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NUREG/CR-5159

:'KEI No.'1559

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Prediction- of- -Check---Valve

Performance -and--- Degradation

in Nuclear Power. P.lant Systems

i .

Final Report

Septem ber 1987 - April 1988

.,.: j , . '.

Prepared by M.S. Kalsi, C.L. Horst, J.K. Wang

Kalsi Engineering, Inc.

.1 :1 I'll

Prepared for

U.S. Nuclear Regulatory Commission' . .

*.
NOTICE

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Government. Neither the United States Government nor any agency thereof, or any of their

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product or process disclosed in this report, or represents that its use by such third party' would

not infringe privately owned rights. -

. ~ ~ -__ ., . - -

NOTICE .

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NUREG/CR-5159

KEI No. 1559

RG, RV

Prediction of Check Valve

Performance and Degradation in Nuclear Power Plant Systems

Final Report September 1987 - April 1988

Manuscript Completed: April 1988

Date Published: May 1988

Prepared by

M.S. Kalsi, C.L. Horst, J.K. Wang

Kalsi Engineering, Inc.

745 Park Two Drive

Sugar Land, TX 77478

Prepared for

Division of Engineering Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555

NRC FIN D2042
ABSTRACr

Degradation and failure of swing check valves and resulting damage to plant

equipment has led to a need to develop a method to predict performance and degradation

of these valves in nuclear power plant systems. This Phase I investigation developed

methods which can be used to predict the stability of the check valve disk when piping

disturbances such as elbows, reducers, and generalized turbulence sources are present

upstream of the valve within 10 pipe diameters. Major findings include the flow

velocity required to achieve a full open, stable disk position, the magnitude of disk

motion developed with these upstream disturbances with flow velocities below full open

conditions, as well as disk natural frequency data which can be used to predict wear

and fatigue damage. Reducers were found to cause little or no performance

degradation. Elbow effects must be considered when located within 5 diameters of the

check valve, while severe turbulence sources have significant effect at distances to 10

diameters.

Clearway swing check designs were found to be particularly sensitive to

manufacturing tolerances and installation variables making them likely candidates

for premature failure. Reducing the disk full opening angle on these clearway

designs results in significant performance improvement.

iii
CONTENTS

Page

EXECUTIVE SUMMARY 1

INTRODUCTION 2

OBJECTIVES - ; 4

TEST PROGRAM -5

SCALING CONSIDERATIONS- 6

TEST SPECIMEN 7

INSTRUMENTATION 9

Design Features 9

TEST RESULTS 10

BASELINE TESTS 10

ELBOW TESTS 11

Disk Fluctuations for Elbow Up, Od.. 11

Disk Fluctuations for Elbow Down, Od 11

Disk Tapping Zones 12

Velocity Margin Over Vmin Required for Elbows. 12

REDUCER TESTS 13

Velocity Margin Over Vmin for Reducers 13

TURBULENCE TESTS 14

Multi-Hole Orifice Plates as a Turbulence Source 14

Disk Fluctuation Data, 3-Inch Valve 16

Disk Fluctuation Data, 6-Inch Valve 16

Summary of Maximum Disk Fluctuations for Elbows, Reducers,

and Turbulence-Sources - 17

DISK NATURAL FREQUENCY AND DAMPING TESTS 17

DISK-TO-STOP IMPACT LOAD TESTS 17

DISCUSSION 19

CLEARWAY SWING CHECK PERFORMANCE PROBLEMS 19

VELOCITY MARGIN REQUIRED FOR ELBOWS AND

REDUCERS - -

THEORETICAL PREDICTION OF TURBULENCE INDUCED

DISK MOTION 20

RESULTS AND CONCLUSIONS - 23

RECOMMENDATIONS 24

REFERENCES 25

V
I

I

LIST OF FIGURES

Page

1 Check Valve Configuration 27

2 Check Valve Test Setup 28

3 Kalsi Engineering, Inc. Flow Test Facility 29

4 Instrumented Valve Cross Section 30

5 Composite Plot of Disk Fluctuations for 6" Valve, Baseline Tests 31

6 Disk Projection Comparison, Clearway Swing Check Test Valve 32

7 Composite Plot of Disk Fluctuations for 3" Valve, Baseline Tests 33

8 Definition of Elbow Orientation 34

9A Composite Plot of Disk Fluctuation for 3" Valve at Various Disk

Stop Angles 35

9B Composite Plot of Disk Fluctuations for 6" Valve at Various Disk

Stop Angles 36

10A Composite Plot of Disk Fluctuations for 3" Valve at Various Disk

Stop Angles So

10B Composite Plot of Disk Fluctuations for 6" Valve at Various Disk

Stop Angles 38

11 Maximum Disk Fluctuation vs. Distance to Upstream

Disturbance 39

12 Elbow and Reducer Test Results for 3" Swing Check Valve 40

13 Elbow and Reducer Test Results for 6" Swing Check Valve 41

14 Composite Plot of Disk Fluctuations for 6" Valve with 8" x 6"

Reducer at Od 42

15 3 and 6 Inch Perforated Plate Holders 43

16 Disk Motion of 3" Valve Due to Turbulence: Perforated Plates at

1.5d Upstream, Stop Angle at 73° 44

17 Disk Motion of 3" Valve Due to Turbulence: Perforated Plates at

2.5d Upstream, Stop Angle at 730 45

18 Disk Motion of 3" Valve Due to Turbulence: Perforated Plates at

4.5d Upstream, Stop Angle at 730 46

19 Disk Motion of 3" Valve Due to Turbulence: Perforated Plates at

10.5d Upstream, Stop Angle at 73° 47

vi
List of Figures (continued)

Page

20 Maximum Disk Motion of 3" Valve Due to Turbulence: Perforated

Plates at 1.5 to 10.5d Upstream, Stop Angle at 730 48

21 Disk Motion of 3" Valve Due to Turbulence: Perforated Plates at

2.5d Upstream, Stop Angle at 53° 49

22 Disk Motion of 3" Valve Due to Turbulence: Perforated Plates at

2.5d Upstream, Stop Angle at 630 50

23 Disk Motion of 6" Valve Due to Turbulence: Perforated Plates at

1.5d Upstream, Stop Angle at 70° 51

24 Disk Motion of 6" Valve Due to Turbulence: Perforated Plates at

2.5d Upstream, Stop Angle at 70° 52

25 Disk Motion of 6" Valve Due to Turbulence: Perforated Plates at

4.5d Upstream, Stop Angle at 70° 53

26 Disk Motion of 6" Valve Due to Turbulence: Perforated Plates at

10.5d Upstream, Stop Angle at 70° 54

27 Maximum Disk Motion of 6" Valve Due to Turbulence: Perforated

Plates at 1.5 to 10.5d Upstream, Stop Angle at 700 55

28 Disk Motion of 6" Valve Due to Turbulence: Perforated Plates at

2.5d Upstream, Stop Angle at 50° 56

29 Disk Motion of 6" Valve Due to Turbulence: Perforated Plates at

2.5d Upstream, Stop Angle at 60" 57

30 Maximum Disk Fluctuation vs. Distance to Upstream

Disturbance 58

31 Disk Response to Disturbance, 6" Valve Natural Frequency

Tests 59

32 Disk Natural Frequency 60

33 Disk-to-Stop Contact Loads, 3" Valve 61

vii
ACKNOWLEDGMENTS

We are greatful for the funding provided by the Small Business Innovation Research

program of the U. S. Nuclear Regulatory Commission which made this research

possible. Special thanks are due William S. Farmer of the NRC for his guidance and

encouragement throughout the project.

Individuals of the Kalsi Engineering, Inc. staff who should be recognized for their

important contribution toward successfully completing the project are Lannie Dietle

for his meticulous work in conducting thousands of tests, Daniel Alvarez for the

development of the load cell concept employed in these tests, and Bobbie Lambert for

patiently working through many revisions and typing this manuscript in a beautiful

form.

ix
EXECUTIVE SUMMARY

The purpose of this Phase I investigation'was to determine'th~e feasibility of a technical

approach in predicting and quantifying the stability or instability of swing check

valve diskshin piping systems which have significant flow disturbances -within 10 pipe

diariieters' 'upstream of the valve. Specifically, -the'effects of elbows; reducers; and

turbulenie_ sources have been investigated. A' series 'of over 2,000 flow tests were

performed to ~ifeasure' the-performance of 3-inch and 6-inch:swing check valves

subjected to these upstream disturbances.

The following is a summary of major findings from the test program:

1. Disk' stability for swing check valves can be predicted if the valve disk angle,

the type and proximity of the'upstrearii disturbance, and the flow velocity-

'through the valve are known. ' -.

2. To account for the presence of an upstream flow disturbance, some velocity

margin over the -minimum velocity needed to suppress disk fluctuations (and

fully open it against the backstop) in an undisturbed, uniform:pipe flow is

required. The necessary margins for various flow disturbances were

developed and are included in this report.

3. Clearway swing check valves are most susceptible to the effects of.upstream

- disturbances such -as 'elbows and'turbulence producing devices. (A clearway

swing check is one in which the disk, at its full-open position is lifted up into the

bound area and nearly out of the path of the flow stream through the valve body).

Stich' valves may never achieve disk stability under certain flow conditions.:

4. One of the important findings from this research was that upstream reducers

have no measurable effect on either velocity required over baseline to fully open

the valve-or on disk stability.' This holds even when the reducer is located

immediately adjacent to the check valve. (A reducer is defined as an element

which reduces the pipe size as one approaches the valve in the direction of flow.)

5. Some specific-results from Phase I tests are that at proximities of 3 to 5 pipe

diameters upstream, elbows require an increase in flow velocity over baseline

of 10 or 15 percent to fully open the valve disk. Proximities of 0 to 1 diameter

will require up to 50 percent higher flow velocity over baseline to fully open the

disk, except for clearway designs which will require velocities more than 100

percenthigherthanbaseline. ' ,: ' - -:

6. The amplitude of disk motion which occurs before fully seating the disk

increases as the'flow disturbance is brought closer to-the valve,. The maximum

- disk fluctuations in the case of a severe turbulence source can reach as'high as

16 degrees' and for elbows up to 9 degrees. The reducers have negligible effect

on disk fluctuations. -

- 7. 'Suitable nondimensional parameters ,were chosen in the test program to check

the validity of the results when scaling from one size to another. 'The 3-inch

and 6-inch test results show that reasonable predictions can be made while

scaling to other sizes. "

1
INTRODUCTION

Accelerated degradation of check valve internals can lead to their failure to perform,

leading to serious safety consequences or causing extensive damage to other plant,

components and systems. The primary cause of premature degradation is continuous.

fluctuation of the disk when it is not firmly held against the backstop by the fluid forces

pushing on it. The severity of the degradation depends upon the type and severity of

disturbances in the fluid system that are present upstream of the valve which disturb the

uniform flow and create velocity skew or fluctuations.

Disk fluctuation can be eliminated by,. ensuring that the valve is sized to create;

sufficient flow velocity through it to fully open the disk against its backstop. However,

many valve manufacturers lack accurate data to make minimum velocity (Vmin)

recommendations concerning their valves. A generalized Vmin formula has recently

been developed (References 5, 23) for swing check valves which has been shown to

correlate well with experimental results covering a wide range of sizes and pressure

ratings. It should be pointed out, however, that this formula is applicable only to

installations in which fully developed pipe flow, unaffected by upstream disturbance,

is entering the valve.

In order to predict the performance of check valves in practical installations at nuclear

power plants, additional data are needed to properly account for the effect of flow

disturbances that are, commonly present in close proximity (less than 10 pipe

diameters) upstream of the check valve. The minimum velocity, Vmnin, for an ideal

pipe flow may not be sufficient to hold the disk fully open under such conditions. One of

the objectives of this research is to develop techniques which can be used to predict

whether disk stability can be achieved in the presence of such upstream disturbances by

increasing the velocity above Vmin and to develop the velocity margins necessary to

suppress disk fluctuations.

