-8000

-10000

-12000

0 5 10 15 20 25

Time (s)

30

S298 KGO-1090-02

Figure 49. Force measurements during a 4-in. valve design basis flow interruption test.

63
40000

30000

20000

D 10000

0

ua O

0

U- -10000

E

2 -20000

Cn

-30000

-40000

-60000

-60000

0 5 10 15 20 25 30 35

Time (s) S211 KGD-0490-75

FIgure 50. Stem force for Valve 5 closing at normal BWR HPCI operating pressure and temperature.

40000

30000

20000

3 10000

a)

o- 0

0

U. -10000

E

+ -20000

-30000

-40000

-50000

-60000

0 5 10 15 20 25 30 35

Time (s) S211 KGD-0490-79

FIgure 51. Stem force for Valve 5 opening to 30% open and reclosing.

64
25

zL

E

(U

0)

L0

4-

20

15

10

5

0

0 5 10 15 20 25 30 35

Time (s) S211 KGD-0490-76

Figure 52. Motor current for Valve 5, closing at normal BWR HPCI operating pressure and temperature.

25

-a

E

4o

4-

I3

ULo

0

20

15

10

5

0

0 6 10 15 20 25

Time (s)

Figure 53. Motor current for Valve 5, opening to 30% open and reclosing.

30 35

S2S8 KGD-1090-01

65
2000

2 1500

I10

0

500

0

50

40

-a

S

co

*0

aa0

I-W

0

4.

30

20

10

0

0 10 20 30 40 50 s0 70

Motor Torque (ft lb)

S211 KGD-0490-85

Figure 54. Comparison of Limitorque's motor performance curves and observed performance during

motor trip.

66
5. CONCLUSIONS

The flexible wedge gate is the predominant valve

design installed in the G1-87 systems of interest.

Five manufacturers provided the largest fraction of

these Installed valves. Four of these

manufacturers, which produce 90% of the installed

valves, were represented in the 01-87 tests. The

tested valves do not represent every variant of the

manufacturer's valve product. However, they do

represent the most popular designs with the

exception of Valve 3, which had a flexible guide

design. They differ slightly in manufacturing

technique, but they are basically the same

functional design. Valve 2 was tested in Phase I

with a hardface disc guide and in Phase 1I with a

normal disc guide. However, there was no

discernible difference in performance between the

two tests. Valve 6, the 10-in. version of 6-in.

Valve 2, performed quite differently so it cannot be

concluded that one manufacturer's product

performs consistently across all sizes. Valve

performance appears to be more valve-specific

depending on stackup manufacturing tolerances,

design clearances, and surface finishes and the way

functional load affects the stackup and finishes.

The valves with machined guiding surfaces had less

guide damage than those with cast surfaces. That

fact did not influence the frequency of seat

damage, however.

Functionally, current industry models that are

used to predict the MOV's performance are

inadequate. Industry has performed very little

testing on each design and apparently has lumped

all of the unknowns into a disc friction coefficient.

Facility limitations usually restrict industry testing

to small valves. Test results from small valves

obscure the internal pressure effects taking place

and, because of the lower disc loadings, would not

be as susceptible to the damage mechanisms we

observed in the Phase I and Phase II tests.

All of the valves required more stem force and

subsequently more torque to operate than what

would have been estimated prior to the test

program. This would indicate that even with the

conservatisms that are added to typical

motor-operator sizing calculations, there is a

potential for undersized motor operators on the

isolation valves of concern to 01-87.

Additionally, we believe that for the predictable

valves used in this test series the actual disc sliding

friction factor Is probably in the 0.5 or 0.6 range at

maximum differential pressure in both the opening

and closing direction. Pressure distributions

throughout the valve during opening and closing

obscure the true value of friction being

experienced in the valve. All of the loads that

affect total disc load must be quantified for correct

operator sizing.

Valves that sustained plastic deformation of the

guide and sealing surfaces cannot be bounded by a

disc factor equation that considers only sliding

friction.

Valve diagnostic testing is in its infancy and many

schemes are being offered to the nuclear utilities

to respond to GL 89-10. The valve measurement

system developed for this G1-87 testing has clearly

identified a minimum diagnostic capability to

ensure the credibility of the measurement and

identify performance margins. At the time of the

G1-87 test programs, none of the major diagnostic

equipment suppliers provides this minimum

diagnostic capability. Since the completion of the

G1-87 test program, several suppliers have been

working toward a more complete system. This will

eventually help the utilities perform better tests.

