df

NUREG/CR-4977

EdG-2505

Vol. 2

SHAG Test Series

Seismic Research on an Aged United States Gate Valve

and on a Piping System in the Decommissioned

Heissdampfreaktor (HDR): Appendices

Prepared by R. Steele, Jr., J. G. Arendts

Idaho National Engineering Laboratory

EG&G Idaho, Inc'

Prepared for

U.S. Nuclear Regulatory

Commission
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NUREG/CR-4977

EGG-2505

Vol. 2

R1

SHAG Test Series

Seismic Research on an Aged United States Gate Valve

and on a Piping System in the Decommissioned

Heissdampfreaktor (HDR): Appendices

Manuscript Completed: July 1989

Date Published: August 1989

Prepared by

R. Steele, Jr., J. G. Arendts

Idaho National Engineering Laboratory

Managed by the U.S. Department of Energy

EG&G Idaho, Inc.

P.O. Box 1625

Idaho Falls, ID 83415

Prepared for

Division of Engineering

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555

NRC FIN A6322

Under DOE Contract No. DEAC07-761D01570
SUMMARY

Volume 2 of this report contains Appendices A through E, which present details and

specifics of the tests and results reported in Volume 1. Appendix A, the valve anomaly

appendix, is a copy of the report that describes the investigation, results, and conclusions

of the INEL effort to determine the cause of the reduced performance of a naturally aged

Crane gate valve with a Limitorque motor operator. Appendix B describes the VKL

instrumentation for the SHAG test series. Appendices C and D provide supplementary

information regarding snubbers and transducers used in the HDR tests. Appendix B

discusses the analyses of the frequency response functions recorded during the HDR testing.

ii
CONTENTS

SUMMARY.......................................................................... ii

APPENDIX A-:VALVE ANOMALY .............. ....................................... A-1

APPENDIX B-INSTRUMENTATION OF THE VKL FOR THE SHAG TEST SERIES ..... ..... B-1

APPENDIX C-INVESTIGATION OF SNUBBER MALFUNCTIONS ...... ................. C-1

APPENDIX D-INVESTIGATION OF SPURIOUS DISPLACEMENT READINGS AT HDR ...... D-1

APPENDIX E-VALVE RESPONSE FREQUENCY ANALYSIS METHODOLOGY ..... ......... E-1

iii
APPENDIX A

VALVE ANOMALY
ABSTRACT

This report describes the investigation, results, and conclusions of the INEL effort to

determine the cause of the reduced performance of a naturally aged Crane gate valve with

a Limitorque motor operator. The motor-operated valve served 25 years in the Shippingport

Atomic Power Station as a feedwater isolation valve before being refurbished and installed

in a piping system in the Heissdampfreaktor (HDR), where valve operability in typical

pressure and temperature environments and during simulated earthquakes was studied.

During the test program it was discovered that under some valve hydraulic loadings the

motor operator failed to reach torque levels high enough to open the closing torque switch.

Failure of the torque switch to open caused the motor to go into a stall. In normal plant

service, stalling an operator motor can cause motor burnout and render the valve inoperable

for subsequent safety functions. The investigation concluded that the poor performance

of the valve was caused by heating of the motor windings and by external circuit resistance,

both of which prevented the motor from developing its rated torque. The investigation

also identified torque spring aging as a problem.

FIN No. A6322-Environmental and Dynamic Qualification Program

and FIN No. A6861-Equipment Survivability Program

A-ii
SUMMARY

During the HDR SHAG test program, we found that

under certain hydraulic loadings the naturally aged

motor-operated gate valve obtained from the Shippingport Atomic Power Station did not operate to design

requirements. The valve is a 900 lb class Crane gate

valve with a Liiitorque -model SMA size 2 motor

operator.

During most of the ambient temperature tests with

two valve cycles, the motor operator failed to open the

torque switch on the second closing cycle of the valve.

As a result, the motor stalled (while still drawing

power). During the elevated temperature tests with

higher system pressure, the motor stalled at the end

of both closing cycles.

We conducted an extensive investigation to try to

isolate the cause of the poor performance at the motor

operator. This investigation included follow-on in situ

tests at HDR, dynamometer testing of the motor

operator at the Limitorque laboratory, testing of the

torque spring at INEL, dynamometer testing of the

motor alone ithe Peerless Motor laboratory, and a

mathematical analysis of the HDR power circuit. The

investigation identified three causes of the motoroperator's poor performance: torque spring aging.

heating of the motor windings, and resistance in the

dc power cabling at HDR. The investigation also

demonstrated that normal plant testing of valves is not

adequate to ensure proper performance under flow and

pressure loads in combination.

During the follow-on tests at HDR, we found that

when the valve was subjected to flow loads and pressure loads in combination, the valve either torqued out

in the partially open position, stalled in the partially

open position, or stalled in the fully closed position,

depending on the load and the torque switch setting.

The valve torqued out in the fully closed position only

when pressure and flow loads were very low.

Dynamometer testing in the LiMitorque laboratory

showed that the operator developed 8000-lb thrust to

the stem at a torque switch setting of 3 at full rated

motor voltage. The motor operator was originally

designed to deliver 12,000 lb of thrust at a torque

switch setting of 3. The motor operator would require

a torque switch setting of about 3.6 to achieve an output thrust of 12,000 lb. Testing of the operator's torque

spring and of other new and aged torque springs at

INEL confirmed that the difference is the result of a

permanent set in the aged torque spring

The 125-Vdc motor is a 40-ft-lb-rated motor. At

motor stall, the motor should develop 62 ft-lb of torque

with a stall current of 115 amps. Analyses of the performance data taken at HDR show the motor developed

about 20 ft-lb at 50 amps at stall. Dynamometer testing

of the motor operator at the Limitorque laboratory

showed that the motor stalled at 38 to 45 ft-lb of torque

with a stall current of 66 to 75 amps.

The motor operator's motor was manufactured by

Peerless Winsmith. Dynamometer testing of the motor

at the Peerless laboratory, with the motor removed

from the Limitorque gear box, showed that when the

motor was at ambient temperature, the locked rotor

performance met the original design requirement.

However, the motor windings tend to heat up during

motor operation. When the motor was loaded with a

50 ft-lb load, its performance degraded at I amp per

second over a 20 8 period.

Dynamometer testing in the Limitorque laboratory

and in the Peerless laboratory provided data that exhibited far better motor operator performance than was

obtained at the HDR. At HDR, voltages and currents

were measured at the motor and, according to those

measurements, voltages did not drop below 107 Vdc

at motor stall. Even when voltages were dropped during dynamometer testing at Limitorque to 80 Vdc, the

HDR performance could not be duplicated.

Coinciding with our investigation Limitorque also

participated in a separate investigation of problems with

dc-powered valves at a U.S. utility. This investigation

pointed out some of the same problems that were experienced at HDR, along with some additional relay

and switch problems. Both investigations point out the

special requirements of dc-powered valves. The results

of the utility investigation prompted us to perform a

mathematical analysis of the power circuit to which

the motor-operator was connected at HDR. This analysis showed that the main cause of the valve's poor

performance at HDR was undersized power cables.

The cables were originally sized (at HDR and at

Shippingport) according to the motor controller ratings

and the continuous full-load current value stamped on

the motor nameplate. However, we now know that this

value is the power level at which the motor can run

without becoming overheated in its 5-min duty cycle;

five times that current level is required for the motor

to reach its potential maximum output.

As the loading slows the motor down, dc motors

develop their high torque, allowing more current to

flow. At motor stall, only the resistance of the motor

windings (approximately I ohm) limits the current

flow, combined with the resistance of the power cables.

Thus, the resistance of the power cables becomes an

important variable. Typically, the motor armature and

the field windings are connected in series through the

motor controller, so four long cables, instead of two,

A-iii
contribute to the resistance of the power circuit. Our

analysis showed that this kind of connection also makes

it impossible, using conventional methods and instruments, to get true measurements of voltage drop

across the motor.

According to our analysis, the I-ohm resistance of

the power circuit at HDR (combined with the 1-ohm

resistance of the motor windings) would reduce the output of the motor by approximately one-half at high

motor loadings such as those experienced during valve

closure against flows. These analysis results correspond well with the measured output at HDR as compared with the output at the Limitorque laboratory,

where the large cabling of the power circuit minimized

the influence of power circuit resistance.

Torque spring aging can be a problem in motor

operators in the field. If the spring has taken a permanent set, the operator might torque out too soon at

recommended torque switch settings and under fluid

load conditions, leaving the valve in a partially open

position.

Under some conditions, heating of motor windings

can be a problem in the field. When a valve subjected

to high loads is cycled more than once in a short period

of time, heating of the motor windings can be the determining factor that causes reduced performance, especially in a marginally powered valve. This was the case

at HDR during the two-cycle ambient temperature tests

in which the operator failed to torque out on the second

cycle of the test.

Undersized power supply cabling has surfaced

as a problem in at least two dc motor operators in

the field. Though the other two factors contributed

to the anomalous performance of the valve at HDR,

undersized cabling was the main cause. The NRC has

recently issued an information notice regarding the

issue.

None of the three problems discovered during the

HDR tests and follow-on investigation would be

detected during the normal in-plant testing where the

valves are subjected to no load or to pressure loads

alone. The problems are detectable only at higher

loadings, that is, flow loads in combination with pressure loads, where the load slows the motor down such

that momentum cannot carry the unit through complete

closure and torque out.

A-iv
APPENDIX A

VALVE ANOMALY

INTRODUCTION

As part of its equipment qualification research

program, the Nuclear Regulatory Commission (NRC)

installed a refurbished, naturally aged, 8-in. motoroperated gate valve in an existing piping system, the

Versuchskreislauf (VKL), at the decommissioned

Heissdampfreaktor (HDR) facility in West Germany.

The VKL was modified to simulate the flexibility of

a nuclear piping system supported by a typical U.S.

stiff piping support system. This 30-year-old, 900-ibclass, Crane gate valve was obtained from the

Shippingport Atomic Power Station, where it served

25 years (from 1957 until the plant was decommissioned in 1982) in the secondary system as an isolation valve. Identified as 53-H-2-S, the valve was used

to isolate the boiler feedwater pump discharge from

the safety injection lines. The valve is powered by a

Limitorque SMA-2 motor operator equipped with a

Peerless 125 Vdc motor.

The valve was included in the SHAG (Shakergebaude) seismic research program conducted by Kernforschungszentrum Karlsruhe (KfK) at the HDR

facility. The SHAG tests consisted of shaking the HDR

containment using a large coastdown eccentric mass

shaker. The objectives and results of the SHAG program are documented in Volume 1 of this report. The

equipment qualification program objectives were to

determine the valve's in situ response and operability

characteristics and compare these to those determined

from bench tests performed to equipment qualification

standards. The results of these comparisons and other

similar test comparisons will be used to supplement

the technical bases for assessing the adequacy of

nuclear equipment qualification standards and to propose any needed improvements. The Idaho National

Engineering Laboratory (INEL) was commissioned to

manage this part of the NRC's involvement in the

SHAG program.

During the SHAG test program, we found that the

valve's motor operator did not develop sufficient torque

to open the closing torque switch when the valve was

exposed to flow and static pressure loads in combination. Failure of the torque switch to open resulted in

motor stall. The investigation that followed concluded

that the poor performance of the valve was caused by

heating of the motor windings and by external circuit

resistance, both of which prevented the motor from

developing its rated torque. The investigation also identified torgue spring aging as a vrnhlzrm.

The potential safety implications of this problem are

significant. Valves powered by direct current, such as

this one, are typically installed in applications where

the valve function is of concern in losses of ac power

and where the valves are needed for control or mitigation of a given operating scenario. The circumstances

when this anomaly could further complicate a transient

are, of course, plant specific. A generic scenario would

feature a normally open valve that is closed on a logic

command, such as containment isolation, and then required to reopen to mitigate a transient. The valve fails

to torque out on closing, and the motor heats up, opening the thermal switch. The motor is then not available

until the thermal switch resets. Even then, the motor

will operate less efficiently until the overheated motor

windings have a chance to cool. If the thermal switch

fails to open, or if it has been bypassed, the motor could

burn out under stall conditions, making the valve inoperable for subsequent safety functions.

This appendix documents the acquisition, refurbishment, bench testing, installation, and in situ testing of

the valve and describes the investigation, results, and

conclusions of the INEL effort to determine the cause

of the motor operator's reduced performance.

DESCRIPTION OF THE VALVE ASSEMBLY

Shippingport provided a data package for the 8-in.

valve. This data package included a photograph of the

Shippingport valve installation, name plate information from the valve and actuator, a copy of the equipment specification, information from Crane Catalog

No. 53, and a Limitorque operation and maintenance

manual. The package also t e mtence

history of the valve. This maintenance history consisted

of summary sheets with one-line entry items identifying the maintenance performed. The history indicated

A-l
that the valve was overhauled in October 1963 and

received various maintenance actions from 1971

through 1984. Details and results of each maintenance

item or inspection were not available. INEL performed

a search of the Nuclear Power Experience Database

for records of the Shippingport facility that may have

recorded problems with the valve. No incident reports

for the valve were found.

Valve

The valve is an 8-in., 900-pound-class, pressure-seal

bonnet, cast steel (ASTM, A217 WC 1), flex-wedge

gate valve with butt weld ends (see Figure A-I). The

valve seats are hard faced with Steilite and are seal

welded to the valve body. The one-piece flexible wedge

(disc) is also hard faced with Stellite on the seating

faces. The valve was manufactured in 1956 by the

Crane Co. The following information was contained

on the valve identification plate affixed to the valve:

interfacing with a reversing controller and position

indication lamps. The motor is compound wound for

high-torque, low-starting-current operation, and

operates on 125 Vdc. The actuator is of weatherproof

construction using Class B insulation, and is capable

of opening the valve in a maximum travel time of

15 seconds under no-load conditions.

The actuator was manufactured in 1956 by Philadelphia Gear Works (now Limitorque Corp.). The

following information was contained on the actuator

identification plate:

Type

Size

Order No.

Serial No.

Valve

An-2917

468817

SMA

2

67910

57400

8

Catalog No.

Pressure

Temperature

Part No.

Size

Body

Stem

Disc

Seat

783 U4

900 psi

8500F

DA B8S713B

8

WC 1 Steel

CR 13

HF

HF

In addition, the motor identification plate contained

the following data:

hp

kW

750 rpm

COMP wound

5 min duty

125 volts

2 3 amp

75*C

Type and Frame D 2026

Serial No. HG 50272

40 ft-lb

Motor Controller

In addition, the following information was either cast

or engraved on the valve body:

8 Crane

WC I Steel

900

NL 326

62987

Actuator

The actuator is a dc motor-driven geared mechanism

that controls the opening and closing travel of the valve

and limits the torque loads applied to the valve stem.

The single torque switch stops the motor and valve

travel when a predetermined motor operator torque

load has been reached. A clutch mechanism and handwheel can be used to override the electrical operating

mode and allow manual operation.

The open limit switch stops the motor and valve

travel in the opening mode only. Both the open and

close limit switches provide additional contacts for

The motor controller is a Westinghouse LifeLinestarter Type N, NEMA V reversing. The controller has components for starting and reversing, and

has overload protection of the actuator motor. A

3-position, remote, panel-mounted selector switch is

used with the controller to select the operating mode

(Open-Auto-Close) of the valve assembly.

The controller was manufactured in 1956 by Westinghouse Electric Corporation. The following information was taken from a tag inside the controller:

Class:

Style:

Volts:

Size:

15-831-NW2

16-E-9874

125

2

A-2
Figure A-l. Eight-inch gate valve and actuator in "as received" condition.

A-3
RECEIVING INSPECTION

The valve was supplied from Shippingport with

8-in., Schedule-80 pipe stubs (safe ends) butt-welded

to the valve body (see Figure A-1). One pipe stub

measured 3.25-in. long; the other measured a minimum

length of 1 in. A weld backing ring was found tackwelded to the inside of the 3.25-in.-long pipe-stub-tobody weld. The backing ring was removed by grinding

out the tack welds. The 1-in. pipe stub was removed

because it was too short to weld a flange to.

The valve was manually opened using the actuator

handwheel. The valve was observed to operate smoothly. However, initial inspection of the valve assembly

indicated the valve actuator handwheel handle was

broken off. No other damage was observed on the outside of the valve.

Initial visual inspection of the valve body internal

condition revealed a black-gray oxide coating with

some scaling present on the fluid surfaces. Foreign

material was also observed on the wedge guide rails

when the valve was open and viewed through the end.

Inspection of the valve seats revealed a flashing or

steam-cut type defect in one of the seat sealing surfaces. The defect appeared as an indentation approximately 0.03 in. deep and 0.25 in. in diameter.

The compartment cover of the valve limit switch was

opened and the components shown in Figure A-2 were

visually inspected. No sign of moisture, corrosion,

wear, or damage was observed. The limit switch and

torque switch shaft penetration seals were also inspected. No sign of grease weepage into the limit

switch compartment from the actuator housing was

observed, indicating the shaft seals were satisfactory.

The torque switch setting, "as received," was 2. Initial discussion with the Limitorque Corp., Lynchburg,

Virginia, determined that a set point of 2 corresponds

to a stem thrust load of approximately 12,000 lb. (After

dynamometer testing at Lynchburg, however, old

documents were found indicating that a torque switch

setting of 3 was required to achieve the 12,000-lb

thrust.)