Another important objective of this effort is to develop techniques to quantify the

magnitude and severity of disk fluctuations under various flow conditions in the

presence of these upstream disturbances when the velocity through the valve is

insufficient to hold it fully open. These techniques will aid in predicting (1) whether a

swing check valve installed with a known upstream disturbance at a defined

proximity will suffer from accelerated degradation, and (2) the severity of disk motion

which causes that degradation of the internals. Eventually, the results from the

successful completion of the research effort envisioned under Phase I and II should be

useful in predicting suitable maintenance intervals during which the degradation is

within acceptable limits. Within these limits the check valve should continue to

perform without compromising the safety or reliability of its operation.

This report summarizes the results of the Phase I effort in which the performance of

swing check valves with certain selected upstream flow disturbances was

systematically evaluated. In a broad sense the types of flow disturbances that can be

found upstream of the check valves fall into two distinct categories:

1. Piping elements such as elbows, reducers, and tees which are quite specific in

their geometry,

2. Turbulence sources which includes a large variety of devices capable of

creating velocity fluctuations in the flow stream. This includes control valves

2
of various types -'eg., single ported globe valve,'doible ported globe valve,

multiple hole cage guided trim valves, butterfly valves - pumps, and orifices,

etc. ; - - .

From the first category of flow disturbances, elbows and reducers were selected for

Phase I testing since these are the most commonly used piping elements upstream of

check valves. Phase I testing also included the investigation of an upstream

turbulence source. In order to generically characterize the effect of the various types of

turbulence sources described above, orifice plates with multiple holes were used. By

varying the size and number of holes, the two important characteristics of turbulence,

e.g., eddy size (or scale of turbulence) and the intensity of turbulence-could be

controlled independently. Multiple-hole orifice plates have the additional advantage

of.creating uniform .turbulence across the pipe cross-section, which provides data of

more general utility than some specific throttling devices, e.g., butterfly valves that

have been used in the past (References 5, 24), which introduce a velocity skew along

with turbulence. Extensive testing was done using several perforated -plates to develop

upper bounds of disk fluctuation under the most severe turbulent conditions.- These

bounding disk fluctuation results can be used to make conservative predictions about

check valve degradation. Additionally, some interesting trends were-found which

show how the size of holes in the orifice plates affects the severity of disk fluctuations.

It should be pointed out that until recently there had been no reported results of research

directed at systematic evaluation of upstream disturbances on check valve

performance, even though some unique configurations-have been investigated

(References 13, 24,25). Reference 5 reports the results of the first such effort towards the

systematic evaluation of upstream disturbances completed by the investigators in the

present research. 'That work included investigation of elbows and a throttled butterfly

valve as a turbulence source. However, the velocity increments used in those tests were

relatively coarse and they did not permit an accurate determination of the velocity at

which the disk begins to impact the backstop and when it finally reaches its fully open

position. Secondly, the butterfly valve introduces a velocity skew along with intended

turbulence which prevented any useful velocity margin data over Vmin to be extracted

from these tests. The Phase I tests were designed to specifically overcome these

limitations and to develop more general, upper bound data for turbulence as well as

reducers.

The tests were performed on two different sizes of swing check valves (3- and 6-inch)

with an adjustable stop to simulate the disk angle variations found in valves made by

different manufacturers. An important consideration in this Phase I effort was to use

methods which can allow the results to be scaled and applied to other sizes. All together,

over 2,000 tests explored the effects of elbows, reducers, and turbulence sources

upstream of the check valve.

Phase I effort also included a thorough review of the current state of the art in

fluid/structure interaction techniques, and their capabilities and limitation to predict

disk response to upstream turbulence if the turbulence characteristics are known.

This report summarizes the results of the experimental and theoretical investigation

completed under Phase I.

3
OBJECTIVES

The objective of the Phase I effort was to determine the feasibility of a technical

approach in predicting and quantifying the stability or instability of the swing check

valve disks in practical piping systems which have a significant'flow disturbance

within 10 pipe diameters upstream of the valve.

More specifically, the objectives were to

1. Develop a general approach which can be used to modify the Vmin formula to

account for the orientation and' oroximity of an upstream elbow;

2.' Develop a general approach to account for the effect of an upstream reducer on

the Vmin formula;

3. Develop'techniques for predicting the amplitudi'and frequency of check valve

disk motion in' the presence of an upstream turbulence source. 'These results

can be used-in a' Phase II program to develop 'quafntitative wear and fatigue

prediction techniques;-

4. Establish velocity margins above Vmin necessary to fully open the disk and

suppress disk oscillations' due to' 1, 2, and 3 above.'

A test program wa's'devised to me'et these objectives. The tests were performed'on 3-inch

and 6-inch swing check valves supplied by MCC Pacific'Valve Company. The valves

are similar to that shown in Figure 1. The tests 'were performed at Kalsi Engineering,

Inc.'s flow test"'facility in Sugar Land, Texadsi(Figures 2 and 3). Water at outdoor

ambient temperature was used as the'flow media at velocities between 1.5 and 25 feet per

second. 'All together, over 2,000 tests explored the' effects of elbows, reducers, and

turbulence upstream of the check valve.

4
TEST PROGRAM

Phase I testing was divided into six major areas as follows:

1. Baseline Tests: To e'stablish check valve performance with uniform, fully

developed flow unaffected'by upstream'disturbances. '

2. Elbo'w Tests: To'determine the effect of elbow orientation and proximity on

check valve performance. Elbow up and down orientations were tested at

upstream distances from 0 to 5 pipe diameters. Based upon earlier test results,

the effects 'of elbows b'eyond 5 pipe 'diameters are known to be insignificant.

3. Reducer Tests: Similar in nature to the elbow tests, a pipe reducer was installed

upstream at distances from 0 to 5 diameters. A reducer is defined as an

element that decreases the pipe size when moving in the direction of flow.

4. Turbulence Tests: Turbulence was developed by flowing through a perforated

plate upstream of the check valve. A selection of plates was -available with

different numbers and sizes of holes. Geometric scaling of plate geometry was

mainrtain'ed between both valves. The turbulence source was located from 1.5 to

10.5 diameters upstream of the check valve:

5. Disk Natural Frequency and Damping Tests: A series 'of tests were run to

measure the disk/hinge arm natural frequency and damping under actual

'flowing conditions. 'The data from these tests will be be used in Phase II to

develop predictive techniques to quantify wear and fatigue of the hinge pin and

disk stud connection, the two known weakest areas in swing check valves.

6. Tapping Force Tests: To quantify the magnitude and duration of disk impacts

with the backstop, a load cell was inco'rporated in''an' adjustable backstop

mechanism. Transient disk contact with the backstop' could then be recorded as

well as the steady-state load when the disk is pegged in the fully open condition.

.~~ . . . . .

5
I

SCALING CONSIDERATIONS

One of the important considerations in Phase I work was to properly take the pertinent

nondimensional scaling parameters into account. Two sizes of valve, 3-inch and 6-

inch, were tested so that data can be used to check the validity of how the results scale up

between the two sizes using the nondimensional parameters chosen.

The following table lists the dimensional and nondimensional parameters that were

used in the tests:

D = Nominal size of the valve; diameter of the upstream pipe

L = Distance between the valve end and upstream disturbance

L/D = Nondimensional distance between the valve end and upstream

disturbance; also referred to as proximity of the disturbance

AO = Disk fluctuation angle

9 = Total fluid impingement angle against the disk (see Figure 1)

= [90° - (the angle swept by disk from fully closed to fully open position)]

V = Flow velocity through the valve, based on valve inlet area

V ' = Velocity required to fully open the valve with an upstream disturbance

Vniin = Minimum flow velocity required to fully open the disk against the

backstop without any upstream disturbance

V'(Vmin = Nondimensional velocity ratio used to express velocity margins

needed over Vmin to stabilize the disk in the presence of an upstream

disturbance

Cup = (Same definition as ratio V'/Vmin) A nondimensional velocity

margin factor to account for upstream disturbances

Standard short-radius 90-degree elbows and concentric reducers manufactured by

Victaulic Company were used in these tests. They were dimensionally inspected to

determine what, if any, deviations exist in the geometrical scaling between the 3- and

6-inch sizes. It was found that the ratio of mean elbow radius and the inside diameter

of the elbow varied by approximately ± 10 percent around the mean value of 1.075. The

angle of the concentric reducers was also found to be consistent: between 14 and 15

degrees for both reducers, which equates to a total included angle of about 30 degrees.

Geometrical similarity was also maintained in the design of various multiple-hole

orifice plates used in tests on the 3-inch and 6-inch valves. The size and number of

holes used in these plates is shown in Table 3. The hole size governs the size of eddies

emitted by an orifice plate. The ratio between the eddy size and physical dimensions of

the check valve disk is an important nondimensional scaling parameter for

fluid/structure interaction caused by turbulence.

6
TESrTSPECMIENS K

The valves used throughout the test program were 3-inch, 300-pound and six-inch, 600-

pound clearway swing check valves. Dimensional and physical data are presented for

both valves in tables 1A and 1B. Both valves were modified to incorporate an LVDT

(Linear Variable Differential Transformer) type of displacement transducer for

measurement of disk/hinge arm motion and an adjustable backstop assembly

(Figure 4). The adjustable backstop takes the place of the fixed backstop which is

normally cast integrally into the bonnet in these valves. The adjustment mechanism

allows the stop position to be varied, thus simulating different full-open disk positions

covering the range of variations found in valves made by different manufacturers. In

addition, the backstop assembly incorporates a load cell which provides direct

measurement of the loads developed whenever the disk contacts the stop.

7
I

Table 1A

VALVE DATA FOR 3-INCH VALVE

Valve Type 3", 300# Swing Disk Check

Manufacturer MCC Pacific Valve Company

Weight and Dimensional Data.

Disk Weight (ind. nut) 1.94 lbs.

Hinge Arm Weight 0.72 lbs.

Hinge Pin Diameter 0.375"

Disk O.D. 3.90"

Seat Bore Diameter 3.0"

Seat Tilt From Vertical 3°

Full Open Angle 73°

Table 1B

VALVE DATA FOR 6-INCH VALVE

Valve Type 6", 300# Swing Disk Check

Manufacturer MCC Pacific Valve Company

Weight and Dimensional Data

Disk Weight (incl. nut) 8.94 lbs.

Hinge Arm Weight 3.38 Ibs.

Hinge Pin Diameter 0.500"

Disk O.D. 6.90"

Seat Bore Diameter 6.0"

Seat Tilt From Vertical 30

Full Open Angle 700

8
INSTRUMENTATION

Design Features ..*.* . -

The load cell apparatus.incorporates a spring to preload the stop rod against the load

cell, thus.taking up-any clearance that might exist. Thisfeature prevents unwanted

load amplification that can be caused by impact of the stop rod against the load cell if

any clearance is present between them. A static seal was used to isolate the load cell

from the valve interior. -This results in a fictitious disk-to-stop load in response to

valve internal pressure. To. minimize this load, the pressure effective area was made

as small as practical resulting in a pressure induced load of 2 pounds at 75 psig.

Disk motion measurement was accomplished by. attaching the LVDT core directly to

the hinge arm,-with a stiff plastic tube capable .of transmitting push/pull motion while

retaining the ability to flex slightly. to accommodate kinematic restraints of the

arrangement. This connecting link transformed the rotary motion of the hinge arm

into linear displacement of the LVDT core. Calibration of the apparatus against disk

angle proved the mechanism to be nearly linear with a worst case error of ± 0.5

degrees., Under certain flow conditions turbulence-in the valve can buffet the

connecting link causing displacement of the LVDT core, not associated with disk

motion. 'This phenomenon was eliminated by placing a protective aluminum tube over

the connecting link to shield it from the turbulence.--:

Flow rates were calculated by measuring the pressure drop across an orifice plate. The

entire orifice meter run is constructed in accordance with the requirements of

ANSI/API 2530. Three different orifice sizes were used to provide flow accuracy of

+ 0.5 percent over a range of 1.3 to 27 feet per~second (water). This flow range exceeds

that required for the present test program. .

