NUREG/CR-5406

EGG-2569

Vol. 1

BWR Reactor Water Cleanup System

Flexible Wedge Gate Isolation

Valve Qualification and High

Energy Flow Interruption Test

Analysis and Conclusions

Prepared by K.G. DeWall, R. Steele, Jr.

Idaho National Engineering Laboratory

EG&G Idaho, Inc.

Prepared for

U.S. Nuclear Regulatory Commission

J
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NUREG/CR-5406

EGG-2569

Vol. 1

R1

BWR Reactor Water Cleanup System

Flexible Wedge Gate Isolation

Valve Qualification and High

Energy Flow Interruption Test

Analysis and Conclusions

Manuscript Completed: September 1989

Date Published: October 1989

Prepared by

K.G. DeWall, R. Steele, Jr.

Idaho National Engineering Laboratory

Managed by the U.S. Department of Energy

EG&G Idaho, Inc.

P. 0. Box 1625

Idaho Falls, ID 83415

Prepared for

Division of Engineering

Division of Safety Issue Resolution

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555

NRC FIN A6857, B5529
ABSTRACT

This report presents the measured data and the analyses performed to date on the full-scale high-energy

qualification and flow interruption gate valve testing to develop technical insights for the United States Nuclear

Regulatory Commission (USNRC) effort regarding Generic Issue 87 (GI-87). The research was sponsored by the

USNRCa and conducted by researchers from the Idaho National Engineering Laboratory. We tested two 6-in., 900-lb

class valve assemblies, which represent a significant percentage of the reactor water cleanup isolation valves installed in

plant applications. These valves were modified before testing by adding a high temperature load cell in the valve stems,

which allowed the direct measurement of valve stem thrust during both opening and closing valve cycles.

Instrumentation installed in the flow loop and on the valve assemblies measured the important valve and system test

responses. Additionally, during the test program, all of the currentlypopularmotoroperatedvalve diagnostictest systems

monitored the performance of the valves. Initially the valves were subjected to the hydraulic and leakage qualification

tests defined in ANSI B 16.41 and then to flow interruption and reopening valve tests at boiling water reactor primary

system water temperature and pressure conditions with downstream line break flows. For the two valves tested, results

show that (a) the disc factor used in current industry motor operator sizing equations underpredicts actual valve thrust

requirements at all high temperature loadings, and for one valve design the equations may require an additional term to

account fornonlinearperformance, (b) the thrusts required to close the valves were sensitive to the fluid temperature, and

(c) the results of testing at lower pressures, temperatures, and flows cannot be extrapolated to design basis pressures,

temperatures, and flows for valve designs that have not exhibited linear performance behavior during design basis

prototypical testing.

a. Work supported by the U.S. Nuclear Regulatory Commission, Division of Engineering, Office of Nuclear

Regulatory Research, under DOE Contract No. DE-AC07-761D01570.

A6857-Equipment Operability

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

iii
EXECUTIVE SUMMARY

Recent testing sponsored by the U.S. Nuclear

Regulatory Commission (USNRC) showed that for at

least some gate valves installed in nuclear applications,

the equations used by industry to size the valve operators

do not conservatively calculate the thrust needed to close

the valves under design basis loadings. The tests also

showed that the results of in-situ valve testing at lower

loadings cannot be extrapolated to design basis loadings.

The testing was conducted by researchers from the Idaho

National Engineering Laboratory (INEL) to provide

technical data for the USNRC effort regarding Generic

Issue 87 (GI-87) "Failure of HPCI Steam Line Without

Isolation." The test program also provides information

applicable to Generic Issue IIE.6.1, "In-situ Testing of

Valves" and a related document, lE Bulletin 85-03,

"Motor Operated Valve Common Mode Failures During

Plant Transient Due to Improper Switch Settings."

Of the three boiling water reactor (BWR) process

lines covered under GI-87, an unisolated break in the

reactor water cleanup (RWCU) supply line was selected

for the first phase of testing because such a break would

have the greatest safety impact. The high pressure

coolant injection steam supply line and the reactor core

isolation cooling steam supply line will be addressed in

subsequent research efforts. All three GI-87 process

lines have common features. All communicate with the

primary system, pass through containment, and have

normally open isolation valves.

IEBulletin 85-03 required the utilities to develop

and implement a program that would ensure that the

switch settings on selected safety related motor operated

valves (MOVs) are chosen, set, and maintained correctly

to accommodate the maximum differential pressures

expected on these valves during both normal and

abnormal events within the design basis. It is also

understood that the USNRC issued Generic Letter No.

89-10 "Safety Related Motor-Operated Valve Testing

and Surveillance," which will expand the coverage of IE

Bulletin 85-03 to a largernumber of safety related valves

in the plants. To meet these new valve operating criteria,

industry developed new MOV diagnostic test equipment

and methods for in-situ valve testing. IE Bulletin 85-03

succeeded in significantly improving the operability of

the selected safety related valves because it caused many

of the utilities to reanalyze the design basis load for the

applicable MOVs and to reset the control switches

accordingly. In many cases, these analyses were more

complete than the analyses in the original procurement,

and the utilities reset the control switches in accordance

with the improved analyses.

However, very little design basis testing of valves has

been conducted outside the plant to verify the analytic

assumptions used to determine valve switch settings.

Analytic assumptions are necessary because in many

cases the utility carmot test valves at design basis loadings

in situ. The GI-87 testing provides some of the first

measured valve responses with which industry's valve

operator sizing equations can be compared.

In this initial test program, two representative

RWCU isolation valves were subjected to the hydraulic

qualification tests described in ANSI B 16.41, the nuclear

valves qualification standard, and then to full flow

RWCU pipe break flow interruption tests. In all, fourteen

flow interruption tests were performed, ten on Valve A

and four on Valve B. In the Valve A tests, the parametric

study included varying both the degree of inlet water

subcooling and the pressure. Break flows were

maintained throughout the 30-second valve closure. The

four Valve B tests were all performed at a normal BWR

100F subcooling, and only the inlet pressure was varied.

The test loop and valves were instrumented to determine

the valve response to flow, including a load cell installed

in the valve stems to measure thrust.

Test results show that for both valve designs

tested, the force required to open and close the valves at,

temperatures above 1 000F were significantly higher than

the force predicted by the valve manufacturers. Only in

the room temperature valve opening tests without flow

did the typical industry valve thrust equation predict the

valve response. Industry has also assumed that for valve

opening thrust requirements, the highest load would be

when the disc lifted offthe seat. This was also determined

for the valves tested not to be true. The highest opening

loads with flow occurred at different degrees of opening

for both valves, but in both cases they were well off their

respective seats when the maximum thrust was measured.

Valve closing thrust at full line break flows were higher

than anticipated. One of the valves exceeded the pretest

calculated closing thrust by one third.

V
The test results provide evidence for two

concerns with MOVs in nuclear power plants. First,

proper sizing of motor operators is complicated by the

fact that the equation used for calculating the stem force

needed to close or open a gate valve does not have terms

for temperature, degree of fluid subcooling, internal

valve clearances, and the differences in the opening and

closing forces not accounted for by the stem rejection

term. Second, effective in-situ testing is very difficult

because (a) the tests cannot be conducted at design basis

conditions and (b) even with the valve loadings properly

quantified during the in-situ tests, the results cannot be

extrapolated to design basis conditions because the final

thrust varies depending on the extent to which disc

friction rather than disc seating causes the torque switch

to be compressed to torque switch trip and because the

stem factor varies with the load imposed during valve

operation.

The disc factor of 0.3 typically used in industry to

calculate disc friction force is not conservative for either

of the valves tested. A disc factor of 0.5 marginally

predicts the forces for one valve during both opening and

closing. The response of the other valve is enveloped by

the 0.5 disc factor during opening but not during closing.

Today's tools for analyzing valve response to fluid

loadings are not sophisticated enough to detect small

design differences that make large response differences.

Temperature also affects the thrust requirements

of these gate valves. These facts justify continued

qualification testing of prototypical valves at design basis

loadings and point out the need for industry to modify the

variables in the sizing equation. It may be necessary to

add new terms to the equation or to increase the disc

factor to a very conservative number to account for the

missing terms.

When tests have determined the thrust needed to

operate a valve at its design basis loading, utilities can use

one of several modem diagnostic systems to

conservatively set the motor operator control switches.

Industry will have to account for the varying stem factor

and for the excessively high thrusts resulting from

seat-induced torque switch trips that occur with valve

operation with low flow or no flow. However, this

method may exceed the allowable thrust on some valve

designs. This job will be easier and the result more

conservative if both the valve torque and thrust can be

measured when the switches are set. If further research

proves that there is a proportional relationship between

stem load and stem factor, the degree of conservatism can

be reduced.

The stem factor is a calculation made to predict

the efficiency of the motor operator torque to stem thrust

conversion. Until recently industry has always

considered the stem factora constant. Procedures used by

two of the more popular in-situ valve diagnostic test

systems are based on this premise. Test results show that

the stem factor changes with stem load, thus making it

very difficult to extrapolate normal in-situ valve testing

to design basis conditions.

vi
CONTENTS

ABSTRACT ..........................................................................

