NUREG-1367

Functional Capability of

Piping Systems

U.S. Nuclear Regulatory Commission

Office of Nuclear Reactor Regulation

D. Terao, E. C. Rodabaugh
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NUREG-1367

Functional Capability of

Piping Systems

Manuscript Completed: October 1992

Manuscript Completed: October 1992

Date Published: November 1992

D. Terao, E. C. Rodabaugh

Division of Engineering

Office of Nuclear Reactor Regulation

U.S. Nuclear Regulatory Commission

Washington, DC 20555

ABSTRACT

General Design Criterion 1 of Appendix A to Part 50 of

Title 10 of the Code of Federal Regulations requires, in

part, that structures, systems, and components important

to safety be designed to withstand the effects of earthquakes without a loss of capability to perform their safety

function. The function of a piping system is to convey

fluids from one location to another. The functional capability of a piping system might be lost if, for example, the

cross-sectional flow area of the pipe were deformed to

such an extent that the required flow through the pipe

would be restricted.

The objective of this report is to examine the present

rules in the American Society of Mechanical Engineers

Boiler and Pressure Vessel Code, Section III, and potential changes to these rules, to determine if they are adequate for ensuring the functional capability of safetyrelated piping systems in nuclear power plants.

iii NUREG-1367

CONTENTS

Page

A B ST R A CT ................................................................................ iii

NOMENCLATURE .................................................................. vii

1 INTRO D U CTIO N ....................................................................... 1

2 BA CK G R OU ND ........................................................................ 3

2.1 Present Code R ules ................................................................... 3

2.2 Nuclear Regulatory Commission's Position on Piping Functionality ........................... 3

2.3 Nuclear Regulatory Commission Piping Review Committee Report ......................... 4

2.4 Relevance of Tests to Piping Functional Capability ........................................ 4

3 BEANEY DYNAMIC LOADING TESTS ON STRAIGHT PIPE ............................... 5

3.1 Relationship Between Accelerations and Moments ........................................ 5

3.2 Comparisons with Theoretical Limit Moments ............................................ 5

3.3 Comparisons with Elastic Analysis ....................................................... 5

3.4 W eight Stresses ......... ............................................................. 6

3.5 Pressure Stresses ............................................................. 7

4 ELECTRIC POWER RESEARCH INSTITUTE, NRC, AND GENERAL ELECTRIC

COMPANY TESTS OF PIPING.COMPONENTS ............................................ 9

4.1 Scope of Tests and Reported Results .... ............................................... 9

4.2 Comparisons with Theoretical Limit Moments ........................................ 9

4.3 Comparisons with Elastic Analysis ....................................................... 10

4.4 W eight Stresses ...................................................................... 13

4.5 Pressure Stresses ..................................................................... 13

4.6 T ests30 and 37 ....................................................................... 13

5 ELECTRIC POWER RESEARCH INSTITUTE, NRC, AND GENERAL ELECTRIC -

COMPANY TESTS OF PIPING SYSTEMS ................................................. 15

5.1 Piping System Configurations and Materials ............................................. 15

5.2 Loadings ............................................................................ 15

5.3 Comparisons with Theoretical Limit Moments ............................................ 15

5.4 Comparisons with Elastic Analysis ...................................................... 16

5.5 W eight Stresses ...................................................................... 17

5.6 Pressure Stresses ............. ............................................... 17

6 OTHER PIPING SYSTEM TESTS ......................................................... 19

6.1 Hanford Engineering Development Laboratory Tests (Reference 14) ......................... 19

6.2 R eferences 15-20 Tests ................................................................ 20

6.3 Summary of Other Piping System Tests ................................................. 21

7 OTHER DYNAM IC LOADS .............................................................. 23

7.1 Fluid H am m er ....................................................................... 23

7.2 Relief-Valve Actuation ............................................................... 23

7.3 Postulated Pressure Boundary Breaks .................................................... 23

7.4 V ibrations ........................................................................... 24

8 SUMMARY AND LIMITATIONS ......................................................... .25

8.1 Reversing Dynam ic Loads ............................................................. 25

8.2 O ther Dynam ic Loads ................................................................. 27

v NUREG-1367
9 CO N CLU SIO N S ......................................................................... 29

9.1 Functional Capability Assurance, Present Code Requirements ............................... 29

9.2 Functional Capability Assurance, Future Code Requirements ............................... 29

10 REFERENCES ................................................................ 31

FIGURES

1 Test A rrangem ent ........................................................................ 33

2 Response Versus Input Acceleration ........................................................ 34

3 Strain at Pipe Midspan Versus Input Acceleration ............................................. 35

4 Deformed Shape and Permanent Strain After Tests ........................................... 36

5 M ean Strain Versus Input Acceleration ...................................................... 37

6 Deformed Shape of Upper Surface of 103-mm Pipe, Test 16 .................................... 38

7 In-Plane Elbow Test Arrangements, Tests 1, 3-8, 13, 19, and 31 ................................. 39

8 4 NPS, Sch. 40 Stainless Steel Pipe, Test 15 .................................................. 40

9 6 NPS, Sch. 40 Carbon Steel Pipe, Test 34 ................................................... 41

10 6 NPS, 9-in. Bend Radius, Sch. 10 Stainless Steel Elbow, Test 3 ................................. 42

11 6 NPS, 6-in. Bend Radius, Sch. 40 Carbon Steel Elbow, Test 13 ................................. 43

12 Elbows: Static In-Plane Closing Moment Capacity and Dynamic In-Plane Moment

C apacity T ests ........................................................................... 44