The running torque required for handwheel rotation

of the valve was measured using a torque wrench. Both

the open and closed valve travel directions required

70-in.-lb of torque on the handwheel. The valve was

observed to operate smoothly. The valve stem stroke

length was measured at 7-9/16 in.

The actuator was removed from the valve bonnet

yoke, and the actuator lubricant was removed for inspection and replacement. Since the actuator uses a

grease lubricant, it was necessary to partially disassemble the actuator. The housing cover and motor were

removed to allow access for grease removal. Approximately 18 lb of grease was removed, with some

residual grease remaining. The lubricant was visually

examined for wear products, metal particles, and

foreign material. No contamination from foreign

material or metal particles was found in the lubricant.

The amount of lubricant recommended by the Limitorque operation and maintenance manual ranged from

13 to 25 lb for the type SMA-2 actuator. Approximately 12 lb of Chevron Heavy Duty Grease-i, equivalent

to the Limitorque recommended grease Exxon Nebula

EP-1, was placed in the actuator housing. A new grease

seal was also installed in the housing cover.

While the actuator motor was removed, the motor

shaft drive gear was turned by hand and was observed

to rotate freely. No lateral or axial play in the motor

shaft was observed, indicating that the motor bearings

were in satisfactory condition.

The motor controller and 3-position panel switch for

the valve were shipped separately from Shippingport.

The cover on the motor controller unit was removed,

and the internal components shown in Figure A-3 were

inspected. No evidence of moisture, corrosion, or

damage was found inside the controller. One mounting screw from a terminal block or contact set was

found loose in the housing and was replaced. The contact sets were operated by hand and found to operate

smoothly. All circuits were resistance-checked and

found satisfactory. The insulation resistance was tested

at 500 V and found satisfactory. The 3-position switch

was also inspected and found to be in good working

condition.

At the conclusion of the visual inspection, the valve

was subjected to nondestructive examination, including

ultrasonic, magnetic particle, and X-ray inspection, as

part of the USNRC Nuclear Plant Aging Research

(NPAR) Program. NPAR objectives were (a) to identify and characterize aging and service wear effects

associated with electrical and mechanical components,

interfaces, and systems likely to impair plant safety,

(b) to identify and recommend methods of inspection,

surveillance, and monitoring of electrical and mechanical components and systems that will be effective in

detecting significant aging effects prior to loss of safety

function so that timely maintenance and repair or

replacement can be implemented, (c) to identify and

recommend acceptable maintenance practices that can

be undertaken to mitigate the effects of aging and to

diminish the rate and extent of degradation caused by

aging and service wear. After the nondestructive examination was completed, repairs were made to the

valve sealing surface, and a new safe end was welded

on to replace the one cut too short during the valve's

A4
Figure A-2. Limitorque actuator limit switch compartment.

A-5
Figure A-3. Limitorque actuafor motor controller.

A-6
removal from Shippingport. Two new 1500-lb flanges

were welded to the safe ends. Selection of materials,

welding, and hydrostatic tests were all done to the intent of the current issue of the ASME Boiler and Pressure Vessel Code, Section IIl, Class 2. The valve was

reassembled and remated with the Limitorque actuator.

All electrical connectors, terminals, and fasteners

were checked to make sure they were tight. All electrical contact sets were cleaned by nonabrasive means.

No other refurbishment was required for the electrical

components.

ANSI Standard B16.41, which presents the guidelines for the qualification of nuclear plant valve assemblies, was used in defining test requirements for

applicable tests performed at the INEL prior to installation at HDR, and for tests performed at HDR prior

to dynamic loading. German DIN (Deutsche Institut

fUr Norming) standards required the valve pressure

boundary to be certified prior to installation. The

ASME code and ANSI B16.41 tests were substituted

acceptably for DIN standard testing. ANSI B16.41 tests

that were performed at the INEL before installation

at HDR included Annex A, Valve Leakage Test;

Annex B, Cold Cyclic Test; and Annex E, Exploratory

Vibration Test for determination of fundamental frequencies. Annex G, Flow Interruption Capability Test

(the subject of another research program at INEL) was

performed at the HDR at both ambient and elevated

temperature within the limits of the VKL piping

system. A modified version of Annex C, Hot Cyclic

Test was performed at the HDR prior to elevated

temperature testing. Annex D, Pipe Reaction End

Loading Qualification Test, was not performed; this

loading is being researched in another program at

INEL. Annex F, Seismic Loading, was not performed

for two reasons: the test could possibly damage the

valve, and the HDR test program was a more realistic

simulation of seismic testing. Figure A4 shows the

valve installed in the pressure test fixture at the INEL.

The valve successfully completed limited qualification

testing and was judged suitable for installation in the

VKL at the HDR.

INSTALLATION AT HDR

A section of pipe was cut out of the VKL and replaced with a flanged spool piece consisting of the refurbished valve, safe ends, and flanges. Fifteen-hundredpound-rated stainless steel flanges were also welded

to the pipe ends. The valve spool was installed with

spiral-wound Flexitallic gaskets. This modification to

the piping was completed by changing out an adjacent

spring hanger to account for the added weight and by

performing a hydrostatic test of the piping system. The

HDR facility provided 125 Vdc to which the Limitorque operator and motor controller were connected. The

original motor controller from Shippingport was used.

Figure A-S details the valve instrumentation. Figure A-6 shows the installed valve and a portion of its

instrumentation. Important parameters recorded by the

instruments include valve stroke time, motor current

and voltage, system pressure, valve differential pressure, valve stem position and strain, and system fluid

temperature. The valve was installed operator down

to avoid structural interference above the pipe run. This

type of installation, while not the most preferred, is

second choice and typifies many industry applications.

SHAG TESTING AT HDR

Valve tests conducted at the HDR included functional

tests (without dynamic loading) at ambient and at elevated temperatures. Dynamic tests also were conducted

both at ambient and at elevated temperatures, usually

with the valve cycled during the dynamic excitation.

The valve was baseline functional tested with and

without internal hydraulic loading (pressure and flow

loads) before the SHAG dynamic tests were conducted.

The same baseline functional tests were performed at

elevated temperatures (2000C) just before the hot

dynamic tests.

The first four SHAG dynamic tests were performed

to check out instruments, equipment, and data acquisition systems and were conducted at less than maximum

dynamic loadings so that building and shaker safety

studies could be verified' Many instrument mounting

and equipment changes were made during these checkout tests to optimize recording of peak responses;

however, the valve instruments and data acquisition

systems worked well and produced a complete set of

data for the valve during the first four tests. Table A-I

relates the valve operational parameters to the dynamic

A-7
a zero,, ' . Ap- -_Xio I 1;4 I I - 0 1 - t

xo

U

qlca -- 07i

.  Vs p. i w. 14, ", ,'I Ift I;

I 11 I">11 '',,

I - - , Fu*Vvi al

Figure A-4. Valve installed in the pressure test fixture.
x

Valve

0 '_ Strain

l <110 guage U~~~tpe

Strain guage

Legend

- Accelerometer and its detection axis

Single strain guage and Its detection axis

.. * Strain guages (see strain guage detail)

m Displacement guage

m Pressure

& P

E Valve position

{J Valve current

Valve voltage

II Temperature

Figure A-S. Valve instrumentation detail.

A-9
Figure A-6. Installed gate valve and part of the installed

instnrunentation.

A-10
Table A-I. Valve function parameters'

.~~~~~~~~~~~~~~Pesr

Current in Amps -,

Stroke Time

(a) Test Temp

T40 (°C) Closing

Aj~b 20

34 20

35 20

36 20

37x 20

Y -

40x 20

Y. -

20x 20

Y -

60x 20

Y -

SOx 20

y _

70x 20

Y -

lOx 20

Y -

30x 20

Y 20

31x 20

y

41 20

21 20

11 20

51x 20

Y -

HFcx 210

Y -

52x 210

Y -

32x 210

Y -

42x 210

Y -

12 210

22 50

12A 210

14 40

16 Valve te

13 Valves t

AFdx 40

Y -

14.4

15.5

17

15.5

22

22

20

19.5

20

Opening

14.4

14.7

1S

15.5

15.5

15.6

17

16

15

16

16

16

16

16

16

16

14

17

12

12

12.5

12.5

12.5

12.5

12.5

12.5

13.5

12.5

14

Closing

Run Peak

12 12+

12 14

12.5 15

11 13.5

12 13.5

11.5

12 17

12 -

13.5 18

12.5 16

14 19

13.5 -

12.5 14

11 -

11 13

11.5 -

12.5 -

12.5 -

12.5 -

12.5 13.5

12 13

12.5 13

12.5 -

12.5 -

17 27

17 27

12.5 25

12.5 25

12.5 25

12.5 26

12.5 25

13.5 25

13 25

13 25

10 17

Opening

Stall Peak Run

- 16.41 10.94

- 17.67 10.6

- 17 11

- 14 9.5

- 17 11

35 13 10

- 17 10

38 20 9

- 24 12

38 25 12.5

- 23 11

_ 20 9

38 20 9

- 19 7

37 18 7.5

- 24 8

38 25 10

- 25 10

38 25 10

- 25 11

38 25 11.5

- 25 11.5

- 25 10.5

- 25 11.5

38 25 11

37 25.5 11

40 12 7

38 12 7

37 10 6

36 10 7

37 10 5

36 11 6

36 10.5 5

37 10.5 6

38 11 6

- 18 7

37 10 5.5

- 17 8

Pressure

(psi)

System AP

73.95 384

73.95 377

15.0 333

55.5 330

67 333

67 259 - 371

74 344 - 351

74 340 - 344

67 333 - 263

67 333 - 263

70 362 - 329

70 362 - 340

70 348

- 333

70 333

70 325

70 329

70 322 - 333

924 313 - 313

924 301

- 301

924 301 - 301

924 305 - 305

924 305

70 359

924 313

50 350

74 359

354

sts were not performed during this dynamic test.

ests were not peformed during this dynamic test.

14.5 14 13 19 - 20

15 14 13 19 41 22

12

11

a. The first and second cycles of two-cycle tests are indicated by x and y.

b. Functional test performed at ambient temperature at the beginning of the test sequence.

c. Functional test performed at elevated temperature.

d. Functional test performed at ambient temperature after the elevated-temperature tests were completed.

A-ll
test series nomenclature, and includes the important

parameters recorded for the valve during the test program. The valve performed flawlessly in the first three

tests (identified as T40.34, T40.35, and T40.36).

Figure A-7 shows valve position and operating current histories for the first two dynamic tests and for

a baseline functional test (with flow). There are only

slight variations in the overlaid current traces; for all

practical purposes, the valve functioned during the first

two transients as it did in the baseline tests.

Torque Out Failure

Normally, the high torque produced by the motor

when the valve reaches the fully closed position causes

the closing torque switch in the motor operator to open

and interrupt current to the motor. If the closing torque

switch fails to open, the motor stalls and overheats.

After the overheating has tripped the thermal overloads, the valve is unavailable until the motor has had

a chance to cool. If the thermal overloads are rated

too high, or if they have been bypassed, the result can

be motor burnout and the functional loss of the valve.

Table A-I shows that the valve performed without

problem also during the third dynamic test. However,

during the fourth dynamic test (test T40.37), the torque

switch failed to open at the end of the second closing

cycle, and the motor went into a stall. Figure A-8

shows the valve position history beginning 4 s before

initiation of the dynamic transient and during the first

56 s of the transient. At the start of the transient, the

valve was set 35% open so the maximum hydraulic

loading (valve fully closed) would occur near the peak

of the dynamic loading. The stroke time from fully

open to fully closed would have been 15 s; the stroke

time from 35% open to fully closed was approximately 6 s. The valve was left closed for 10 s, reopened

to 68% open, reclosed for 6 s, then reopened.

Figure A-9 shows the valve motor current history

for this same open-close sequence. As shown in

Figure A-9, the motor current for the valve opening

and closing cycles were normal until the second valve

closing cycle, where the operator failed to torque out

on closing and the motor went into a stall, with the

current rising to 43 amps and holding until the motor

was manually reset for reopening. VKL system pressure control requirements mandated that the valve not

be left closed longer than 10 s, so manual setting of

the motor controls for reopening was required for each

valve cycle. The resetting of the motor cancelled the

closing signal, and this prevented early detection of

the failure to torque out on closing and kept the motor

from overheating.

This pattern of the valve failing to torque out on the

second cycle repeated itself on most of the subsequent

ambient temperature tests where the valve was cycled

twice, until the last ambient test prior to elevated

45

co

E

4-

C

4)

II40

35

30

25

20

15

10

5

0

-5

0 10 20 30

Time (s)

40

UGHO 1625

Figure A-7. Composite time histories for valve position and current.

A-12
9

7

a

0

Q._

0

cc

5

3

6- 9.

-1 - _

69.92 81.92 93.92 105.92 117.92

Time

Figure A-8. Valve position versus time during the fourth dynamic test.

129.92

MHO013S6

C.

E

CZ

-

c

ID

-5 --

69.92 81.92 93.92 105.92 117.92

Time

Figure A-9. Valve current versus time during the fourth dynamic test.

A-13

129.92

MCF01354
temperatures testing, when the valve failed to torque

out on both closing cycles in test T 40.51. The valve

failed to torque out on closing in all of the elevated

temperature tests, including the functional testing

without dynamic loading and on both single and multiple cycle tests with dynamic loading. In the subsequent

ambient temperature testing, the valve returned to its

behavior prior to the elevated temperature testing, failing to torque out on the second cycle of a two-cycle

test. This pattern is shown in Table A-I in the "Stall"

column.

Figures A-10 and A-Il show the valve position and

valve current histories for two-cycle hot test T40.52;

Figures A-12 and A-13 show corresponding histories

for a single-cycle hot test. Both sets of figures are

typical examples showing the valve's failure to torque

out on closing during hot testing.

.2 9, . .

7 -

0

>5-

@ 89.93 101.93 113.93 125.93 137.93 149.93

Time

Figure A-10. Valve position versus time during a two-cycle,

elevated temperature valve test.

Basic valve function, i.e., the ability to open and

close under normal hydraulic loadings, was affected

by the magnitude of the loading. As expected, higher

system pressure, imposed during elevated temperature

testing, and higher WP across the gate affected stroke

times and motor current demands. The range of readings is reasonably small. The measured data indicate

that the dynamic loading (from the eccentric mass

shaker), in combination with the normal loadings, did

not have an adverse impact on valve performance. In

fact, the valve closing current was lower during the

elevated temperature dynamic testing than during the

elevated temperature functional tests (which were run

without dynamic input). The vibration during dynamic

testing may have helped break friction, thus causing

the closing running currents to be lower in the dynamic

tests than in the functional tests.

9 , . .

7

.Z3 5 0

94.99 106.99 118.99 130.99 142.99 154.99

Time

Figure A-12. Valve position versus time during a singlecycle, elevated temperature valve test.

45

4-

C

0

35

25

'lII X

15

5

45

35

:S 25

9 15

5

-5

94.

l l

I I

-5

89.93 101.93 113.93 125.93 137.93149.93

Time

9-0648

Figure A-I1. Valve current versus time during a two-cycle,

elevated temperature valve test.

99 106.99 118.99 130.99 142.99 154.99

Time 9-0841

Figure A-13. Valve current versus time during a singlecycle, elevated temperature valve test.

A-14
Failure of the valve motor operator to actuate the

closing torque switch, which led to subsequent motor

stall, appeared to be a motor operator problem rather

than a valve problem. This anomaly was not detected

during the test portion of the program, because of the

format of the quick-look plots and the necessity (mentioned earlier) to manually reset the control to reopen

the valve to accommodate system pressure control

requirements. The short length of time that the valve

was left closed prevented the motor from overheating

during stall current, making it impossible to determine

if the thermal overloads would have tripped before permanent damage occurred to the motor. Since thermal

overload trip was not an objective of this task, no

testing was performed to address this capability.

Failure of the valve to torque out on closing was

detected during posttest data reduction.

Posttest Investigation

A posttest investigation was initiated to determine

the probable cause of the valve motor operator anomaly. Limitorque was contacted, and probable causes

were discussed. Limitorque suggested at that time that

the most likely cause was that the torque switch was

set too high. A brief parametric study of the valve's

operability was conducted at the HDR facility to determine torque switch setting limits. That study showed

that the torque switch was not set too high. in those

tests, conducted without flow or pressure loads, a

torque switch setting of 4 was found to be too high,

but a setting of 3.8 was low enough to allow the motor

to trip the switch. A setting of 2.9 had been used

throughout the SHAG program. All torque switch locking features were found to be secure.

Further discussions were held with Limitorque and

other motor operator experts, and no positive scenarios

were developed that could explain the anomalous

performance of the valve. According to the voltage

readings, the motor voltage even at stall currents did

not fall below 107 Vdc. The motor should have performed much better at these voltages. In a subsequent

analysis of the power circuit, we demonstrated that

those voltage readings were unreliable, and that with

conventional methods and instruments, it was impossible to take true measurements of motor voltage on

that circuit. In the meantime, our investigation focused

on possible causes other than inadequacies in the power

supply or the power circuit.