EXNECUTIVE SUMIM{ARY ........................................................ . v

ACKNOWLEDGMENTS ....................................................... . xi

1. INTRODUCI ION ................................................................. 1

1.1 Background .................................... 1

1.2 Motor Operator Sizing .................................... . 2

2. TEST OBJECTIVES ................................... 5

3. APPROACH TO TESTING ................................................. . 7

3.1 Test Design ...................................... 7

3.2 Test Procedure ............................................................... 10

4. TEST RESULTS ANND INTERPRETATION ......... ................................... 15

4.1 Data Analyses ............................................................... 15

4.2 Stem Force-Static Components ........................ .......................... 24

4.3 Stem Force-Dynamic Component ....................... ........................ 27

4.4 Valve Opening Versus Closing Tests ...................... ........................ 30

4.5 Stem Factor .................................................................. 34

4.6 Operator Torque Switch Trip Anomaly ..................... ....................... 37

5. CONCLUSIONS ................................................................... 41

6. REFERENCES ................................................... 43

FIGU RES

1. Typical motor-operated valve, similar to the two valves used, which were modified by installing

a load cell in each valve stem . ................................................ 3

2. Instrumentation installed in the test loop to monitor temperature (T), pressure (P), pressure

differential (AP), stem force (F), and flow; motor current and voltage, valve stem position,

and other important variables were also monitored ........................................ 8

3. In this Valve A test, peak thrust encountered during opening was measured not while the disc

was being lifted off the seat, but well after flow was established ......... .................... 16

vii
4. The response of Valve B is similar to that of Valve A (Figure 3); the absence of a spike at the

hammer blow is because the valve was not fully seated at the end of the previous closing cycle .... 16

5. In this Valve A test, the loads measured during closing were greater than the loads calculated

using 0.3 and 0.5 disc factors ............... ......................................... 17

6. The response of Valve B is more linear than that of Valve A; the calculation using a 0.5 disc

factor marginally envelopes the measured load, but the 0.3 calculation is not conservative .... .... 17

7. During closure, the Valve A body guides should provide proper disc alignment until the disc

contacts the entire seat ring surface .............. ...................................... 19

8. The Valve A disc, which shows sealing surface damage on the downstream face ..... ........... 20

9. Close-up of the Valve A lower right guide surface, which shows yielding, plastic deformation,

and significant galling on the lower edge of the disc ...................................... 21

10. The Valve A body guide, which mates with the right disc guide (shown in Figure 8), shows

galling near the fully closed position ............. ...................................... 22

11. Close-up of the Valve A lower left guide surface, which shows yielding and plastic deformation

similar to the right side of the disc ............. ....................................... 23

12. A linear curve fit using the industry equation closely approximates the Valve A running stem

forces for opening without flow ............... ....................................... 25

13. A linear curve fit using the industry equation closely approximates the Valve A running stem

forces for closing without flow ................ ....................................... 25

14. A linear curve fit using the industry equation closely approximates the Valve B running stem

forces for opening without flow ............. ......................................... 26

15. A linear curve fit using the industry equation closely approximates the Valve B running stem

forces for closing without flow ........................................... 26

16. With an increase in the fluid subcooling, Valve A returns to more linear behavior and stem forces

are bounded by the 0.5 disc factor calculation ........................................... 28

17. At 1400 psig and 100F subcooled, Valve A shows the same nonlinear response as in Figure 5 ...... 28

18. At 1400 psig and 130TF subcooled, Valve A returns to more linear behavior and stem forces are

again bounded by the 0.5 disc factor calculation .29

19. Flashing occurred upstream of Valve A during the early part of the closure stroke, but recovered

to subcooled prior to the start of the nonlinear behavior (compare Figure 5) .29

20. With Valve A opening against differential pressure only (no flow, cold fluid), the measured thrust

is enveloped by the 0.3 disc factor calculation .31

21. With Valve B opening against differential pressure only (no flow, cold fluid), the measured thrust

is enveloped by the 0.3 disc factor calculation .31

viii
22. With an increase in fluid temperature, the Valve A measured thrust is not enveloped by the

03 disc factor (compare Figure 20) .................................................... 32

23. With an increase in fluid temperature, the Valve B measured thrust is not enveloped by the

0.3 disc factor (compare Figure 21) .................................................... 32

24. The Valve B AP opening test at normal BWR operating conditions, like the high flow interruption

test (see following figure), appears to isolate the disc factor at or near 0.5 ..................... 33

25. For the Valve B pipe break flow interruption test at normal BWR operating conditions, the

relationship between measured and calculated stem thrust is very similar to that seen in the

AP opening test (see previous figure) .................................................. 33

26. At the torque switch trip, the worm may either turn the worm gear and drive the disc deeper into

the seat, or (if the disc will move no further) climb the worm gear, overcompressing the torque

spring and producing additional thrust in the valve stem ................................... 35

27. Though the torque switch was set to trip at 18,000-lb thrust in the absence of operator momentum,

this lightly loaded valve achieved a significantly higher final seating thrust ..... ............... 36

28. As stem factor and operator momentum are affected by increased loadings, the final thrust is less

even though the torque switch setting is the same ......................................... 36

29. The stem factor increases with load, reducing the available thrust at a given torque ..... ........ 37

30. Below-normal stem forces at torque switch trip were observed for three test series with Valve A ... 38

TABLES

1. Test parameters measured during blowdown tests ............... ......................... 9

2. Valve diagnostic equipment used for subcooled blowdown tests ............................. 11

3. Test step matrix for qualification and blowdown tests ..................................... 12

4. A design comparison between Values A and B ........................................... 18

ix
ACKNOWLEDGMENTS

The authors wish to acknowledge the following companies for their assistance and cooperation in this research

effort. Anchor-Darling Valve Company, Velan Incorporated, and Limitorque Corporation contributed to the testing by

providing representative test hardware at significantly reduced cost and schedule and information concerning valve

motor operator sizing procedures. Liberty Technology, Limitorque Corporation, MOVATS Incorporated, and WYLE

Laboratories contributed by providing valve assembly diagnostic support representing typical in-plant test capabilities.

Three nuclear utilities, Philadelphia Electric Company, Duke Power Company, and Toledo Edison Company, provided

information concerning in-situ equipment configurations and test capabilities. The valve, motor operator, and diagnostic

equipment manufacturers and utility representatives all cooperated in the early test planning, critically reviewed test

procedures and data analysis and participated in NRC sponsored test results review meetings.

The views expressed in this report are the authors' andmay not be the opinion ofthe valve, motoroperator, and/or

diagnostic equipment manufacturers who cooperated in this effort.

xi
BWR REACTOR WATER CLEANUP SYSTEM

FLEXIBLE WEDGE GATE ISOLATION VALVE

QUALIFICATION AND HIGH ENERGY FLOW

INTERRUPTION TEST

VOLUME I

ANALYSIS AND CONCLUSIONS

1. INTRODUCTION

The Idaho National Engineering Laboratory (INEL),

under the sponsorship of the United States Nuclear

Regulatory Commission (USNRC), is performing

research to provide technical input for the resolution of

specific generic issues and to provide information to

develop and improve industry mechanical equipment

qualification and operating and maintenance standards.

This overall research effort includes aprogram that tested

the operability (opening and closing) of two full-scale

motor-operated gate valves typical of those installed in

boiling water reactor (BWR) reactor water cleanup

(RWCU) process lines for containment isolation. The

valves were parametrically tested at, above, and below

the pressure, temperature, and flow conditions of a

worse-case downstream pipe break in the RWCU supply

line outside containment. The purpose of the test

program was to provide technical input for the USNRC

effort regarding GenericIssue 87 (GI-87), "Failure ofthe

HPCI Steam Line Without Isolation." The test program

also provides information applicable to the

motor-operated valve portion of another highly visible

generic safety issue, Generic Issue H.E.6.1 (GI-ll.E.6.1),

'In-situ Testing of Valves," and its related documents,

IE Bulletin 85-03, "Motor Operated Valve Common

Mode Failures During Plant Transients Due to Improper

Switch Settings," and Generic Letter No. 89-10

"Safety-Related Motor Operated Valve Testing and

Surveillance," which expands many of the

IE Bulletin 85-03 requirements to other safety-related

motor-operated valve testing and surveillance.

The analyses performed to date on the measured data

obtained during the first phase of the GI-87 valve test

program andconclusions arediscussedinVolume Iofthis

report. Volume II contains the measured data taken in the

more significant test sequences of the test program. The

data is also available in IBM PC compatible format, for

those who wish to analyze the data, and can be obtained

through our DOE Technology Transfer Office, at

(208) 526-8318. Volume m is a review of the BWR

containment isolation valve designs and piping

configurations, qualification methods, and previous

research

1.1 Background

GI-87 applies to the BWR process lines that

communicate with the primary system, pass through

containment, and contain normally open isolation valves.

Two steam supply lines, the high-pressure coolant

injection (HPCI) and the reactor core isolation cooling

(RCIC) lines, and one hot water supply line, the RWCU

line, meet these criteria. GI-87 addresses whether the

isolation valves in these lines will close in the event of a

downstream pipe break outside containment.

a. Mention of specific products and/or manufacturers in this document implies neither endorsement or preference nor

disapproval by U.S. government, any of its agencies, or EG&G Idaho, Inc., of the use of a specific product for any

purpose.

1
The project began with a review of the valves installed

in these applications (see Appendix B, Volume IE1),

specifically their sizes, manufacturers, previous testing,

and risk significance. A survey identified the flexible

wedge carbon steel gate valve with a Limitorquea motor

operator as the predominate valve in the three systems

(HPCI, RCIC, and RWCU) addressed by GI-87. The

most common valve size for the RCIC system is 4 in.,

6 in. for the RWCU system, and 10 in. for the HPCI

system. Valve manufacturer Anchor-Darling had the

largest share of the installed valves, followed by Velan,

Crane, Powell, and Walworth (with about equal shares).

It was also determined that a downstream break in the

RWCU system would represent the highest risk to the

plant, so it was decided that the initial flow isolation

testing should provide information on valve operability

questions associated with the RWCU environment.

To avoid duplication, we reviewed previous applicable

test programs. The reviewed test programs included the

Electrical Power Research Institute (EPRI)

power-operated relief valve/block valve testing at Duke

Power in 1980. This program had three shortcomings:

(a) the block valves were stainless steel as opposed to

carbon steel; (b) the tests were go/no-go type tests where

neither motor-operated valve thrust nor torque were

measured; and (c) the EPRI test medium was steam,

which would be more applicable to the HPCI and RCIC

systems than to the RWCU system. Kraftwerkunion

(KWU) of West Germany had tested a 3-in. stainless steel

parallel disc gate valve at blowdown flows for the Central

Electric Generating Board (CEGB), United Kingdom.

Mechanical interference on the downstream disc

prevented closure. Kraftwerkunion has also performed

full flow interruption testing on a large number of valve

types; however, our initial contacts indicated that the

information is proprietary. Since that time, Bechtel and

KWU have formed an alliance and have indicated that the

information may become more available in the future.

The results from the survey of previously completed

work determined that adequate and sufficient technical

test information was not available for the USNRC effort

on GI-87 and that additional testing was required.

The USNRC effort regarding IE Bulletin 85-03

required the utilities to develop and implement aprogram

that would ensure that switch settings on selected

safety-related motor-operated valves (MOVs) are

chosen, set, and maintained correctly to accommodate

the maximum differential pressures expected on these

valves during both normal and abnormal events within

the design basis. The follow-on generic letter win

expand IE Bulletin 85-03 requirements to a large number

of safety-related MOVs, including those that may be

mispositioned.