13 Test 20 Configuration: 4 NPS Nozzle in 12 NPS Vessel ........................................ 45

14 Test Arrangement, Tests 30 and 37 ......................................................... 46

15 Piping System 1 Configuration [Material: Carbon Steel (A106-B)] ............................... 47

16 Piping System 2 Configuration [Material: Stainless Steel (Type 316)] ............................. 48

17 Load Measurement Device at Sleds 2 and 4 .................................................. 49

18 Piping System 1, W eight Stresses, Swr ...................................................... 50

19 Piping System 2, W eight Stresses, Sw r ...................................................... 51

20 Hanford Engineering Development Laboratory Piping System ................................... 52

TABLES

1 Beaney (Refs. 7, 8, 9, and 10) Straight Pipe Tests: Materials, Yield Strengths, Dimensions,

Sinusoidal Input Test Frequencies, Pressures, and Test Planes ................................... 53

2 Beaney (Refs. 7, 8, 9, and 10) Straight Pipe Test Results Evaluated in Relation to Elastic

Analysis, 2% or 5% Dam ping .............................................................. 54

3 Reference 11 Pipe Tests: Limit Moments and Measured Moments .............................. 56

4 Reference 11 Elbow (6 NPS, 900) Tests: Limit Moments and Measured Moments ................. 57

5 Reference 3 Static and Dynamic In-Plane Moment Capacity Tests on Elbows ...................... 58

6 Reference 11 Pipe Tests: Comparisons with 2Sy Limit ......................................... 59

7 Reference 11 Elbow (6 NPS, 900) Tests: Comparisons with 2Sy Limit ............................ 60

8 Reference 11 Tests on Other Components: Comparisons with 2Sy Limit .......................... 61

9 Reference 13 and Other Piping System Tests: Comparison with Elastic Analyses ................... 62

NUREG-1367 vi
NOMENCLATURE

B 1, B2, Ba,, B2 r = Code stress indices

D = pipe mean diameter

Do = pipe outside diameter

Mi = resultant moment, used in Code Equation (9)

ML = calculated limit moment

Mm = maximum measured moment

P = internal pressure

S = calculated stress based on elastic response

spectrum analysis with +/-15% peak broadening and with either 2% or 5% damping

Sh

Sm

Sp.

= Code allowable stress, Class 2 piping

= Code allowable stress intensity, Class 1 piping

= stress due to internal pressure PDo/2t

Sw = stress due to weight

Sy = yield strength of material

t = wall thickness

Z = nominal section modulus of piping component

Other symbols are defined where used in text or tables.

vii NUREG-1367

1 INTRODUCTION

General Design Criterion 1 of Appendix A to Part 50 of

Title 10 of the Code of Federal Regulations (10 CFR)

requires that structures, systems, and components important to safety be designed to withstand the effects of

earthquakes without a loss of capability to perform their

safety function. The function of a piping system is to

convey fluids from one location to another. Sizing of the

pipe usually involves a compromise as size increases

between increasing installed costs and decreasing pressure drop. Functional capability of a piping system might

be lost if, for example, displacements were large enough

to "crimp" a pipe cross section and thus reduce the flow

area..

The Code* does not address the functional capability of

piping systems; rather, it addresses pressure boundary

integrity. Accordingly, it does not necessarily follow that

meeting Code rules will ensure functional capability.

The objective of this report is to examine present Code

rules, and potential changes to these rules, to see if they

are sufficient to ensure maintenance of functional

capability.

"Code" as used in this report refers to the American Society of Mechanical Engineers (ASME) Boiler and Pressure Vessel Code (Ref. 1).

Portions of the Code are identified as they appear in the Code (e.g.,

NB-3652). For the purpose of this report, NC-3600 (Class 2 piping)

and ND-3600 (Class 2 piping) are identical; hence, reference is to

NC-3600 for Class 2 piping.

1 NUREG-1367

2 BACKGROUND

2.1 Present Code Rules

Primary loads, such as internal pressure and weight, in

combination with other loads such as those due to earthquakes are controlled* in the Code by Equation (9) in

NB-3652 (Class 1 piping) and Equations (8) and (9) in

NC-3652 and -3653.1 (Class 2 piping). These Code equations are

B 1PD 0 /2t+B 2Mi/Z < lesser of XxSx or YxSy (1)

The symbols are defined in the "Nomenclature" section

of this report. The values of X and Y are

Class 1 Piping Class 2 Piping

Condition X Y X Y

Design 1.5 --- 1.5 ---

Level A --- --- 1.8 1.5

Level B 1.8 1.5 1.8 1.5

Level C 2.25 1.8 2.25 1.8

Level D 3.0 2.0 3.0 2.0

In Equation (1), Sx = Sm (allowable stress intensity) for

Class 1 piping, and Sx = Sh (allowable stress) for Class 2

piping. Values of Sm are usually greater than those of Sh.

For example, for SA106 Grade B carbon steel at 500'F,

Sm = 18.9 ksi, while Sh = 15.0 ksi. However, for austenitic stainless steels at elevated temperatures, Sm is

almost the same as Sh; for example, for SA312 Type 304

stainless steel at 650 OF, Sm = 16.2 ksi, and Sh = 15.9 ksi.

The material yield strength, Sy, is 17.9 ksi; thus, SmISy =

ShISy = 0.9.