First Review Group Meeting

A review group was established consisting of

representatives from the NRC staff, INEL, Oak Ridge

National Laboratory (ORNL), and Limitorque, and a

meeting was called to assist in identifying the root cause

of the valve functional anomoly and to plan a course

of action for its timely resolution. The review group

was briefed with all the information available at that

time, including a motor power curve for a motor

similar to the one installed at HDR. As shown in

Figure A-14, the curve indicated that the locked rotor

current for the motor is approximately 120 amps.

The probable causes of the valve's failure and contributing factors were discussed. The following list

contains items that were identified as probable causes

either singularly or in combination:

1. Torque switch malfunction

a. Damaged parts in torque switch assembly

2. Insufficient motor torque to compress the

torque spring pack in combination with the

valve loads

a. Internal motor operator friction

b. Excessive motor power circuit voltage

drop

c. Inadequate electrical power supply

d. Torque switch set too high

e. Hydraulic lockup in torque spring package

f. Internal motor operator damage

g. Motor operator motor, rotor, stator, or

brush degradation.

The meeting provided a panel of knowledgeable and

informed people who suggested ideas and provided pro

and con arguments for each of the possible causes and

contributing causes. The group requested additional

information to determine the root cause. This information, subsequently obtained from the HDR, did not

substantially change the measured data obtained during the SHAG testing or provide an obvious answer

to the motor operator anomaly. Therefore, the group

recommendation for further testing in the HDR at

SHAG hydraulic conditions was pursued. The intent

of this additional testing was to duplicate the motor

operator anomaly and then to perform tests and inspections to determine the root cause. Limitorque agreed

to support this activity from its European office at no

charge to the NRC.

A-IS
150

0

C

i

-100

- 80

- 60

- 40

- 20

-0

E

0.

I125

100

0

E

50

25

0

0 10 20 30 40 50 60

Load (lb-ft)

mMG000o9

Figure A-14. Motor power curve provided by Linitorque.

ADDITIONAL IN SITU TESTING AT HDR

INEL personnel returned to the HDR test facility to

perform the recommended additional testing. The valve

was reinstrumented at the motor/limit switch junction

strip in the limit/torque switch compartment of the

motor operator, and instrument power supplies and

recorders were located within a few feet of the valve.

A parametric study was conducted with the objective of obtaining more information on the relationship

of the valve's performance to the following parameters:

Torque switch setting

Torque spring travel

Valve disc position

Motor current

Motor voltage

Valve load

This study showed that at the lower torque switch settings, the valve did not completely close at higher flows

and pressures, while at higher torque switch settings

the valve operator motor stalled. The valve closed farther with the higher settings, but the motor stalled

before the valve closed completely.

The Limitorque representative reviewed the results

of the parametric study and inspected the motor

operator. It was recommended that additional lubricant be added. This recommendation was followed, but

subsequent testing showed no improvement in valve

performance.

The motor was removed for disassembly and inspection. The visual appearance of the motor internals was

good, but the front motor bearing was rough. One field

coil showed a high resistance. After discussions with

Limitorque USA, it was decided to rewind the field

coils. The field coils were rewound and both motor

bearings replaced. The repaired motor was reinstalled

on the operator. (Later discussions with the motor

manufacturer identified the coil with the high resistance

as a shunt field coil, which is a speed control coil that

would not affect the performance of the motor under

a load.)

In the tests that followed, the valve still showed no

noticeable improvement in performance, and after

discussion with Limitorque USA and the NRC technical monitor, it was decided that further in situ testing

at HDR would be unproductive and that the motor

operator should be removed from the valve and returned to the USA. Table A-2 is a summary of the

in situ HDR valve testing.

Results and Analysis of

In Situ Tests at HDR

The results of these tests indicated that the valve

functional problem is hydraulic load sensitive; however, the valve appears to be more sensitive to flow

loading in combination with static loading than to

A-16

increased static pressure load alone. During qualification testing, the valve closed successfully with a full

design pressure static load of 2300 psig, which is

significantly higher than any static pressure load at

HDR. With a static pressure load alone, the load on

the valve is constant throughout the closing cycle, and

motor speed is not affected. When a flow load is applied along with the static load (as was sometimes the

case during subsequent testing), motor speed is affected

as the gate closes and as WP across the valve increases.

The slowing of the motor speed affects the total

momentum of the motor operator and its ability to open

the torque switch. This, of course, is true only in a

marginally powered valve and is dependent on the

torque switch setting.

Figures A-15 and A-16 show this valve disc velocity relationship for the valve closing stroke at three

pressures with and without flow. With pressure alone

(no flow), the stroke time increases with pressure

(Figure A-15). Note in Figure A-16 the decrease in

the disc velocity, indicated by the slope of the trace,

for all three pressure cases (with flow) as the disc approaches the closed position. In the cases with flow,

the torque switch failed to open. The extent to which

the valve remains open is increasingly greater with each

increase in pressure combined with the flow load.

Figure A-17 shows a 147 psig static pressure case with

flow, along with the three pressure cases shown in

Figure A-16. The disc velocity near closing for the

147 psig pressure case does not decrease as much as

it does for the higher pressure cases. I -n the

valve torque switch opened. The valve stem area is

2.77 square inches. The difference between the torque switch opening and failing to open is the 407 Ibf

reduction in the stem rejection load.

Review of the data obtained from the in situ HDR

tests (see Table A-2) shows that when the valve was

subjected to the higher flow and pressure loads in

combination (>80 t/hr flow and >30 bar), the motor

operator consistently failed to completely close the

valve, and at torque switch settings greater than

2.9 or 3.0, the operator also failed to open the torque

switch (3.0 was the setting recommended by Limitorque). Normal in-plant valve testing would not detect

this type of marginal valve performance. Typically,

testing is performed with no load or with a static load

alone, and assessment of performance is based on

changes in valve current and stroke time. This type

of marginal performance would not even be detected

by MOVATS or similar motor operator performance

equipment that monitors operation of a valve with a

static load alone.

Second Review Group Meeting

Another review group meeting was held to discuss

the results of the in situ tests and to plan a course of

8

'I,_

c 6

%-.

c

0

._.

G 4

0

n Z

0

0 5 10 15 20 25

Time (sec) HE0020t

Figure A-15. Valve position during closing stroke under three different static pressure loads with no flow.

A-18
8

(-'r

0

6

4

2

0

0 5 10 15 20 25

Time (sec) Figue Valv stro poitin e wth drngcloing low (360gpm un er treediferet stticpresur l~oad.20o

Figure A-16. Valve position during closing stroke with flowv (360 gpm) under three different static pressure loads.

v -43 4 I3 IT-48)

C

*0

0 5 10 15 20 25

Time (lec)

Figure A-17. Valve position during closing stroke with flow

(360 gpmn) and with a fourth static pressure

load.

action for the timely resolution of the anomaly. All of

the necessary background information was reviewed,

and new information from the in situ tests was discussed

with the review group. Additional questions and comments concerning particular items were discussed; the

highlights of that discussion are given in the following list.

1. Concern was expressed over the increase in

the technical specification limits.

2. For many valve assemblies in safety systems,

the thermal overload protective devices are

either set too high or bypassed to prevent

motor cut out while performing its safety function. If the operator stalls, motor burnout is

likely to occur, making the valve unavailable

for subsequent operation.

3. Testing of valve assemblies at less than design

loads may not be adequate.

The general consensus of the meeting was that the

motor operator anomaly was probably the result of a

marginally sized motor; the motor achieves successful

operation only when allowed to seat the valve at full

speed, with momentum compressing the torque spring

to the switch set point.

The review group recommended the following:

1. Perform a dynamometer check of the motor

operator to determine its torque/speed and

torque/current characteristics. Compare the

results to the HDR test data.

2. If motor is undersized, determine if it is the

result of a manufacturing problem or some

type of aging effect. Determine what in the

motor has degraded.

3. If the motor operator is not the problem, reevaluate the HDR power supply.

Limitorque agreed to dynamometer test the motor

operator assembly on its arrival from Germany.

closing times at higher loads. This change may

be significant for plant safety concerns since

demonstrations of stroke times are typically

performed at ambient conditions. The stroke

times under higher loads may not fall within

A-19
LIMITORQUE DYNAMOMETER TESTS

Dynamometer testing was performed at the Limitorque facility with representatives from INEL present

to witness the results. The Limitorque Laboratory

motor controller and amp meter were different enough

in performance from those used in earlier testing to

make comparison difficult. The motor controller

dropout time was longer, and the amp meter lock-in

feature did not consistently give an accurate reading

of the highest current drawn during torque out. On request from the INEL, the control circuit was rewired

to include the Shippingport motor controller and a standard shunt to measure current. The dynamometer

loading was applied slowly to the operator in an attempt to simulate the increasing hydraulic loading on

an in-service gate valve during the closing stroke.

Table A-3 shows the results of that testing.

During the first seven tests, chatter was observed

in the operator torque switch, and arcing was observed

across the torque switch contacts. There are redundant

contacts on this model torque switch, and the circuit

was reinstalled on the rear contacts for the remainder

of the testing.

Results and Analysis

The original design specification stem thrust requirement for the valve and operator at Shippingport was

12,000 lb. The Limitorque manual for this operator

specifies a maximum torque switch setting of 2. The

operator was set at 2 when originally received from

Shippingport. At the time of the dynamometer testing,

Limitorque had not found any records on this unit to

show that a torque switch setting over 2 was required

to meet the original design requirements. During functional testing at the INEL after valve refurbishment,

it was necessary to change the torque switch setting

to slightly less than 3 to obtain tight shutoff. For

25-year-old sealing surfaces this was not considered

unusual, and the opening and closing currents were

well within motor ratings. After the dynamometer

testing at the Limitorque facility was completed, old

records were found at Limitorque that show that a setting of 3 was necessary to obtain an output thrust of

12,000 lb. Figure A-18 shows the original design

torque spring compression versus torque output. The

1-3/4 in. dimension is the radius of the torque switch

arm, and the numbers 1 through 5 around that radius

indicate the torque switch settings. Figure A-19 shows

the type of torque switch used in this application. Item

number five in this drawing shows the torque switch

tripper arm, which has the 1-3/4-in. radius shown in

Figure A-18. Item 4 is the torque switch dial showing the settings from 1 to 5. The torque switch is adjusted by moving Item 14 in and out. Figure A-20 is

a simplified diagram of part of a motor operator. The

splined output shaft drives the worm, which turns the

worm gear and the stem nut. The stem nut drives the

threaded valve stem. As the valve seats and the worm

gear resists motion, the worm climbs the worm gear,

floating on the splined shaft and compressing the torque

spring until the shoulder of the worm contacts and

rotates the arm of the torque switch and opens the

torque switch. The higher the torque switch setting,

the further the worm must compress the spring before

the torque switch opens and interrupts current to the

motor.

Review of the dynamometer tests results (Table A-3)

shows that a setting of 3 increased the stem thrust to

only about 8,000 lb. The table also shows that the

operator required a torque switch setting of 3.75 to

develop an output thrust of 12,000 lb. The most likely reason for the higher torque switch setting required

to produce the rated torque is a change in the torque

spring constant or length. The spring must be compressed further to produce the same force. This type

of aging in coil springs is not unusual, but it does create

a problem. The problem would not be discovered in

normal in-plant valve testing where a change in stroke

time is the primary go or no go criterion and where

valves are typically tested without hydraulic loading.

Gate valves such as this one are typically never used

in normal service to control flow or to close against

flow and could go for years without such a problem

being discovered. If the valve were needed to mitigate

an accident, a flow load might make the motor operator

torque out early and leave the gate partially open.

Dynamometer testing results also show a marked improvement in motor operator motor performance over

the results obtained at the HDR. Stall currents as shown

in Table A-3 are between 66 and 75 amps, producing

30 to 32 ft-lb motor torque. This is below the performance specified by the motor manufacturer in

Figure A-14. The stall current should be 115 amps

with 62 ftlb torque. The stall currents at HDR were

near 50 amps with a maximum of 21 ft-lb torque.

Third Review Group Meeting

The results of the dynamometer testing at Limitorque

were presented to the review group and are summarized as follows:

1. It was determined that under ideal laboratory

conditions the motor operator performance

was better than at HDR.

A-20
Table A-3. Results of motor operator testing at Limitorque Laboratory

Torque

Test Switch

No. Setting

1

2

3

4

.5

6

6.1

7

8

9

10

11

12

13

14

15

16

17

18

19

20

21

22

23

24

25

26

27

28

29

30

31

32

33

34

35

36

2.00

2.00

2.50

2.50

2.90

2.90

2.90

2.90

3.10

3.10

3.10

3.50

3.50

3.50

3.50

3.75

3.75

3.75

4.00

4.00

4.00

4.20

4.50

4.50

4.50

4.50

4.75

4.75

3.50

3.50

3.50

3.50

3.50

3.50

3.50

3.50

Set

Voltage

(Vdc)

110

125

125

110

110

110

125

125

125

110

100

100

100

110

125

125

110

100

100

110

125

100

100

100

110

125

125

110

90

80

80

80

90

100

110

125

Minimum

Voltagea

(Vdc)

108

110

121

107

105

105

120

120

119

104

95

94

94

104

119

118

103

94

93

103

118

93

92

91

102

116

112

99

85

75

73

79

89

99

109

124

Plotter

Peak

Current

(amp)

9.0

15.0

15.0

21.0

21.0

25.0

25.7

26.0

29.5

34.0

35.0

34.0

41.0

39.0

40.0

38.0

39.0

42.0

42.0

46.0

46.0

46.0

50.0

58.0

56.0

54.0

51.0

75.0

66.0

39.0

40.0

41.0

Peak

Torque

(ft-lb)

99

88

125

132

154

185

169

176

220

229

231

242

275

242

264

277

264

245

286

286

264

308

352

363

330

308

484

407

231

255

220

Peak

Thrust

(lb)

4480

3982

5656

5973

6969

8371

7647

7964

9955

10362

10453

10951

12444

10951

11946

12534

11946

11086

12942

12942

11946

13937

15928

16426

14933

13937

21901

18417

10453

11539

9955

Comments

No load/baseline

Stall

Stall

No-load

No-load

No-load

No-load

No-load

5.5

5.3

6.0

6.0

6.0

a. Derived from visual readings.

A-21
Motor Operator

Torque OutPut

Spring comp. Torque In Inches Output

Switch from preload Torque

Position position ft-Ibs

(Preload:

1 equals .25) 67.75

1.5 .1 11-i

2 .32 155

2.5 .484 199

3 .648 244

Wire dia. 3/o.

Spring Characteristics

j|- Nominal Free Length = 451

Nominal Installed Length = 411t.

< Max. Compression = 314' + preload

III I s Working Compression = "i' + prepoad

*UG0044

Figure A-18. Torque switch setting versus spring compression and torque output (HDR valve).

A-22
Single torque switch

Parts Est PC PC P _- Parts __t ____ | Name No Name Na Name

PC Name 7 Spring 15 Cotter pin 23 Oil seal

No. 8 Roller 16 Washer 24 set screw

1 TS. Bracket 9 Roller pin 17 Miper pin 25 jam nut

2 Actuating arms 10 Sta. terminal block 18 Hex nuts 26 Bushing lock screw

3 Dial holder 11 Movable term block 19 Lockwashers

4 Dial 12 Term. studs 20 Capscrew

5 Tipper arm 13 Threaded bushing 21 Contacts

6 Actuating arm shaft 14 Sock HD.C'SCR (knurled) 22 Contactor _

74903

Fig=ue A-19. Liitorque torque switch cross section from the LMI-157 operation and maintenance manual.

A-23
Torque switch

Figure A-20. Limitorque SMA-2 motor operator cross section from the LMI-157 operation and maintenance manual.

A-24
2. It was determined that the dc motor performance was less than originally specified.

3. It was determined that the torque output versus

torque switch setting was lower than originally

specified.

The review group recommended removing the torque

spring and dc motor for further testing to complete the

investigation and to have a final meeting when the

results become available.

At the completion of the investigative dynamomnr testing at Limitorque, the dc motor was removed and a new ac motor was installed. The motor

operator was returned to the dynamometer and its

performance characterized. The results of these

dynamometer tests will be compared to in situ performance when the motor operator is remated with

the valve at the HDR. The dc motor was sent to the

motor manufacturer, Peerless Winsmith, for dynamometer testing, and the motor operator returned to

the INEL for installation of a new torque spring and

torque switch. The recommended spring testing would

be completed at the INEL during torque spring change

out.

SPRING TESTING

After its removal from the operator, the old torque

spring was measured for length and compared with a

new spring purchased from LUmitorque as a replacement. Both springs had the same wire size and number

of coils. The new spring measured 4.461 in., and the

old spring measured 3.981 to4 lf in. inlenadh. The

old sprin than the other.Ue

average length of the old spring (3.994 in.) was

.468 in. less than the length of the new spring. According to information obtained from Limitorque, the

nominal free length of a spring should be 4.312 in. The

shorter free length of the old spring explains the torque

switch position versus torque output obtained during

the Limitorque dynamometer testing, where a higher

torque switch setting was necessary to obtain the design

output.

Both torque springs were then load tested. Figure A-21 shows the results of those tests. The old

spring was approximately 1/2 in. shorter, so the curve

for the old spring does not start showing a loading for

the first 1/2 in. displacement. Because of a misunderstanding between Limitorque and the INEL about

it z allowable spring travel, the springs were taken to coil

touch during load testing. As can be seen In Figure A-21, at about one inch travel on the first loading

td.,4 pf the new spring, yielding took place and the spring

developed a permanent set of 22 hc n the second

loading cycle no sig ifitykiding h took place. The

old spring was also load cycled twice to coil touch.