The following parameters were measured and recorded during each of the tests

performed: disk.motion, backstop load, flow orifice AP (for flow velocity), and pressure

upstream of the check valve. In addition, during the turbulence tests, a second pressure

measurement was added to monitor pressure drop across the perforated plate. Each of

the analog signals was acquired through a 12-bit analog-to-digital converter, then

stored on disk:' A test log was also printed out documenting important test parameters

and summary test data.

-~ ~ F-.''

* ' .- 

9
I

ESTR rSUuTS

Identical tests were run on both 3-inch and 6-inch swing check valves. The flow

velocity range spanned 1.5 to 25 feet per second, using water. Flow velocity increments

were varied to adequately characterize disk motion through the velocity range.

Therefore, in the velocity range where the disk begins to tap against the stop, relatively

fine velocity increments were taken, typically 0.1 to 0.2 fps. Once the disk was pegged,

the velocity increments were increased to 2 fps. This methodology' allowed a more

precise determination of the velocity at which tapping begins and the velocity at which

the disk becomes pegged against the stop. On average, 20 separate tests were needed to

cover the 1.5 to 25 fps range at one disk stop angle.

Data were taken at stop angles of 53 degrees, 63 degrees, and 73 degrees for the 3-inch

valve and 50 degrees, 60 degrees, and 70 degrees for the 6-inch valve. The 73- and 70-

degree positions represent the as-manufactured full-open angle for'the 3-inch and 6-

inch valves respectively.

BASELINE TESTS

Baseline tests establish performance data when' the valve is presented with, fully

developed pipe flow. To provide these flow conditions, 20 diameters of straight pipe

incorporating a section of flow straightening vanes was installed upstream. This

upstream configuration exceeds ANSI/API Standard 2530 requirements for a flow

metering piping run. Results of the baseline tests for the 6-inch valve are presented in

Figure 5, which shows peak-to-peak disk motion reported as ± 3cs (standard deviations)

from the mean disk position plotted against flow velocity. Data for all three stop

positions are shown. Inspection of Figure 5 shows that the highest levels of disk motion

reaching 1.5 to 2 degrees, occur between 4 and 6 fps. We can conclude that disk

fluctuation values of about 2 degrees (peak to peak) are the lowest that' can be expected,

even with the best upstream conditions we can create. This minimum amount of disk

motion must be expected whenever Vznl ,has not been reached. Figure 5 can be used to'

point out other important aspects'of valve behavior. First; disk miotion occurs over a

much wider range of velocities for the 70-degree'stop angle than at 50 degrees and 60

degrees. This can be seen by measuring the velocity range over which disk

fluctuations exceed 1 degree*. Table 2 shows the velocity range and span in fps for the

three stop angles. This range is also indicated in Figure 5.

Table 2

6-INCH VALVE DISK MOTION

Stop Velocity Velocity

Angle Range Span, fps

50° 3.8 - 4.9 1.1

60° 3.8 - 6.5 2.7

70° 3.8 - 13.8 10.0

0 Because of mechanical clearances in the LVDT assembly are variable, we have

established a 10 fluctuation level (± 0.50) as the threshold below which we consider

the disk actually seated against the backstop.

10
The wide velocity span oveii'which disk'motion occurs alt7ehe70-degree stop angle' is

inherent in the' design of a clearway swing' check'valve: As' was reported in Reference

5, this is due in large part to the shieldin'g effect'of the valve body/bonnet area when the

valve is opened to the clearway position. Figure 6 shows in cross-section how the

clearway design causes the disk to be lifted very nearly out of the normal flow stream

through the valve. Measurements of the twovalves show'thatlthe 6-inch disk protrudes

0.8 inches into the flow stream while the 3Winch disk has no'protrusidn. Data from the

3-inchbvalve baseline tests are presented'in Figure 7. Again, the maxir'mum' fluctuation

levelsiare'in the 2-degree range and the range over'which disk motion occurs is much'-

wider at the 73-degree stop angle than at 63 degrees or 53 degrees.:

ELBOW TESTS

Tests were run with elbows at proximities of 0, 1, 3, and 5 diameters upstream of the

check valves and oriented both up and down. Tests were not-run at distances beyond 5d

because results reported in Reference 5 showed that at distances of 5 diameters and,

beyond, the effects of elbows become very small, approaching the performance achieved

with baseline conditions. 'To avoid confusion, Figure 8 shows sketches representing

the definitions used in this text for elbow up and elbow-down orientations.

A detailed review and comparison will be made of the elbow test results at Od for the 3-

and 6-inch valves. Remaining data will be covered later in this section.

Disk Fluctuations for Elbow Up, Od .

Figures 9A and 9B show the results of disk fluctuation vs. flow velocity tests for the 3-

and'6-inch valves. Maximum disk motion for both valves occurred when the stop was

positioned at the clearway positions of 73- and 70 degrees for 3- and 6-inch valves

respectively. Values of approximately 9- and 6 degrees'were recorded for the 3- and 6-

inch valves respectively. The levels 'of fluctuation' at the 600/630-and 500/530 stop

positions were lower than those reached at the clearway position. As seen in the

baseline data, the velocity range over which disk motion occurs is much wider at

70°/730 than at the lower angles. Most interesting, however, is the fact that for the 3-inch

valve a stable, fully open disk position 'was never achieved at the 73-degree clearway.

stop position even at velocities of 24 to 25 fps. This phenomenon did not occur during

elbow tests reported in Reference 5 which showed the 3-inch valve reaching a fully open

condition at 18 fps. This demonstrates the extreme sensitivity of the clearway design to

minor changes in flow conditions or actual pbsitioningiof the disk near the-nominal

73-degree position.- The 3-inch valve, with no disk projection into the flow stream, is

very sensitive to physical changes in the piping setup or conditions inside the valve.

The 6-inch valve, with some disk projection into-the flow stream, is less sensitive to

these changes. ' ,-. It -. , : , .

DiskFluctuations forElbowDown, Od . ,

Figures 1OA and 10B show the elbow down test results for the 3- and 6-inch valves. The

same conclusion's are reached as for the elbow up case, except that the maximum disk

fluctuations'are not as high. This is due to the fact that the highest levels of turbulence:

exist along the inner radius of an elbow (Reference 6); thus, the elbow down orientation

positions this more turbulent portion of the flow further from the valve disk than the

elbow up orientation. Once again the 3-inch valve exhibits the inability to reach a

stable, fully open position.

Figure 11.compares the maximum disk fluctuation for both valves at each- locati6n'

tested. To normalize the data, both horizontal axes haveibeen'offset to account for what

11
I

amounts to a built-in length of straight pipe between the flange face and the disk region

inside the valve. This distance measures 2d and 1.5d for the 3- and 6-inch valve

respectively. The highest levels of disk motion occur at elbow proximities of Od and

generally decrease as the elbow is moved further from the valve.

Disk Tapping Zones .-.

Another important aspect of check valve behavior is disk tapping. Disk tapping occurs

when the flow velocity is not quite sufficient to firmly seat the disk against the stop, so

that fluctuations in disk motion cause the disk to repeatedly tap. against the stop.

Tapping stops when flow velocity, is raised to the point where the disk becomes firmly

seated or pegged. By carefully determining the flow velocity at which tapping begins

and the velocity at which the disk becomes pegged, the tapping "zone" can be defined in

feet per second.

By examining the elbow test data, the following observations about disk tapping can be

made:

1. The onset of tapping begins at lower flow velocities when-the elbow is oriented up

than when oriented down. This results in wider tapping zones for elbow up.

installations compared to elbow down. This is particularly evident at Od and id

and less so at 3d and 5d.-

2. The tapping zone is widest with the stop angle set at 700/730, typically spans a

magnitude of 2 to 4 fps.

3. The tapping zone is very narrow when the stop position is decreased to 600/630 and

500/530.. At these stop positions, the onset of tapping very nearly coincides with the

velocity at which maximum disk motion occurs. A very small increment in flow

velocity is then required to peg the disk; typically less than 1 fps at 600/63° and less

than 0.5 fps at 500/53°.

Disk tapping plays an important role in fatigue and impact wear of the disk stud and

wear of the hinge pin. Once again, the clearway design is seen to be most susceptible to

damage because of the wider tapping zone and the fact that the zone lies at velocities in

the. 12 to 15 fps range commonly encountered in check valve installations.

A predictive wear and fatigue model should be developed by combining this data with

wear data for specific material combinations gathered as part of a Phase II program.

Velocity Margin Over Vmii Required for Elbows

Figures 12 and 13 show the results for 3- and 6-inch valves for the velocity required for

full disk opening at each stop angle plotted against the nondimensional distance to the

elbow, L/D. For both valves, it can be seen that at distances of Od to id, the velocity

required to fully open the disk can be as much as 75 percent higher than baseline in the

clearway position, and, as noted earlier, may never actually become stable in the case

of the 3-inch valve at Od." At 3d and beyond, the velocity margin required drops to 10

percent over baseline conditions.

The onset of tapping was defined as four disk-to-stop contacts in ten seconds.

12
At the lower disk openinzgtaingles, the velocity margins required decrease'markedly.

At 600/630, a 35 percent margin is required at Od and id and io percent elsewhere. At'

50°/530, a 15 percent margin is required at'Od'and id and 10 percent elsewhere.

REDUCERTESTS -.

A common piping system element found 'upstream of check valves is the pipe reducer.

Marny reducer geometries'exist with'various ratios of inlet-to-outlet 'size and overall

length, as well as being fabricated or forged, concentric or eccentric; The sizes chosen''.

for 'these tests were 4 'x 3 for the 3-inchbvalve ind 8 x 6 for the 6-inch valve. This'

selection provides a reduction of one nominal pipe size in both cases, and maintaiins'

geometric similarity in both cases by providing a reduction ratio of 0.75. The reducers

were'3;5 and 5 inches long in their ove'rall -length 'for the 3-inch and -6-inchxlves

respectively. The actual 'slope of the iwall in'the 'reduction section' wVas '6lose to"15

degrees (30 degrees total included angle) for both reducess. ; ' ' '

Tests were run with the reducer placed at'the same distances from'the valve 'as the

elbows: 0, 1 3, and 5 diameters. Figure 14'shows disk fluictuation'data compiled from

tests of the 6-inch valve with reducer at'Od." Interestingly, it looks very much like the'

baseline test data of figures 5 and 7 with maximum disk fluctuations in' the 2-'to 2.5- -

degree range occurring between 4 and 6 fps. Velocities at which the disk reaches the

fully open position are very close to baseline'levels. Data from the other proxiriiities (1,'

3 and 5d)1as well~a]' the 3-inch tests sh6w':veriy similar performance. 'Unlike-elbows,'

there is no characteristic increase 'in'disk'ifistability as the'ifeducer is'br'ought clos'er to

the valve?'" This is easily'seen by'reexarmining figures 11'- 13, which shbo`w summary

data for the 'rducer tests.' In Figure 11, it can be seen that maximnium- disk fluctuations

for bih valves at a distance'are well within'the levels of baseline fluctuations shown

by the dashed lines. ' ' - ''' - : ' ' ' '

Figures 12 and 13 show the velocity required 'to fully open the disk at the different stop

angles. In each case, for both valves and at all stop angles and distances, the required

velocity is within+± 5 percent'of baseline.- These results are-in agreement with expected

behavior based on theoretical formulations in Reference .7 where it is shown that, in

fact, a reducer will tend to derease the level of axial velocity variations by 'an 'amount

approximately equal to X2 where' = ratio of the contraction. For the reduce'rs tested

here, the axial velocity variation could'be expected tobe reduced by-X2 = (0.75)2, or 56

percent of the-original velocity'fluctuations. IThis example applies'.to' an ideal.case,

neglecting viscous effects: and compressibility and allowing no- geometric

discontinuities which will otherwise disturb the flow field. This also assumes no

contribution of radial velocity'variations to disk instability. Although 'reducion of this

level will not be realized in a real application, it does show that an upstream reducer

will tend to reduce axial flow fluctuations or turbulence intensity and resultant disk

motion as observed in our-tests.:

Velocity Margin over Vmin for Reducers ii - ' ' __. ' -:

To summarize the reducer test results, it can be stated that the reducers do not have any

detrimental effect on the stability of the swing check valve disk. There is essentially

no vel6city'marginir'equired 'above Vmin to'suppress disk 'oscillations due to this type'of

flow' disturbance. i In' fact ethe reducershave' 'a beneficial effe'ct'-of-decreasing the

turbulence intensity that isresponsible for' causing disk'fluctuations> ' '

.. .. ' * .. ' ' , i,: : ,', ' . . . : ' . '; ' -- . .. . .