New MOV diagnostic test equipment and methods for

in-situ valve testing have been developed to meet these

new operating criteria. One of the new requirements was

that the valve control switches be set correctly for the

design basis loading for each valve. However, very little

design basis testing of valves has been conducted outside

the plants, andin many cases the utilities could not test the

valves at in-situ design basis conditions. This situation

left the utilities relying on valve motor operator switch

settings that were based on analyses of the design basis

loadings. Utilities typically verified the torque or thrust

levels for each valve through seat or back seat type

loadings, with very low hydraulic loadings. To determine

if one could extrapolate the results of the testing

performed at typical in-situ test conditions to design

basis conditions, we invited the manufacturers of the

more widely used valve diagnostic test systems to join us

inthe GI-87 testprogram. The insights gained from this

testing would be applicable to both GI-87 and

TE Bulletin 8543.

1.2 Motor Operator Sizing

The gate valve is a high recovery positive shut-off

valve and is usedin systems where minimal pressure drop

is desired when the valve is open. It is ideally suited to

situations where isolation of one part of a system from

another is required and control of the dynamic properties

of the fluid (throttling) is unnecessary. When the disc (or

gate) is in the open position, the run of the valve is free of

any obstruction with approximately the same head loss as

in the adjacent piping. When the disc is lowered into the

seat, the upstream pressure forces it against the seat,

creating a seal and isolating the downstream system from

the upstream fluid.

2
Figure 1, a cutaway drawing of a typical

motor-operated gate valve, shows the components

important to this discussion. The forces needed 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 foruse 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.

where

F. = total stem force

Yd = disc factor

Ad = disc area

AP = differential pressure

A,= stem cross-sectional

area

P = stem presssure

F =acingdrag load

Via constant).

I

Dynamic

component

Static

component

Ft =PdAdA.P ± AP + F (1)

Electric

thrust

(load cell)

Yoke

Seat ring

Valve

stem -

Disc A-A

9s3007

Figure 1. 'Typical motor-operated valve, similar to the two valves used, which were modified by installing a load

cell in each valve stem.

3
The disc factor (It, ) normally used for wedge-type

gate valves in Equation (1) is 0.3. Note that in this

equation the stem rejection load can be either positive or

negative depending on 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 a constant that depends on the packing design, gland nut

torque, and direction of operation. The equation is shown

divided into two components, which will be referred to in

the analysis found later in this report: (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

plant.

Motor operators control output torque, not valve stem

thrust. Thus, in sizing the operator and determining the

torque switch setting formotor-operated gate valves, one

must consider the conversion ofoperator output torque to

valve stem thrust. This conversion of torque to thrust is

one of the equations in the set of motor operator sizing

equations. The torque to thrust relationship normally

used in sizing motor operators depends on a stem factor

calculation, given by

T = Xt (2)

where

T = operator torque

As = stem factor

Ft = total stem force [from Equation (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. Most in industry use a 0.15 or 0.20 coefficient of

friction for this parameter. Normally it is assumed that

only damage and lubrication of the stem/stem nut threads

can significantly alter the stem coefficient of friction.

Limitorque 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.

4
2. TEST OBJECTIVES

As discussed previously, the gate valve qualification

and flow interruption testing was performed to provide

information to assist in resolving the uncertainties in gate

valve operator sizing and torque switch setting. Specific

objectives of the testing included the following:

1. Determine the valve stem force required to close

a typical RWCU gate valve at typical operating

test conditions and under full flow blowdown

conditions.

2. Compare valve closing load to opening load at

various system operating conditions.

3. Measure valve closure force components such

as disc drag, stem rejection, and packing drag

loads.

4. Make limited assessments of the effect of

temperature, pressure, and valve design on

valve closing and opening loads.

5. Evaluate the validity of using present industry

standard equations for determining valve stem

force.

6. Provide detailed technical information for the

above steps to assist in the USNRC effort

regarding GI-87.

An additional goal of the INEL testing was to provide

information so that specific guidelines might be

developed to improve valve qualification and operating

and maintenance standards.

5
3. APPROACH TO TESTING

3.1 Test Design

Two full-scale, representative nuclear valve

assemblies were cycled under various design conditions

and design basis RWCU pipe break conditions. The

valves were manufactured for this test program by

Anchor-Darling Valve Company (Valve A) and Velan

Incorporated (Valve B), using nuclear design and

materials, without third-party inspections. Both valves

were modified to incorporate extended yokes (4 in.

longer than normal) and the stems were cut in half and

threaded to allow installation of a special stem force

measurementdevice. Flanges and safe ends were welded

to both sides ofeach valve for mating with the test system

piping.

The first test specimen, Valve A, was a 6-in., 900-lb

standard rated, cast steel, flexible-wedge gate valve with

apressure seal bonnet andbutt weldends. The valve seats

were hard faced with Stellite and seal-welded to the valve

body. The one-piece flexible wedge (disc) was also hard

faced with Stellite on the seating faces. The disc guides

were carbon steel. The valve was powered by an

oversized Limitorque SMB-2-40 electric motor

operator. The basic valve design, without the oversized

operator, is representative of 40% of the isolation valves

installed in BWR RWCU systems.

The secondtestspecimen,ValveB,wasa6-in., 900-lb

standard rated, forged steel, flexible-wedge gate valve

with a bolted bonnet and butt weld ends. The valve seats

were hard faced with Stellite and seal-weldedto the valve

body. The one-piece flexible wedge (disc) was also hard

faced with Stellite on the seating faces. The valve was

powered by a Limitorque SMB-0-25 electric motor

operator. Representing one of the newer valve

assemblies delivered since 1970, the Valve B design

incorporated hardfaced disc guide wear surfaces.

B oth valves utilized 460-Vac, 3-phase, 60-Hz electric

motor operators. To ensure valve closure and data

collection at the anticipated greater-than-normal

loadings, Valve A utilized a larger, greater-capacity

motor operator than would normally be used. The motor

operator used with Valve B was sized in accordance with

current practices to represent a typical MOV assembly

used in nuclear power plants today. Because of their

differences in internal design and friction bearing surface

design, the two valve assemblies represented a large

number of the MOVs used in nuclear powerplants today.

The test system used for the subcooled water

qualification and flow interruption testing featured a

large water tank, heated and pressurized so that various

thermal hydraulic conditions could be established and

regulated, replicating actual BWR conditions. The water

was propelled by high-pressure gaseous nitrogen during

the high energy flow interruption testing. The water

heating system consisted of a heating section and a

high-pressure, high-temperature water pump. The

heating section contained an electrical heater, which

heated the water as the pump recirculated water from the

pressure vessel, through the test section and test valve,

and back to the pressure vessel. The test section was a

6-in. pipe with the test specimen mounting flanges and

appropriate fittings for obtaining temperature and

pressure measurements. The test system also featured a

fast-acting (approximately 300-msec opening stroke),

hydraulically operated valve, positioned so that when the

valve was actuated, the system's fluid was abruptly

dumped to the atmosphere, resulting in high-flow

(blowdown) conditions through the test specimen. The

system is shown schematically in Figure 2.

To accomplish the functional testing, the system

contained bleed valves, which provided the means to

reduce system pressure on both sides (upstream and

downstream) of the test specimen. In this manner,

differential pressure conditions could be established

across the test valve's disc.

The test system was instrumented to monitor flow,

pressure, and temperature at various locations, including

the test valve upstream and downstream positions. Motor

operator electrical characteristics were also recorded.

Valve stem force was monitored using the previously

described high-temperature load cell installed between

the two halves of the valve stems. The test parameters

measured are listed in Table 1.

7
Test specimen

11/12 /13 /El/E2

ST/LS1/LS2/TS Bleed

A4/AS/A 6

Pump -

Heater

Water tank

Vent valve Flowmeter

MGH02167

Figure 2. Instrumentation installed in the test loop to monitor temperature (T), pressure (P), pressure differential

AP, stem force, (F), and flow; motor current and voltage, valve stem position, and other important

variables were also measured.

8
Table 1 . Test parameters measured during blowdown tests

Data X-Y

Transducer Measurement FM Tape Oscillograph Logger Plotter

Ti System water temperature X

T2 Test valve inlet water temperature X

T3 Test valve body temperature X

T4 Heating section water temperature X

T5 Load cell temperature X

P1 System water pressure X X X X

P2 Test valve inlet water pressure X X X

P3 Test valve outlet water pressure X X X

P4 Discharge section water pressure X X X

ANP Test valve differential pressure X X X

AP2 Venturi differential pressure X X X

AP3 Pump differential pressure X X

I1 Actuator current X X

12 Actuator current X

I3 Actuator current X

El Actuator voltage X X

E2 Actuator voltage X

ST Valve stroke-LVDT X X X

LS1 Open limit switch X X

LS2 Close limit switcha

TS Close torque switch X X

F Valve stem force X X X

Al Acutator acceleration Y X X

A2 Actuator acceleration X X

A3 Actuator acceleration Z X

A4 Valve body acceleration Y X X

A5 Valve body acceleeration X X

A6 Valve body acceleration Z X

a. Control room light indicator only.

9
A secondary objective of the qualification and flow

interruption test program was to determine if normal

utility in-situ valve testing, using available diagnostic

equipment, could be extrapolated to provide assurance of

a valve's operability at design basis loadings. Several

MOV diagnostic system manufacturers supported this

objective by participating in the testing, as listed in

Table 2. The manufacturer participation was not a

competition but, rather, an attempt to determine what

factors need to be considered to provide reasonable

assurance of valve operability using each of the

diagnostic systems.

3.2 Test Procedure

Upon installation in the test system, each valve

assembly was subjected to a typical ANSI B16.411

functional qualification test, including the valve leakage

test (Annex A), cold cyclic test (Annex B), and hot cyclic

test (Annex C). These tests provided a baseline

characterization of the valve assembly operation for

comparison with the information obtained from the later

testing. The valve leakage test established the mainseat

valve leakage rate and the packing leakage rate of the test

valves, while the cold cyclic test demonstrated the

capability of the test valve assembly to open and close

under adverse combinations of motive power and system

pressure with the assembly at room temperature, not

exceeding 1000F. The hot cyclic test sequence was

performed to demonstrate the capability of the test valve

assemblies to open and close under adverse combinations

of motive power and system pressure with the assembly

at design temperature, in excess of 1000F. Annex G, flow

isolation, was the subject of this test program and thus

was not performed as part of the pretest qualification

series. Table 3 lists the valve cycles performed during

both the qualification tests and the subsequent blowdown

test series.

Once baseline qualification testing of each test valve

assembly was completed, several test series were

performed to address the questions of GI-87. Each test

series included leakage tests, cyclic tests without flow,

cyclic tests at normal system flow, and cyclic tests at full

line break flow conditions. A wide range of design

upstream pressures and temperatures were maintained

throughout the valve closures, with line break flow

limited only by flashing and choked flow in the test loop.