Figure A-21 shows no yielding took place on either

cycle. Note that in Figure A-21 the second cycle for

each spring was moved slightly to the right of the first

cycle for clarity. It can also be seen from the figure

the spring constant for both springs (K-factor) was the

same (within the accuracy of the equipment) for all

loadings. The final length of the new spring after

testing was .1 in. shorter than the nominal length

(4.312 in.) of a new spring.

During the disassembly of the actuator to remove

the old spring, it was noted that there was no preload

on the old spring. The preload should have been

.25 in., as can be seen in Figure A-18. The information contained .in this figure was not obtained from

Limitorque until after the spring testing had been

completed. As shown In Figure A-18, the maximum

allowed compression was 0.75 in. plus 0.25 in.

preload. This explains why the new spring showed

significant yielding after one inch of travel during the

first test. It also shows why there was no preload on

the old spring. The correct installed length for the

spring is 4 1116 in., and the old spring was slightly

less than 4 in. long. This confirms the hypothesis

developed during the dynamometer testing at Limitorque that the torque spring was too short, requiring

a higher torque switch setting to achieve the torque output versus torque switch setting shown on Figure A-18.

Another spring, similar to the spring in the HDR

Shippingport valve assembly, was obtained from Oak

Ridge National Laboratory (ORNL). This spring was

removed from Shippingport valve Number 54475D

with a Limitorque motor operator type SMA-1 S/N

08531A. The free length of this spring was 3.745 in.

Figure A-22 shows the relationship of torque switch

setting versus spring compression for the ORNL valve,

as Figure A-18 does for the INEJLHDR Shippingport

valve. The information obtained from Limitorque used

to draw Figure A-22 specified two springs with differcnt free lengths: 3-7/8 in. and 3-13/16 in., as shown.

Limitorque did not know from the old records which

of the two springs was installed in this valve. This

spring is short by either slightly more than 3/16 in.

or slightly more than 1/16 in., depending on which

11

To A-25
2,000

1,500

0

, /// DX ExtCompension/

/ ./6tX ~~~~~Rate =0.625 in./min

0.0 0.5 1.0 1.5 2.0

Displacement (in.)

HEA00203

Figure A-21. Load testing of old and new torque springs.

A-26
* Motor Operator

Torque OutPut

Spring comp.

Torque In inches Output

Switch from preload Torque

Position position ft-lbs

(Preload:

1 equals .50) 415

I11, .167 555

191 8 .207 587

2 .335

21IS .502 ____

21214 .563 ___.

Wire dia. $/a

Spring Characteristics

Nominal Free Length= 37/ or 313/ee

Nominal Installed Length . 33J'

Max. Compression = Iiei + preload

1 Working Compression o I",1 + preload depending on date of manufacture

5eo00073

IiI'

CI

Figure A-22. Torque switch setting versus spring compression and torque output (ORNL valve).

ffi I * * ! ' 1 i-l * spring was used. Fige A-23 shows the load test for

tot-@ as," * this spring. The total allowable elastic travel for this

spring should be 1.188 or 1-3/16 in. plus or minus a

small margin for tolerances. As can be seen from the

load test, the spring went nonlinear at 1.01 in. The

difference between the allowable installed travel

(1.188 in.) and the load test travel (1.01 in.) provides

a good indication the spring was originally 3-7/8 in.

// W long and has nearly 3/16-in. permanent set. The ORNL

valve, like the Shippingport valve, was a normally

closed valve, which means the spring was normally

e *ts o *o of; 1eo 1 t5 in the compressed position for the life of service

Dssplcement (In) (>25 years). It is believed this near full time compression may accelerated spring aging. Figure AL-23. LaAd testin7 of ORNL toroue sorine.

A-27
MOTOR TESTING AT PEERLESS

The Limitorque dynamometer testing showed a

marked improvement in motor performance over that

obtained at HDR; however, it still did not meet the

motor performance curve (Figure A-14) provided

earlier by Limitorque for a motor similar to the one

on the SMA actuator. The motor manufacturer, Peerless Winsmith, agreed to perform dynamometer testing

on the motor alone to determine if the apparent reduced

performance of the motor actuator was in the motor.

Figure A-24 shows the results of the dynamometer

testing of the motor. At motor loads of less than

30 ft-lb, the actual performance follows the predicted

performance very well for both rpm versus load and

current versus load. Above 30 ft-lb, the measurements

of rpm versus load are fairly close to predicted values,

but the current falls off, indicating that the motor is

not meeting predicted load requirements. The higher

torque loadings resulted in significant motor heating,

with smoke coming from inside the motor. The motor

was not permanently damaged by overheating during

these high loadings, and time was allowed between

each test for the motor to cool. At the request of the

INEL, the dynamometer loadings were applied slowly

to simulate the closing of a valve with a 15 second

stroke time.

Another test was conducted to evaluate the effects

of motor heating. Before the test, the motor load was

adjusted to 45 ft-lb, and then the motor was cooled to

ambient temperature. The motor was brought to speed

without a load, and the preset load was then applied.

The motor speed slowed immediately and stabilized

within 2 seconds. Figure A-25 shows the results of

this test. The load was applied for 20 seconds. Internal motor heating caused the motor to drop in

current demand at 1 amp per second for 20 seconds.

In round numbers, for each 2-amp reduction in

demand, motor output is reduced by one ft-lb of torque.

The motor output performance was reduced by 10 ft-lb

in 20 seconds.

The engineers at Peerless supplied a motor performance curve for the subject motor (Figure A-26). Performance characteristics shown on this plot are similar

to those shown on the one Limitorque had supplied

earlier in the program (Figure A-14). Figure A-26 also

shows the peak torque obtained in each of the

parametric test programs.

RPM versus load _ .

* ran _,_ _ " .owu

3,000

2,500

2,000

5

C. 1,500 I1.000

500

0

. ,- I . I . I . a,,

a Actual performance

iK ~ ~~~~ -Predicted performance

ok .I.I

0 10 20 30 40 50 6o

Load (ft-lb) Load (ft-lb)

Cutrent versus load

I-1 50 . .. . . . .

ii6. , * ' ' ' 130 a Actual performance

la 12 Predicted performance

10

0 10 20 30 40 50 60

Load (ft-lb)

Figure A-24. Motor dynamometer testing at Peerless.

130

fij1200

*100 S 90 - so

C 70

50

30

2019

O Aalua t Styano

- Fredictd pu~rmanofn

unt + 20 se

u

0 10 20 30

Load (ft-lb)

40 50 60

MeA

Figure A-25. Effects of motor heating on motor output.

A-28
3.500

3,000

2,500

2,000

E

L.

1,500

1,000

600

0

150

125

100

0t

75 4°C

E

50

25

0

0 20 40 60

Load (lb-ft) 3M000042

Figure A-26. Motor performance curve obtained from Peerless Motors and peak torque values obtained during testing.

ANALYSIS OF THE HDR POWER CIRCUIT

Concurrent with Limitorques assistance to the INEL,

with the HDR investigation, a utility called upon

Limitorque to assist with similar dc motor-actuated

valve problems. The utility was a two unit station, and

all of the dc powered valve problems were associated

with one unit. Several control circuit problems were

found during the investigation, and the root cause of

the dc motor failures had been attributed to those

problems. Interestingly, the power cables in the unit

having problems were smaller than those in the other

unit. However, this difference was not highlighted in

the investigation.

Shortly afterwards, Limitorque published a maintenance letter that contained an electrical circuit

calculation basis for dc motor operated valves. This

calculation was developed from Limitorque's investigations of dc motor problems at HDR and at the utility

mentioned above. In most dc motor operated valve control circuits, the armature coil and the series fields are

interconnected through the motor controller, and four

power cables must be considered in the calculation,

not two as one might assume. Limitorque recommended using this calculation to determine voltage drop

instead of trying to measure the voltage drop in the

circuit; in attempts to measure the voltage drop, it is

very difficult to load the circuit and sum the various

voltage drops.

From this new methodology we developed an

analytical model for the HDR valve power circuit

to determine the influence of external circuit resistance on the HDR valve anomaly. The HDR in situ tests

and the Limitorque tests were chosen for comparison,

as the motor operator was operated a comparable

number of times over a short period of time at both

locations and the size of the large external circuit

cabling at Limitorque should eliminate the influence

of any external circuit resistance. Two different

calculations of the HDR power conductor resistance

were made: one was based on the size and approximate

length of the cables, and the other on measurements

of the current through the circuit and the voltage drop

across the portion of the circuit that was measured.

The calculations were comparable. Four times the

calculated resistance for a single cable was inserted

in the analytic model to account for the armature and

the field being connected in series through the motor

controller. The measured parameters from the HDR

in situ testing and the Limitorque dynamometer testing

were then analyzed and actual motor resistances

calculated. Very good comparisons were found. The

results demonstrate that the differences between the

motor operator performance at HDR and the performance at Limitorque were caused by external

circuit resistances.

A-29
These results point out the significance of the external circuit resistance and motor heatup in reducing the

safety margins of motor operator torque output. The

motor operator operated successfully with no pressure,

static pressure alone, and with static pressure and very

low differential pressure. When the differential pressure from the flow load was increased, the valve either

failed to close all the way or the operator motor stalled,

depending on the torque switch setting.

To understand this phenomenon, one must understand dc motors. These motors generate a back electromotive force (EMF) when turning. This back EMF acts

like a bucking voltage in the circuit. Stated in very

simple terms, this back EMF limits the current in the

circuit much the way a resistance would. The back

EMP is proportional to the motor speed and current.

These high-torque compound-wound motors have a

weak shunt field and act very much like a series wound

motor in application. As the load is increased the motor

slows down, reducing the back EMF. This allows more

current to flow, thus producing higher torque. This

behavior continues down to motor stall, where the dc

resistance of the motor is the only internal resistance

to current flow. These motor resistances at or near

locked rotor are in the 1.0 to 1.5 ohm range for a 40 to

60 ft-lb output torque motor operating on 125 Vdc.

Because the field and armature cabling are connected

in series through the motor controller, four long cable

runs, not two, contribute to the resistance of the external circuit. External circuit resistance of just 0.5 to

1.0 ohm can reduce the motor output torque by 1/3

to 1/2.

During normal valve testing or operation, the resulting motor loadings are in the 20 to 30% running torque

range. The back EMF or effective motor resistance

would be approximately 5 ohms, and a 0.5 to 1.0 ohm

external resistance would not significantly degrade

motor performance. This conclusion was supported by

the results of the HDR static pressure testing. However, during a transient or a line break where isolation is required, the high differential pressure across

the valve disc would require a higher motor output,

and the external circuit resistance in series with the

motor resistance could reduce the output torque of the

motor significantly, as it did at the HDR during tests

with flow loads.

At the end of our investigation of the valve's

anomalous performance at HDR, we had identified

three separate problems. The potential safety implications of these problems are significant. Spring aging

can result in a valve that torques out early and, depending on the extent of the degradation, can leave the valve

partially open. Aged coil springs in the older SMA

motor operators may go undetected, as there are fewer

diagnostic test systems adaptable to these units. Motor

heating can reduce motor operator output if valve

is cycled more than once without time for the windings to cool. If a marginally powered valve is subjected to high loads on closing, this reduced output

can result in motor stall, probably with the valve

partially open. External circuit resistance also can

cause the motor to stall before the valve fully closes.

Motor stall can cause the thermal overload switches

to open and render the motor operator temporarily

unavailable for use. If the thermal overload switches

have been bypassed or set too high, or if they malfunction, the motor will burn out. Both the SMA and

the newer SMB operators with dc motors and high

external circuit resistance can go undetected in normal

in situ valve testing.

The HDR anomaly would not have been discovered

in normal plant testing, where valves are tested without

hydraulic loading and usually stroke time change is the

go or no go criterion. The problems are detectable only

at higher loadings when the motor is slowed down and

momentum cannot carry the unit through torque out.

Flow loads in combination with static pressure stem

rejection loads present a more demanding load on

motor operators than do stem rejection loads alone. The

data and the design of a Limitorque operator tend to

cast some doubt on the validity of using opening loads

to predict closing loads. When the valve opens, the dog

that creates the hammer blow allows the motor to come

to speed prior to engaging the stem drive gear; the hammer blow then assists in breaking friction, and the stem

rejection load then assists in opening the valve. Adding two times the stem rejection load to this load in

calculations will not account for the motor slowdown

and the decreased efficiency of the unit. Motor speed

and total actuator momentum appear to be important

factors in predicting valve closure. The concerns expressed here for dc motors are not limited to the SMA

design. The newer SMB have the same motor design

as the SMA.

A-30
SUMMARY OF RESULTS AND ANALYSIS

The initial failure of the valve at the HDR to close

completely, even at stall currents, when exposed to the

higher hydraulic loadings was caused by the combined

influences of undersized power cables, motor heating,

loss of motor momentum due to motor slowdown, and

loss of gear efficiency due to actuator slowdown near

the fully closed position.

A good example of undersized power cables and

motor heating can be seen from tests 42 and 42A (see

Table A-2). Test 42 was the first test run with pressure

and flow loads after motor field rewind, and the motor

was at ambient temperature. The torque out current was 42.5 amps, and the valve closed to within

0.2 inches of being fully closed. The 42.5 amps is less

than it should have been and is a result of the voltage

drop in the undersized cables. During the very next

cycle, test 42A, the torque out current was 29 amps,

and the valve closed to within 0.3 inches of being fully

closed. The 25 amp closing shows the combined affect

of undersized cables and motor heating. Test 42 was

the only test where an ambient-temperature motor was

known to be subjected to a maximum flow and pressure

load. The degradation shown in test 42A is significant.

Two full cycles on the valve are equivalent to one

minute of motor operating time with a 15 second

average stroke time.

The motor test at Peerless shows the motor performs

very close to design, except at the higher loads, where

heating starts to reduce performance. The peak performance differences between the Li.mitorque dynamometer tests and the Peerless motor dynamometer tests

are due to motor heating, and they are fairly significant: 75 amps versus 105 amps at motor stall.

Because of motor slow down under load, there was

no momentum to assist in compressing the torque

spring. This total unit loss of efficiency could be very

significant in marginally powered units in the field. The

HDR static pressure loadings during the elevated

temperature test were equal to the original valve design

pressures, but static pressure loadings during all other

tests were much lower than design. The maximum

VKL flow of 360 gpm is a very small flow foran 8-in.

valve. The valve performed poorly even when pressure

and flow loads were well below design loads (see

Figures A-15, A-16, and A-17).

The degraded torque output versus torque switch

position was caused by a short torque spring. Maximum torque switch settings during the HDR tests were

not high enough to cause the permanent set in the

torque spring. The initial dynamometer test at Limitorque isolated the problem. The highest torque switch

settings in subsequent dynamometer tests did enter into

the permanent set travel range; however, tests following the high torque switch tests do not show further

degradation.

It is unknown how many SMA operators there are

installed in nuclear plants. The SMA operator was

manufactured from the mid-1950s to the mid-1960s,

when it was replaced by the SMB design. Four plants

are known to have SMA operators in 33 applications:

Haddam Neck (16), Oyster Creek (9), Nine Mile Point

(2), and San Onofre (6).

Discussions at the review group meetings indicated

that in the 20 plus years since the SMA design has

been manufactured, considerable experience and information acquired in the field and at Limitorque

on this design has been lost. A large part of the work

in trying to determine the root cause of the valve

functional anomaly was spent by the INEL and

Limitorque learning the design characteristics of

these operators. Additionally, the dc motor is not

as common as the ac motor, and less is understood

about its performance characteristics under extreme

loadings.

Other work is underway that will address some of

the issues of this report. The current work and research

being performed for El Bulletin 85-03 and Generic

Issue 87 are programs designed to determine valve

closure requirements under extreme loads. In the

follow-on HDR testing, the dc motor on the HDR SMA

motor operator was changed out to an ac motor of the

same horsepower rating. Those tests will allow us to

compare the performance of ac- and dc-powered motor

operators.

A-31
CONCLUSIONS AND RECOMMENDATIONS

Failure of the HDR valve to close under certain

hydraulic and pressure loadings was due to a combination of problems, including a) undersized power

cables, b) motor heating, c) torque spring aging, and

d) loss of motor operator momentum at load.

One can analyze power cables for possible undersizing using a procedure recommended by Limitorque

that considers the length of the four cables and the peak

power of the motor. The effects of motor heating can

be estimated. Those effects should be included in the

motor operator motor sizing calculation.

Torque spring aging can be detected on some SMA

and SMB type motor operators by a simple torque

wrench handwheel test. Others will require that the

springs be removed, measured, and compared to the

length specified in the original drawings.

Motor operator momentum should not be a problem

in a correctly sized and adequately powered motor

operator. However, it is something that must be considered in testing and analysis of dc-powered motor

operators.