It should be emphasized that the resultant moment amplitude, Mi, includes both steady-state loads, such as

weight, and dynamic loads, such as those caused by earthquakes. In Level D applications, the dynamic loads have

usually been the major contributor to Mi. However, increasing the Level D stress limits is being considered. This

possibility, along with the use of higher (e.g., 5%) damping in evaluating the response of piping systems to dynamic loads, makes it more important to recognize that

Mi represents combinations of steady-state loads with

dynamic loads.

*In NB-3658 and NC-3658, rules are given for the analysis of flanged

joints. These rules are based on the prevention of excessive leakage at

the joints. Because loss of functional capability of a flanged joint (without loss of pressure boundary integrity) is deemed to be incredible, the

rules for flanged joints are not considered any further in this report.

2.2 Nuclear Regulatory Commission's

Position on Piping Functionality

In the early 1970s, the stress limit of 3Sm was considered

to be quite high, relative to prior stress limits used in

piping design. For example, the industrial piping code,

USAS B31.3-1967 (Ref. 2), permitted stresses of 1.2xSh

for loadings acting not more than 1% of the time. (Earthquake loadings fit in this category.) The concerns of the

U.S. Nuclear Regulatory Commission (NRC) related to

functional capability of piping with the 3Sm limit resulted

in the preparation of NUREG/CR-0261 (Ref. 3).

Reference 3 includes summaries of available data on

static load capacities of straight pipe, elbows, branch connections, tees, and other piping components. In this reference, several changes in B-indices were suggested:

(1) Restrict application of B-indices to Do/ t < 50 (because of the buckling of straight pipes with

Do/t > 50).

(2) Decrease B1 for elbows from 1.0 to 0.5.

(3) Decrease B2 for elbows from 0.75xC2 to 0.67xC2 .

(4) Decrease B•, for branch connections from 0.75XC2b

to 0.50xCb.

However, the data in Reference 3 indicated that Equation

(1), 3Sm limit, as applied to straight pipe (B1 = 0.5, B2 =

1.0), was the least defensible from the standpoint of static

load capacity. For straight pipe, limit load theory (confirmed by cited tests) gives the bending moment, ML, at

zero pressure of

ML = (4/kr) ZSy (2)

For austenitic stainless steels at elevated temperatures,

Sm = 0.9Sy. Equation (1) with a 3Sm limit would permit

application of a moment of 2.7/(4/1r) = 2.1 times the

static bending limit moment. In Reference 3, it was suggested that the Level D limit be made the lesser of 3Sm or

2 Sy.

From the standpoint of functional capability, the 2Sy limit

is not defensible if Mi in Equation (1) comes from static

loads such as weight or steady-state relief valve thrust.

Thus, Reference 3 indicated that even the 2Sy Level D

limit was not clearly defensible for assurance of functional capability.

NRC Standard Review Plan Section 3.9.3 (Ref. 4), Appendix A, states:

2.3 Functional Capability

The design of Class 1, 2, and 3 piping

components shall include a functional capability

3 NUREG-1367
assurance program. This program shall demonstrate that the piping components, as supported,

can retain sufficient dimensional stability at service conditions so as not to impair the system's

functional capability. The program may be based

on tests, analysis, or a combination of tests and

analysis.

The Mechanical Engineering Branch of NRC's Office of

Nuclear Reactor Regulation prepared an interim technical position on the functional capability of essential piping

systems* to serve as a guide for applicants in preparing

their functional capability assurance programs. In the

interim technical position, the staff indicated that meeting Equation (1) with Level C limits was sufficient assurance of functional capability for components with

Do/t < 50. The applicant was to provide additional demonstration for components using Level D limits and for

components with Do/t > 50.

During specific plant reviews, applicants submitted other

methods of demonstrating functional capability to the

NRC. Among these was "Functional Capability Criteria

for Essential Mark II Piping" (Ref. 5), which included

guidance for components with Do/t > 50.

It is apparent that functional capability assurance requires, in addition to the Code rules, another set of evaluations. An ideal solution to the problem would consist of

evidence that meeting the Code rules for piping (with

modifications discussed later) would also ensure the functional capability of piping systems.

2.3 Nuclear Regulatory Commission

Piping Review Committee Report

Starting in early 1983, the NRC Piping Review Committee reviewed nuclear power plant piping in the context of

current regulations, regulatory guides, standard review

plans, and other pertinent documents. The results of the

review were published in late 1984 and early 1985 in

NUREG-1061 (Ref. 6), which consists of five volumes.

*Essential piping systems are piping systems that are necessary (1) for

safe shutdown of the plant and for maintaining the plant in a safe shutdown condition or (2) for preventing or mitigating the consequence of

an accident that could result in potential offsite exposures exceeding the

guidelines of 10 CFR Part 100. Piping systems that are not essential do

not require a functionality evaluation.

Functional capability of piping is discussed in Section

2.8.5 of Volume 2 of NUREG-1061. By 1984, sufficient

earthquake-type-loading-test data were available to indicate that earthquake loadings on piping systems, in the

absence of high static loads, would not cause "collapse"

(large plastic deformations) of piping systems. A staff

consultant suggested that functionality capability could

be ensured by meeting Equation (1) with Level D limits

(lesser of 3Sm, 2Sy), provided at least one-half of the

stress in Equation (1) came from earthquake-induced

loadings.

The Piping Review Committee, at that time, was not

ready to endorse the consultant's recommendation and

recommended the following:

The functionality criterion for piping will be

maintained. Current ASME Code Class 1 or

Class 2 stress evaluation procedures, not to exceed Level C limits, will be used. These limits are

similar to those now being used on a case-by-case

basis to satisfy the functionality criterion. It is

recommended that the upcoming EPRI [Electric

Power Research Institute]/NRC pipe tests be

evaluated to confirm that position and to determine whether it is appropriate to use the current

higher Level D stress limits.