Required nonflow data were collected during the

preparation period for full-scale flow and postfull-scale

flow tests.

Fourteen line break flow tests (see Table 2) were

accomplished, ten on Valve A and four on Valve B. The

ten tests on Valve A with the oversized operator included

a parametric study in both pressure and the degree of

water subcooling. Pressures were varied from 600 to

1400 psig valve inlet pressure and the coolant

temperatures ranged from 10 to 1300F subcooled. The

four tests performed on Valve B with the normal sized

operator were performed to demonstrate expected

in-service performance with the operator motive power

closer to normal. In these tests the pressure was varied

from 600 to 1400 psig at a constant 101F subcooling.

10
Table 2. Valve diagnostic equipment used for subcooled blowdown tests

Valve Test Series Description Diagnostic Equipmenta

A

A

A

A

A

A

A

A

A

A

A

1

3

2

4

6

5

7

9

8

10

11

Qualification test

Blowdown, 1000 psig, 480°F

Blowdown, 1000 psig, 530°F

Blowdown, 1000 psig, 400°F

Blowdown, 1400 psig, 530°F

Blowdown, 1400 psig, 580°F

Blowdown, 1400 psig, 450°F

Blowdown, 600 psig, 430°F

Blowdown, 600 psig, 480°F

Blowdown, 600 psig, 350°F

Blowdown, 1000 psig, 530°F

Qualification test

Blowdown, 1000 psig, 530°F

Blowdown, 1400 psig, 580°F

Blowdown, 600 psig, 480°F

Blowdown, 1000 psig, 530°F

MCSA

MCSA

None

V-MODS

V-MODS

MOVATS

MOVATS

None

None

MAC, VOTES

MCSA

MCSA

MCSA, MOVATS

V-MODS

MAC

V-MODS

B

B

B

B

B

1

2

3

4

5

a. MAC

MCSA

MOVATS

V-MODS

VOTES

Limitorque Motor Actuator Cbaracterizer

ORNL Motor Current Signature Analysis

MOVATS, Inc. (MOV Analysis and Test System)

WYLE Laboratories Yalve Motor operator _iagnostic System

Liberty Technology Yalve Qperator Test & Evaluation _ystem

11
Table 3. Test step matrix for qualification and blowdown tests

Step Number Valve Cycle

Qualification tests (Test 1, Valves A and B)

Description

1

2

3

4

5

6

7

8

9

10

11

12

13

14

15

16

17

18

19

20

21

22

23

24

25

26

O-C

C -PtO

Pt 0 -> C -4 0

0 t->CO

C-*0

0-4C-40

0-+C

C-40

C-)O

O-C

C->O

0-4C

C->0

C-> O

0->C

Close valve at 2220 psig, cold

Seat leakage test

AP opening at 1700 psig, cold

Packing leakage test

Cycle valve at 0 psig, cold

Close valve at 2220 psig, cold

AP opening at 1700 psig, cold

Cycle valve at 0 psig, cold

Cycle valve at 0 psig, cold

Cycle valve at 0 psig, cold

Close valve at 2220 psig, cold

AP opening at 1700 psig, cold

Close valve at 2220 psig, cold

AP opening at 1700 psig, cold

Close valve at 2220 psig, cold

AP opening at 1700 psig, cold

Cycle valve at 0 psig, cold

Close valve at 1650 psig, 600°F

AP opening at 1650 psig, 600°F

Close valve at 1650 psig, 600°F

AP opening at 1650 psig, 600°F

Close valve at 1650 psig, 600°F

AP opening at 1650 psig, 600°F

Close valve at 1650 psig, 600°F

AP opening at 1650 psig, 600°F

close valve at 1650 psig, 600°F

12
Table 3. (continued)

Step Number Valve Cycle Description

Blowdown tests (Valve A, Tests 2-11, and Valve B, Tests 2-5)

0 O-C -O

1 O-4C

2 C- O

3 O*C

4 C- O

5 O- C

6 C-+ptO

7 ptO -C

8 C+OO

9 O0-C-4O

Final cold testing (Valve A, Test 11, and Valve B, Test 5)

Cycle valve at test pressure and temperature

Close valve at test pressure and temperature

AP opening at test pressure and temperature

Close valve at 100 gpm, test pressure and

temperature

Open valve at 100 gpm, test pressure and

temperature

Blowdown at test pressure and temperature

Blowdown at test pressure and temperature

Blowdown at test pressure and temperature

Open valve at 0 psig and test temperature

Cycle valve at 0 psig and test temperature

10

11

12

13

14

15

16

17

18

19

20

21

22

23

C ->O0

0-4C

C-~pto

PtO0-> O

O -+C

C-+O

O -~C

C -)

0-*C

C-+O

0-4C

C->O

AP opening at 1000 psig, cold

Close valve at 1000 psig, cold

Seat leakage test

AP opening at 1000 psig, cold

Packing leakage test

Open valve at 1000 psig, cold

Close valve at 1000 psig, cold

AP opening at 1000 psig, cold

Close valve at 0 psig, cold

Open valve at 0 psig, cold

Close valve at 0 psig, cold

Open valve at 0 psig, cold

Close valve at 0 psig, cold

Open valve at 0 psig, cold

0

C

Pt

open

close

partial valve stroke

13
4. TEST RESULTS AND INTERPRETATION

A torque switch setting of 2.0 was selected for the

Valve A motor operator so that the stem thrust capability

was maximized without exceeding the valve and

instrumentation capacity. (The torque switch was reset to

2.5 aftertest lO to compensate foran observed torque-out

anomaly, discussed later.) Valve A closed satisfactorily

during all tests; however, the measured stem loads were

significantly higher than the stem loads predicted by the

valve manufacturer.

A torque switch setting of 1.75 was selected for the

Valve B assembly to provide the needed closure thrust for

the given test conditions. However, this setting resulted

in delivered stem thrust (as determined by the

stem-mounted load cell) below that specified by the

valve manufacturer for the highest pressure flow

interruption test. Therefore, the torque switch setting was

raised to 2.0 before the first flow interruption test

Valve B performed satisfactorily during the lower

pressure testing; however, during the 1400 psig test, the

operator torqued out before the disc reached the fully

closed position (1/4 in. of travel remaining). During this

test the valve had closed far enough to produce a seal,

with no leakage observed. Higher-than-predicted stem

loads during flow isolation and a reduction in delivered

stem thrust accounted for the valve not completely

closing at its design basis loading.

4.1 Data Analyses

Both valves exhibited higher opening and closing

forces at normal operating temperature than would have

been predicted using Equation (1) and the 0.3 disc factor

typically used in the past by industry. Examples of these

higher forces for both valves are shown in Figures 3

through 6. The predicted forces in these figures were

calculated using both a 0.3 and a 0.5 disc factor and the

actual measured pressures, exposed disc areas, packing

drag loads, and stem rejection loads throughout the

opening and closing cycles. We used this calculational

technique and these plots to look for any deviations from

the predictions, and (if they occurred) to determine at

what specific point in the opening or closing cycle they

occurred. The opening and closing cycles shown in

Figures 3 through 6 are all at normal BWR operating

temperatures, pressures, and line break flows.

The measured stem force history shown in Figure 3 for

the Valve A opening under high flow starts out with the

valve closed and the valve stem in compression. We see a

decrease in stem compression as the opening cycle

begins, and then the stem goes into tension due to the

operator hammer blow. The stem force history, for the

remainder of the cycle, reflects the combined effects of

the disc drag load, the stem rejection (assists opening),

and the packing drag load. Figure 5 illustrates this same

comparison of actual measurements to calculated values

during Valve A closure. The figure shows the measured

stem compression (negative values) increasing as the

valve closes, until the compression reaches a peak when

the flow path is finally blocked. Then the stem

compression decreases to a value representing the force

required to slide the disc on the full seat ring to the final

seating position. Finally, the measured force increases

sharply through torque switch trip to the final stem

compressive load (at approximately 40,000 lb, not shown

in Figure 5). This additional stem force beyond torque

switch trip is due to the circuit dropout time and the

momentum of the operator motor.

Valve B's measured forces, shown in Figures 4 and 6,

follow the general shape of the calculated forces quite

well, and we describe this performance as linear. We

describe the performance of Valve A, which did not

follow the shape of the calculations, as nonlinear. For

both valves the measured opening forces are bounded by

the 0.5 disc factor calculation; however, they occur much

later in the valve cycle than is predicted. The measured

closing forces for Valve A are not bounded by either

calculation and they do not follow the shape of the curves.

Because of its larger port and stem size and a packing

design with greater drag, Valve B would be expected to

need about a three percent higher force during closing

than Valve A. A comparison of Figures 5 and 6 shows

this to be true for the first half of the closing stroke and

during the seating period at the end of the valve cycle.

However, during the last half ofthe valve's cycle, down to

flow isolation, Valve A required much higher forces than

Valve B. Even though both valves were 6-in., 900-lb

class, flexible wedge gate valves, they responded to

similar thermal hydraulic loadings quite differently.

15
->

-0

E2

0

U/)

20.0

10.0

0.0

-10.0

-20.0

-30.0

0 4 8 12 16

Time (sec)

20 24 28

MGH02 130

Figure 3. In this Valve A test,. peak thrust encountered during opening was measured not while the disc was being

lifted off the seat, but well after flow was established.

20.0

° 10.0

C$

x

-D 0.0

I-

.- 10.0

E

Cl)

"'-20.0

-30.0

Valve B, Test 2, Step 6, 1000 psig, 530 OF

...........-

| Actual

Calculated (disc factor = 0.3)

Calculated (disc factor = 0.5)

0 2 4 6 8 10

Time (sec)

12 14 16 18

MGH02131

Figure 4. The response of Valve B is similar to that of Valve A (Figure 3); the absence of a spike at the hammer

blow is because the valve was not fully seated at the end of the previous closing cycle.

16
Valve A, Test 2, Step 5, 1000 psi, 530OF (100 F subcooled)

x

.M

0

0

E

C/,

6

4

2

0

-2

-4

-6

-8

-10

-12

-14

-16

-18

-20

-22

-24

-26

4 8 12 16 20

Time (sec)

24 28 32 36

MGH02128

Figure 5. In this Valve A test, the loads measured during closing were greater than the loads calculated using 0.3

and 0.5 disc factors.