- . ? .--.r -.-;~~~ ' ' '

13
I

TURBULENCE TESTS

Test results 'reported in Reference 5' concluded that using a partially closed butterfly

valve as a throttling device generated high levels of turbulence. However, flow over the

canted disk of the butterfly valve resulted in a highly skewed velocity'profile entering

the check valve. The superposition of velocity skew and turbulence masked the

individual effects of the- two independent phenomena, making it 'impossible todetermine the velocity margin over Vmin due to turbulence effects alone. To provide

controllable levels of turbulence without velocity skew for this series of tests, a special

set of perforated plates were constructed which could be placed at a variety of upstream

locations.

When considering turbulence sources which might affect.check valve performance,

the search quickly becomes overwhelming. Many piping system elements are sources

of turbulence. Control valves of all types, pumps, pipe fittings such as tees and elbows,

all contribute to turbulence which will affect check valve performance. Characterizing

the effects of many different variations of control valves, pumps, or other throttling

devices on an individual basis would be an impossibly large task. Instead, we took the

approach of using a relatively simple but generic source of turbulence which could be

easily modified or scaled..

Multi-Hole Orifice Plate as a Turbulence Source

The device settled upon for generating turbulence is a plate of stainless steel with a

pattern of drilled holes captured in a holding fixture fitted with pipe stubs. This

allowed for easy axial positioning of the turbulence source with respect to the check

valve as well as interchangeability of the perforated plates. Figure 15 shows the plate

holder and an assortment of plates used for the 3- and 6-inch valve tests, respectively.

Data regarding perforated plate geometry are presented in Table 3.

Table 3

PERFORATED PLATE DATA

3-Inch Valve 6-Inch Valve

No. of Holes x No. of Holes x

Hole Diameter Area Ratio*, Hole Diameter Area Ratio*,

(inches) a (inches) a

36 x 3/16 .13 36 x 3/8 .13

61 x 3116 .23 67 x 3/8 .26

3 x 7/8 .24 3 x 13/4 .24

4 x 7/8 .33 4 x 13/4 .33

* Area Ratio = Total Area of Holes/Pipe Flow Area

One important characteristic of turbulence which we wanted to simulate is the size of

the turbulent eddies developed by different devices. Caged control valves can have

numerous small holes through which the flow will pass. The smaller(3/16 inch and

3/8 inch) holes were chosen as representative of that class of device and most likely'the

smallest eddies to be encountered. At the other end of the spectrum, we chose the largest

hole size for a three-hole orifice geometry. Three seemed to be the least which could be

14
used without introducingtoo .much bias or unevenness'ifito the flow field across the

pipe. 'Selection of the -total flow area in' the orifice, which isldethmined by the number

and size of holes, was guided'by the pump's ability' to deliver the 'required flow while

still preventing cavitation at the'perforated plate. 'To prevent excessive cavitation,-the.

cavitation index was continuously monitored during the tests. The cavitation index

was calculated according to the formula:

Pdown- Pvap .' ;

Pup- Pdown

where ? .: ' ;

c = Cavitation Index

Pups Pdown = pressures upstream and downstream of the perforated plate, psig

Pvap = Vapor pressure of fluid, psig

For water at 1 O00F, Pvap 1 psig

So the formula simplifies to ;

P down- 1 - f

Therefore, a smaller cavitation index, a, means higher levels of cavitation.

Work 'by Tullis with single and multi-hole orifices, skirted control valves;'' and

butterfly valves (References'8,9) has helped define -a range'of operating regimes with'

differing 'cavitation'characteristics ranging from incipient cavitation at'the' mild'end

to choked flow at the severe end. Incipient cavitation can be characterized by sounds

similar to the' crackling of frying'bacon or popping corn. This is a very mild level of'

cavitation.' -For the perforated plates used in these tests, this level of cavitation was

typically reached with cavitation indices of about a = 0.5. All tests were'performed with

a = 0.7 as a minimum to guarantee no influence of cavitation.

Turbulence tests were run at flow velocities from 2.5 to 18 fps. Given the pump

characteristics and cavitation criterion, higher flow velocities could not be developed.

Nevertheless, 18 fps was enbugh to fully open the disk against the backstop in all cases.

Tests were'run at 'all three stop' angles and at'proximities of 1.5, '2.5, 4.5i'and 10.5

diameters upstream. In general, the 3-hole and 36-hole plates were used to cover the

flow ranges from 2 to 9 fps, and the 4-hole and 61/67 hole plates were used to cover the

range from 7 to 18 fps. It should be remembered that pre~ssure drop -across any

perforate'd plate increases withb thesquare of the flow velocity.- Thus, at the low end of

the tw~o velocity ranges just mentioned, plate'P was in the 2 to 4 psi range while at the

high' end of the' ranges the piessure' drop was in the '25 to 35 psi range. "Detailed

discussionrof results for the 3-inch' valve at;1.5d and 73-degree stop angle will be

covered, with the remaining data summarized later in this section.., l

;15
I

Disk Fluctuation Data, 3-inch Valve

Figure 16 shows the disk fluctuation data for the 3-inch valve with the stop angle at the

73-degree clearway position and turbulence source at 1.5d. The overall plot represents

data from all four of the perforated plates used in the test program. Some important

trends should be noted:

1. Peak disk fluctuations are considerably higher than for any of the elbow tests,

reaching levels of over 15 degrees (peak to peak, 3a).

2. The highest levels occur in the 4 to 6 fps range with the 36-hole plate. As shown

in Table 3 data for various orifice plates, this plate has the smaller (3/16")

diameter holes.

3. The velocity required to fully open the disk is in the 17 to 18 fps range based upon

disk fluctuation levels dropping to under 1 degree. This agrees closely with the

baseline data.

4. The darkened symbols represent those tests where tapping of the disk against

the stop begins as flow velocity is increased. Compared to the elbow tests, the

velocity zone over which tapping occurs is wider. This is the zone over which

accelerated fatigue and impact wear occurs.

Figures 17, 18, and 19 show the fluctuation data for the remaining proximities of 2.5,

4.5, and 10.5d at the 73-degree stop angle. On each plot an upper bound curve has been

drawn for disk fluctuations at each proximity.

The upper bound data for the most severe disk fluctuations as a function of both flow

velocity and proximity of the turbulence source shown in previous figures are presented

in a more comprehensive manner in Figure 20. As one would expect, at any given

velocity, maximum fluctuations decrease as the turbulence source is moved away from

the valve. Also, at all proximities the maximum fluctuations occur at around 6 fps.

To complete a set of turbulence data at all stop angles, the 53-degree and 63-degree

fluctuation data at 2.5d is shown in Figures 21 and 22 respectively. We see that peak

fluctuations are lower than at the 73-degree stop angle, and, the velocity range over.

which disk motion occurs gets; progressively, narrower as the stop angle is decreased.

Additionally, the velocity range between onset of tapping (darkened symbols) and disk

seating (fluctuations less than 1 degree) is much narrower as well. These trends were

also present in the elbow data. Even though disk fluctuations are much higher than

were seen for the elbow tests, the velocity margin required over baseline to firmly seat

the disk is very similar, approximately 30 percent at 63 degrees and 15 percent at 53

degrees.

Disk Fluctuation Data, -inch Valve

Figures 23 through 29 represent the disk fluctuation results for the 6-inch valve with

various orifices and their proximities. All the trends noted for the 3-inch valve data

still apply to the 6-inch valve; with the exception that at 2.5d the peak fluctuations are

slightly higher at velocities above 6 fps than at 1.5d. This can be seen in Figure 27

where the data are summarized. A comparison between Figure 27- and Figure 20 shows

that the overall magnitudes of the disk fluctuations are approximately the same for both

the 3-inch and the 6-inch valves, confirming the validity of scaling up of the results in

these tests.

16
Summary of Miuimurn Disk Fluctuations for

Elbows, Reducers, and Turbulence Sources e

Figure 30 is an overall iuxiim'ary and comparison of maximu-m disk'fluctuation data

for all of the upstream disturbances tested for both valves at each proximity. Clearly,

the turbulence source has the most severe effect on disk stability, with fluctuation

magnitudes reaching as high as 16 degrees at the closest proximity. As expected, the

disk fluctuations created by the turbulence source continue to decrease as the proximity

from the check valve is increased, but even at 10.5d the magnitude is as high as with

elbows at 5d. One can also see that both the 3-inch and the 6-inch valves show

overlapping data which provides some confirmation of the fact that disk motion

expressed in nondimensional terms of fluctuating angle AO is a valid scaling

approach. Comparisons made against published data by others show that our results

bound the reported magnitudes of disk fluctuations with other turbulence sources which

include control valves, throttled butterfly valves, and pumps as well as the effect of

larger check valves' (References 5, 10, 13, 24, 25). 'Therefore, the upper bound'data

developed from' Phase I experiment'.could be"used in -pre'dicting'maximum wear or

fatigue that can be caused by disk motion.

The elbow data for both' valves shows that the' overall disk motion is considerably less'

than with a' turbulence source and is typically in'the range' of 3 to 5 degrees.- The'

maximum fluctuations' at 'the closest proximity-of Od 'are between 6 and 9 degrees range

for the elbow up condition, which is the most severe elbow orientation. As'reported in'

previous elbow results (Reference 5), the disk' fluctuations become negligible 'beyond

5d.

One of the most interesting results was the'performance of the check valve with

reducers as an upstream disturbance. 'As showni in' this figure,' the disk fluctuations

with the' reducers'even at the closest proximity to the valve' are within the baseline

fluctuations that were measured ivith 20d length of upstream straight pipe which

included a flow straightening section.' The 'restults' show that the 'reducer employed to

install a' smaller size check'valve'in'a large'r diameter upstream pipe is not a '

detrimental disturbance in such configuration. One'should be cautioned, however, that'

the'reverise situation where a reducer is being used to install a'larger size check valve

in a smaller size upstream pipe (which in fact makes the reducer function as a flow

expander) is known to be a definite source of turbulence. Fortunately, for swing check

valves such installations are rare in typical power plant systems.'

DISK NATURAL FREQUENCY AND DAMPING TESTS '

By, combining data for amplitude of 'disk motion with disk natural frequency data,

estimates for hinge pin' wear and disk stud impact wear and fatigue can be made. -To

measure 'the 'disk natural frequeiic' tinder flowing c'onditi6ns, a simple test was run

on each valve where steady state flow condition's were established with'no' upstream

disturbance. 'A pushrod was'introduced through 'the; bonnet and used to quickly disturb

th'e disk from"'it's equilibrium 'position.' A-strip chart recording was-made of the' re'sulting response 'to te'distiiiban~ce. ''Figurie 31 chrs t 'a' s i '-f a ical recordingw T ~ ~ ~ ~ ~ ~~~iue sows'a section'o atypiclredig

As' is evident from this 'recording, the disk response is relatively'highly damped. 'By'-'

measuring successive peaks from a nurmber' of tests, the damping ratio -is calculated as"

0.5 ±0.1 for the 6-inch valve and 0.3 0.1' for the 3-inch.' The naturai frequency tests'

were performed at a number "of' flow velocities which 'correspond to' different

equilibrium'n positions' of the disk. !The' resuIts are presented in Figure 32. ' '

17
I

An analytical formulation for to predict natural frequency was also developed by considering the disk as a single degree of freedom spring-mass system. This is presented in Appendix A and the results from it for the 3-inch and the 6-inch valves are plotted in Figure 32 for comparison. Another comparison was made against work by Sununu (Reference 10) which suggests that the disk natural frequency is proportional to the eddy frequency generated in normal turbulent pipe flow. This frequency can be calculated as follows:

0.04 V.,g

'eddy Rpipe

Test data published in Refe'rence 10 for 2-inch and 4-inch valves is also plotted in Figure 32 for comparison.' Two trends are apparent. First is a trend of increasing disk natural frequency with decreasing valve size. Second is an" increasing natural frequency with increasing velocity.' Unfortunately, not enough data are'available to draw firm conclusions about the trends. Other factors such as effective 'disk weight and valve internal geometry may have some effect on the natural frequency. In fact there does not appear to be alinear relationship between natural frequency and velocity for majority of the data as the previous ~formula suggests. Nonetheless, these

measurements give a reasonable quantification of disk natural frequency and form a basis for performing wear and fatigue estimates.