The HDR anomaly would not have been discovered

in normal plant testing, where valves are tested without

hydraulic loading and usually a stroke time change is

the go or no go criterion. Flow loads in combination

with static pressure stem rejection loads appear to present a more demanding load on motor operators than

do stem rejection loads alone. The data and the design

of a Limitorque operator tend to cast some doubt on

the validity of using opening loads to predict closing

loads. When the valve opens, the dog that creates the

hammer blow allows the motor to come to speed prior

to engaging the stem drive gear; the hammer blow then

assists in breaking friction, and the stem rejection load

then assists in opening the valve. Adding two times

the stem rejection load to this load in calculations will

not account for the motor slowdown and the decreased

efficiency of the unit. Motor speed and total actuator

momentum appear to be important factors in predicting valve closure. The concerns expressed here for dc

motors are not limited to the SMA design. The newer

SMB have the same motor design as the SMA.

A-32
APPENDIX B

INSTRUMENTATION OF THE VKL

FOR THE SHAG TEST SERIES
APPENDIX B

INSTRUMENTATION OF THE VKL

FOR THE SHAG TEST SERIES

Appendix B provides details with respect to the

VKL instrumentation for the SHAG test series.

Figure B-1 indicates instrument locations for the U.S.

stiff configuration; Table B-1 identifies the type and

number of instruments at each location. INEL and

KfK divided instrument responsibilities on the VKL.

Staatliche Materialpriifungsanstalt (MPA) of the

Universitat Stuttgart assisted INEL in installing strain

gages and instrumentation cables and recording test

data from a portion of the instruments. INEL installed the U.S. instruments and recorded test

data for a portion of the instrumentation. HDR installed

all KfK instrument cables and recorded all of their

data. KfK was responsible for providing a single

set of data tapes with a unified time base for all

participants.

B-I

Table B-1. Detailed U.S. and German Instrumentation Table B-I. Detailed U.S. and German Instrumentation

Location

Number

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

27

28

29

30

31

32

33

34

35

36

37

38

39

40

41

42

43

44

45

46

47

48

49

50

Instrument

Type

Strain

Strain

Strain

Strain

Displacement

Displacement

Acceleration

Acceleration

Acceleration

Acceleration

Acceleration

Was not installed

Strain

Strain

Acceleration

Acceleration

Acceleration

Acceleration

Acceleration

Acceleration

Acceleration

Acceleration

Strain

Strain

Strain

Strain

Strain

Strain

Strain

Strain

Strain

Strain

Strain

Strain

Acceleration

Acceleration

Strain

Strain

Strain

Acceleration

Acceleration

Acceleration

Valve position

Valve current

Valve voltage

Pressure

Differential

Temperature

Video camera

Displacement

German U.S. Remarks

6 _

-6 -

6 _

6 _

2(XZ) -

2(XZ) -

3(XYZ) -

3(XYZ) -

3(XYZ) -

3(XYZ) -

3(XYZ) -

owing to discontinuities _ 6 _ 1

_ 2(XZ)

_ 3(XYZ)

_ 1(X)

- 2(YZ)

_ 2(YZ)

- 2(YZ)

- 1C)

- 3(XYZ)

- 1(X) _ l(Z)

_ 1(X)

_ 1(Y)

- 1(Z - 1(X)

- 1(Z) - 1(Y)

3(XYZ) _

- 2(YZ)

_ 6

- 6

- 2

- 3(XYZ)

_-. 3(XYZ)

_ 3(XYZ) _ 3

_ 1

_ 1

- 1

_ 1

_ 1

_ 1

_ 1(Y)

Bending elbow

Bending nozzle DR202

Bending transition between valve and spherical tee

Bending nozzle DR109

-HDU rocking

-HDU rocking

HDU response bottom

HDU response top

-Spherical tee response

-Elbow response

-Piping response

in piping and nozzle.

Nozzle loads DR201

Threaded rod

Impact/load input HDU

Mass response normal tee

Support location U.S. H-1

Support location U.S. H-2, H-3

Support location U.S. H-7, H-8

Support location U.S. H-9 and threaded rod

Support location U.S. H-10

Manifold input mass and other piping (DF16)

Support H-5

Support H-4

Support H-I

Support H-2

Support H-3

Support H-6

Support H-8

Support H-7

Support H-9

Support H-10

Support H-il

Support H-12

DF16 nozzle

Support location U.S. H-11, H-12

Valve nozzle loads

Valve body loads

Valve shaft axial forces

Valve body centerline

Valve/actuator center of gravity Valve Actuator

Variable/limit stops

Motor current

Motor voltage

Valve inlet pressure

Pressure across valve pressure

Fluid temperature at valve

Visual valve monitor

Thermal growth at DF16

B-3
Table B-1. (continued)

Location

Number

51

52

53

54

55

56

57

58

59

60

61

62

63

64

65

66

67

68

69

70

71

72

73

74

75

76

77

78

79

80

81

87

88

700-1-2

703-4-5

619

629

639

649

659

Instrument

Type

Accelerometer

Temperature

Temperature

Pressure

Temperature

Acceleration

Acceleration

Acceleration

Acceleration

Acceleration

Acceleration

Acceleration

Acceleration

Acceleration

Acceleration

Acceleration

Acceleration

Force

Force

Force

Force

Force

Force

Force

Force

Displacement

Displacement

Displacement

Displacement

Displacement

Displacement

Force

Force

Acceleration

Acceleration

Force

Force

Force

Force

Force

German U.S.

3(XYZ)

3(XYZ)

1(Y)

1(Y)

1(Y)

1(Y)

1(Y)

I

1

I

1

1

1(X)

1(Y) 1(X)

1(Y)

1(X)

1(Y)

1(Z)

1(Z)

1(Y)

1(-X)

1(-Y)

1(-Y)

1(+Z)

1( -Z)

l(+Y)

1(-Y)

1(-Y)

1(X)

1(Y)

1(Y)

1(Y)

1(Z)

1(Y)

1(+Y)

1(+Y)

Remarks

Valve noise

System thermal response normal tee

System thermal response H-7

Valve internal measurement

System thermal response H-1l

HDR structure (H-1)

HDR structure (H-2)

HDR structure (H-3)

HDR structure (H4)

HDR structure (H-5)

HDR structure (H-6)

HDR structure (H-7)

HDR structure (H-8)

HDR structure (H-9)

HDR structure (H-10)

HDR structure (H-li)

HDR structure (H-12)

Cloud impact (H-i)

Cloud impact (H-2)

Cloud impact (H-6)

Cloud impact (H-7)

Cloud impact (H-7)

Cloud impact (H-7)

Cloud impact (H-7)

Cloud impact (H-12)

Pipe displacement at snubber location (H-l)

Pipe displacement at snubber location (H-2)

Pipe displacement at snubber location (H-6)

Pipe displacement at snubber location (H-8)

Pipe displacement at snubber location (H-7)

Pipe displacement at snubber location (H-12)

Cloud impact (H-12)

Cloud impact (H-6)

HDR structure near HDU bottom

HDR structure near HDU top

Spring hanger UD 20-1-6

Spring hanger UD 20-5-8

Spring hanger UD 50-1-1

Spring hanger (H-13)

Spring hanger UD 20-3-2

B-4
APPENDIX C

INVESTIGATION OF

SNUBBER MALFUNCTIONS
APPENDIX C

INVESTIGATION OF

SNUBBER MALFUNCTIONS

INTRODUCTION

The HDR experiments provided a very good

place to test snubbers in an in situ environment.

The results address some of the functional problems

encountered by snubbers in response to seismic

loading. The information produced by the HDR

tests can provide additional technical input to the

NRC effort to finalize snubber test requirements.

This appendix describes the snubbers used in the HDR

tests and reports the poor performance of the INC

snubbers.

SNUBBER DESCRIPTIONS

Snubbers have two proper functions: (a) they should

move freely at low accelerations and (b) they should

resist motion (a state commonly referred to as "lock

up") at higher accelerations. They are frequently used

in areas where piping must be allowed to undergo free

thermal expansion movement, but where clearance is

limited or where inertial movements caused by seismic

or hydrodynamic events must be restrained because of

stress considerations. Mechanical snubbers can malfunction either by failing to lock up when they should,

such as during a dynamic event, or by locking up when

they should allow movement to accommodate thermal

expansion of the piping. Hydraulic snubbers can

develope leaks, lose their fluid and fail to lock up when

they should.

Snubbers are of two basic types: hydraulic and

mechanical. The hydraulic type includes a doubleacting cylinder, a flow control device, and a hydraulic

fluid reservoir (Figure C-i). During slow movements,

such as when piping experiences free thermal expansion, hydraulic fluid passes from one side of the piston

to the other. However, at higher load rates the flow

of fluid is restricted and motion is limited. Hydraulic

snubbers are velocity controlled. Release rate velocities

are controlled by bypass flow circuits or through leakage inherent in the flow control device. An external

reservoir accommodates the thermal expansion of the

fluid and volumetric changes during snubber motion.

It also provides reserve fluid.

Figure C-2 illustrates a Pacific Scientific mechanical

snubber, which is the most popular mechanical snubbers. Any movement of the pipe to which the snubber

is connected causes a compressive or tensile load into

the left end of the snubber. This force is converted from

linear motion to rotary motion by means of a ball

screw. The ball screw is keyed to a torque transfer

drum that surrounds a capstan spring and causes it to

rotate also. The two ends of the capstan spring protrude through a slot in the torque transfer drum and

contact shoulders located on the inside of the inertia

mass. When the torque transfer drum turns, it bears

against one end of the capstan spring and moves it.

The opposite end of the capstan spring bears against

the inertia mass, which turns on a bearing. If an

acceleration greater than approximately 0.02 g is imposed, the torque carrier (the ball screw and the torque

transfer drum) attempts to accelerate also. Because of

its inertial resistance, the inertial mass will resist acceleration. This causes a relative angular displacement

between the torque carrier and the inertial mass, which

tightens the capstan spring around a mandrel that is

an integral part of the main housing; thus, the constricting action creates a braking force. When the

acceleration drops below 0.02 g, the inertia mass no

longer exerts sufficient force to limit motion, and the

capstan spring relaxes. V

Figure C-3 illustrates a second mechanical snubber

design, an International NuclearSafeguards Company

(NC) snubber. The function of this snubber is similar

to the Pacific Snubber; however, the lockup is not as

soft. .Tbe INC snubber incorporates a nonrotating

screw that moves through a rotating nut, integral

adapter, and friction plate. The friction plate is maintained between fixed flanges, at a minimal clearance,

by a centering spring. For normal operation, the

thermal growth of the piping or attached equipment

is accommodated by the linear translation of the

screws, which rotates the nut-adapter assembly and

causes a change in the effective strut length. For this

operating mode, the frictional areas are in a maintained

C-1
RET z Minimum Rod Position

EXT = Maximum Rod Position

l -~ Maximum

e- D Rodu

"A" Position

0, z Maximum Stroke

Figure C-1. Schematic of hydraulic snubber.

Telesoping Ball screw

cylinder ohsft

Torque transfer

drum

7*8906 Figure C-2. Schematic of Pacific Scientific mechanical snubber.

C-2
Figure C-3. Schematic of International NuclearSafeguards Company (INC) mechanical snubber.

position owing to the reactive force of the centering

spring. When subjected to a shock loading, the inertia

of the nut-adapter causes the centering spring to compress, resulting in a mating of the friction surfaces

and a consequent resistance to the externally applied

force.

Both types of snubbers were used in the U.S. stiff

pipe support system-five mechanical snubbers and one

hydraulic snubber. Originally, four of the mechanical

snubbers were of INC design, and the fifth was manufactured by Pacific Scientific. Because of their erratic

performance during the preliminary tests, the INC

snubbers were replaced by Pacific Scientific snubbers

for the remaining tests. The hydraulic snubber was produced by Bergen-Patterson. Table C-1 defines the

snubbers used in the tests.- The INC snubbers were

located in positions H-6, H-7, H-8, and H-12 during

the preliminary tests.

Table C-1. Snubber descriptions

Support

Location Direction

Snubber Snubber

Model Type

H-1 Horizontal and perpendicular to horizontal run Pacific Scientific PSA-I

H-2 Vertical

H-6 Vertical

H-7 Horizontal and perpendicular

H-8 Vertical

H-12 Vertical

Bergen-Patterson 2525-10

Pacific Scientific PSA-I

Pacific Scientific PSA-3

Pacific Scientific PSA-1

Mechanical

Hydraulic

Mechanical

Mechanical

Mechanical

Pacific Scientific PSA-1 Mechanical

C-3
EVALUATION OF INC SNUBBER PERFORMANCE

Because the INC snubbers were installed only during

the preliminary shakedown tests, acceleration measurements are not generally available; thus, only measured

displacement and force are reported and analyzed here.

The four snubbers manufactured by INC, an affiliate

of Bergen Pipe Support Corporation, U.S.A., were

originally NRC-owned spares for the Loss of Fluid

Test (LOFT) Facility at the INEL. The four snubbers

were all model MSVA AS, one a size 1, and three a

size 2. The four units were all N-stamped and met all

requirements for installation in a nuclear plant.

The snubber locations suspected of reduced performance were H-6, H-7, H-8, and H-12. Snubbers at H6,

H-7, and H-8 were MSVA-SA-size 2; the snubber at

H-12 location was size 1.

Figures C4 through C-7 compare force and displacement at the four snubber locations for test T40.34.

The displacement time histories indicate that in each

case there was sufficient motion at each snubber location to activate the snubber. The histories also indicate

that the displacements got larger when snubbers quit

operating. The force histories show when a snubber

was resisting motion and when it was not. They also

provide evidence of the maximum force at which each

snubber reacted. These forces are all less than normal

and upset design loads specified by the manufacturer.

Figure C-8 compares the force time history recorded

at the Pacific Scientific snubber located at H-I with

the force history recorded at the INC snubber located

at H-6 during test T40.34. The snubber at H-1 resisted

motion, as shown in the force history, whereas the

snubber at H-6 performed erratically.

The displacement histories provide evidence there

was sufficient motion of the piping to cause the snubbers to operate beyond the time at which they failed.

The displacement histories associated with the erratic

snubber performance show that displacement increases

during the time that snubber performance decreases.

Acceleration time histories from later tests, recorded

at the same locations and shaker starting frequencies,

verify that sufficient motion was available to actuate

the snubbers.

The posttest visual examination and manual actuation of the snubbers provided no insight into the

malfunctions of the snubbers. The snubbers that per- formed erratically when installed in the VKL system

performed erratically also when tested manually after

removal from the system.

During our investigation, we discussed the snubber

problem with snubber vendors and staff from the NRC

Office for Analysis and Evaluation of Operational Data

(AEOD). The discussions verified our conclusions.

The company is now out of business, and apparently

most, if not all, of the INC snubbers in nuclear plants have been replaced. AEOD also said that EE had issued

several reports and bulletins concerning snubbers in

general, but none specifically warning of problems with

INC devices. Further notice may be needed to ensure

all INC snubbers have been replaced.

Comparison of the responses of the hydraulic snubber at H-2 and the Pacific Scientific mechanical snubber at H-1 with the responses of the INC snubbers

provided evidence that the INC snubbers were not performing to design requirements. The snubbers were

removed and manually checked for operation. Each

performed erratically when exercised by hand. The

snubbers were replaced by mechanical snubbers provided by Pacific Scientific Company.

C-4
1400

1000

600

200

2

-200

-600

-1000

0 20 40 60 80 100

Time (s)

.3

.2

aE 0almMbi1a11.

CL

5

-. 3

0 10 20 30 40 50 60

Time (s) 7-8929

Figure C-4. Comparison of force and displacement at support location H-6. Note the erratic behavior of the snubber from

45 to 80 s. The snubber is locking up only in the negative direction throughout most of this period.

C-S
600

' 400

D. 200

0

0

U. 0

-200

-400

''I_~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~ I

.~~~~~~~~ . . . ."M

0 20 40 60

Time (s).

80 100

.25

.2

.15

-d .1

E .05

E 0

0

D-.15~f

-. 2

.25 10 20 30o4 50 60

Time (s)

7-8923

Figure C-5. Comparison of force and displacement at support location H-S. This snubber resisted motion for only 2 or 3 s.

C-6
1000

800

600

4000

20 0

.0

25

i2 -200

-400 -

-800

-1000

0 20 40 600 0

Time (s)

.25

.15

.0 5

E0

-. 25

0 lo 20 ~~~30 40 50 60

Time (s) 78922

Figure C-6. Comparison of force and displacement at snubber location H-7. This snubber performed reasonably well from

40 to 80 s. Note that as the snubber performance becomes more erratic, the displacement increases.

C-7
a0

-50

-100

-150

-200

0 20 40 60 80 100

Time (s)

.25

E0

0 10 20 30 40 50 60

Time (S) 7.8921

Figure C-7. Comparison of force and displacement at snubber location 11-12. This snubber operated erratically for a few

seconds then failed to resist motion for the remainder of the transient.

C-8
z

0

U60

Time (s).

a. Location H-1

0

U60

Time (s).

T.6901 b. Location H-2

Figure C-S. Comparison of force time histories for the snubber located at H-1 with the snubber at H-6.

C-9
APPENDIX D

INVESTIGATION OF SPURIOUS

DISPLACEMENT READINGS AT HDR
APPENDIX D

INVESTIGATION OF SPURIOUS

DISPLACEMENT READINGS AT HDR

Displacements recorded during SHAG testing at

HDR indicated that the five Pacific Scientific mechanical snubbers and the Bergen-Patterson hydraulic snubber allowed displacements greater than 0.1 in. peak

to peak, the maximum specified by Pacific Scientific

for the mechanical snubbers. Figure D-1 shows the

four responses recorded at snubber location H-12 during one of the tests. These responses are typical of those

measured at all snubber locations.