Valve B, Test 2, Step 5, 1000 psi, 530 OF (10 OF subcooled)

.0

-0

I0

E

0

C,,

6

4

2

0

-2

-4

-6

-8

-10

-12

-14

-16

-18

-20

-22

-24

-26

0 5 10 15 20 25 30 35

Time (sec) MGI02 129

Figure 6. The response of Valve B is more linear than that of Valve A; the calculation using a 0.5 disc factor

marginally envelopes the measured load, but the 0.3 calculation is not conservative.

17
In the initial analysis for the adequacy of Equation (1)

it appeared there were two problems: (a) in the case ofthe

linearvalve performance, a term maybemissing from the

equationorthe typical 0.3 disc factormay be too low and

(b) the equation appeared inappropriate forpredicting the

performance of a nonlinear valve.

A summary of some of the important design features of

the two valves is presented in Table 4. Two design

differences may account for the valve's response. The

first difference is in the disc guide face and body guide

materials. Valve A is typical of most nuclear valves with

guide surfaces made of cast carbon steel, type

A216-WCB, while Valve B uses similar base materials

but hardfaces the disc guide face with Stellite 6. The

second difference is in the disc guide to body guide

clearances. Valve A had 1/4-in. clearances as opposed to

Valve B's 1/8-in. The other differences in valve design,

stem diameter and packing drag, would serve to increase

the loadings on Valve B.

Disassembly and inspection of the two valves after

completion of the test program provided some insight

into the nonlinear behavior of Valve A. Inspection of the

disc guide surfaces showed a wearpattern, indicating that

the disc had tilted downstream as it closed, with a very

small bearing area of the disc guide riding on the valve

body guides. These small bearing areas show signs of

yielding, galling, and plastic deformation. As mentioned

the disc to guide clearances on Valve A are twice as large

as the clearances on Valve B. We believe that the

nonlinear performance of Valve A is the result of the

greater disc-to-guide clearance, which allowed the disc

to tilt in this valve design. This hypothesis is further

confirmed by the fact that at flow isolation, when the disc

entered into full contact with the seat ring and the disc

sealing surfaces became the primary guiding surface, the

closing forces dropped, as shown in Figure 5. The final

seating forces are slightly less than the calculation using

the 0.5 disc factor.

Figure 7 is a view of the inside of the Valve Abodywith

the two seat sealing surfaces and body guide surfaces.

Figure 8 shows the Valve A disc. Sealing surface damage

can be seen and the lower right disc guide shows

indications of the small bearing area that was engaged

during closing. Figure 9 is a view of the disc rotated,

looking down the guide surface. The right guide surface

shows evidence of yielding, plastic deformation, and

significant galling on the lower edge of the disc.

Figure 10 provides a close-up of the body guide that

mated with the right disc guide shown in Figure 9. This

guide also shows signs of galling. Figure 11 shows the

left disc guide, showing the same small bearing area and,

again evidence of plastic deformation. The damage to

Valve A is of course magnified after undergoing ten

design basis loadings; however, the wear patterns do

provide evidence that the disc tilted in the guides,

resulting in the nonlinear performance.

Judging by the wear patterns shown in Figure 8, it is

unlikely that Valve A could have produced a tight seal

using the downstream face alone. A seal on the

downstream face would be necessary to isolate flow if the

valve torqued out before full travel but with the disc on

the seat. However, the valve maintained its leak integrity

throughout testing, indicating proper sealing on the

upstream face of the disc-the result of using the

oversized operator with a higher-than-necessary seating

thrust.

Table 4. A design comparison between Valves A and B

Valve Design Valve A Valve B

Disc guide face

Valve body guide

Disc sealing surface

Seat

Disc guide to body guide clearances

Stem diameter

Estimated maximum packing drag

A216-WCB

A216-WCB

Stellite

Stellite

1/4 in.

1 1/2 in.

1500 lb

A216-WCB Stellite

A-36

Stellite

Stellite

1/8 in.

1 3/4 in.

5000 lb

18
Figure 7. During closure, the Valve A body guides should provide proper disc alignment until the disc contacts the

entire seat ring surface.
Figure 8. The Valve A disc, which shows sealing surface damage on the downstream face.
Figure 9. Close-up of the Valve A lower right guide surface, which shows yielding, plastic deformation, and

significant galling on the lower edge of the disc.
FIgure 10. The Valve A body guide, which mates with the right disc guide (shown in Figure 8), shows galling near

the fully closed position.
Figure 11. Close-up of the Valve A lower left guide surface, which shows yielding and plastic deformation similar

to the right side of the disc.
Disassembly and inspection of Valve B's internal

surfaces showed only very minor galling on the body

guides. The disc guide surfaces were hardfaced and

showed no sign of wear.

It must be noted that a survey of the valves installed in

these BWR isolation functions shows that the majority of

the valves now in use have nonhardfaced guide surfaces.

The manufacturer of Valve B started hardfacing disc

guides after 1970, and the other manufacturers only

hardface guides on special request. We do not believe the

nonlinear performance of Valve A is solely a function of

whether the disc guide surfaces are hardfaced, but rather

the nonlinear performance is primarily a function of the

large disc-to-body guide surface clearances, which

allowed the disc to tilt and thus reduce the contact bearing

area.

As more and more test data were analyzed, the

adequacy of Equation (1) appeared to depend on many

fluid parameters. The only tests where Equation (1) with

a 0.3 disc factor predicted the response of either valve

were the ambient temperature tests with upstream

pressure loads. All tests at normal operating temperature

required a higher disc factor. In order to assure ourselves

that we were not adding conservatism to an inappropriate

equation, we performed a detailed analysis, as shown in

subsequent areas of this report on each term in the

equation to determine if there was aterm missing or if the

disc factor had been underestimated by industry.

It must be also noted that we were aware of the

Westinghouse work performed after the EPRI

power-operated relief valve (PORV) and PORV block

valve tests conducted at the Duke Power, Marshall

facility. Westinghouse found that their disc factors

needed to be increased significantly to account for the

added friction in the disc-to-body guide surfaces. The

initial differences we saw comparing our data with the

Westinghouse data were that the Westinghouse data were

for stainless steelvalves in a steam environment and their

problems did not go away at ambient temperature. Our

carbon steel valves seemed to be more sensitive to

temperature and to the fluid properties as the fluid

approached saturation.

4.2 Stem Force-Static

Components

We initially divided Equation (1) into static and

dynamic components. In order to investigate how well

Equation (1) models the actual behavior of the valve, the

conservatism needs to be removed from the calculation.

If the disc load is eliminated from Equation (1), such as

would be the case without flow, what remains is a linear

equation in slope-intercept form (y = mx + b), namely

Fs = -As(P) ± Fp - (3)

Note that this equation has been written so that the stem

rejection load is always negative (compression), while

the packing load is either negative orpositive depending

on whether the valve is closing (compression) or opening

(tension).

Figures 12 and 13 show the test data for Valve A that

apply to the above equation. The dataplotted are the stem

forces measured at mid-stroke (running load) for tests at

varying temperatures and pressures but without flow.

The line fit through the data points has a slope equal to the

stem cross-sectional area and provides an indication of

the true packing load for each case. The data show a

packing load of 835 lb for opening and 430 lb for closing

for Valve A. Both values are well below the 1500-lb

maximum packing load used by the manufacturer in the

sizing equation. The difference between the two values

can be partially accounted for by the weight of the disc

and lower half of the stem. This difference also provides

evidence that the packing load is affected by direction of

travel, possibly caused by water carried with the shaft

changing the lubrication of the packing/stem surfaces or

by other phenomena associated with stem travel through

packing.

Figures 14 and 15 show the Valve B test data that apply

to Equation (3). Here again, the data plotted are the stem

forces measured at mid-stroke for tests at varying

temperatures ind pressures but without flow. The line fit

through the data points has a slope equal to the stem

cross-sectional area and provides an indication of the true

packing load for both the opening and closing strokes.

The increase in packing forces over those found for

Valve A is believed to result from the different packing

design and greater stem diameter of this valve.

24
1.0

C>

-o

0

C)

0

x

.4-,

.0X,

.6

.2

-.2

-.6

-1.0

-1.4

-1.8

-2.2

0. .2 .4 .6 .8 1.0 1.2 1.4 1.6

System Pressure (psig x 1000) MqHO2 151

Figure 12. A linear curve fit using the industry equation closely approximates the Valve A running stem forces for

opening without flow.

0.

0)

C0

-D

-0

W

E

4-

-.5

-1.0

-1.5

-2.0

-2.5

-3.0

-3.5

-4.0

-4.5

0. .4 .8 1.2 1.6 2.0 2.4

System Pressure (psig x 1000)

Figure 13. A linear curve fit using the industry equation closely approximates the Valve A running stem forces for

closing without flow.

25
x

-

X

0-

cI

.4-

-4

0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 2.0

System Pressure (psig x 1000))

MGH02153

Figure 14. A linear curve fit using the industry equation closely approximates the Valve B running stem forces for

opening without flow.

x

- -4

Cl)

4-

E 6 -T -3

-o -

ai)

E -6

4-

Uf)

-8

0. .4 .8 1.2 1.6 2.0 2.4

System Pressure (psig x 1000) MGH02 154

Figure 15. A linear curve fit using the industry equation closely approximates the Valve B running stem forces for

closing without flow.

26
The line fit through the data points for Valve B shows a

packing load of 1610-lb opening and 1632-lb closing.

The maximum force calculation used by the

manufacturer for this case included a 5000-lb packing

load. The difference between the opening and closing

values is less than expected, given the weight of the disc

and lower stem half (approximately 50 lb). This may

indicate a directional relationship for packing load,

believed to be a characteristic of the packing type used

and its orientation.

The measured packing loads for both valves were

below the maximums estimated by the manufacturers,

and the stem rejection loads were indeed linear with

pressure as predicted in Equation (1) for these no-flow

tests.

4.3 Stem Force-Dynamic

Component

Throughout the testing, bothvalves repeatedlyisolated

flow, although at a higher than anticipated thrust. In

generating the analytical thrust calculations for

comparison to measured valve stem forces, we used

actual response histories for all the variables in

Equation (1) except the stem factor. This comparison of

measured stem forces with values calculated using both

the 0.3 and 0.5 disc factors provides a common basis for

discussing valve operating characteristics. In the

previous section on the static component forEquation (1)

we found good agreement with the measured loads versus

the calculated loads. This would indicate that if there is a

problem with the equation, it is in the dynamic

component.