DISK-TO-STOP IMPACT LOAD TESTS

To calculate impact fatigue life of the disk stud/nut assembly when the disk is tapping against the stop, estimates of the'impact force can be made based upon the transfer of kinetic energy of'the moving disk and hinge arm' as'sembly into strain energy absorbed by the contacting elements. With the load cell apparatus incorporated in the

stop mechanism, actual load- measurements were made for the 3-inch valve to characterize the load signature' when tapping. It was found that the data 'acquisition rate had to be increased' considerably - from 40 hz (which is sufficient for the

measurement of'displacement' and pressures) to 4 khz - to properly capture the load/time history.

Figure 33 shows the results' of one'such measurement'made on the 3-inch valve under the following conditions: flow velocity of 17.5'fps, an'upstream elbow oriented up at a proximity of id, the valve stop angle set at 73 degrees. A number of such measurements

were made under the same flowing conditions, with the highest recorded loads being 35 pounds. These measurements' are,'of course, with' the clear'plastic bonnet and the adjustable'stop mechanism which was being'used in the test instead of the actual steel bonnet with'integral 'stop. In all cases,; the contact duration was between 0.003'to 0.0035 seconds. To use-this load data, it is necessary to account for the relative stiffnesses of the test valve.bonnet/stop assembly and the actual valve. These stiffnesses can be, determined by'testing or analysis. One method would be to apply'known loads to each assembly and measure'the displacements. Appendix B documents an analysis method and a calculation in wliich the'relative stiffhesses are estimated. Based'on the 35- pound load measured'and this analysis approach, it is estimated that the' actual impalct load with: steel bonnet having integral stop arrangement will be approximately'270 pounds for the 3-inch valve. The impact force results thus obtained can be used in predicting impact wear between the disk/stud nut and the stop as well as fatigue life of this assembly.

18
DISCUSSION

CLEARWAY SWING CHECK PERFORMANCE PROBLEMS

A very important result of the elbow tests conducted in this project is' the 'apparent

discrepancy' between these'tests and those reported in Reference 5 for the minimum

velocity requirements for the 3-inch valve with' stop angle at the' 73-degree clearway

position and elbows at Od. The earlier tests reported values of 17 fps for elbow down and

18 fps for elbow up. The current test program did not fully seat the disk at velocities as

high as 25 fps. Possible reasons for that difference can be categorized as follows:

1. Differences in flow velocities/conditions;

2. Variations in inlet piping setup;

3. Minor differences in' physical location of stop position(s). -

The first item is unlikely because no evidence of velocity based errors shows up in

other tests. All eequipment was calibrated and, in fact,'was the same in both tests.

The second item could conceivably contribute if, for example, the elbow was tilted

slightly' as it was bolted to the valve.' This would cause the flow to enter the valve at a

different angle relative to the nominal flow axis, in effect changing the full open angle

of the valve. This possibility exists-because of the flexibility of the elbow/valve

connection and the difficulty in establishing :the lelbow orientation (using a bubble

level) at this Od proximity. Finally, physical changes made to the valve between test

programs could have affected the absolute position of the adjustable stop. Since

different assemblies were"iised for the adjustable-stop-devices,' there-could have been

some offset introduced at each stop position;. Although great care was taken to prevent

this, a difference between assemblies of 1 degree is a possibility.

The second or third items could create' enough of a difference in this valve to account

for the observed phenomenon. The large 73-degree opening angle of this valve,

combined with the clearway design in which the disk is completely shielded from

direct flow impingement, makes it very susceptible to these physical changes. It is

important to realize that there are real-world analogies to both of these test related

items. First, a real piping installation' will be as likely to'have small angular

alignment variations between valve and pipe as the test setup, if not more so. A pipe

could very easily be welded to a valve or flange in such a way as to introduce a couple of

degrees of misalignment between the pipe and valve centeilines. Second, valve

manufacturing tolerances are such that the full open angle of the disk can vary by as

much as 5 degrees, which'has been observed to be' a typical variation in product lines

made by most manufacturers. This is considerably more than the variation between

test setups. Even a comparison between the 3-inch and the 6-inch valve (both products of

MCC Pacific Valve Company) in the full open clearway position for the valves used in

our tests shows a significant difference in the velocity required to-fully open'the disk:

For these reasons, sufficient velocity margins should be used for clearway designs. To

-avoid possible problems, it is best to lengthen the backstop to ensure the disk'protrusion

into the flow stream by at least 10 degrees, which has been found to provide 'much better

disk- stability under various flow conditions and upstream' disturbances.

19
VELOCITY MARGIN REQUIRED FOR ELBOWS AND REDUCERS

One of the project objectives was to develop appropriate modifications to the minimum

velocity formula to account for upstream elbow and reducer effects and develop the

velocity margins necessary to fully open the. disk under these conditions. The data

presented in Figures 12 and 13 form the basis for these modifications. The appropriate

margins for upstream disturbances are most conveniently expressed as a multiplying

factor:

V'min = Cup Vmin

where Vmin = minimum velocity required to fully open the disk with an

upstream disturbance

Vmin = Minimum velocity required to fully open the disk with no

upstream disturbance

Cup = The velocity margin factor accounting for an arbitrary

upstream disturbance.

Table 3 presents values for the factor Cup for the different upstream disturbances.

Keeping in mind some of the test variables, we have chosen to use factors we feel

represent the' worst case conditions. Although this may result in somewhat higher

velocities being required in certain instances, the error must be conservative to ensure

safe performance.

Table 3

VELOCITY MARGIN FACTORS FOR

UPSTREAM DISTURBANCES, Cup

Elbows Reducers

50. 60' 70° 50 60 70°

Otold 1.2 1.4 1.8 1.05 1.05 1.05

1to3d 1.2 1.2 1.4 1.05 1.05 1.05

3 to 5d 1.1 1.1 1.1 1.05 1.05 1.05

Beyond 5d i.0 1.0 1.0 1.0 1.0 1.0

THEORETICAL PREDICTION OF

TURBULENCE INDUCED DISK MOTION

.Discussion of the turbulence test results always involves the issue of relating

laboratory results to the* many different real-world situations which exist. ' One

approach, ofcour'se, would be to test all the different types of turbulence producing

devices individually, for their'effect on check valve performance. - An effort this

extensive would"involve' many man-years of test and analysis'-time' and would

invariably remain incomplete as new devices arrive on the scene. An alternate

method would be to use turbulence characteristics of the flow to predict check valve disk
response.- Iesearch in the area of fluid-structure'interaction has resulted in a general

formulation of the type given below which can be used to 'predict disk response to a

turbulent flow'stream (References 11', 12, and 14 through 22).

1/2

irSF(f)2 X:MJI {X 2 XfflxQXJ

i . , ._..K.. -

whereXrms='= Root mean square displacement response, inch

SF(f) = Power spectral density of the fluctuating forces over the

disk area, lb2Ihz. This can be obtained by multiplying

the power spectral density of pressure field, expressed

in psi2/hz, by the disk area, in 2

K= Slope'of the fluid dynamic force on the disk vs. linear

displacement, lb/in (see Appendix B)' -

fn = Natural frequency of 'the disk/arm assembly 'dangling

in the flow stream' ''

Q = Transmissibility at resoi'ance = 1/2 Cc' -

Cc= Critical damping ratio

J(f) - Joint acceptance for the generalized force power spectra

at frequency, f

The disk natural freque'ncy and damping were measured, the results of which are

discussed in this report. An' alternate analysis method 'to 'calculate the natural

frequency of the disk is shown in Appendix A,'which also includes a method to

determine effective stiffness, K, for the disk fluid dynamic force vs. position curve.

Thus,' analytical. methods have been developed which can provide a reliable

quantification for'calculatinj fn, K, as well as Q; and the remaining variables are SF

and J(f). These pertain to the turbulence power spectral density and how the pressure

fluctuations in the turbulent flow stream are time correlated across the disk area. Joint

acceptance factor, J(f), can have a possible maximum value of 1 for perfect span-wise

correlation of pressure fluctuations and a minimum value of zero for random

fluctuations having no phase correlation whatsoever.

Preliminary analysis of pressure transducer data to determine the turbulence power

spectral density has been performed, and a general trend of higher disk motion with

increasing power spectral density of the pressure field in the turbulent flow has been

found to exist. However, considerably more effort is needed in this area to better

correlate disk motion to turbulence phenomena by using this power spectral density

approach.

A thorough review of literature in the area of fluid structure interaction revealed that,

even though the theoretical formulations based on power spectral measurements for

predicting mechanical response of an elastic system due to turbulent flow have been

generally well established, considerable work is still under way to develop reliable

21
correlation between theoretical predictions and actual tests (References 11, 12 and 14

through 22). State of the ,art in this very active area of research still relies heavily upon

sound experimental mneasurements which'then form the basis to refline the theoretical

models that predict structural response. Vibration resp'onse of cylindrical tubes

subjected to cross flow and axial flow is one subject that has been studied extensively by

many different researchers because of its obvious importance to the nuclear power

reactors, heat exchangers, and condensers. This has resulted in significant

advancements in the development of analytical models,' but they are limited to the

cylindrical tube structures. Even for this well-studied structure, not much data

regarding power 'spectral density of the pressure field are available in the open

literature.

Therefore, recognizing the limitations of the current state of the art in this field and the

long-term research effort needed to develop reliable analytical methods to predict disk

motions, the effort under Phase I put more focus on developing the needed information

by systematic experimentation using a turbulence source of generic nature. In order to

ensure that the results have general applicability, suitable: nondimensional scaling

factors were used in our test program.

As presented in detail in the previous sections, many tests were conducted with multihole orifice plates in which different sizes and'number of holes were used to control the

two important characteristics of turbulence: eddy size and turbulence intensity.

Proximity of the turbulence source was varied up to 10.5d from the valve. Extensive

data were generated for both 3-inch and 6-inch valves which are presented in Figures

16 through 26, 28, and 29. The most important results from these turbulence tests,

however, are presented in the family of bounding curves which relate the maximum

disk fluctuation experienced as a function of flow velocity and proximity of the

turbulence source and are'shown in Figures 20, 27, and 30. 'Any data available from

published sources quantifying the disk response due to various turbulence sources

(References 5, 10, 13, 24, 25) were reviewed and compared against these bounding

curves. It was found that in no case did the maximum disk fluctuation exceed the

values plotted in this family of bounding curves.

Thus, these bounding curves, which were produced from our Phase I effort, can be used

to predict maximum swing check valve disk response with a severe turbulence source

present at a defined proximity upstream of the valve.
I 1<11

RESULTS AND CONCLUSIONS

* Phase I has resulted in the development of extensive data and techniques that can be

.used to predict the stability of swing check valves under three different types of

upstream disturbances: elbows, reducers, and turbulence sources.' The major

conclusion from Phase I is that the swing check valve disk performance can be

predicted. The specific predictions include: '

* Whether disk stability can'be achieved for a swing check valve' of known disk

angle with an upstream disturbance at a defined proximity.