Although the displacement readings were greater

than they should have been, forces measured at the

snubber pins indicated that the snubbers resisted

motion. Measurements of piping acceleration also indicated that the snubbers functioned properly, though

the acceleration responses were large enough that if

the major frequency content was at low enough frequencies, the large displacements were theoretically

possible.

The apparent excessive motion allowed by the snubbers, as indicated by the displacement measurement,

is of concern. Piping systems in nuclear plants are

designed and analyzed for small dynamic displacements, and larger displacements allowed by snubbers

could cause higher stress, or even impacting, where

piping systems are installed close together. Pipe impacting with valve operators is of even greater concern, where structural damage could prevent valve

operation.

The displacements were measured by Celesco

PTI1O1-20A position transducers. The transducers were

a

LOL

C

09

0

0

-8001 1 1 1 1 1 1 1 1

104.99 116.99 128.99 140.99 152.99 164.99

Time

(a) Force

104.99 116.99 128.99 140.99 152.99 164.99

Time

(c) Pipe acceleration

-1.04

C

0

E

0

U

0

0.52

0

-0.52

1.04

0.30000

'X 0.20000

E 0 0.10000

<D 0

I I - ' I I I I

11 01 J~~~~~ I I I - I r'

104.99 116.99 128.99 140.99 152.99 164.99

Time

(b) Displacement

-0.10000 I I-l II I I I I I 1

-0.20000 1 lI I I I I I I I I I

104.99 116.99 128.99 140.99 152.99 164.99

Time

(d) Wall acceleration V.UbA4

Figure D-1. Responses recorded at Pacific Scientific snubber location H-12 during Test T40.30, 8-Hz starting frequency.

D-1
mounted solidly to the HDR structure perpendicular

to the run of pipe at each snubber location. The cable

from the transducer was then fastened to the pipe. The

transducer function, simply described, is a potentiometer, turned by displacement of the cable, that is

continually preloaded by an internal spring. The

potentiometer output wiper voltage is calibrated to read

cable travel in inches. Where extra long distances

between the structure and pipe existed, an additional

length of cable was added to the transducer cable.

Transducer calibrations performed at HDR prior to

testing and during the test program indicated that static

deflection of the transducer cables provided accurate

measurements.

We conducted a posttest investigation at the INEL

dynamics laboratory to determine whether the Pacific

Scientific snubbers, the most widely used mechanical

snubber used in U.S. nuclear plants, were in fact allowing displacements larger than the 0. 1-in. peak-to-peak

dead band specified by the manufacturer. The investigation indicated that the large displacement readings

were a product of the Celesco transducers rather than

the actual movement of the piping. The results of that

investigation are reported in the following paragraphs.

One way to check measured displacements is to calculate displacements from the measured accelerations.

For each of the snubber locations, we double integrated

the acceleration histories to determine how well the

calculated displacements would correlate with the

recorded displacements. The results of this work confinned that the displacements allowed by the snubbers,

as determined from integrated acceleration, were within the tolerance specified by the manufacturer. As can

be seen from the examples shown in Table D- 1, displacements recorded by the Celesco transducers were

generally an order of magnitude greater than those

calculated from the double-integrated accelerations.

The displacement transducers, excitation power supplies, accelerometers, recording equipment, and other

associated equipment used in the HDR work were set

up in the INEL dynamics laboratory. Using a small

electromechanical shaker, we simultaneously subjected

the accelerometers and displacement transducers to

various vibration histories, including random and sine.

The accelerometer histories were double-integrated and

compared to displacement histories recorded by the

transducers. Figure D-2 is an example of this work

with a random signal comparison. The comparison

shows that the recorded displacement and accelerationderived displacement correlate very well.

Four of the snubbers were returned to Pacific Scientific for testing. The snubbers met all performance

requirements.

The Pacific Scientific snubber from location H-I at

HDR was installed on a test machine at INEL. The

Table D-1. Comparison of recorded

displacements and doubleintegrated accelerations

Zero-to-peak Displacement

(in.)

Double

Location Celesco Integration

H-I 0.7 0.064

H-8 0.45 0.013

H-12 0.42 0.035

snubber was subjected to random, sine, and sine with

random noise superimposed. The random and random

sine signals were used to more closely duplicate natural

earthquakes and the SHAG spectra. Pacific Scientific

is not equipped to subject snubbers to these types of

signals. The snubber met or exceeded the manufacturer's dead band tolerances for all signals imposed

with loadings up to ASME code Level C rating.

A seventh Celesco displacement transducer was used

at HDR to measure thermal growth at the DF 16 manifold, instrument location 50 (see Figure 5, Vol. 1).

This transducer also measured the manifold's vertical

displacement during dynamic excitation. Adjacent to

this displacement location was an accelerometer, instrument 35. We double integrated the acceleration

history for test T40.30 (U.S. stiff piping configuration) at 8-Hz starting frequency and compared it to the

displacement history. Figure 61, Vol. 1, shows this

comparison. The correlation is very good. Both plots

show approximately 0.2-in. double amplitudes.

All seven of the transducers were excited and recorded at HDR with the same equipment. The data

inconsistencies cannot be attributed to difference in

measurement technology. The following theory is

offered as a possible explanation. The manifold was

restrained in both horizontal axes by large circumferential restraints, allowing only vertical motion. The

piping at the snubber locations responded in all three

axes, including axial motion parallel to the pipe. This

axial motion is offered as the primary difference in the

specimen response between the six transducers at snubber locations and the one at the manifold.

The logic for the transducers' anomalous response

is as follows. The transducer system consisted of a

cable wound up on a preloaded drum attached to a

multiturn potentiometer. We believe the spurious

transducer readings are a result of cable whip. The

mass of the cable and the rotational inertia of the

D-2
E

C-

_5x10-21

Time (s)

(a) Displacement transducer time history.

_X... 2

CL

-5X10' 2

o 5.11

Time (s)

7-8902

(b) Accelerometer double-integrated acceleration history.

Figure D-2. Laboratory test comparison of a displacement time history and an accelerometer double-integrated acceleration

history.

D-3
preloaded drum were excited by transverse and

parallel vibration simultaneously. This resulted in a

resonance situation that caused the potentiometer to

read one or more cable drum system harmonic vibrations and add them to the actual displacement of the

piping.

The Celesco Transducer Company worked very

closely with INEL Engineers to resolve the displacement measurement anomaly. This work included considerable work at their expense. In the end, agreement

could not be reached and the following is a quote from

Celesco:

"Celesco transducers were operated in a condition

exceeding their published specifications, i.e., g forces.

Operating under these conditions, as well as with the

use of external attachments, i.e., extension cables,

could provide an adverse effect on the performance of

the transducers."

INEL research does not support this conclusion

offered by Celesco, for the following reasons. The

transducers installed without cable extension, i.e.,

where extended to midstroke (H-i and H-7), yielded

some of the most inconsistent data. Pipe accelerations

of 1 to 3 g are not outside the published envelope for

this model of Celesco transducer. The actual response

of the transducer is outside its published operating

envelope, and, as previously stated, we believe to be

a result of cable-drum resonance.

In later tests conducted at the HDR, the same type

of snubbers were subjected to dynamic loads created

by servohydraulic shakers attached to the VKL. Those

tests used Linear Variable Displacement Transducers

(LVDTs) instead of Celesco transducers to measure

displacements allowed by the snubbers. The results of

those tests confirmed that the snubbers performed

within manufacturer's specifications.

D4
APPENDIX E

VALVE RESPONSE FREQUENCY

ANALYSIS METHODOLOGY
APPENDIX E

VALVE RESPONSE FREQUENCY

ANALYSIS METHODOLOGY

INTRODUCTION

Frequency analysis of time-history transducer records sometimes yields additional information that is

not apparent in the original time records. Perhaps the

most useful frequency domain method of data analysis

is the computation and examination of frequency

response functions (FRFs). However, this procedure

requires one or more time-synchronized reference

signal(s) that may be interpreted as input(s) to the

system. In the present study, there exist a number of

inputs to the system such that use of FRFs would

become a difficult and costly task. An alternate, but

less powerful, method of frequency domain analysis

is the study of individual response signals. The primary

disadvantage of this method is that phase information

is lost; relative behavior of response signals is unattainable. The most useful form of frequency domain

function, when individual time responses are transformed into the frequency domain, is termed autospectrum, mean-square spectral density, or power

spectral density (PSD). Owing to the complexity and

number of force inputs to the VKL system, the PSD

was the method of choice for frequency domain

analyses.

This appendix presents discussions of structure, PSD

computation techniques, and theory used in the frequency analysis.

All of the data reduction reported herein was performed on a GenRad 2508 minicomputer system using

the MODAL PLUS software package from Structural

Dynamics Research Corporation (SDRC). Standard

MODAL PLUS time data files consist of sequential

frames of data, each of which contain 1024 discrete

values of digitized data with constant time-step. Data

acquired by MPA was digitized with time-step equal

to 4.6 ms, whereas data acquired by KfK was digitized

with time-step equal to 8 ms. Thus, each data frame,

in the time domain, represented 4.71 s and 8.18 s for

MPA and KfK acquired data, respectively.

Data Analysis

Standard procedures were used to calculate PSDs of

the acceleration time records. Briefly, for each time

record studied, this consisted of (a) multiplying the time

frame by a window function (Hann function) to reduce

distortion or leakage of the resultant PSD, (b) Fourier

transforming the windowed time frame, (c) multiplying the transform by its complex conjugate, and, finally

(d), multiplying the result by a scale factor to account

for loss of power during the windowing process and

to scale the squared transform by its bandwidth. This

operation is symbolically described as follows:

Let S(co) = Power Spectral Density (PSD)

= Bw 7 [w(t)g(t)]2

where

Bw

y I

w(t)

g(t)

= effective bandwidth

I = Fourier transform of time function

= window function

= time domain acceleration.

Owing to the differing time-steps employed in data

acquisition, net (upper bound) frequency band widths

were 108.7 Hz and 62.5 Hz for MPA and KfKacquired data, respectively. The effective band width

was taken to be 0.8 times the net band widths, or 87 Hz

and 50 Hz, respectively, to account for finite roll-off

of antialiasing filters. Thus, for subsequent calculations

and comparisons, a common effective band width of

50 Hz was used. Since 1024 point time frames were

used in the Fourier transforms, frequency resolution

was net band width divided by 512, or 0.212 Hz and

0.122 Hz for MPA- and KfK-acquired data,

respectively.

A standard technique used in frequency analysis is

to average the PSDs obtained from several sequential

time frames in order to minimize the effect of noise

in the measurements. However, for the measurements

studied herein, several of the resonant peaks in the

PSDs were changing in frequencyas a function of the

time frame studied due to shaker coastdown. Thus,

averaging several sequential frames had the effect of

smearing these variable PSD peaks. In order to reduce

this smearing effect, it was decided to study and compare the data on a frame-by-frame basis. However,

owing to the differing time frame lengths between

MPA- and KfK-acquired data, and since the MPAacquired frames were approximately one half the length

of the KfK frames, two adjacent PSDs resulting from

E-1
MPA data were averaged together for comparison with

an individual frame PSD resulting from KfK data.

Thus, each PSD frame from KfK acquired data

(spherical tee and top of the HDU) represents 8.18 s

and each resultant averaged PSD frame from MPAacquired data (valve body, valve actuator, and standard

tee) represents 9.41 s of acquired time data. The start

time of the first frame of time data chosen for transformation was taken to be approximately 1 s before transient initiation. It was found that the major response, for each location studied, was contained in either of

the first two frames, or before approximately 16 s after

initiation of the transient. Thus, subsequent discussion

of analysis and results deals with the first two frames

of KfK data, or the first two composite frames of MPA

data.

In order to refine and allow more quantitative evaluation of the PSDs, integration of the raw PSDs was

performed. Three integration procedures were used:

first, integration and subsequent square root of the

entire PSD; second, integration and square root of

selected frequency bands through significant peaks in

the PSDs; and, third, cumulative sums (as a function

of frequency) of the band integrated PSDs were found.

A symbolic description of each of these reductions may be given as follows:

The average power of the measurement (mean square) between frequencies ac1 and w2 is,

C C2 (2 =| S (O)di

At

then cumulative acceleration from 0 to f Hz is

f } 1/2

G~m,(f) = f S(4)df

n f 112

i=i (Jo

The first reduction technique may be interpreted as

the root mean square (rms) average acceleration, from

zero to a given frequency, of each frame of data. The

second technique may be interpreted as the rms acceleration contribution of each significant resonance peak contained in the PSD. The third technique represents the sum, as a function of frequency, of all of the

resonance peaks contained in the PSD.

Analysis Results

Figures E-1 through E-60 reproduce PSDs found

from frequency analysis of the 30 acceleration time

records (X, Y, and Z directions for each of the five

locations and for each of the two tests). The first two

frame PSDs are shown in all cases. The most signifi- cant aspect of the PSDs is the relatively large magni- tude high-frequency content of valve operator PSDs, especially for excitation in the Z direction. Examination of PSDs for the other four locations studied does

not indicate nearly the degree of high frequency content. Due to the relative difficulty of quantitative examination of the PSDs, the band integrated represen- tation of the data results is discussed.

Tables E-1 through E-3 present rms values of the

resultant band integrations performed for all signifi- cant peaks in the data studied (PSD frames 1 and 2 of the 30 acceleration time records). Each table represents

a direction, and each horizontal group of values repre- sents a location. The first two columns represent frames

one and two of test T40. 10, and the last two columns

represent corresponding frames of test T40.30. The

data are given as a number (the integrated rms acceleration in G) followed by a bracketed range (the band over

which the integration was performed in Hz). In gen- eral, each row of data represents common frequency bands for significant peaks found in the PSDs. Thus, where blanks occur in the data tables, there are no corresponding significant peaks in the PSD. For each

frame of data, the sum of all of the major contributing bands is given. This sum may be interpreted as the rms acceleration of the corresponding time frame in the case

where frequency components are additive. In general, comparison of the summed rms accelerations with the

or

GM, ={f S(Q)dJ.

Average acceleration from 0 to f Hz is

( f 112

Grmf = {fr S()d }4 I

If S(co) is concentrated in discrete bands, [coil, ci21, let:

SM(O) = S(O), sil 4 co 4 COU

= O. co < coill IaC > COU

and

n

S(w) = S(aw) - I Si(c),

i=1

E-2
corresponding time records does indicate that the frequency components are additive.

The tabulated data are graphically represented in

Figures E-61 through E-96. These figures are threedimensional bar chart representations of the band

integrated data where the height of each bar indicates

rms acceleration, and the width indicates the band over

which each integration was performed. The first set

of bar chart figures, E-61 through E-90, graphically

reproduce the numerical values given in Tables E-I

through E-3, where results of both data frames are

shown on a figure. These figures illustrate variability

of some of the component frequency bands as a function of time, where the shift to lower frequencies of

some of the bands is evident. The bands that exhibit

this shift are, generally, in the lower frequency region,

from 5 to 15 Hz. If all of the bands contained in a PSD

are classified in two categories-those that are associated with natural modes of vibration of the VKL

system, and those that are associated with the input

driving force-then the bands that do not vary in frequency are natural modes and those that do vary reflect

the driving force. The second set of bar chart figures,

E-91 through E-96, graphically illustrates the degree

of acceleration amplification from the valve body to

the valve actuator (first frame). Note that this amplification is especially significant for higher frequencies

in the horizontal directions. Unfortunately, data that

were acquired by KfK was low-pass filtered at 30 Hz

during the acquisition process. Thus, potential high frequency (35 Hz to 50 Hz) information present at the

top of the HDU vessel and at the spherical tee has been

removed from the acceleration histories recorded for

these locations.

Figures E-97 through E-102 illustrate the average

rms acceleration, as a function of frequency, for all

five location first frame PSDs. Figures E-103 through

E-108 illustrate corresponding cumulative accelerations

for the five locations as continuous functions of frequency. Note that the cumulative functions are similar

to a running sum of the band integrated data, except

that PSD information not included in the individually

integrated bands is shown in the continuous cumulative

functions. The average and cumulative functions represent, respectively, lower and upper bounds of time

domain rms acceleration. The average data would more

accurately represent the time data in the case where

there is little direct superposition of frequency components, whereas the cumulative functions would more

accurately reflect the case where there is significant

superposition. If the rms values are multiplied by a factor of 1.414 to approximate the change from rms to

peak, comparison of these frequency distributions with

corresponding time records indicates that the cumulative distributions more accurately represent the peak

measured accelerations.

E-3
CPSMVALVE RESPONSE

To further verify the significant high-frequency valve

operator acceleration response observed during the

HDR SHAG tests, we performed a limited examination of the results of a similar but related study, the

Containment Penetration System (CPS) dynamic test

(reported in NUREGICR-4734).

The CPS test program was designed to examine the

response of typical containment penetration piping

systems and associated components for a variety of

loads. Of the systems tested in the CPS program, two

contained large valves: one was an 8-in. system with

gate valves, representing a containment spray system,

and one was an 8-in. system with butterfly valves,

representing a containment purge and vent system. One

of the loadings imposed on the system was a simulated

SSE three-dimensional dynamic shake. The CPS

dynamic tests were designed so that the piping systems

studied were supported from a large stiff integral

framework (see Figures E-109 and E-110). For the

SSE tests, this framework was excited through a

number of hydraulic actuators so that a full threedimensional base-motion SSE excitation was accurately

simulated.