The EPRI/NRC tests have now been completed; see Sections 4 and 5 of this report. The remainder of this report

consists of evaluating the EPRI/NRC tests, along with

other dynamic loading test data, to determine whether it

is appropriate to use the current Level D stress limit for

ensuring the functionality of piping systems.

2.4 Relevance of Tests to Piping

Functional Capability

A significant aspect of the test data is that, with one

exception discussed in Section 4.6, none of the tests resulted in loss of functional capability. Thus, the staffs

evaluations are based on the premise that the test data

provide lower bounds on combinations of steady-state

(e.g., weight) and dynamic loadings that will not cause loss

of functional capability. This lower bound premise may

introduce conservatisms in the staff's recommendations.

But, as will become apparent in the following discussions,

the premise leads to a significant relaxation of the present

NRC position on functionality.

NUREG-1367 4
3 BEANEY DYNAMIC LOADING TESTS ON STRAIGHT PIPE

E. M. Beaney of the Berkeley Nuclear Laboratories in the

United Kingdom has conducted a series of dynamic loading tests on straight pipes, on straight pipes with stress

concentrations, and on straight pipes with discrete components. The reports by Beaney of particular relevance to

functional capability are References 7, 8, 9, and 10.

3.1 Relationship Between

Accelerations and Moments

Figure I illustrates the test arrangement used by Beaney.

Table 1 is a summary of the material types, material yield

strengths, and pipe dimensions. A sinusoidal dynamic

input was applied to the pipe ends as indicated in Figure

1. The tests of direct interest herein were run with the

sinsusoidal input frequency equal to the first mode natural frequency of the pipe; that frequency is shown in Table

1. Some tests were run with internal pressure in the pipes,

as indicated in Table 1.

Figure 2 shows the test results from Reference 8. The

input amplitude was increased to about 3g; the response

acceleration, gr, at the midspan of the pipe (see Figure 1)

was measured. The relationship between moment and

response acceleration derived by Beaney is

M = [386EI/(4L2 f

2 )]gr (3)

where M = moment at midspan of pipe, in.-lb

E = modulus of elasticity, psi (3E + 7 psi used

herein)

I = moment of inertia of pipe cross section, in.4

L = length of pipe, in. (see Figure 1)

f = frequency of input during testing, Hz

gr = response acceleration

For example, Equation (3) as applied to Test 1 of Reference 8 gives

M = 386x3E7x0.01263/(4x147.2 2

x52)

= 67.5 in.-lb per unit gr

Having a relationship between M and gr, the gr corresponding to the theoretical limit moment, ML, can be

calculated as follows. The limit moment (in.-lb) for

straight pipe is

ML = D2tSy (4)

where D = pipe mean diameter, in.

t = pipe wall thickness, in.

Sy = yield strength of pipe material, psi

and g rL , the response acceleration corresponding to ML,

is

g rL = ML/(M/g r) (5)

For example, Equations (3), (4), and (5) as applied to Test

1 of Reference 8 give

grL = 0.9642 2x0.03583x43200/67.5 = 21.3

(g units)

3.2 Comparisons with Theoretical

Limit Moments

Figure 2 showsg rL for each of the five tests of Reference

8. It can be seen in this figure that, sincegrL corresponds

to ML, the limit moment is an approximate upper bound

to the moment that could be sustained in these dynamic

loading tests. (Test 5 is anomolous in that the applied

moment did not exceed about 65% of the limit moment.)

Equation (1), for zero pressure, 2Sy limit, permits the

application of a bending moment that is about 1.6 times

the limit moment. Thus, the results shown in Figure 2

present a paradox: If applied moments in a piping system

are accurately calculated, then Equation (1), with a 2Sy

limit, does not place any limit on input accelerations.

Of course, to accurately calculate moments due to dynamic loads that are high enough to cause gross plastic

response, an elastic-plastic analysis would be required.

3.3 Comparisons with Elastic Analysis

An elastic-plastic analysis of piping systems is within the

state-of-the-art. However, in the past and, the staff believes, foreseeable future, for piping system analysis, an

elastic analysis has been and will continue to be used and,

for earthquake loadings, an elastic response spectrum

analysis with +1-15% peak broadening and not more

than 5% damping. Thus, it is pertinent to evaluate

Beaney's test results in relation to elastic analysis, as

described below.

Beaney's tests were run with an essentially constant frequency input. The input "response spectrum" is a singlevalue acceleration at the test frequency; peak broadening

is meaningless. Because the pipe response is similar to

that of a single-degree-of-freedom dynamic structure,

5 NUREG-1367
g, = gI2ý (6)

where gr= acceleration at pipe midspan

g = input acceleration

= damping factor

Figure 2 shows lines representing 1/2%, 1%, 2%, and 5%

damping. It can be seen in this figure that, for low-level

input, the responses correspond to about 1% damping.

However, for high-level input, the response is much less

than that indicated by an elastic analysis, even for 5%

damping. It is this aspect of an elastic analysis that makes

Code Equation (9) [Equation (1) herein] highly conservative for reversing dynamic loads.