Previously in Figures 3 through 6 we saw that the

forces required to close the valves were above those

calculated using the 0.3 disc factor, however, using a 0.5

disc factorwe came close to bounding the linearbehavior

of Valve B. We were also able to explain the nonlinear

behavior of Valve A, which was apparently due to its

design (large guide tolerance). However, as we increase

the fluid subcooling, Valve A returns to more linear

behavior, as shown in Figure 16. The inlet pressure is the

same pressure (1000 psig) as in Figure 5; however, the

temperature is 130°Fless than in Figure5. Figures 17 and

18 show this same relationship at the 1400 psig inlet

pressure. The actual stem force in Figure 17 shows the

same nonlinear response as shown in Figure 5 (both 100F

subcooled cases). Figure 18 shows the same actual versus

calculated relationship as shown in Figure 16, which are

both 1300F subcooled tests; note that the behavior is

linear.

The upstream temperature and pressure measurements

(see Figure 2, measurement locations P-2 and T-2) in the

100F subcooled tests were evaluated. Figure 19 shows

that upstream flashing occurred from about the 5%

through the 55% closed position, in the 1000 psig, 101F

subcooledtest. The nonlinearbehavior showninFigure 5

does not start until about the 26-s time line, which

equates to about 66% closed. While flashing and

two-phase flow may have occurred upstream ofthe valve

in the 100F subcooled cases, the upstream liquid is

apparently recovered to subcooled before the start of the

nonlinear behavior. The actual differential pressures

across the valve disc in these four tests are incrementally

part of the calculations and are thus accounted for,

however, the peak of the nonlinear behavior was at flow

isolation and the differential pressure for all four tests at

isolation is near 100% orequal to the upstream pressure.

The fluidproperty differences between the 10 and 1301F

subcooled cases are not significant. The density is higher

in the 1300F subcooled case but the velocity is lower, in

the 100F subcooled case the density is lower and the

velocity is higher. The small differences in density and

velocity do not account for the differences in valve

response.

The single major difference in the tests is fluid

temperature. The valves are preheated before the flow

interruption test to the initial temperature for that test. As

the valve disc closes, lowering the pressure downstream

of the disc, the valve is much warmer than the coolant

saturation temperature downstream of the valve; this

temperature and the lower downstream pressure could

dry out the sliding surfaces between the valve disc andthe

valve body. This would change the interface between the

disc guide and valve body guide from a film of water to

steam. Steam is much less effective as a lubricant than

water, therefore the friction factor on the guide surfaces

could be much greater after dryout. This phenomenon

(drying out) would occur sooner during the tests starting

at 10F subcooled.

27
Valve A, Test 4, Step 5, 1000 psi, 400 OF (1300F subcooled)

.0

-0

U,

6

4

2

0

-2

-4

-6

-8

-10

-12

-14

-16

-18

-20

-22

-24

-26

4 6 8 10 12

Time (sec)

14 16 18 20

MGH02 155

Figure 16. With an increase in the fluid subcooling, Valve A returns to more linear behavior and stem forces are

bounded by the 0.5 disc factor calculation.

Valve A, Test 5, Step 5, 1400 psig, 580 OF (10 0F subcooled)

20.0

C)

x

.0

0

E

0)

4-

10.0

0.0

-10.0

-20.0

-30.0

0 2 4 6 8 10 12 14 16 18

Time (secd

20 22 24 26 28 30

MFHg2 162

Figure 17. At 1400 psig and 10O0Fsubcooled, Valve A shows the same nonlinear response as in Figure 5.

28
Valve A, Test 7, Step 5, 1400 psig, 450 OF (130 0F subcooled)

C.)

Li,

0

E

C,,

20.0

10.0

0.0

-10.0

-20.0

-30.0

0 2 4 6 8 10 12 14 16 18 20 22 24 26

Time (sec)

MGH02 163

Figure 18. At 1400 psig and 1300F subcooled, Valve A returns to more linear behavior and stem forces are again

bounded by the 0.5 disc factor calculation.

Valve A, Test 2, Step 5, 1000 psi, 530 'F

560 I I II I I I I I I I 6 I I I * i

550 _

LL

0)

E

0)