A striking similarity is evident when CPS gate and

butterfly valve acceleration responses are compared

with the corresponding HDR SHAG test gate valve

acceleration response. This observation is even more

significant when the complete dissimilarity between

CPS and HDR VKL system geometries and excitation

methods are taken into account. We conclude that the

observed frequency-dependent behavior of the three

valves studied is primarily a function of the valves'

dynamic characteristics rather than the piping system

geometry or method of input loading excitation. However, the observed behavior may depend on the valve

being installed in a prototypical piping system, versus

bench or analytic frequency determination.

The CPS dynamic testing provides two additional

references for valve in situ response. Both valves are

modem nuclear N-stamped valves from cancelled

nuclear facilities, Hope Creek and WPPSS 4. The valve

manufacturers have determined the valve fundamental frequencies are greater than 33 Hz, but, like the

HDR valve, both show amplification under 33 Hz, as

well as the high-frequency response. HDR and CPS

are two of the largest full-scale simulated seismic test

programs performed. There are some distinct similarities in response, that point out possible shortcomings

in current valve qualification standards for seismic

qualification and for single effects testing. The results

of the test programs will be presented to the ASME

valve qualification subcommittee for review.

Data and Test Description

ANCO engineers supplied CPS system dynamic

tests, instrumentation, and data acquisition. For. both

of the valves in the simulated SSE tests, ANCO digitized and recorded full triaxial acceleration of the valve

body and actuator. The duration of each SSE excitation was approximately 30 s, during which all data

were acquired and simultaneously digitized with sample rate and antialiasing filter frequency set at 200 sps

and 45 Hz, respectively. Thus, the useful bandwidth

of the data was 50 Hz, as was the case with HDR

SHAG data. All data were available via magnetic tape

in standard ANCO ASCII format.

Data Analysis

Data analysis of CPS valve acceleration data was

performed on the GenRad 2508 minicomputer-based

systems in a manner similar to HDR SHAG data

analysis. In the present case, strong motion valve

acceleration data was of 30-s duration. Thus, for the

CPS test sample rate (200 sps or time step equal to

5 ms), five data frames were available, each consisting

of 1024 data points with 5.1 s duration. Each data

frame was transformed, as previously discussed, to a

PSD. In the case of CPS data, no frequency shifts of

major PSD peak values appeared when individual

frame PSDs were compared. Thus, to minimize noise

inherent in the recorded acceleration data, all five PSD

frames were averaged together. Figures E-111 and

E-112 give the resultant average PSDs (X, Y, and Z

directions for each of the two valve bodies and actuators). Figure E-111 represents response of the gate

valve; Figure E-112 represents response of the butterfly valve. As opposed to the HDR SHAG PSDs, there

are few distinct major peaks in the CPS valve PSDs.

Thus, band-integrated rms accelerations were obtained

with fixed integration frequency bands, each with a

5-Hz bandwidth, resulting in 10 integration bands per

PSD. Tables E-4 and E-5 give the results of these band

integrations. For each band, valve body and operator

rms accelerations are compared through use of an

amplification factor (operator acceleration divided by

valve body acceleration). Finally, frequency-dependent

average mis accelerations were found through integration of each PSD across the entire 50 Hz band. Figures E-113 through E-116 show the results of these

integrations, where each figure represents triaxial acceleration for a given location and valve, with all three

average mns acceleration functions superimposed.

E-4
Analysis Results

Both CPS valve acceleration responses exhibited

significant high-frequency response and amplification,

as was observed for response of the valve installed in

the HDR VKL system. This is most significant in the

horizontal response directions (X and Z directions),

which are normal to the axes of the valve operators.

In the case of the gate valve, the greatest response

and amplification occur for response in the X direction and in the higher-frequency range (35 to 45 Hz).

However, for gate valve response in the Z direction,

the greatest response and amplification occur in the

intermediate-frequency range (20 to 30 Hz). In addition, a small degree of intermediate-frequency acceleration amplification may be observed for response

parallel to the valve actuator stem axis (Y direction).

Butterfly valve results are similar to results found

from the gate valve response. However, for the butterfly valve, greater response and amplification in the

higher frequency range are found in the Z direction

than the X direction. We note that for response in

the X direction, amplification occurs across a widefrequency range (8 to 45 Hz).

E-5
X

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so

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0.5

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0 to 20 30 40 50

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0.05. F .5

0 t0 20 30 40

Frequency (Hz)

Figure E-69. Band Integrated Acceleration: X direction, top HDU, Test T40. 10.

50

a1

L

a

.0

a

.4

a

uU

UC

II

O

II

II

II __s1%- -- - - -

K I

0 t0 20 30 40

Frequency (Hz)

Figure E-70. Band Integrated Acceleration: X direction, top HDU, Test T40.30.

E-40
L

e

-; O .- C

L0

V

p4

U.

0 to 20 30

Frequency (Hz)

Figure E-71. Band Integrated Acceleration: Y direction, valve c

0.30

0.25

;2 0.20o

LE

a

0.15

.a

v

eL

V4

U

O 010.

0.05

_Wqt~~ - - - _ - _ v

I

20 30

40 50

operator, Test T40.10.

- - -e - - _ _ _ _~ ... . _ . s _~

I

= -

0 t0 40 SO

Frequency (Hz)

Figure E-72. Band Integrated Acceleration: Y direction, valve operator, Test T40.30.

E-41
a

C

8

.4

41

.4U

Frequency (Hz)

Figure E-73. Band Integrated Acceleration: Y direction, valve body, Test T40. 10.

0.30

0.25

i 0.20

C

.0 0.15

4*

-4

U4

u.4 0.10

0.05

_W S~C,-__-_-X- -- XC B ls-C:- - - I t0 to 20 30 40 5-

Frequency (M)

Figure E-74. Band Integrated Acceleration: Y direction, valve body, Test T40.30.

E42
'; 0. 4

r.E

.0 o .3 ~

L0

.4

a1

E 0.065

0

5

.4

0.02

U

-C

0.02-

0 t0 20 30 40 50

Frequency 1Hz)

Figure E-75. Band Integrated Acceleration: Y direction, spherical "T", Test T40.10.

s. x, '. '..-

0 t0 20 30 40 50

Frequency 1Hz)

Figure E-76. Band Integrated Acceleration: Y direction, spherical "T", Test T40.30.

E-43
0.

a

L

.

4.

U

L

a

Fe

a

-4

0. F

0 so 20 30 40 50

frequency (Hz)

Figure E-77. Band Integrated Acceleration: Y direction, standard 'T", Test T40.10.

*; o.'

0.

at°

44

0.

0 t0 20 30 40 50

Frequency (Hz)

Figure E-78. Band Integrated Acceleration: Y direction, standard "T", Test T40.30.

E-44
o to 20 so 40 so

Frequency (Hz)

Figure E-79. Band Integrated Acceler0tion: Y .10on, top HDU, Ted T40.10.

0.06

0.06

0.0

0.02

0 to 20 so 40 so

Frequency IHz)

Figure E-80. Band Integrated Acceleration: Y direction, top HDU, Test T40.30.

E-45
*.I

-0.4

L

S

I- I a

~f

a

U

1 0.2-

0 t0 20 30 40 s0

Frequency (Hz)

Figur E-81. Band Integrated Acceleration: Z direction, valve operator, Test T40. 10.

0 10 20 30 40 50

Frequency (Hz1

Figure E-82. Band Integrated Acceleration: Z direction, valve operator, Test T40.30.

E-46

0.S
0.30.

0.25

* 0.20

E

C.6

C

° 0. O5

94

£

U< *1U O. 10.

I

I

0.05

0

! i~~~--

- 20 - 30 -- ~ -S- 4-

20 30 40 50

Frequency (HzJ

to

Figure E-83. Band Integrated Acceleration: Z direction, valve body, Test T40. 10.

0.30

0.25

7 0.20

LO

a

._ _ _ _

0 t0 - 20 30 40

Frequency 1Hz)

Figure E-84. Band Integrated Acceleration: Z direction, valve body, Test T40.30.

E47
S£

C.

.4

U

U0

4

0 .

- - - - -

0.

C

a

.4

.4

4

0 t0 20 30 40 50

Frequency (Hz)

Figure E-85. Band Integrated Acceleration: Z direction, spherical "T", Test T40.10.

0 t0 20 30 40 50

Frequency (Hz)

Figure E-86. Band Integrated Acceleration: Z direction, spherical "T", Test T40.30.

E-48
tl C.,

a

C

a

-I

LI

re 0.1

U

C

: - _

0 t0 20 . 30 40

Frequency (Hz)

Figure E87. Band Integrated Acceleration: Z direction, standard "T", Test T40.10.

So

T

c

L

a

C

a

.e

'I

S

c

LI1

49

0 to 20 30 40

Frequency (ez

Figure E-88. Band Integrated Acceleration: Z direction, standard "T", Test T40.30.

-E-49

50
0.7

I

a

a

C.a

4

0.

_____ - s - - - - -_- 

% % 1%~~~~~~~~~

0 t0 20 30 40 50

Frequency tHz)

Figure E-89. Band Integrated Acceleration: Z direction, top HDU, Teat T40. 10.

0.7

0. a

0.5

C

L 0. 3

0.2

-~ ~ ~ ~ ~ ~ ~ ~ ~ ~ -t

0 10 20 30 40 s0

Frequency (Hz)

Figure B-90. Band Integrated Acceleration: Z direction, top HDU, Test T40.30.

B-SO
ax

c

e 0.2

c

C

~i0.2.

0 t0 20 30 40

Frequency (Hz)

Figure E-91. Band Integrated Acceleration Amplification: Frame #1, X direction, Test T40.10.

50

0.4-

I

L

C

.4a

I.

.4

S

L

<

U

Al

0.{

_.14~4 s

I .

0 to 20 so 40

Frequency (Hz)

Figure E-92. Band Integrated Acceleration Amplification: Frame C1, X direction, Test T40.30.

-51
0.

C.1

C

a 0.

L

.4

U

0.4

0. V

L

a

C

.4

U

A

0.0

0 10 20 30 40 50

Frequency (Hz)

Figure E-93. Band Integrated Acceleration Amplification: Frame #1, Y direction, Test T40. 10.

5

5

0 t0 20 30 40 50

Frequency (Hz)

Figure E-94. Band Integrated Acceleration Amplification: Frame #1, Y direction, Test T40.30.

E-52
0.6

'; 0. 4'

C

C.0 S

U1

C20 .2.

I

C .;

C

a

41

L0

A 0.1"

U1

-U

0 t0. 20 30 40 so

Frequency (HZ)

Figure E-95. Band Integrated Acceleration Amplification: Frame #1, Z direction, Test T40.10.

3

I~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~ _

0 140 20 - 30 40 50

Frequency Piz)

Figure E-96. Band Integrated Acceleration Amplification: Frame #1, Z direction, Test T40.30.

E-53
0.S.

'a

0.

a

-I

U1

U OA .4

U0 0

4.1

O.Z

0

Valve Operator

Standard TT'

_ - Z~~~~

/ ~ ~~~ V--alve body

j t~~~~~~~~~ ~~Spherical 'T'

0 to 20 30 40 50

Frequency tHzi

Figure E-97. Average Integrated Acceleration: Frame #1, X direction, Test T40.10.

0.1

C.1

'a

C.

.4

.4

C.0

IS1

C

Valve Operator

0.;

Standard IT'

0

0 to 20 30 40

Frequency (iHz)

Figure E-98. Average Integrated Acceleration: Frame #1, X direction, Test T40.30.

E-54

50
0 t0 20 30 40

Frequency (Hz)

Figure E-99. Average Integrated Acceleration: Frame #1, Y direction, Test T40.10.

0.5

0.4

E

0.3 0

L ~~~~~~~~~~Valve Operator _ ~ ID

C8J

< 0. d/ Valv Eoy

WC / Sphe~ridal 'T'

< O . t Standard 'T'.<9

U 0.4

0, ~~~~~~~~~Top HDU

~0.

o t 20 30 40 so

Frequency (Hz)

Figure E-100. Average Integrated Acceleration: Frame #1, Y direcion, Test T40.30.

E-ST
1.0o

Ra

c 0.6

a

.4

U

.2

C.o

0

.1

< 0.4

L

C.a

.4

0.2

0

0. B'

0.'

a

.4

C.

U0 4

C.a

.4

0.2,

0

0 t0 20 30 40

Frequency (Hz)

pure E-101. Average Integrated Acceleration: Frame #1, Z direction, Test T40.10.

so

0 t0 20 30 40

Frequency (HzJ

Figure B-102. Average Integrated Acceleration: Frame #1, Z direction, Test T40.30.

E-56
- 5.5 E

c

10i

0 1.0~

CU

I11

.4

U

Figp

2.00

*.-; 1.50

C.1

C

.4,

a* 1.00 .4

.3

.4

0. 50

a

Valve Operator

Body

Frequency (Hz)

ire E-103. Cumulative Integrated Acceleration: Frame 11, X direction, Test T40.10.

Valve Operator

Valve Body

0

0 to 20 30 40

Frequency tHz)

Figure E-104. Cumulative Integrated Acceleration: Frame #1, X direction, Test T40.30.

E-57

50
0.75-

W *5l Valve Operator=

Vaive Body

2

0.50 < I 4 ~~~~~~~~~~~~~~Standard 'T'

i0.25 gggTop NDU

0

U

0 10 20 30 40 50

Frequency (HZ)

Figure El105. Cumlative Integrated Acceleration: Frame #1, Y direction, Test T40.10.

.00

5.75

L

0

4 Y~~~~~~~~~~~~~~~alva Operator

< ~~~~~~~Valve Body St=

v ~~~~~Spherical 'T' tndr

0.°25

0

0 t0 20 30 40 50

Frequency (Hz)

Figure E-106. Cumulative Integrated Acceleration: Frame #1, Y direction, Test T40.30.

E-58
E

LI.a

e 2.1

C

C

v

40

a" !

U

.4

0

C

O.q

Fi

2.(

U

O.~

CD

a

44C.

U

U

.4

0.I

U

Valve Operator

Standard 'T'

I

. Frequency IHZI

;ure E-107. Cumulative Integrated Aceleration: Frame *1, Z direcdon, Test T40.10.

Valve Operator

Top WDu

0

0 10 20 30 40

Frequency 1Hz)

Figure E-108. Cumulative Integrated Acceleration: Frame #1, Z direction, Test T40.30.

-59

50
Us 7+ t.,,> -t . q -t 'a' , 'o %7- oh' . . ;.'- .- V.'k , Y.;i*- P V -ss$ v, l " . . -i l "I-''v - mal .w w.vwv.r^ .*. r . .,.a >9 -- r.e " : - ..-. . VJ,, ,.-,. .? *-

8-in. gate valve

Box beam ///N

hangers,_ \ll>

Figure E-109. Sketch showing the location of the 8-in, piping system with gate valve installed in the test fixture.
Y

Gate vdve 2(2)

1

3

A - Penetration

1

3

rn

3

Key

1-. Strain gages on pipe.

2- Strain gages on support

3- Accelerometers

4- Static pressure

For Each Valve-22 Channels

3 Accelerometers (9 channels)

1 Stem position (1 channel) (6 channels) 6 Uniaxial strain (6 channels)

(1 channel) 2 Strain rosette (6 channels)

(3 channels)

(1 channels)

MHO 1546

Figure E-110. Piping system configuration and location of instrumentation used for testing the 8-in. piping system with gate valve.


1.0

0.8

a

E

CD

< 0 .4

0

0 ~ to 20 30 40 50

Frequency (Hz)

Figure E-113. CPS average acceleration, gate valve body.

2 .0

E

U

C.

U

0 /

CD

0.2

0

0 10 20 30 40 50

Frequency (Hz)

Figure E-114. CPS average acceleration, gate valve actuator.

E64
0.

Ap

0.2

. to 2'0 30 40 to

Fr-equency JHz)

Figure E-I1IS. CPS average acceleration, butterfly valve body.

9.0.

T

LF

S

U

L

0

0 to 20 30 40 50

Frequency (Hz)

Figure E- 16. CPS average acceleration, butterfly valve actuator.