Equation (1), for zero internal pressure, B2 = 1.0

(straight pipe), in conjunction with Equations (3) and (6),

can be written as

MIZ = S = (M/g,)g/2Z (7)

For example, Equation (7) as applied to Test 1 of Reference 8 for the highest test level ofg (= 3.6), 2% damping,

gives

S = 67.5x3.6/(2x0.02526x0.02) = 240,500 psi,

amplitude

For comparison with the Code Level D limit, the 2 Sy limit

(not 3Sm) will be used because Sy relates directly to limit

load theory. The ratio of S to 2 Sy is thus a direct indication

of the test dynamic loadings to the dynamic loadings permitted by the Code with a Level D stress allowable of 2Sy.

For example, for Test I of Reference 8, Sy = 43,200 psi:

S/2 Sy = 240500/(2x43200) = 2.78

Thus, for Test 1 of Reference 8, the maximum input of

3.6g is equivalent to 2.78 times the Code Level D allowable.

Table 2 gives values of S/ 2 Sy for all of References 7, 8, 9,

and 10 tests. Because collapse did not occur in any of

these tests, the S/2Sy values in Table 2 suggest that, for

piping functional capability assurance, Equation (1) with

limits of about the following is appropriate, provided the

moment used in Equation (1) is almost entirely a reversing dynamic moment.

(2% damping) or 4 Sy (5% damping) only if the stress due

to weight or other steady-state stress does not exceed

about 0.15Sy.

3.4 Weight Stresses

Table 1 includes a column headed "Test Plane." A "V" in

this column indicates that the dynamic loading is in a

vertical plane as indicated by Figure 1. With this test

arrangement, the weight stress adds to the maximum

dynamic moment in the downward-displaced position. An

"H" in this column indicates that the actuators were rotated 90' from the plane indicated by Figure 1. With this

test arrangement, the maximum weight stress is 900 from

the location of maximum dynamic moment.

Figures 3 and 4, which show strain at pipe midspan and

deformed shape and permanent strain after tests, respectively, are from Reference 7. As indicated in Table 2, the

weight stress at the pipe midspan was 0.1 1Sy.. This weight

stress was sufficient to induce biased strains (Figure 3)

and a post-test deformed shape (Figure 4). This magnitude of deformation is well below that which will impair

functional capability.

Figure 5, which shows mean strain as a function of input

acceleration, is from Reference 8. The column in Table 2

headed "Sw/Sy" indicates the weight stresses at pipe

midspan. Other than Test 4, which showed the largest

SwISy and the highest mean strain, there is no obvious

correlation between SwISy and mean strain. In Test 4 high

mean strains of about 1.9% were developed; however,

these were not sufficient to indicate any significant loss of

functional capability. Figures 3, 4, and 5 serve as a warning that weight and other steady-state stresses must be

appropriately limited if Code Equation (9) with limits

such as lOSy (2% damping) or 4 Sy (5% damping) is to be

clearly defensible.

In Reference 9, Beaney mentions that "the pipe sags due

to the one sided effect of gravity," but does not give

quantitative data on the magnitude of the sagging.

Figure 6, which shows deformed shape of upper surface of

pipe, is from Reference 10. The buckling indicated in this

figure apparently occurred only in Test 16, during which a

pipe with Dolt = 103/1.5 = 69 was tested. The pipe was

filled with water. The combination of large Do/t, relatively high weight stress, and relatively low dynamic load

input (1.9g) led to the incipient buckling as depicted in

Figure 6. There is a bit of a mystery that Beaney noted but

did not explain: Why did signs of buckling occur in Test 16

but not in Test 15?

The onset of buckling could pose a challenge to maintenance of functional capability. Therefore, the staff recommendations in this report will be "hedged" by limiting

the applicability to Dolt < 50, that is, the same Do/t limit

imposed by the Code on the applicability of B-indices.

Analysis Damping, % Stress Limit

2 lOSy

5 4Sy

Specifically, the column in Table 2 headed "Sw/Sy"

(weight stress/yield strength) supports the use of lOSy

NUREG-1367 6
3.5 Pressure Stresses

Table 2, column headed "Sp/Sy," shows the ratios of circumferential stress due to internal pressure to the pipe

material yield strength.

Although internal pressure is significant with respect to

pressure boundary evaluation, Beaney's tests suggest that

internal pressure has little, if any, significance with respect to functional capability. Pressure stresses are, of

course, uniform around the circumference of the pipe

and, thus, do not bias the displacement direction, in contrast to weight stresses, which may bias the displacement

direction.

Code Equation (9) [Equation (1) herein] includes the

pressure term B1PDo/2t; its continued use is expected in

any foreseeable Code rule changes for pressure boundary

evaluations.

7 NUREG-1367

4 ELECTRIC POWER RESEARCH INSTITUTE, NRC, AND GENERAL

ELECTRIC COMPANY*TESTS OF PIPING COMPONENTS

4.1 Scope of Tests and Reported

Results

A total of 41 component tests were run. The types of

components included elbows, tees, reducers, straight

pipe, and fabricated branch connections. The tests are

described and the test results are given in Reference 11.

Reference 11 results were supplemented by data provided in a letter from H. Hwang (General Electric Company) to E. C. Rodabaugh dated October 16, 1991 (available in the author's personal file).

Figure 7 shows a representative test arrangement with an

elbow as the test component. Dynamic loadings were

applied by motions applied to the sled. Numerous runs

were made in each test. The run of main interest to

functional capability is (in most tests) an earthquake time

history applied to the sled, scaled up to the highest magnitude used in the test.

Reference 11 contains the results of measured moments

acting on the components. The measured moments were

derived from strain gages placed on the inertia arm (see

Figure 7). The inertia arm was sufficiently strong so that it

responded elastically in all tests. Thus, comparisons can

be made between measured moments. and theoretical

limit moments.