I540

530

520

~~~~~~~~~~~~II

- Upstreamn temp.

- -- Saturation temp.

I I I . I I I I I I 1 I I I I I

I aly r

500

0

. . .

20 40 60 80 100

Stem Position (% close) MGH02156

Figure 19. Flashing occurred upstream of Valve A during the early part of the closure stroke, but recovered to

subcooled prior to the start of the nonlinear behavior (compare Figure 5).

29
A second example of temperature affecting stem thrust

requirements is shown in Figures 20 through 23.

Figures 20 and 21 show the stem force requirements to

open Valves A and B respectively, against upstream

pressure with the downstream side vented (opening AP

test) at ambient temperature. The calculated forces are

again shownwith a 0.3 and 0.5 disc factor. As can be seen,

both valve opening histories are enveloped by the 0.3

calculation. Figures 22 and 23 show the same opening

tests with a slightly lower starting pressure, but at 580RF.

Note the actual stem force requirements have increased

and are no longer enveloped by the 0.3 disc factor

calculation. These tests were performed during the initial

valve qualification tests. The slightly lowerinletpressure

is due to temperature considerations on the 900-lb class

valves. The primary difference in test conditions is

temperature.

Review of the dynamic component of Equation (1)

shows that the total force was measured, the disc area at

anytime during the opening or closing cycle was known

from measured disc positions, and the differential

pressure was measured. The unknown value is the disc

factor. A disc factor of 0.3 has been used to size most gate

valve motor operators in the past. That numberappears to

be unconservative for the valves tested in this program at

conditions above ambient temperature. The 0.5 disc

factor appears to marginally bound most linear valve

responses at temperature. Nonlinear valve responses

may be caused by design problems, thus Equation (1) is

inappropriate for predicting those responses.

4.4 Valve Opening Versus

Closing Tests

Valve opening stem forces and the comparison of

opening loads versus closing loads were evaluated for the

following reasons: (a) the safety function of some valves

is to open against the full system AP and (b) opening tests

have been used by industry in the past to predict closing

loads. If this practice can be substantiated, it might in

some cases provide utilities the ability to determine disc

factors in situ.

As discussed in the previous section under temperature

effects, we performed a cold opening AP test at 1700 psig

and less than 1000Fduring the qualification tests required

by ANSI B 16.41, Annex B. The valve was closed under

static pressure in this test and the downstream piping was

vented. The valve was then opened with the system

differential pressure across the disc. Figure 20 shows the

data from this test with Valve A. The stem force history

starts out with the valve closed and the valve stem in

compression. As the opening cycle begins we see a

decrease in stem compression and then the stem goes into

tension with the operator hammer blow. Next, the stem

force shows the drag load as the disc slides on the seat and

then a decrease in tension as the valve unseats and system

pressure is equalized. Once pressure on both sides of the

valve is equal, the stem force shows the sum of the

packing drag load (tension) and the stem rejection load

(compression). Figure 21 shows the same test for

Valve B. The figures also show a comparison between

the measured stem forces and calculations using the 0.3

and 0.5 disc factors. As previously stated, the 0.3 disc

factor calculation envelopes the forces measured for both

valves during the cold opening AP test.

Reviewing Figures 22 and 23, which show the results

of the Annex C hot opening AP tests against a slightly

lower pressure differential and with a fluid temperature

of 580'F, we again see the forces needed to open the

valves are significantly higher than those measured

during the cold tests, and the calculation with a 0.5 disc

factor only marginally envelopes the responses. The

temperature effects on the disc factors show that a correct

disc factor cannot be determined from cold testing.

The hot opening AP test may identify the correct disc

factorforvalve closing requirements, which have shown

linear disc friction characteristics through complete

qualification, including flow interruption testing.

Figure 24 shows the results from the Valve B opening AP

test performed at BWR normal operating conditions just

before the full flow isolation test. The relationship

between actuals and calculated is very similar to the

Valve B Annex C test (the results of which are plotted in

Figure 23) and also to the pipe break flow isolation test

for the same fluid conditions shown in Figure 25. The

data from all three tests appear to support the use of adisc

factor very near 0.5.

A similar analysis for Valve A was not successful as

none of the opening or lightly loaded closing tests

providedinsights intothe nonlineardisc friction behavior

of Valve A seen in the 1 O0Fsubcooled flow isolation tests,

shown in Figures 5 and 17.

30
10.0

-o l_ __0.0

° -10.0 - Actual - Calculated (disc factor = 0.3)

E 7--- Calculated (disc factor = 0.5)

-20.0

-30.0 ,, ,, ,, ,, , 0 5 10 15 20 25 30 35

Time (sec)

MGH02132

Figure 20. With Valve A opening against differential pressure only (no flow, cold fluid), the measured thrust is

enveloped by the 0.3 disc factor calculation.

Valve B, Test 1, Step 12, 1700 psi, -100 'F (no flow)

20

-o

0)

P,

10

0.0

-10

-20

0 10 20 30

Time (sec)

40

MGH02 133

Figure 21. With Valve B opening against differential pressure only (no flow, cold fluid), the measured thrust is

enveloped by the 0.3 disc factor calculation.

31
Valve A, Test 1, Step 19, 1650 psig, 580 OF (no flow)

20.0

o 10.0

x 0.0

-0

C)

> -10.0

E

un -20.0

-30.0

a 5 10 15 20 25 30 35

Time (sec)

MGH02134

Figure 22. With an increase in fluid temperature, the Valve A measured thrust is not enveloped by the 0.3 disc

factor (compare Figure 20).

Valve B, Test 1, Step 19, 1650 psi, 580 OF (no flow)

20

C>

C>

x

Q>

0

E

CD1-,

10

0

-10

-20

0 10 20 30

Time (sec)

40

MGH02139

Figure 23. With an increase in fluid temperature, the Valve B measured thrust is not enveloped by the 0.3 disc

factor (compare Figure 21).

32
Valve B, Test 2, Step 2, 1000 psi, 530 OF (no flow)

20

C)

x

a,

C0

E

a,

Co

10

0

-10

-20

-30

0 10 20 30

Time (sec)

40

MGH02157

Figure 24. The Valve B AP opening test at normal BWR operating conditions, like the high flow interruption test

(see following figure), appears to isolate the disc factor at or near 0.5.

Valve B. Test 2, Step 5, 1000 psi, 530 OF

Z0

-0

.0

0

ECY0

6

4

2

0

-2

-4

-6

-8

-10

-12

-14

-16

-18

-20

-22

-24

-26

0 5 10 15 20 25 30 35

Time (sec)

MCM0215S

Figure 25. For the Valve B pipe break flow interruption test at normal BWR operating conditions, the relationship

between measured and calculated stem thrust is very similar to that seen in the AP opening test

(previous figure).

33
The test results indicate that valve opening and opening

versus closing comparisons are dependent on the

following factors, which unfortunately may make the test

infeasible for some valve locations.

1. The valve and fluid have to be at operating

temperature and pressure since cold water tests

do not produce the same results.

2. Type testing must have been performed to verify

that the valve design exhibits linear disc friction

characteristics.

3. The inlet pressure source must be large enough

for in-situ testing so that the increasing leakage

as the valve opens does not drop the inlet

pressure significantly, before maximum disc

load is determined.

4.5 Stem Factor

The test results were also analyzed to determine the

adequacy of the torque to thrust conversion

[Equation (2)], specifically the stem factor term. As

previously stated, valve and motor operator

manufacturers use one of two basic coefficients of

friction in these calculations, 0.15 or 0.20. The 0.15

constant coefficient of friction was found to be

conservative. The 0.20 value would be considered very

conservative unless the thrust, which might be obtained

from a motor operator sized for a 0.20 coefficient and

operating with a 0.10 coefficient, overstressed the valve.

While many people believe the stem factor is a constant,

we found that the stem factor actually varies with valve

loading.

We believe stem factor will surface as a problem as

maintenance and motor operator diagnostic testing is

performed. It is well known that as a motor operator ages

and as maintenance is performed, the correspondence

between the torque switch setting and the delivered

output thrust becomes less reliable. Modern diagnostic

test equipment for valve motor operators has allowed the

utilities to recalibrate the motor operator torque switch

in situ. However, the variability of the stem factor and the

deceptively high thrusts ofvalve seat-inducedtorqueouts

could result in improperly set motor operator control

switches. Test results and the following analysis point out

some of the ways in which the stem factor can be a

problem in recalibrating motor operator output torque.

Valve B was selectedforthis analysisbecause of itslinear

performance during both the opening and closing tests

and because the Valve B motor operator was normally

sized for the loadings. This allowed adequate torque

spring deflection and good operator output torque to stem

thrust comparisons.

Analysis of our test results showed that the final thrust

in the valve stem varied depending on how the motor

operator was loaded before and at torque switch trip.

Initially we believed this variability to be a function of

motor operator momentum; however, the measured

motor operator parameters did not bear this out. The

measured parameters did show that when the valve was

lightly loaded before the moment the disc wedged in the

seat, the stem factor was low. When the valve was highly

loaded before seating, the stem factor was higher,

resulting in a poorer conversion of the torque to thrust.

With high loads prior to valve sealing, the torque spring

was deflected by the disc load almost to the point of

torque switch trip before the disc first contacted the seat.

The initial contact with the seat, combined with the disc

loads, was enough to trip the torque switch. From this

point to the time the motor controller drops out and the

motor operator momentum is spent, the worm acts like

the input to a planetary gear where the remaining

revolutions of the motor are split between the worm

turning the stem nut and the worm climbing the worm

gear and compressing the torque spring past the torque

switch trip point (see Figure 26). With light loads,

however, the disc is already wedged very tightly at torque

switch trip, and the remaining revolutions of the motor

are not split but all go into overcompression of the torque

spring, and thus the resulting final stem forces are higher.

The test results discussed below are consistent with this

explanation.

Before the start of the qualification test, we set the

torque switch to deliver 18,000 lb of thrust, as specified

by the valve manufacturer for a full flow closure at

1,400 psig. In setting the torque switch, we used the load

cell installed in the valve stem to measure the thrust, and

we manually turned the handwheel to close and seat the

valve, so there was no motor momentum involved with

the determination of the torque switch position versus

output thrust relationship.

34
Figure 26. At the torque switch trip, the worm may either turn the worm gear and drive the disc deeper into the

seat, or (if the disc will move no further) climb the worm gear, overcompressing the torque spring and

producing additional thrust in the valve stem.

Figure 27 shows the forces measured as the valve

closed againstpressure only. This test is typical of what a

utility might be able to do. Note the final thrust

(22,000 lb) with the valve lightly loaded and with torque

switch trip induced by the disc wedging in the valve seat

When the disc contacts the seat and the torque switch is

tripped, power continues to be supplied to the motor until

the motor controller drops out (typically a time lag of 15

to 60 ins). At this time, the valve disc is wedged deeply in

the seat. After the motorelectricalpower is broken, there

is a period of deceleration of the motor operator

components. This deceleration is proportional to the

speed and mass of the motor operator, primarily the

motor. (With fast acting valves, there can also be

significant momentum in the valve internals.)

Both the dropout time of the motor controller and the

motor operator momentum show up as additional force

after torque switch trip; these additional revolutions of

the motorproduce little if any additional movement ofthe

disc and instead result in overcompression of the torque

spring. In seismically qualified valves with very stiff

yokes, the motion is divided between overcompression of

the torque spring and compression of the stem. In

addition, the low stem factor that accompanies these

relatively low valve loadings allows a better conversion

of torque to thrust, producing a higher measured force in

the valve stem at torque switch trip.

In Figure 28, we see the forces measured as the same

valve closed against three different pressures at high

flows. Note that with the same torque switch setting, the

force when the torque switch tripped in the 600 psig test

with high flow is less, at 18,100 lb, than the force when

the torque switch tripped in the no-flow static pressure

test, at 19,900 lb (see Figure 27).

35
10.0

-Q

a)

CIO

0

:5

E -20.0

CD

-30.0

Valve B, Test 2, Step 1, 1000 psig, 530 'F (no flow)

Start of closing cycle

Load at torque switch_--

trip 19,900 lbs. thrust

Final seating thrust 22.000 lbs.

0 5 10 15 20

Time (sec)

25 30 35

MGH02136

Figure 27. Though the torque switch was set to trip at 18,00-b thrust in the absence of operator momentum, this

lightly loaded valve achieved a significantly higher final seating thrust.

Valve B, Step 5, Tests 4, 2, and 3 (line break flows)

5

- 0.0

-5 x

-10

a)

L. 0.

-15

E

-20

-25

0 5 10 15 20 25 30 35

Time (sec)

MGH02137

Figure 28. As stem factor and operator momentum are affected by increased loadings, the final thrust is less even

though the torque switch setting is the same.

36
The valve closing at 1,000 psig shows a significantly

higher load before isolation of flow. Just before this test,

the valve stem was lubricated, and a slightly higher thrust

(I 8,600 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 force was a reflection more of

closing load than of seating load. During the closing at

1,400 psig inlet pressure (the design basis for operator

sizing and torque switch setting), the valve marginally

isolated flow but did not seat; the operator tripped on disc

friction. The thrust when the torque switch tripped was

lower, at 16,500 lb, a 17% reduction in the thrust at torque

switch trip and a 25% reduction in final thrust, as

compared to the lightly loaded case, shown in Figure 27.

The diagnostic equipment monitoring the operator

performance showed that in contrast to the varying thrust,

the operator output torque varied less than 3% for all

valve loadings.

The variability of the stem factor under changing valve

stem loads is shown in Figure 29. This stem factor history

is derived from measurements of stem force and

measurements of torque spring deflection

mathematicallyconvertedto operatortorque. This figure