E65
Table E-1. Band-integrated X acceleration

Acceleration (G rms) [Bandwidth (Hz)]

Test T40.10 Test T40.30

Location

V. Oper.

Frame #1

0.31 [ 4.9-8.1 ]

0.10 [11.6-14.0]

0.17 [17.2-19.21

0.35 [19.4-22.11

0.22 [22.3-25.51

0.23 [26.2-28.6]

0.19 [28.9-31.1]

0.12 [32.0-37.2]

0.22 [39.244.71

Frame #2

0.48 [ 4.0-6.1 1

0.09 [ 7.6-11.0]

0.12 [14.0-16.7]

0.21 [18.0-22.61

0.28 [22.6-34.2]

0.12 [37.2-43.7]

Frame #1

0.23 [ 5.3-7.4 1

0.32 [18.3-20.91

0.14 [21.8-24.21

0.18 [25.0-30.0]

0.13 131.4-33.4]

0.25 [38.0-42.2]

0.09 [44.1-48.2]

1.34

Frame #2

0.40 [ 4.1-5.9 1

0.16 [ 6.4-7.9 1

0.16 [13.8-16.6]

0.29 [17.4-21.01

0.18 [22.3-24.4]

0.22 [27.3-32.2]

0.10 [41.0-44.41

1.91 1.30 1.51

V. Body 0.27 [ 4.9-6.9 ] 0.38 [ 3.9-6.1 ] 0.16 [ 5.3-7.1 ] 0.37 [ 4.0-5.9 1

0.12 [ 6.3-7.9]

0.09 [23.5-24.91 0.07 [19.1-20.71 0.06 [22.5-24.11

0.08 [26.4-27.9] 0.07 [25.4-27.5] 0.08 [25.0-30.0] 0.05 [24.6-26.1]

0.11 [29.3-31.1] 0.12 [28.1-32.8] 0.10 [27.3-30.91

0.09 [32.1-36.5] 0.09 [30.8-34.9] 0.08 [31.4-34.0]

0.10 140.5-44.9] 0.07 [39.3-43.5] 0.09 [38.0-42.1] 0.07 [41.0-43.8]

0.74 0.64 0.49 0.85

Sph. Tee 0.11 [ 6.4-7.7 ] 0.27 [ 4.6-5.9 ] 0.23 [ 5.6-7.5 ] 0.42 [ 4.6-6.1 ]

0.07 [ 9.8-11.1]

0.06 [15.3-16.7] 0.06 [15.5-17.0]

0.07 118.5-21.6] 0.05 [20.7-22.0] 0.05 [19.6-21.0]

0.09 [24.4-28.5] 0.07 [25.4-28.0] 0.09 [24.6-28.2] 0.04 [27.3-28.0]

0.04 [31.8-33.4] 0.06 [29.0-31.2]

0.31 0.58 0.37 0.52

Std. Tee 0.33 [ 4.9-8.1 ] 0.30 [ 4.1-5.9 ] 0.11 [ 5.3-7.1 1 0.35 [ 4.0-6.0]

0.04 [11.5-12.5] 0.07 [ 9.0-10.4] 0.04 [11.0-12.0] 0.07 [ 6.2-7.9]

0.07 [15.2-16.6] 0.11 [14.9-17.0]

0.09 [17.4-19.5]

0.08 [23.7-25.0] 0.13 [23.6-25.4]

0.06 [26.6-28.1] 0.07 [24.9-27.2] 0.08 [24.7-28.0]

0.07 [30.0-31.0]

0.07 [34.4-37.41 0.05 [32.2-35.1]

0.06 [37.6-41.6]

0.65 0.57 0.33 0.70

Top HDU 0.20 [ 5.6-8.0 ] 0.06 [ 4.6-5.6 ] 0.19 [ 5.8-8.0] 0.07 [ 4.6-6.4 ]

0.03 [14.8-16.2]

0.08 [18.4-21.2] 0.08 [18.3-21.2]

0.05 [26.0-27.5] 0.05 [24.6-27.4] 0.03 [25.6-27.11

0.06 [31.2-34.2] 0.06 [31.0-35.0]

0.39 0.06 0.38 0.13

E-66
Table E-2. Band-integrated Y acceleration

Acceleration (G rms) [Bandwidth (Hz)]

Test T40.10 Test T40.30

Location

V. Oper.

Frame #1

0.23 [ 4.9-7.6 1

0.06 [ 8.8-10.11

0.14 [17.4-18.9]

0.14 [19.8-22.0]

0.06 [23.0-24.6]

0.07 [33.1-36.51

0.08 [40.5-44.5]

0.78

Frame #2

0.03 1 4.9-5.6 1

Frame #1 Frame #2

0.05 [14.1-15.01

0.06 [18.3-19.6]

0.04 [23.8-24.6]

0.03 t33.5-34.4]

0.08 [12.5-14.41

0.07 [16.5-17.9]

0.20 [18.5-21.1]

0.07 [31.9-35.6]

0.09 [13.2-14.6]

0.10 [18.1-19.5]

0.06 120.0-21.6]

0.05 [32.0-33.61

0.21 0.42 0.30

V. Body 0.23 [ 5.5-7.4]

0.06 [ 8.8-10.0]

0.04 [14.0-14.91 0.08 [12.5-14.41 0.09 [13.3-14.61

0.13 [17.3-18.9] 0.06 [16.5-17.9] 0.08 [18.1-19.5]

0.12 [19.8-21.6] 0.06 [18.3,22.0] 0.17 [18.4-21.1] 0.05 [20.0-21.5]

0.06 [23.0-24.6] 0.05 [23.8-25.1]

0.05 [26.6-31.1]

0.05 [33.8-36.8] 0.05 [28.5-35.0] 0.05 [31.9-34.8] 0.04 [31.9-33.6]

0.05 [40.5-44.5]

0.75 0.20 0.36 0.26

Sph. Tee 0.07 [ 4.4-5.5 1 0.18 [ 5.0-5.9 ] 0.03 [ 5.0-5.6 ]

0.53 [ 5.6-7.5 ] 0.07 [ 6.0-7.2 ] 0.08 [ 5.8-7.6]

0.17 [ 8.0-10.1] 0.06 [10.0-11.5]

0.04 113.0-14.2] 0.06 [12.5-13.5] 0.05 [11.8-13.5]

0.08 [13.6-14.9] 0.05 [15.1-16.9]

0.07 [19.6-21.51

0.04 [31.3-33.2] 0.05 [31.4-34.2]

0.85 0.31 0.34 0.13

Std. Tee 0.11%[ 5.4-7.0] 0.04 [ 4.2-5.6] 0.07 [ 5.1-7.0 0.05 [ 4.1-5.4]

0.03 [ 9.1-10.1]

0.04 [13.5-14.6]

0.05 [19.0-20.9]

0.09 [23.0-24.8] 0.05 [22.6-24.1]

0.20 [26.2-29.51 0.08 [27.7-30.2] 0.21 [24.7-29.51 0.08 [27.1-29.51

0.08 [29.6-31.51 0.04 [30.4-31.3] 0.08 [30.0-34.31

0.07 [33.4-36.0] 0.06 [32.7-35.0]

0.07 [40.6-43.1] 0.05 [41.0-43.4]

0.62 0.19 0.39 0.35

Top HDU 0.06 [ 5.7-7.4 ] 0.03 [ 4.9-5.5] 0.06 [ 5.8-7.2 ] 0.02 [ 4.6-5.5]

0.03 [12.1-14.1] 0.04 [12.1-14.1]

0.03 118.8-21.0] 0.04 118.8-21.01

0.09 [24.4-29.4] 0.08 [24.0-29.2]

0.21 0.03 0.22 0.02

E-67
Table E-3. Band-integrated Z acceleration

Acceleration (G rms) [Bandwidth (Hz)]

Test T40.10 Test T40.30

Location

V. Oper.

Frame #1

0.49 [ 5.1-7.6 ]

0.26 [11.6-12.7]

0.10 113.1-14.4)

0.26 [19.4-22.0]

0.20 [23.4-24.9]

0.33 [26.3-29.1]

0.21 [29.2-31.61

0.55 [32.3-37.6]

0.24 [39.445.61

Frame #2

0.26 1 4.4-5.7 ]

0.13 [13.1-16.2]

0.17 [18.8-21.9]

0.12 [23.3-25.2]

0.14 [28.4-31.1]

0.27 [31.8-35.0]

0.15 [37.0-39.9]

0.16 [40.0-46.4]

Frame #1

0.29 [ 5.1-7.2 ]

0.07 [10.6-12.01

0.07 [12.5-14.2]

0.20 118.8-21.01

0.16 [25.0-29.4]

0.43 [31.6-36.41

0.33 [38.0-41.6]

Frame #2

0.20 [ 4.1-5.5 1

0.09 [ 6.4-7.6 ]

0.14 1 8.9-11.0]

0.10 [18.3-19.91

0.13 [22.3-24.2]

0.23 [24.6-29.0]

0.16 [30.0-32.1]

0.28 [32.3-34.1]

0.27 [35.0-40.01

0.16 [40.9-43.4]

0.10 [45.8-47.71

2.64 1.40 1.65 1.76

V. Body 0.14 [ 5.5-7.2 ] 0.11 [4.3-5.7 ] 0.16 [ 5.3-7.1 ] 0.12 [ 4.2-5.6 ]

0.09 [ 9.1-11.21 0.11 [ 8.9-11.0]

0.21 [11.6-14.4] 0.11 [13.1-16.4] 0.04 [12.5-14.11 0.13 [13.3-16.2]

0.05 [18.8-20.61 0.04 [18.9-19.5]

0.07 [30.3-31.2] 0.07 [28.4-32.6]

0.06 [33.8-37.1] 0.05 [30.6-35.6] 0.06 [31.9-33.5]

0.05 [40.4-43.9) 0.04 [39.3-41.61

0.53 0.38 0.34 0.46

Sph. Tee 0.07 [ 5.9-7.1 ] 0.08 [ 4.8-5.7 ] 0.17 [ 5.6-7.1] 0.15 [ 4.6-5.7 l

0.05 [10.9-11.7] 0.11 [ 9.5-11.1] 0.08 [10.0-11.0]

0.19 [11.9-14.5] 0.04 [12.6-14.0] 0.05 [12.8-14.0]

0.06 [15.4-16.6] 0.08 [15.1-17.0]

0.05 [19.4-21.6] 0.04 [20.8-22.0] 0.05 [18.5-21.5]

0.06 [25.8-28.5] 0.04 [26.4-28.0] 0.05 [24.6-28.6]

0.07 [29.0-31.1] 0.04 [27.5-29.5]

0.42 0.44 0.32 0.35

Std. Tee 0.42 [ 5.4-7.9 ] 0.19 [ 4.1-5.8 ] 0.20 [ 5.2-7.4 ] 0.14 [ 4.1-5.6 ]

0.14 [ 8.7-10.9]

0.08 [11.6-12.6] 0.05 [11.0-12.0] 0.08 [13.6-14.61

0.06 [16.2-17.91 0.11 [15.9-17.0]

0.09 [18.9-21.0] 0.11 [17.8-19.5]

0.16 [23.1-25.31 0.21 [23.6-25.4] 0.06 [22.7-23.9]

0.07 [26.6-27.9] 0.14 [25.0-28.0] 0.10 [24.6-27.0]

0.15 [30.0-31.6] 0.11 [28.1-30.3] 0.07 [28.1-29.6]

0.17 [33.7-37.4] 0.13 [32.2-35.0]

0.09 [37.7-41.2]

1.05 0.65 0.83 0.60

E-68
Table E-3. (continued)

Acceleration (G rms) [Bandwidth (Hz)]

Test T40.10 Test T40.30

Location Frame #1 Frame #2 Frame #1 Frame #2

Top HDU 0.63 [ 5.5-8.0 ] 0.23 [ 4.6-5.9] 0.64 [ 5.6-7.9] 0.24 [ 4.3-6.4 ]

0.06 [15.1-17.0] 0.06 [15.3-16.5]

0.14 [18.0-22.4] 0.05 [20.5-21.6] 0.14 [18.1-22.4] 0.04 [20.5-21.5]

0.13 [24.5-28.1] 0.13 [24.3-28.0]

0.90 0.34 0.91 0.34

Table E-4. CPS gate valve acceleration

Acceleration (G rms)

X Direction Y Direction Z Direction

Band

(Hz) Body Oper. Amp. Body Oper. Amp.

0-5 0.250 0.305 1.22 0.383 0.382 1.00

S.10 0.196 0.257 1.31 0.184 0.182 0.99

10-15 0.140 0.252 1.80 0.208 0.198 0.95

15-20 0.142 0.338 2.38 0.259 0.267 1.03

20-25 0.195 0.118 0.61 0.188 0.356 1.89

25-30 0.053 0.116 2.19 0.091 0.134 1.47

30-35 0.040 0.216 5.40 0.083 0.109 1.31

35-40 0.036 0.384 10.67 0.135 0.146 1.08

40-45 0.039 0.315 8.08 0.105 0.115 1.10

45-50 0.040 0.126 3.15 0.125 0.149 1.19

Body Oper. Amp.

0.260 0.276 1.06

0.368 0.385 1.05

0.312 0.398 1.28

0.169 0.398 2.36

0.206 1.040 5.05

0.071 0.276 3.89

0.102 0.211 2.07

0.116 0.165 1.42

0.196 0.105 0.54

0.179 0.088 0.49

Sum 1.131 2.427 2.15 1.761 2.038 1.16 1.979 3.342 1.69

E-69
Table E-5. CPS butterfly valve acceleration

Acceleration (G rms)

X Direction Y Direction Z Direction

Band

(Hz) Body Oper. Amp. Body Oper. Amp. Body Oper. Amp.

0-5 0.510 0.530

5.10 0.321 0.429

10-15 0.203 0.469

15-20 0.326 0.769

20-25 0.223 0.537

25-30 0.090 0.467

30-35 0.071 0.533

35-40 0.064 0.417

40-45 0.064 0.304

45-50 0.057 0.252

Sum 1.929 4.707

1.04 0.530 0.548 1.03 0.448 0.456 1.02

1.34 0.307 0.329 1.07 0.408 0.402 0.99

2.31 0.245 0.266 1.09 0.211 0.363 1.72

2.36 0.464 0.516 1.11 0.229 0.471 2.06

2.41 0.265 0.323 1.22 0.149 0.527 3.54

5.19 0.145 0.156 1.08 0.155 0.759 4.90

7.51 0.208 0.188 0.90 0.171 1.002 5.86

6.52 0.213 0.238 1.12 0.151 0.678 4.49

4.75 0.142 0.145 1.02 0.192 0.397 2.07

4.42 0.159 0.162 1.02 0.221 0.370 1.67

2.44 2.678 2.871 1.07 2.335 5.425 2.32

E-70
NRC FORMI 335 U.S. NUCLEAR REGULATORY COMMISSION 1. REPORT '7U'.BEA

(2- 91 IAssigned by NRC. Add Vol.. Sup,.. Ae,.. NFICM 1102. 8"d Addendumn Numbears. itany~l 3201.3202 BIBLIOGRAPHIC DATA SHEET NUREG/CR-4977

(See instructions on the reverse) EGG-2505

2. TITLE AND SUBTITLE Vol. 2

SHAG Test Series

Seismic Research on an Aged United States Gate Valve and on a 3 OATEREPORTPUBLISHED

Piping System in the Decommissioned Heissdampfreaktor (HDR): AugI 1989

Appendices 4. FIN OR GRANT NUMBER

l_____________________________________________________________________ A6322

S. AUTHOR(S) 6. TYPE OF REPORT

R. Steele, Jr., J.G.Arendts

Technical

7. PERIOD COVERED (lnclsisev Dates,

8. PERFORMING ORGANIZATION - NAME AND ADDRESS lIi NRC. provide Divsson. Ofi(ceor Region, U.S. Nuclearfiegulatorv Comnmission, and nailing address:id contractor,. provio,

name and mailing adudims)

Idaho National Engineering Laboratory

EG&G Idaho, Inc.

Idaho Falls, ID 83415

9. SPONSORING ORGANIZATION -NAME AND ADDRESS (If NRC. typee saewasabowe& if contractor. provide NRC Osvision, Office or Region, US Nuckar Regularory Commoision.

and maing addrerss

Division of Engineering

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, D.C. 20555

10. SUPPLEMENTARY NOTES

11. ABSTRA The Idaho National Engineering Laboratory (INEL) participated in an internationally

sponsored seismic research program conducted at a decommissioned experimental reactor

facility, the Heissdampfreabtor (HDR), located in the Fedgral Republic of Germany (PRG).

The research program included the study of the effects of excitation, produced during a

simulated seismic event, on (a) the operability and integrity of a naturally aged 8-in. motoroperated gate valve installed in the Versuchskreislauf (VKL), an existing piping system

in the HDR, (b) the dynamic response of the VKL and the operability of snubbers, and

(c) the dynamic responses of various piping support systems installed on the VKL. The

INEL work, sponsored by the U.S. Nuclear Regulatory Commission (USNRC), contributes

to earthquake investigations being conducted by the Kernforschungszentrum Karlsruhe

(KXK) and is part of the general HDR Safety Program performed in behalf of the FRG,

Federal Ministry for Research and Technology. This report presents the results of the

KfK-designated SHAG (Shakergebiude) test series; these are the first in situ experiments

involving an actual nuclear power plant and a full scale piping system under simulated

seismic loading. Volume I presents a summary of the tests and results, and Volume 11

contains appendices that present details and specifics of the tests and results of

Volume 1.

12. KEY WORDSIDESCR!PTORS IList words or phrases that wil assist reeachers inltocating ther spon. ) 13. AVAILABILITY STATEMENT

unlimited 14. SECURITY CLASSIFICATION

SHAG test series (Thi, Page)

seismic research program unclassified

aged gate valve {rtit Repr

piping system unclassified

Heissdampfreaktor (HDR) 15. NUMBER OF PAGES

simulated seismic loading

16. PRICE

NRCOM 335 (2408 -

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UNITED STATES

NUCLEAR REGULATORY COMMISSION

WASHINGTON, D.C. 20555

OFFICIAL BUSINESS

PENALTY FOR PRIVATE USE, $300

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