Of the results given in Reference 11, the most significant

with respect to functional capability consists of the calculated stress in the component at the highest magnitude of

sled input. These stresses were calculated using an elastic

response spectrum analysis. The response spectrum was

derived from the time-history input to the sled, using 2%

or 5% damping. The analysis is based on +/-15% peak

broadening of the so-derived response spectrum and

gives the moment acting on the component. The calculated stresses can be compared with the Code Equation

(9) stress limit of 2Sy. If the ratio of calculated stress to

2Sy is greater than unity, the test indicates that, for functional capability, an Equation (9) stress limit greater than

2Sy is defensible.

4.2 Comparisons with Theoretical

Limit Moments

4.2.1 Tests on Straight Pipe

Table 3 is a summary of Reference 11 results for what the

staff deems to be essentially straight pipe tests. Only

*Subcontracted by EPRI to evaluate test results.

Tests 33 and 34 were, in fact, tests on straight pipe. However, in the tests of tees and reducers, the plastic response

was essentially in the pipes, not in either the tees or the

reducers.

The column in Table 3 headed "M,,/ML" indicates that, as

in Beaney's tests on straight pipes, the theoretical limit

moment is about as much moment as could be applied

when pipes are subjected to very high level, simulated

earthquake-type dynamic loads.

For Tests 15 and 34, Reference 11 gives data for several

runs. Figures 8 and 9 are plots of Tests 15 and 34. The

calculated moment represents a measure of the magnitude of the input, analogous to the g-input of Figure 2.

The measured moment represents the response analogous to gr of Figure 2, convertible by means of Equation

(3) to a response moment. As in Figure 2, Figures 8 and 9

indicate a rapid increase in response moment at lowmagnitude inputs and a leveling off of response moment

at high inputs. Appendix A of Reference 11 includes a

column headed "DYN MOM/LIM MOM," where

DYN MOM = maximum measured dynamic

moment

LIM MOM = calculated static limit moment

It might seem that DYN MOM/LIM MOM should be the

same as MmIML in Table 3. This is approximately so,

except for the tee tests. For the tee tests, LIM MOM was

calculated in Reference 11 as

LIM MOM = D2tSy/B 2b (8)

where B2b is defined by the Code as 0.4(D/2t)2/3

The largest discrepancy exists for Test 11; DYN MOM/

LIM MOM = 2.4 compared with Mm../ML = 0.65. For

Test 11, D = 6.491 in., t = 0.134 in., and Sy = 39.7 ksi.

B2b = 3.35 and Equation (8) gives

LIM MOM = 6.4912x0.134x39.7/3.35 = 67 in.-kip

The DYN MOM used was that calculated at an imaginary

location defined as the "tee center"; DYN MOM = 158

in.-kip. Thus, in Appendix A of Reference 11, DYN

MOM/LIM MOM = 158/67 = 2.36. In Test 11, essentially all plastic response was confined to a narrow band of

the Schedule (Sch.) 10 branch pipe at its juncture with the

tee. Thus, in the staff's view, the ratio of 2.4 shown in the

appendix is misleading. From a functional capability

standpoint, however, the important aspect is that displacements were not sufficient to cause any loss of functional capability.

9 NUREG-1367
4.2.2 Tests on Elbows

Table 4 is a summary of Reference 11 elbow tests. The

limit moment was calculated as follows:

ML = 0.8hO'6Do2tSy for P = 0 (9)

ML = 0.96h 0 6 Do2tSy for P > 0

where h = elbow parameter = tR/r2

t = elbow wall thickness

R = elbow bend radius

r = elbow cross-section mean radius

The basis for Equation (9), for P = 0, is discussed in

Reference 3. It is based on an in-plane bending limit

moment, P = 0, theory developed by Spence and Findlay

(Ref. 12). The coefficient of 0.96 for P > 0 was suggested

by the staff (used in Reference 11) to approximate the

increase in moment capacity due to internal pressure.

Insofar as the staff is aware, no closed-form theory exists

for an elbow limit moment with P > 0, or for an out-ofplane or torsional moment. Existing elastic-plastic, finiteelement computer programs might be used; however, to

pick up the pressure effect and to distinguish between

in-plane closing and in-plane opening, such programs

would have to include finite displacement effects. Static,

in-plane moment tests show that the moment capacity for

in-plane closing is much less than for in-plane opening.

For Tests 3 and 13, Reference 11 gives data for several

runs. Figures 10 and 11 are analogous to Figures 8 and 9,

which, in turn, are analogous to Figure 2.

Figure 11 shows responses that are similar to those of

Beaney's straight pipe tests, that is, rapid rise in response

at low-magnitude input, followed by a leveling off of response at high-magnitude input. However, the leveling

off occurs at about two times ML, rather than in the

vicinity of ML. For dynamic equilibrium, the moment

capacity of the elbow cannot be exceeded in either the

closing direction or the opening direction. Thus, a paradox seems to exist.

For Test 13, the paradox is resolved by considering the

actual wall thickness of the elbow that was tested. In a

letter from H. Hwang (General Electric Company) to E.

C. Rodabaugh dated April 21, 1989 (available in the

author's file) regarding dimensional measurements of

Reference 11 test components, H. Hwang provided wall

thickness measurements of the elbows used in the component tests. The Test 13 elbow was nominally Sch. 40,

0.280-in. nominal wall thickness. The measured thicknesses ranged from 0.327 in. to 0.520 in. with an average

wall thickness of 0.425 in. Using the average wall thickness changes the calculated limit moment from 189 to 380

in.-kip. Then, Mm/ML = 400/380 = 1.05.