* The-increase in velocity required over Vmin to achieve stability:when a

disturbance is present upstream of the valve;

* The maximum disk fluctuations that can be expected due to these upstream

disturbances.' To predict the worst-ca'se disk fluctuations caused by a severe

upstream turbulence source, upper bound curves have been developed;-

* The range of velocity over which'the ^disk fluctuations can cause tapping against

the backstop which results in impact wear and fatigue of the disk-to-hinge arm

connection. The zone of velocity over which the disk is free-swinging without

impact has also been defined for various configurations of valves and upstream

disturbances. In the free-swinging zone, only hinge pin wear is of concern.

In addition to these major conclusions which relate to the specific objectives of Phase I,

the important results obtained are summarized below:

1. Stability of clearway swing check valves can be very sensitive to manufacturing

tolerances and the physical alignment of piping to the valve. Clearway designs

with closely coupled elbows may not attain a stable fully open disk position with

any flow velocity.

2. Decreasing the fully open disk angle, *, by 10 degrees in clearway designs

greatly enhances disk stability at the fully open position. It also reduces the

velocity required to fully open the valve;

3. Upstream reducers have little or no disturbing effect on disk stability or velocity

required to fully open the valve. Reducers are used in the sense that it reduces the

flow area as flow approaches the valve;

4. Maximum disk motion due to an upstream turbulence source can be much higher

than for upstream elbows, depending upon the severity of the source. When

calculating wear, turbulence sources out to 10d upstream must be considered;

5. The minimum velocity margins over Vmin for elbows depend upon the disk

opening angle, elbow orientation, and the proximity. Table 3 shows the

minimum margins necessary.

6. The minimum velocity margin required to fully open the disk for non-clearway

valves with an upstream turbulence source is 10 percent higher than baseline

conditions, assuming the velocity field remains uniform across the pipe section.

Higher velocity margins might be required in the case of turbulence combined

with velocity skew effects.

23
RECOMMENDATIONS

To achieve the eventual goal of quantifying degradation trends when swing check

valves are used in less than fully open conditions, which is the case with many valves

in nuclear power plants,' further'research is'necessary. 'Phase I has resulted in

providing definite information regarding flow conditions and the upstream

disturbances which can cause accelerated wear.n It has also produced data regarding

maximum disk fluctuation, disk tapping zone, and free-swinging zone, which provide

the basis for developing quantitative predictions of degradation.

Research should be continued to use the disk'stability information from Phase I to

develop quantitative fatigue and wear life predictions. This can serve as a basis to

make objective decisions regarding suitable maintenance intervals over which check

valve degradation is within acceptable safe'limits and- which will 'not jeopardize its

operation. Phase II research, aimed at developing such predictive techniques,

verifying them, in controlled' laboratory, conditions, and correlating them with actual

plant data should be conducted. The eventual goal of this research effort is to improve

the safety and reliability of check valves in nuclear power plant service properly

accounting for the expected degradation trends.

24
,r,. ,s,^

M USCES. - -. ,

1. Check Valve Failures or Degradation, Significant Operating Experience Report

No. 86-3, Institute of Nuclear Power Operations, Atlanta, GA, October 1986.

2. Aging and Service Wear of Check Valves Used in Engineered Safety-Feature

Systems of Nuclear Power Plants, NRC NUREG/CR-4302, Vol. 1, December 1985.

3. Loss of Power and Water Hammer Event at San Onofre Unit 1 on November 21,

1985, NRC NUREG-1190.

4. Chiu, Chong, and M. S.: Kalsi. Failure 'Analysis of Swing Check Valves - San

Onofre Nuclear Generation Site - Unit 1, NRC Docket No. 50-206, April I986.

5. Application Guidelines for Check Valves in Nuclear Power Plants, EPRI NP5475, January 1988. - -

6. Idelchik, I.E. Handbook of Hydraulic Resistances, 2nd Edition, Hemisphere

Publishing Corporation, New York, 1986. ' -

7. Ward-Smith, A. J. -Internal Fluid Flow: The fluid Dynamics of Flow in Pipes

and Ducts, Clarence Press, Oxford, 1980.. - -

8. Tullis, J. Paul, et al. Perforated Plates 'as Hydraulic Energy Dissipators, Utah

Water Research Laboratory, Utah State University, no date.

9. Tullis,- J. Paul.- Valve Sizing to Avoid Cavitation, Utah Water.Research

Laboratory, Utah State University, no date.

10. Sununu, John E. -Wear in Check Valves Due to Flow Induced Motion, Master's

Thesis, Massachusetts Institute of Technology, May 1987.

11. Blevins, R. D. Flow Induced Vibration, Van Nostrand-Reinhold,1977.

12. Flow-Induced Vibration of Power Plant Components, edited by M. K. Au-Yang,

ASME Paper No.,PVP-41,1980.

13. Tullis, J.; P., and W. J. Rahmeyer. Performance Tests on Feedwater Check

Valves for San Onofre Nuclear Power Plant, Phase A and Phase'B, Utah State

University Foundation Report, April 1986.

14. Flow-Induced Vibration Design Guidelines, edited by P. Y. Chen, ASME No.

PVP-Vol. 52,1981.

15. Advances in Fluid-Structure Interaction - 1984, edited by G. C. Everstine and

M. K Au-Yang, ASME PVP-Vol. 78, 1984.

16. Advances in Fluid Structure Interaction Dynamics, edited by F. J. Moody and

Y. W. Shin, ASME PVP-Vol. 75, 1983.

17. Dynamics of Fluid-Structure Systems in the Energy Industry, edited by M. K.

Au-Yang, S. J. Brown, and J. Ray McDermott, ASME PVP-39,1979.

25
18. Volume 1: Excitation and Vibration of Bluff Bodies in Cross Flow, Symposium on

Flow-Induced Vibrations, edited by M. P. Paidoussis, 0. M. Griffin, and

M. Sevik, papers presented'at The' ASME Winter Annual Meeting, New Orleans,

Louisiana, December 1984.

19. Volume 2: Vibration of Arrays of Cylinders in Cross Flow, Symposium on FlowInduced Vibrations,' edited by M. P. Paidoussis, M. K Au-Yang,'and S. S. Chen,

papers presented at The ASME Winter Annual Meeting, New Orleans,

Louisiana, December 1984.

20. Volume 3:Vibration in Heat Exchangers, Symposium on Flow-Induced

Vibratio'ns, edited by M. P. Paidoussis, J. M. Chenoweth, and M. D. Bernstein,

papers presented at The ASME Winter Annual Meeting, New Orleans,

Louisiana, December 1984.

21. Volume 5: Turbulence-Induced Noise and Vibration of Rigid and Compliant

Surfaces, Symposium on FIlow-Induced Vibrations, edited by M. P. Paidoussis

and A; J. Kalinowski, papers presented at The ASME Winter Annual Meeting,

New Orleans, Louisiana, December 1984. :

22. Volumee 6- Computational Aspects of Flow-Induced Vibration, Symposium on

Flow-Induced Vibrations, edited by M. P. Paidoussis and S.' S. Chen,' papers

presented at The ASME Winter Annual Meeting, New Orleans, Louisiana,

December 1984.

23. Chiu, C.. and M. S. Kalsi. Plant Availability improvement by Eliminating Disc

Vibrations in Swing Che'ck Valves, ASME Paper No. 86-JPGC-NE-6, October

1986.

24. Rahmeyer, W. J. The* Effect of Upstream Piping on Swing-Check Values,

Journal of Authority of Water Work Association, pp. 553-555, November 1983.

25. Statton, J. A. 'Modification Based on Stability and Wear Studies. Presented at

INPO Check Valve Workshop, October 1986.

26. Kalsi, M.S. Analysis of the 4-Inch Tilting Disc Check Valve Problems and

Proposed Modifications, Kalsi Engineering, Inc. proprietary report to Southern

California Edison dated 8/11/86.
-HINGE ARM

OC SEAT. PLANE TILT WITH RESPECT TO VERTICAL

DISC ANGLE WITH RESPECT. TO HOR ZONTAL

G: (+ ,$ ) TOTAL IMPINGEMENT ANGLE

d>r DISC OPENING ANGLE

CHECK VALVE CONFIGURATION

FIGURE 1

27
L, Wuml %t "-j I. I. -

CHECK VALVE TEST SETUP

FIGURE 2

28
KALSI ENGINEERING, INC.

MECHANICAL CIESKGN & ANALYSIS

>-1.. A

PRIMARY

PC TI IOU Co fW BYPASS RETURN FLOW - ADJUSTABLE

BYPASS VALVE aF

I& --

w -

V

1

TlrT i-r~

I I

PIPING SECIION-

-- I 1,STRAIGHTENINO ; *, 

UPSTREAM ELBOW 'K=3[1 3⁄4) KU(VERTICAL UP)

SIRAIGHTENING VANES

StRAIGHTENINO VANEy

I,' / / / / / / / / / / / Z/ / / / / / / / / / / / - / / / / / / / / / - / / / / / / 'I

KALSI ENGINEERING, INC. FLOW TEST FACILITY

FIGURE 3
ADJUSTABLE BACKSTOP

WITH LOADCELL -

LVDT FOR MEASURING

DISC MOTION

INSTRUMENTED VALVE CROSS SECTION

FIGURE 4

30
DISK FLUCTUATION vs FLOW VELOCITY

Baseline, 20d upstream (a Inch)

4.e

I,1

am

to

0.

I

3.

2.8 -

2. -

2.4 -

22

1.6

1.6

1.4

12

1

0.6

o0.

0.4

02

0

0 4 6 12 I1 20 24

Flow Velocity. fps

COMPOSITE PLOT OF DISK FLUCTUATIONS FOR

6" VALVE, BASELINE TESTS

FIGURE 5

31
D, IN. It Y. IN.', Y/ |

3 1 0 I 0 .

6 1 0.8 1.133

DISK PROJECTION COMPARISON

CLEARWAY SWING CHECK TEST VALVE

FIGURE 6

32
3

2.5

2.6

10

I"

Q

11

2.4

2.2

2

1.8

7.6

1.4

12

0.x

0.6

0.4

DISK FLUCTUATION vs FLOW VELOCITY

Basellne (20d pipe upstream) (3 Inch)

I j T I [ II 1 '-1 111 X I

73 eg I

6S -i -

_ = 53 D -zg

02

0

0 4 8 72 16 20

Flow Velocity. fps

~: , ~ . , , .

J~~~~~~~, . I ' ' .- , ."- .' -" :- ,- I e . - - - , <s

COMPOSITE PLOT OF DISK FLUCTUATIONS FOR

3" VALVE, -BASELINE TESTS

FIGURE 7

33
ELBOW UP

DEFINITION OF ELBOW ORIENTATION

FIGURE 8

34
DISK FLUCTUAT ION vs FLOW VELOCITY . Elbow Up C Od I

a

a

4J1

a

+

7

a

4

31

2

1

0 f

0 4 8 72 la 20 24

Flow Velocty, tps

COMPOSITE PLOT- OF DISK FLUCTUATIONS FOR

3" VALVE AT VARIOUS DISK STOP ANGLES

FIGURE 9A

35
DISK FLUCTUATION vs FLOW VELOCITY

Elbow Up 0 Od (d Inch)

a

El

'a

U

:4.

0

5

1.2

*-0

I

4

3

2

I

0

0 4 . 8 12 1a 20 24

Flow Velocity. t

COMPOSITE PLOT OF DISK FLUCTUATIONS FOR

6" ,VALVE AT VARIOUS DISK STOP ANGLES

FIGURE 9B

36
!' !