The one significant problem that may not be

solved easily deals with motor-operator electric

motor performance data. Motor performance

curves may not be obtainable for older motors and

projecting accurate performance on more modern

motors at higher loads may make it difficult for

utilities to establish motor margins.

67
6. REFERENCES

ANSIASME (American National Standards Institute/American Society of Mechanical Engineers), 1983, B16.41,

Functional Qualification Requirements for Power Operated Active Valve Assemblies for Nuclear Power Plants.*

ANSI/ASME (American National Standards Institute/American Society of Mechanical Engineers), OMA8,

American Society of Mechanical Engineers Nuclear Codes and Standards Operation and Maintenance of Nuclear

Power Plants, Section &*

ANSI/ASME (American National Standards Institute/American Society of Mechanical Engineers), OM-10,

American Society of Mechanical Engineers Nuclear Codes and Standards Operation and Maintenance of Nuclear

Power Plants, Section 1O.

EG&G Idaho, 1990, Generic Issue-87 Fleidble Wedge Gate Valve Test Program Phase If Data Report, Vols. 1-8,

EGG-SSRE-8970, March 1.

EG&G Idaho, 1989, BWR Reactor Water Cleanup System Fleible Wedge Gate Isolation Valve Qualification and

High Energy Flow Intermption Test, Vols. 1-3, NUREG/CR-5406, EGG-2569, October.

NRC, "Motor Operated Valve Common Mode Failures During Plant Tkansients Due to Improper Switch

Settings, Office of Inspection and Enforcement, Bulletin 85-03.

68
NRC FORM 335 U.S. NUCLEAR REGULATORY COMMISSION I. REPORT NUM8ER

WAM12.89 eA~ned I yNC i VoI.. Sup.. Rev..

3201.3202 BIBLIOGRAPHIC DATA SHEET NurEG7Z9.,Wpr-'"

ISe. instxuctions on the reverse) EGG-2600

2. TITLE AND SUBTITLE

Generic Issue 87: Flexible Wedge Gate Valve Test Program

Phase II Results and Analysis 3. DATE REPORTPUBLISHE

January 1991

__________ _________ _________ _________ _________ ________ 4 FI g GRAT TUMA ER 4AF68 g7 RB529MBER'

S. AUTHORIS) 6. TYPE OF REPORT

Robert Steele, Jr. Technical

Kevin G. DeWall

John C. Watkins 7. PERIOD COVERED anciuw oasu

U. ~ UPEROLP. .'UAUESti RFofeDvso.Oliee eia, . jia euatr e 7.IPIMI4 o.ed.aln COVEAI dRED IFilorcbh~ O nse r 

B. PERFORMING OGANIZI ZAT- NAEAD ADDKE55 fflRr#<o.,*1V0 "ofRow#on Gfk U..Nlzo*rtlCommnio.Nit $~t

V e and s ai add'esLI

Idaho National Engineering Laboratory

EG&G Idaho, Inc.

Idaho Falls, ID 83415

0.SONOIN N ADES RGNZAIN- lIA.lye~'ae AE mab~:Uotaco. rvieNR ivso. iic r ei..U. m.. egltoyCo~,Jdn

.Sand mailing ddvnmoJ

Division of Engineering

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington D.C. 20555

1G. SUPPLEMENTARY NOTES

11. ABSTRACT I200 wqr* risvu

Qualification and flow isolation tests were conducted to analyze the ability of

selected boiling water reactor process valves to perform their containment isolation

functions at high energy pipe break conditions and other more normal flow conditions.

Numerous parameters were measured to assess valve and motor-opeirator performance at

various valve loadings and to assess industry practices for predicting valve and

motor operator requirements. The valves tested were representative of those used in

reactor water cleanup systems in boiling water reactors and those used in boiling

water reactor high-pressure coolant injection (HPCI) steam lines. These tests will

provide further information for the U.S. Nuclear Regulatory Commission Generic Issue87, "Failure of the HPCI Steam Line Without Isolation," and Generic Letter 89-10,

"Safety-related Motor Operated Valve Testing and Surveillance."

-

12. KEY WORDS/DESCRIPTORS (Lnit t o phrn asm t atuiut fseamie ln 1loeadn the eort. . 1. AVAILABILITY STATEMENT

Generic Issue 87 Unlimited

Generic Letter 89-10 14. SECURITY CLASSIFICATION

isolation valve Un- PAsf

motor-operated valve Unclassified

flow interruption test 111*9wo"I

stem force equation Unclassified

I5. NUMBER OF PAGES

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