Thus, Test 13, evaluated using actual average wall thickness rather than nominal wall thickness, indicates that

Equation (9) is a good indicator of in-plane dynamic moment capacity for a carbon steel elbow.

The average actual wall thickness of the Test 3 elbow was

0.156 in. compared with the nominal wall thickness of

0.134 in. Using the average wall thickness changes the

calculated limit moment from 52.3 to 67 in.-kip. Then,

Mm/ML = 163/67 = 2.43. Obviously, the use of actual

wall thickness does not explain the seeming paradox for

Test 3.

However, evidence that the Mm/ML ratio for Test 3 is

credible can be seen in Figure 12. This figure includes

(1) static in-plane closing moment capacity test data

from Reference 3; see Table 5 herein

(2) dynamic in-plane moment capacity test data from

Reference 3; see Table 5 herein

(3) dynamic in-plane moment capacity test data from

Reference 11; see Table 4 herein

Figure 12 shows that Test 3 results are on the high side of

static test data, but are consistent with prior dynamic test

data; thus, the Test 3 results are credible.

At the other extreme of Mm/ML in Table 4, for Test 37,

Mm/ML = 1.31. This is consistent with static test data for

pipe elbows with zero internal pressure. Test 3 and Test

37 elbows had the same nominal dimensions and were

made of the same heat of stainless steel material.

Code Equation (9) [Equation (1) herein], for zero pressure, 2Sy limit, permits the application of a bending moment of about 1.6ML. Thus, the results for Test 37 indicate that,. if the applied moments are accurately

calculated, Code Equation (9) with a 2Sy limit does not

place any limit on dynamic (e.g., earthquake-induced)

loads.

4.3 Comparisons with Elastic Analysis

In Section 4.2, measured dynamic moments and limit

moments were compared. The staff will now compare

calculated stresses with a 2Sy stress limit. Calculated moments and/or stresses are given in Appendix B of Reference 11. These calculations are based on elastic response

spectrum analyses using either 2% or 5% damping and

+1-15% peak broadening. The response spectra used

were derived from the time-history inputs to the sled; see

Figure 7.

NUREG-1367 10
4.3.1 Tests on Straight Pipe

Table 6 is a summary of the results of Reference 11 tests,

which the staff deems to be equivalent to straight pipe

tests. The staff will use Test 9 as an example to illustrate

the significance of Table 6.

For Test 9, 2% damping, the calculated stress amplitude

is 589 ksi. The material yield strength is 40.8 ksi. Thus,

S/2Sy = 589/81.6 = 7.22. Bypassing, until later, the question of weight stress and pressure stress, Code Equation

(9) could be written as

B 1PDo/2t+B2M/Z < 14.44Sy

That is, looking only at Test 9, the Code limit of 2Sy could

be increased to 14.44Sy, and, since no loss of functional

capability occurred in Test 9, the increased stress limit

would ensure functional capability.

If the moments were to be calculated using 5% damping,

Test 9 indicates the Code Equation (9) limit could be

increased to 8.0Sy, but not necessarily any higher.

A salient point is that the defensible stress limit for Code

Equation (9) is highly dependent on how the moments

acting on the component are calculated. For Test 9:

Defensible Code

Moments Calculated Equation (9) Limit

Accurately, e.g., by elastic-plastic

analysis 1.3Sy

By elastic analysis, 2% damping 14.4Sy

By elastic analysis, 5% damping 8.OSy

Appendix A of Reference 11 includes a column headed

"INPUT X/LEVEL D,"

where INPUT X = calculated stress using linear

response spectrum analysis, 2%

damping, + /-15% peak broadening, and actual sled input.

Stress = B2M/Z.

LEVEL D = 3Sm = 60 ksi.

In the following, for brevity, this ratio is designated as

XID.

In Table 6, the analogous ratio is S/2Sy, 2% damping. In

the context of a meaningful evaluation of the tests, the

staff deems that use of Level D = 3Sm = 60 ksi is

inappropriate. A more meaningful ratio is obtained by

using Level D = 2Sy, where Sy is the yield strength of the

material used in the tested component. Because in these

tests 2Sy > 60 ksi, the staff's S/2Sy ratio is always less than

X/D.

For example, in Test 34 (pipe test), X/D = 731/60 = 12.2,

which agrees with the "12" shown in Appendix A of Reference 11. But, for 2% damping, S/ 2Sy = 731/(2x44.5) =

8.2 as shown in Table 6.

In addition, for those tests that involved tees, INPUT X

= B2M/Z, where, for example in Test 11, B2 (= B2b ) =

3.34 was used to calculate the X/D = 16 shown in Appendix A of Reference 11. Also, in calculating X/D = 16, the

calculated moment at the imaginary point at the centerline intersections was used. In its evaluations, since the

plastic response was confined to a narrow band of the

branch pipe at its intersection with the tee, the staff used

B2 = 1.0 for straight pipe with the calculated moment at

the branch-pipe-to-tee intersection weld. It thereby obtained S = 269 ksi and S/2Sy = 269/(2x39.7) = 3.4 as

shown in Table 6.