DISK FLUCTUATION vs FLOW - . I Elbow Down e Od

VELOCITY

I4. J

'a

e

C.,

*-Z

* - . '~Flow Velocfty. fps

COMPOSITE PLOTOF DISK FLUCTUATION FOR

3" VALVE AT-VARIOUS DISK"STOP ANGLES

FIGURE 1OA

' 37
DISK FLUCTUATION vs FLOW

Elbow Down 0 Od (6 Inch)

VELOCITY

4

3.5

0

4.d

z11

Q

*1

3

2.5

2

1.5

1

0.5

a

0 4 8 12 16 20 24

Flow VelocIty. fps

COMPOSITE PLOT OF DISK FLUCTUATIONS FOR

6" VALVE AT VARIOUS- DISK STOP ANGLES

FIGURE 10B

38
cf,

w

0

0

2-

-

U,

en

x

m:

1.513

-,. 0

. LID, 6 INCH (DARK SYMBOLS)

2 .4' . -. 6, 8 10 12 : 14

16

14-

12-

. . . . . . . . . . I................................ . .

: ,;: . . - -A -ELBOW UP

v v ELBOW DOWN

S O REDUCER

lo+ :

8-f

6-

4.

2

.

BASELINE

FLUCTUATIONS

_ p _~c~ _ 0...__ _ _ _ _

. I

u . l l l

.

o ' '1

I .; ., I i2D i,

.

MAXIMUM

2 . 4 . . . . - .- 6 .8 -10 - *- 12 . 14 .. . , r

L/D, 3 INCH (OPEN SYMBOLS) I I ~ ., I * - I , *' - a - .I .- !- : .' , ., - 1 ' ,. , * : ._ . .,

DISK FLUCTUATION VS. DISTANCE TO

UPSTREAM; DISTURBANCE '

FIGURE 11

39
£

LU

<25 -

C ELBOW UP = _=0

c. \ ' v ELBOW DOWN

20- - REDUCER -

z

-A BASELINE: 17.4 FPS, 73°0 a.

0 15--

in

,10-_

U.

U. c 2 BASELINE: 7.2 FPS, 630

F 5- . ~~~~BASELINE: 5.1 FPS, 530

O0

-J

C] O-. I : i I I. | | i § § I I _

U ° 1 2 3 .4 5 6 7 8 9 10 11 12 13

L/D, DISTANCE TO UPSTREAM DISTURBANCE aU .. .- t .

ELBOW AND REDUCER TEST RESULTS

3 INCH SWING.CHECK VALVE.

FIGURE 12

40
cc:

w

F2- c 25 - * ,

in :&ELBOW UP =

&20- - ,( \ ELBOW DOWN

Z - 0 REDUCER

LU

° 15-BASEUNE: 13.8 FPS, 700

10

0m

BASELINE: 6.3 FPS, 600 -

o - BASELINE: 4.7-FPS, 500

Lu

-0 t01 -2 3 4 5- 6_7. 8 9 10 11 12 13

LID, DISTANCE TO, UPSTREAM.:DISTURBANCE

ai -

ELBOW AND REDUCER TEST RESULTS' - 6 INCH SWING CHECK- VALVE

'FIGURE 13

^~~~ - *. ~~ /. 1'

41
DISK FLUCTUATION vs FLOW VELOCITY -t Reducer 0 Od (a Inch)

C.I

Q

C!,

*.2

3-

2.61

2.0

2.4

2.2

2

1.8

1.4

1.2

1

0.6

0.6

0.4

0.2

0

0 4 8 12 i1 20 24

Flow Velocity, fps

COMPOSITE PLOT OF DISK FLUCTUATIONS FOR 6" VALVE WITH 8" x 6" REDUCER AT OD

FIGURE 14

42
3 & 6 INCH PERFORATED

PLATE HOLDERS

FIGURE 15

.. - . . . I

d . ..

43
2 16- 0=36 x 3/

w ca 14 A- ° \ . 0~~~~~~~~0 61 x 3/1

012- - - / X v 7e = 4 x 7/8'

0103

<8 A

u. 4- /°vv

en 2-

0 2 4 6 8 10 12 14 11

FLOW VELOCITY, FPS

DISK MOTION OF 3 INCH VALVE

DUE TO TURBULENCE

PERFORATED PLATES AT 1.5 DIAMETERS UPSTREAM

STOP ANGLE AT 73 DEGREES

FIGURE 16

6 18

44
1 8 -

I16 0 = 36 x 3/16" DIA

c 0 = 61 x 3/16"'DIA

14 = 3 x 7/8" DIA

Q 12 i _ V = 4 x 7/8" DIA

0 6 o,

-jin 2

. , , ~~~; .1 . .. I .1 . . . .. I . . ...

o 2 4 6 8 10 12 14 16 18

FLOW VELOCITY, FPS

'DISK. MOTION OF 3AINCH VALVE

PRRT DUE-TO TURBULENCE

PERFORATED;.PLATES AT 2.5 DIAMETERS UPSTREAM

STOP 'ANGLE AT 73 DEGREES

FIGURE 17

45
I

18 -

UJ 16- 0 = 36 x 3116" DIA

en14-- 0 = 61 x 3/16" DIA

cJ 1 = 3 x 7/8" DIA

12 V = 4 x 7/8" DIA

o 10--

sol

06-

-J~~~~~~1

0

0 2 4 6 8 10 12 14 16 18

FLOW VELOCITY, FPS

DISK MOTION OF 3 INCH VALVE

DUE TO TURBULENCE

PERFORATED PLATES AT. 4.5 DIAMETERS UPSTREAM

STOP ANGLE AT 73 DEGREES

FIGURE 18

46
18 , , , , , I , .

U16 0 = 36 x 3/16" DIA

us 0 = 61 x 3/16" DIA

a 14 - -,3 x 7/8" DIA

w

D12- V=4x7/8" DIA

0lol

6--

06

0 2 4 '6 ;8 10* 12 14 16 18

FLOW VELOCITY, FPS -

DISK MOTION OF'3 INCH VALVE

DUE TO TURBULENCE

PERFORATED PLATES AT 10.5 DIAMETERS UPSTREAM

STOP ANGLE AT 73 DEGREES

FIGURE 19

47
20- .

U,

w

w 1.5D

0

~10

us 0~~.5D*_

D -

4.51)

IL 5

0

0- I , I , I I I§ I , , I , _

0 1 2 3 4 5 6 7 8 9 10 11 12 13

FLOW VELOCITY, FPS

MAXIMUM DISK MOTION OF 3 INCH VALVE

DUE TO TURBULENCE

PERFORATED PLATES AT 1.5 TO 10.5 DIAM.UPSTREAM

STOP ANGLE AT 73 DEGREES

FIGURE 20

48
wi

Lu

i

.21

C-,

IL

C'd

2 4 6 8 10- 12 14 I

FLOW VELOCITY, FPS

DISK MOTION OF 3 INCH VALVE

DUE TO TURBULENCE

PERFORATED PLATES AT 2.5'DIAMETERS UPSTREAM

STOP -ANGLE AT 53 DEGREES

FIGURE 21

49
u1616 .LU1

014 .

D 12 .

:10 2 4 6 8 10

FLOW VELOCITY, F

DISK MOTION OF 3 -NC

DUE TO TURBULE

PERFORATED PLATES AT 2.5 DIAM

STOP ANGLE AT 63 DE

FIGURE 22

12 14 1

?Ps

CH VALVE

ENCE

IETERS UPSTREAM

GREES

50
18 tCD

Lu

0

-j

LL

U,

05

8 . ' . I O

FLOW VELOCITY, FPS

DISK:MOTION OF,6 INCH VALVE

' DUE TO TURBULENCE

PERFORATED PLATES AT 1.5 DIAMETERS UPS'

STOP ANGLE AT 70 DEGREES

TREAM

FIGURE 23

51
16 0 = 36 x 3I1

w 0 =67 x 3/i

w 14 =3 x1-3,

° 12 \V4 x 1-3 i~~~~~~

10-

0 6 i Ba

A=4 OA a 111I <

<8-

yn

0

0 2 4 6 8 10 12 14 11

FLOW VELOCITY, FPS

DISK, MOTION, OF 6 INCH VALVE

DUE TO TURBULENCE

PERFORATED PLATES AT 2.5 DIAMETERS UPSTREAM

STOP ANGLE AT 70 DEGREES

FIGURE 24

52
1 8 .

cu 1 = 36 x 3/8" DIA..

w16

m 0 = 67 x 3/8" DIA

014 = 3 x 1-3/4" DIA

12 to|V = 4 x 1-3/4" DIA

<8-

6-

LL 4 0

0

0 2 4 6 8 10 12 14 16 18

FLOW VELOCITY, FPS

DISK MOTION-OF 6 INCH VALVE

DUE TO TURBULENCE

PERFORATED PLATES AT- 45, DIAMETERS -UPSTREAM

"'STOP ANGLE AT 70 DEGREES

FIGURE 25

53
u 16 0=36 x,3/8" DIA

14 0 = 67 x 3/8" DIA

ws = 3 x 1-3/4" DIA

cl 12- V= 4 x 1-3/4" DIA

"~8

C.)6

u. 4-

cl_21 0 o a l

0 2 4 6 8 10 12 14 16 18

FLOW VELOCITY, FPS

DISK MOTION OF 6 INCH VALVE

-DUE TO TURBULENCE

PERFORATED PLATES AT 10.5 DIAMETERS UPSTREAM

STOP ANGLE AT 70 DEGREES

FIGURE 26

54
20- . . . . . .

w

w

CID 1 25 1.55-

.0~~~~.1

o 2~~~.51D

10

- ~FLOW VELOCITY,, FOS

MAXIMUM DISK MOTION, OF 6 INCH VALVE

DUE',TO TURBULENCE

PERFORATED,.PLATES'AT 1.5 TOJ10.5 DIAM .UPSTREAM

STOP ANGLE ,AT-,70,DEGREES'- C.,, Fa : E..2.

U,~~~~-, ;,; iGR 2, -D.

-55
us

w

cr. w0

2F0

I-

.pC.)

-J

-

Ca

5

18

16-

14-

12-

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

0 = 36 x 3/8" DIA

o = 67 x 3/8" DIA

a = 3x1-3/4" DIA

v=4 x 1-3/4" DIA

10+

6+ a

0 0 - la

po~~~etA.

/0 . . mp

.4

2-

0 U WA Up .1,& 'VP \Ai .

0 2 4 .6 8, 10 12 14 16 18

FLOW VELOCITY, FPS

DISK MOTION -OF 6: INCH -VALVE

DUE TO TURBULENCE

PERFORATED PLATES AT- 2.5 DIAMETERS UPSTREAM

STOP ANGLE AT 50 DEGREES

FIGURE 28

56
18 4 . . . . . . . . . . . . . . .

I . . . . . . r U

c,, w

w

w

0

2F

0

C.)

-J

IL

U)

16t

14-

12-

10-

8-

o = 36 x 318" DIAo = 67 x 3/8"DA D

a = 3 x 1-3/4" DIA

v= 4 x 1-3/4" DIA

0 0

0-

A

MA' an'-

A --

A 0. 3 

. 0

£1 &Z ,Q 

4+

~~ jr .v~~', V., x

0 . .

. . . . os. VAv %XV - I . . . s . _ . _

0 2 4 6 8 10 12 14

11. 8

16 is

*. us FLOW. VELOCITY, FPS

- DISK.MOTION OF 6 INCH VALVE

DUE TO TURBULENCE

PERFORATED PLATES-AT 2.5 DIAMETERS UPSTREAM

STOP ANGLE AT 60 DEGREES

FIGURE 29

57
fl) w

w

0

en

a:

zT2F0

I--

C.)

-Um

U3

x)

18 ' C1.51

i8, I is

L/D, 6 INCH (DARK SYMBOLS)

2 4 6 8 10 12 14 . . . . . . . . . . . . I

I I a I a a ~~~~~~~~~~~~~f I

16 +

12

10

*A ELBOW UPci }

vv ELBOW DOWN 1m

I 0 0 REDUCER

* OTURBULENCE

at ~~~BASELINE t

FLUCTUATIONS L

-_- ... _=-

8-4

64

41

2-

0 i i

0

I- -1 2D '

I I * I I I I I I

-* 4 a & I I I & I

2 4 6 8 10

L/D, 3 INCH (OPEN SYMBOLS)

12 14

MAXIMUM DISK FLUCTUATION VS. DISTANCE TO

UPSTREAM DISTURBANCE

FIGURE 30

58
I I I I-t I I I A I I I .I I I I I ** I -I i *-r I I I* I I II

.i::

11'd in U

I I

.. A. .. . . . . . . .