shows that the stem factor increased with load, resulting

in a less efficient conversion of operator torque to valve

stem thrust. In this and other stem factor versus load

comparisons, it appears there may be a proportional

relationship between the increase in load and the increase

in stem factor.

4.6 Operator Torque Switch Trip

Anomaly

During the flow interruption testing of Valve A, there

were three incidences of anomalous operator torque

switch trip behavior. It is believed these incidences

occurred in conjunction with installation and removal of

the MOV diagnostic test equipment. The valve stem

forces associated with the torque switch trip were normal

in the numerous tests performed with diagnostic devices

installed. The anomaly appeared in the form of

abnormally low values of the torque-out stem forces

during the tests immediately afterremoval of two types of

diagnostic equipment (see Table 2 for installation

and removal sequence of diagnostic equipment). The

Valve B, 1400 psi, 580 OF

0

o -5

.0

-0

-10

1D

0

E

a -15

-20

.014

.013 50

.0

.012 o

Q

a)

.

ULE

0 .011 4-I

U,

-J .010

18 20 22 24 26 28 30

Time (sec)

MGH02138

Figure 29. The stem factor increases with load, reducing the available thrust at a given torque.

37
investigation that followed the discovery of the low stem

forces showed that incorrect installation of the motor

operator spring pack lock ring was the problem. The

removal of the diagnostic test equipment and the

subsequent incorrect installation of the lock ring

invalidated the findings of the diagnostic test. A recent

problem and investigation at Brunswick

(LER 87-023-01)2 identified a similar lock ring

installation problem and illustrates the potential for

invalidated diagnostic testing and out-of-calibration

torque switch positions.

The point at which the torque switch contacts open

depends only on the setting of the torque switch, spring

constant of the torque spring, and spring pack preload

and/or gap. No matter what causes the stem force to

increase, whether flow loads, valve reaching full stroke,

oreven an obstacle inthe discpath, the switch will always

open when the torque spring compresses to the

predetermined point. The force at torque switch trip was

used to trace the function of the operator from one test to

installations and removals. Figure 30 shows the average

stem compression at torque switch trip for each of the

eleven tests performed on Valve A, arranged in

chronological order. The force measurements were made

using the INEL load cell installed as an integral part of the

valve stem.

During tests I through 7 the valve operator functioned

consistently, with a stem compression at torque switch

trip of approximately 33,000 lb. Tests 9 and 8,

accomplished without operator diagnostic monitoring,

showed consistent torque-out forces, but at a

significantly reduced level. Here a drop of approximately

10,000 lb appeared in the torque-out stem compression.

Two different sets of valve operator diagnostic

equipment were installed to monitor test 10, and the valve

stem torque-out compression returned to about the same

level as tests 1 through 7 (relubrication of the valve stem

threads increased loads slightly). The diagnostic

equipment was removed after test 10, and the results of

test 11 show a similar reduction in force, even after the

another through various diagnostic equipment torque switch setting was increased from 2.0 to 2.5.

Average Stem Force at Torque-out

40

35 -

30'

X

-0

0

0

E

Wl

C',4

25

20

15

10

/' Z/  /'/ z/Z O/ Z 0

0 X

I

5

0

1 3 2 4 6 5 7 9 8 10 11

Test Number

MGH02159

Figure 30. Below-normal stem forces at torque switch trip were observed for three test series with Valve A.

38
After completing testing for Valve A, the Limitorque

motor operator was removed and partially disassembled

by Limitorque representatives, with INEL personnel

attending. The spring pack cover was removed and the

internal configuration inspected. The lock ring that

retains the torque spring and its locking set screw

appeared to be properly installed. The set screw was

removed and a special tool was used to attempt to further

tighten the lock ring. The ring was tightened almost one

full turn before it reached its proper position.

Limitorque design documents were used to correlate

the loosening of the lock ring to torque switch setting and

torque-out thrust. One full turn of the lock ring is

equivalentto 19 degreesrotationof the torque switch; one

full torque switch setting is about 21 degrees. This

loosening of the lock ring had the effect of backing off the

torque switch from 2.0 (the actual setting) to 1.1 (the

equivalent setting). From the torque spring curve the loss

of thrust was estimated at 10,600 lb, very close to the

discrepancy in the measured data.

Improper positioning of the lock ring sometime after

test 7 but before the next test would explain the reduction

in stem force after test 7. How it happened is not

completely understood. None of the diagnostic devices

installed before test 9 required the removal or adjustment

of the torque spring lock ring; in fact, several of the

devices are designed to detect spring pack gap, the result

of improper lock ring installation. Review of the data

taken by the various diagnostic devices shows no

indication of spring pack gap. Also, none of the devices

are designed suchthat theirinstallation would correctthis

problem, with the exception of the Limitorque motor

actuator characterizer (MAC) device, which was

installed for test 10.

Installation of the MAC device requires the removal of

the torque spring lock ring to facilitate the installation of

its spring pack load cell device. According to the

Limitorque technician, the position of the lock ring was

marked before removal and the number of turns during

removal was noted. The load cell device was installed

and tightened to the proper position to provide the design

spring preload. After testing, the load cell was removed

and the lock ring was installed the appropriate number of

turns to the previously marked position. We believe this

explains the similar reduction in stem force before and

after test 10.

A similar problem at Brunswick was evaluated using

the information found in LER 87-023-01. In this case,

the HPCI steam line isolation valve (a GI-87 valve) had

successfully undergone several diagnostic tests using the

MAC system. Later, the valve motor failed on opening

for an unrelated reason. During the subsequent motor

operator check, greatly reduced torque-out forces were

measured. Investigating personnel discovered that a burr

on the threads of the spring pack housing cover had

prevented the lock ring from being fully installed after

diagnostic testing and had caused the

lower-than-expected torque readings.

Both the GI-87 testing and Brunswick instances of

improper lock ring positioning could have been easily

diagnosed. A simple measurement of the lock ring

position can be compared with both the position of the

torque spring transducer during testing and the

manufacturer design position in order to validate

post-testvalve operation. Apparently this procedure was

not completed for the tests described above.

39
5. CONCLUSIONS

The typical industry sizing equation using the standard

variables did not conservatively estimate the total thrust

needed to close the tested valves; disc factors higher than

the normal 0.3 disc factor (pu) were encountered. The

valve thrust equation [Equation (1)] needs to better

model the behavior of valves exposed to slightly

subcooled fluid conditions in BWRs. The disc factor

needs to be increased for both the opening and closing

direction to account for the higher loads associated with

high temperature operation. The thrust sizing equation is

not applicable to valves that sustain damage (such as

galling and plastic deformation of the sliding surfaces) at

design basis loadings. Flashing and two-phase flow

appearto add ayet unquantified factorto the closing load.

The design basis hot water blowdown testing has

shown that, given enough thrust, typical gate valves will

close against the high flow resulting from a line break.

Proper operator sizing depends on correct identification

of the values forthe sizing equation. Evidence exists that

values used in the past may not be conservative for all

valve applications, especially at design basis loadings.

The following items need to be considered during sizing

of gate valve operators:

1. Gate valve guide design and clearances canhave

a significant effect on the operator stem thrust

requirements at design basis fluid loadings.

2. The degree of subcooling at the valve inlet can

greatly influence valve closure forces. Valve

operator force requirements increase as inlet

fluid conditions approach saturation

temperatures.

3. Industry trends toward using 100% system

pressure for all pressure terms in the sizing

calculation are justified for high-flow

applications.

Tests have shown that some form of valve type testing

outside the plant might be necessary to establish specific

valve design thrust requirements and verify that a given

valve design exhibits linear characteristics when

subjectedtodesignloads. Forthe valvesthathavealinear

thrust response, valve opening tests (with a full pressure

drop and no flow) at normal operating temperatures

performed with valve diagnostic test equipment can

provide insights for the valve disc factor and therefore

degradation in valve performance for both opening and

closing. Diagnostic test equipment that measures both

thrust and torque will provide a better set of measured

values for analysis.

Contrary to common belief, the ratio of operatortorque

to stem thrust [stem factor (y, )] is not a constant but

changes with valve loading. The effect of a changing

stem factor on in-plant testing is significant. In-situ

testing loads (the thrust developed when the process of

valve seating causes the torque switch to trip) typically

result in low stem factors and high stem thrusts. A utility

might set the torque switch to a lower value to protect the

valve from excessive thrust at normal loadings due to

both the high stem thrust at torque switch trip and large

increase in force due to momentum loading after torque

switch trip. Then, if the valve is subjected to design basis

loadings, the stem factorwill increase, and the thrust may

be too low. Diagnostic systems that measure operator

torque and thrust would provide stem factor information

for extrapolation, thus assuring that, if the valve is

subjected to design basis loadings and the stem factor

increases, the valve will have sufficient thrust to close.

Although not conclusive, these test results suggest that

the change in stem factor may be predictable. If the

relationship between stem factor and load can be

established, the change in stem factorcould be quantified

for design calculations and in-situ testing.

Improper operator lock ring installation following test

or maintenance can invalidate in-situ test results and

render the valve unable to perform its design function.

Thisisimportantinlightofthepresent trendbyutilitiesto

perform diagnostic testing of safety-related valve

assemblies to answer regulatory concerns such as those

expressed by IE Bulletin 85-03. A final quality check

following diagnostic testing and maintenance must be

made to ensure correct lock ring installation.

41
6. REFERENCES

1. ANSI B16.41, Functional Qualification Requirementsfor Power OperatedActive Valve Assembliesfor Nuclear

Power Plants, 1983.

2. Brunswick Steam ElectricPlantUnit 1, Inoperability ofHigh Pressure lnjection (HPCI) System (E41 ) Due to Failure

of HPCI Turbine Steam Inlet Isolation Valve E41-FOOI, LER 87-023-1, June 27, 1988.

43
NRC FORM 335 U.S. NUCLEAR REGULATORY COMMISSION 1. REPORT NUMBER

1291g (Aisgnod by NFIC. Add Vol.9, uPc., Flv..

NF; cm 1102. Ol~~~~~~~~~~~~~~~~~~~~~~~~d Addomdu- Numbqrg f ny 3201.3202 BIBLIOGRAPHIC DATA SHEET

fSee instructions on the ,ersee NUREG/CR-5406

EGG-2569

2. TITLE AND SUBTITLE VolE . 1

BWR Reactor Water Cleanup System Flexible Wedge Gate Isolation

Valve Qualification and High Energy Flow Interruption Test 3. DATEREPORTPUBLISHED

MONTH 'fEAR

Analysis and Conclusions October 1989

4. FIN OR GRANT NUMBER

A6857 and B5529

5. AUTHOR(S) 6. TYPE OF REPORT

Kevin G. DeWall Technical

Robert Steele, Jr. 7. PERIOD COVERED tincssi.ie otcsi

B. PERFORMING ORGAN IZATION - NAME AND ADDR ESS JI NRC. provide Diision, Ofset or Re9;on. U.S. Nuclear Rellawry Commission and mailing adrs:iF contractor. pro ir

name and mailing ddrs.,)

Idaho National Engineering Laboratory

EG&G Idaho, Inc.

P.O. Box 1625

Idaho Falls, ID 83415

9. SPONSORING ORGANIZATION - NAME AND ADD R ESS Wtt NFC. type sne as abowe- if contracror. provide NRC Disi~on, Office or Regon. US. Nuclear Regfusaowv Commirson,

and maoting address.J

Division of Engineering and Division of Safety Issue Resolution

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555

10. SUPPLEMENTARY NOTES

11. ABSTRACT 1200words or less)

This report presents the measured data and the analyses performed to date on the full-scale high-energy

qualification and flow interruption gate valve testing to develop technical insights for the United States Nuclear

Regulatory Commission (USNRC) effort regarding Generic Issue 87 (GI-87). The research was sponsored by the

USNRCa and conducted by researchers from the Idaho National Engineering Laboratory. We tested two 6-in., 900-lb

class valve assemblies, which represent a significant percentage of the reactor water cleanup isolation valves installed in

plant applications. These valves were modified before testing by adding a high temperature load cell in the valve stems,

which allowed the direct measurement of valve stem thrust during both opening and closing valve cycles.

Instrumentation installed in the flow loop and on the valve assemblies measured the important valve and system test

responses. Additionally, during the testprogram, all of the currently popularmotor operated valve diagnostic test systems

monitored the performance of the valves. Initially the valves were subjected to the hydraulic and leakage qualification

tests definedinANSIB 16AI and then to flow interruption and reopening valve tests at boiling waterreactorprimary

system water temperature and pressure conditions with downstream line break flows. For the two valves tested, results

show that (a) the disc factor used in current industry motor operator sizing equations underpredicts actual valve thrust

requirements at all high temperature loadings, and for one valve design the equations may require an additional term to

account fornonlinearperformance, (b) the thrusts required to close the valves were sensitive to the fluid temperature, and

(c) the results of testing at lower pressures, temperatures, and flows cannot be extrapolated to design basis pressures,

temperatures, and flows for valve designs that have not exhibited linear performance behavior during design basis

_ prototypical testing. 12. L Y wulHu utbumi:r I urs (List worits or phrases MA t will asst Oewarphan in locating the rapotj * 13. AVAILABILITY STATE MENT

valve testing Unlimited

BWR reactor water cleanup 4. SECURITY CLASSIFICATION

Generic Issue 87 (GI-87) (This

Unclassified

(This Report)

Unclassified

15. NUMBER OF PAGES

16. PRICE

NRC FORM 335 (2-891
UNITED STATES

NUCLEAR REGULATORY COMMISSION

WASHINGTON, D.C. 20555

OFFICIAL BUSINESS

PENALTY FOR PRIVATE USE, $300

SPECIAL FOURTH-CLASS RATE POSTAGE 6 FEES PAID USNRC

PERMIT No. G-57

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