.. ... ... ... 1 ...... 1. 1.

. ... . . . . . . . . . . . . .. . . . . . .

....... ... .... .. .... .

1.

*I I I I

IT

* II i *.

ii Ii

B.

.., ::'1 ';in.@.w @@*

w w*s.!: ., . ::

i' ''' ..... .. .j;.. _ i :::: :'? ,. ri.. .: ::::

.. ?.' *:-- I... .

:: :.:. .... :.:. ;:;.

*',0- .'-, I'.e . *- ..': - ;;..

,. '.. :;.: _'i.':;

I . I . .l_ _I .. I

..- .:.". ..- .'. ,; ;:. ;'. .-. .... .... .

.:. rlos ; ~~~~~. .,.. .s'... ... a}. ...

/.. * ... . .... .... ... .... .... .... .... ....

:' - .. : :': ... .. ... l I. ._. . .......

I ~. .' , .:.. '.A\/ '-;...:.

I I I I I I I - I I

. ....... .::: .... ...:.1 :..:1

... .. : . .: .. . . ... :: .. ..

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

. . ..'- .. ... .... -1.... .. ... ....

'/ . 1: : :'....::....:.. 1

....

....

...

I

1::,.

I'

i

I ~. I I I I I I I I I I I I II I I III I I . I- l-~ I

I: i

.;!i

I:::

:.1.

:T

:1!

: V:

 i,: . i.

.r.,

...

': 7; ::I:

L:;!

,:I;

.::: I

.a 

1II

'!Ill

....

::.I

.7

..i.

j I!:

U' I I~~~~~~~~~~~~~~~~ Nw

.. .... ....... . .. ... .. ... ... . .. .. . .... .... .... .... ... ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ I..

... .... . .... .. . ...... ...... ~~~~

--I -I - I -1 L" ..i.I:~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~ D N I S S E R E TS P O L I T L _S _R B _ _A, 6 " V A I J A - L-L.-I .... L . -I- I .1_;.1__AJL.

DISK RESPONSE TO DISTURBANCE, 6" VALVE

NATURAL FREQUENCY TESTS

FIGURE 31
N 4

Z: 1:

z

1u0 3INC

v 4 INC

0 6 INC

0

0 2 4 6 8 10

FLOW VELOCITY, FPS

: FROM REF 10

DISK NATURAL FREQUENCY

FIGURE 32

60
Topping, Stop @ 73 deg, 17.5 fps

4000 Sanples/Second (3LUT7D004 sec. set) 30

28

26

24

22

20

-J.0

18

16

14

12

10 ---

8~1--

41-

0 4

0.03200 0.04000 0.04800 0.05000 0.400 0.07200 OD.80

Time, seconds

DISK-TO-STOP CONTACT LOADS, 3" VALVE

FIGURE 33

61
I

APPENDX A

FORMULATION OF SWING DISC CHECK VALVE

DISC NATURAL FREQUENCY
-APPENDIX A

FORMULATION OF SWING DISC CHECK VALVE

DISC NATURAL FREQUENCY

The natural frequency of a single degree of freedom spring-mass system is given by:

_ 1 '/ K-tiff

fn 2 e

where fn = natural frequency in hertz

Kjtiff = system spring rate, force/displacement

M = system mass

For the check valve, effective spring rate can be expressed as the force required to open

the disc to some equilibrium angle, 0. An-expression for this force as developed ir

Reference 5 is:

F = kApV2 sinO sin0/2 -

where F = fluid force acting on the disc

k = an empirical constant dependent on valve geometry

A = disc area

V = fluid mean velocity

0 = fluid impingement angle (see Figure 1)

K~tiU = 1 d.F where R = radius from hinge pin to disc I

R dOd

Ktiff= 1 kApV2 d (sin nsine)

R dO 2 iG

K~tifr= kR (ose sin2 +-2 c sing)

Finally,

1 .:/kApV2 Z f =[ = n R~M s

where Z = cosO sin 0 + 2cos 2sinG]

63
From the results of baseline tests reported earlier in this document, the relationship

between velocity and 0 can be measured. These values are tabulated below for both

valves, along with the resulting natural frequencies. Pertinent valve data are:

A (ft2)

p ('b/ft 3)

R (ft)

M (lbs)

k

3'

0.067

62.4

0.23

2.3

2.0

6"

0.26

62.4

0.42

9.5

2.0

(water, 70'F)

(1/2Mhinge arm + Mdjsc)

3" VALVE

Velocity

fp S

0

d e z fn

hertz I -c I.

4

6

8

10

12

48

30

25

22

20

0.61

0.47

0.40

0.36

0.33

1.98

2.60

3.20

3.80

4.36

6" VALVE

Velocity

fps

0

d e L

z fn

hertz I J. I

4

6

8

10

12

53

34

27

24

22

0.63

0.51

0.43

0.39

0.36

1.44

1.94

2.38

2.83

3.27

64
APPENDIX B

M-AETHOD FOR ESTIMATING IMPACT FORCE

FROMTESTRESULTS

65
APPENDIX B

METHOD FOR ESTIMATING IMPACT FORCE

FROM TEST RESULTS

Check valve testing uses a plexiglass bonnet for test observation of disc motions. The

stiffness of this plexiglass bonnet is different from the actual steel bonnet stiffness;

therefore, the time history of impact force measurements from the tests will be different

from the actual disc impact in service for the same amount of disc energy before

impact. However, this difference can be adjusted based on the following estimations:

ACTUAL VALVE ASSEMBLY

I DI~SC

TEST ASSEMBLY

/1M7PA67~ 1WrFAS41RCM,;AJ7-

PL PX6,LA&S

The assemblies can be described in the following math models:

ACTUAL VALVE ASSEMBLY TEST ASSEMBLY

.EK,

COA818/JED

480/u,4cF7 AC.S.

.S ri FAIZ;S ~'AI A I

-

KM D/5SC,

I "/uFT/C KE 

4PNsfls Y,

or 04rC

5

KIrCOAIN/MED0

K, oMN~ver ASJ.

0p1E-X16CA34 00,-Al)

/<2 D IsCr

M4 -DIC MAlS,

'-1A(e A'I.V;n7C 19AG'6 Y

C.C DISC

The stiffnesses can be estimated from the design dimensions and material properties.

Then from the test results we can estimate the impact in the actual valve assembly as:

66
Let IKE = Kinetic energy before impact

KA = Combined stiffness for actual valve assembly

1 1.

K, K2

KT = Combined stiffness for test assembly

1

1 1

At peak of impact force, the KE is completely converted to the potential energy, PE,

(assuming no permanent deformation as:

PE =KE K 2=2 K( P ) I 1

2 2 'K/ 2 K

where P = Peak force

K = Combined stiffness

If PA is the actual impact force and PT is test measured peaX force, then we have

-~~~ 2 ~2

PE = KE 1 PA 1 PT

2 KA 2 KT

PA= PTASJ 1 F

KT ,

I.~~ ~ ~ U

67
Bonnet Stiffness Calculations

3" Test Valve Assembly

z 2

EA 30 x 106 x X(3/8) K, = LC = - 7 L ~~~7

= 473,343 lb/in +-

30xl1x -(3/16)2 EA 4 =i L - 0.5

= 1,656,699 lbrin -

K3 = 200= 200,000lb/n 4- .001

6 EA 30x10x

K4 =LL1

= 70,685,835 lb/in 4-

30 x106 x(3/4 3/82) < 7K- K =EA 4 L 2.5

= 3,976,078 lb/in

I

IF _EA _

L

30 x 106 x - (2 1/4 _ 13/16) 4

1.5

= 69,151,854 Wb/in

68
K7: vCentral Load Assumed

C i, E 0 --1I

7.25^+~~125

CastAcryicE = 0.4 x IOC psi

-W a2

Ymx= 16xD a radians

-_____ I 

i ( -' 2 1)

100(7.25 )2

16 xc 30 x 10 1~

12(1 -0.3 2)

= 0.00004872 in

.: K7 = 20,524,968 Wlmin

KT = 1

K 2 K 2 K7

1

10-' (2.1126 + 0.6036 + 5 +0.014 +0.252 + 0.0145 + 0.049)

= 10-6x8.0457

KT = 124,2901Min -

AE 4-(0.6875) x 30 x 10 KDisZ = 1.5 = 7,424,467 lbin

>> KT

If KDisc is close to the actual total stiffness of the disc and bonnet assembly (i.e.

KBonnet >> KDisc) then

Formual - F~easured is e

7,424,467 0lb 3124,290

69
NAC FORM 234 U S. NUCLEAR REGULATORY COMMISSION I REPORT NUMBER IA.s+.d v nIoC Ja" vo N0* R*YI

NACM 1102. N~~~~~~~~~~~~~~IUREG/CR-5 159 0RCM 1302 BIBLIOGRAPHIC DATA SHEET KEI No. 1559

SEE INSTRUCTiCNS ON THE REVERSE

2 TITLE ^ND SUBTITLE 3 LEAVE BLANK

Prediction of Check Valve Performance and

Degradation in Nuclear Power Plant Systems . OATEREPORTCOMPLETED

5 AUTORMSI April 1988 4 DATE REPORT ISSUED

M. S. Kalsi, C. L. Horst, J. K. Wang MONTH YEAR

May 1988

? P(RPORMiNG ORGANIZATION NAME AND MAILING ADORESS FZ Cc W I PROJECTITASKAVORK UNIT NUMBER

Kalsi Engineering, Inc. SPIN OR GRANT NUMBER

745 Park Two Drive

Sugar Land, Texas 77478 FIN D2042

1 SPONSORING ORCGAN1ATION NAME ANODMAILINGADODRESS $ft.d.Zd oCJI 1 I TYPE OF REPORT

Division of Engineering Office of Nuclear Regulatory Research Technical/Final

U.S. Nuclear Regulatory Commission b PERIOD COVEREO (tIIu- /

Washington, DC 20555 9/87 - 4/88

12 SUPPLEMENTARY NOTES

I} ABSTRACT (2OOWEI.l .hj

Degradation and failure of swing check valves and resulting damage to plant equipment

has led to a need to develop a method to predict performance and degradation of

these valves in nuclear power plant systems. This Phase I investigation developed

methods which can be used to predict the stability of the check valve disk when

piping disturbances such as elbows, reducers, and generalized turbulence sources

are present upstream of the valve within 10 pipe diameters. Major findings include

the flow velocity required to achieve a full open, stable disk position, the magnitude

of disk motion developed with these upstream disturbances with flow velocities below

full open conditions, as well as disk natural frequency data which can be used to

predict wear and fatigue damage. Reducers were found to cause little or no performance

degradation. Elbow effects must be considered when located within 5 diameters of

the check valve, while severe turbulence sources have significant effect at distances

to 10 diameters.

Clearway swing check designs were found to be particularly sensitive to manufacturing

tolerances and installation variables making them likely candidates for premature

failure. Reducing the disk full opening angle on these clearway designs results

in significant performance improvement.

14 DOCUMENT ANALYSIS -* KEYWOROSIDESCRIPTORS I5 AVAILABILITY

Swing check valves, check valve degradation, disk fluctuations, upstream STATEMENT

disturbance effects Unlimited

S6 SECURITY CLASSIFICATO%

IT... pf.

B IOENTIPIERS OPEN ENDED TERMS Unclassified

Unclassified

1 NUMBER O, PAGES

iB *0RCE

*U.S.GOVERNMCNiT PRINiTINGC OVICE.¶9a8-202.292MO0SS
-

UNITED STATES

NUCLEAR REGULATORY COMMISSION

WASHINGTON, D.C. 20555

OFFICIAL BUSINESS

PENALTY FOR PRIVATE USE, $300

POSAGEEr EESPAID

USNRC 1

PERMT No.G-6 J

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