For all Table 6 tests, X/D and S/2Sy are as follows:

Test No. 9 10 11 12 14 15 16. 33 34 40

X/D 21 21 16 27 18 13 30 -- 12 22

S/2Sy 7.2 7.4 3.4 9.0 6.5 11 20 -7- 8.2 18

It is apparent that the staff's evaluations of Table 6 tests

are significantly more conservative (and, it believes, more

realistic) than the X/D ratios in Appendix A of Reference

11. Even so, the staff's evaluations support a significant

increase in the present Code Equation (9) limit insofar as

functional capability is concerned; for example, the lowest S/ 2 Sy of 3.4 suggests that the Code Equation (9) limit

could be increased from 2Sy to 6.8Sy, provided the applied moments are calculated using not more than 2%

damping.

4.3.2 Tests on Elbows

Table 7 is a summary of the Reference 11 tests on elbows

in the same format as that of Table 6.

The stress was calculated using

S = B2M/Z (10)

where M was calculated using elastic response spectrum

analyses, 2% or 5% damping and +/-15% peak broadening, and

B2 = 1.3/h 2/3 (11)

11 NUREG-1367
Equation (11) is from NB-3683.7 of the Code; h is the

elbow parameter as tabulated and defined in Table 4. The

B2 for each h involved in the tests is as follows:

h 0.41 0.25 0.17 0.11

B2 2.37 3.27 4.29 5.51

The maximum elastic stress in an elbow (with h > 1.0)

depends on the moment direction:

Moment Direction Multiplier of (1/h)2/3

In-plane 1.86

Out-of-plane 1.59

Torsion 1.00

The B2 of 1.3/h 2/3 is intended to represent a conservative

estimate of the moment capacity of an elbow subjected to

an in-plane closing moment. It is conservative for both

out-of-plane and torsion moments. However, an important aspect with respect to the staff's recommendations

for functional capability criteria is that they are based on

the B-indices as prescribed in the present Code. Any

future Code revision that would decrease any of the

B-indices might invalidate the staff's recommendations.

The ratios in Appendix A of Reference 11, column

headed "INPUT X/LEVEL D," are higher than those in

Table 7, S/2Sy, 2% damping, because Level D = 3Sm =

60 ksi is less than 2Sy. Although the staff's evaluations are

more conservative (and, it believes, more realistic) than

those in Reference 11, they still suggest that the present

limit on Code Equation (9), for functional capability

evaluation, can be increased significantly. The lowest

S/2Sy in Table 7, 5% damping, is 5.2. This suggests that

the 2Sy limit can be increased to 10.4Sy, even when using

5% damping in calculating the applied moments.

4.3.3 Tests on Other Components

Table 8 is a summary of the Reference 11 tests on other

components in the same format as that of Tables 6 and 7.

Reference 11 includes the results of 41 component tests.

Tables 6, 7, and 8 contain the results of the staff's evaluations of 33 of these tests. The eight tests not included in

Tables 6, 7, and 8, and comments concerning them, are

the following:

Test Comment

21, 22 Tests of lugs on pipe-relevant to pressure

boundary integrity but not to functional

capability

23 Test of elbow with strut restraint--relevant

to support loads but not directly to functional capability

24, 32 Static limit moment tests of elbows-results

more or less consistent with Reference 3

static limit moment tests

27 Midfrequency and sinesweep tests of a teeresults for this test not given in Appendix B

of Reference 11

28, 29 Water hammer tests -discussed in Section 7

herein

4.3.3.1 American National Standards Institute (ANSI)

B16.9 Tees

Tests 36, 38, and 39 in Table 8 are tests of 6x6x6 ANSI

B116.9 tees, the same type of components included in

Table 6, Tests 9, 10, 11, 12, and 14. However, the Table 8

tests are significantly different from the Table 6 tests, as

illustrated by the following sketch.

A Test 12

M -..- 0 - 10 M

[M

Table 8

Test 36

Table 6

Tests 9, 10, 11, 12,

14

Table 8

Tests 38, 39

In Tests 36, 38, and 39, plasticity and eventual fatigue

failure occurred in the body of the tees. Thus, it is deemed

appropriate to evaluate these tests using the B2 , (Tests 38

and 39) or B2 r (Test 36) specified in the Code:

Bb = 0.4(R/T)2/3 = 2.02;B 2r = 0.5(R/T)2/3 = 2.52

where R = mean radius of attached pipe (3.1725 in.)

T = nominal wall thickness of attached pipe

(0.280 in.)

4.3.3.2 Tests 18 and 20, Fabricated Branch

Connections

The staff's evaluations of Tests 18 and 20 require a more

detailed explanation. The Code-specified B2 , index for

branch connections per NB-3643 (see NB-3683.8) is

NUREG-1367 12
B•b = 1.5x3(R/T) 2/3(rlR)112(tl7)(r/rp) (12) For Test 20, Run 7, Appendix B of Reference 11 gives

where R = mean radius of run pipe

T = nominal wall thickness of run pipe

r = mean radius of branch pipe

t = nominal wall thickness of branch pipe

rp '! radius to outside of nozzle

For use in Code Equation (9),

M = 724 in.-kip, 2% damping

M = 410 in.-kip, 5% damping

These moments are used as Mb in Equation (13) to give

S = 3.416x724/3.21 = 770 ksi for 2% damping

S = 3.416x410/3.21 = 436 ksi for 5% damping

For Test 20, Bat = 7.79 and Zb = 5.9 in.3 were used in

Reference 11 without an explanation of their basis. They

are obviously not in accordance with the Code.

XID in Appendix Aof Reference 11 and S/ 2Sy, 2% damping, in Table 8 are as follows:

S = B2b (Mb /Zb ) (13)

where Mb = moment applied to branch

Zb = section modulus of branch pipe

Test 18

Test 18 of a pad-reinforced fabricated tee poses a problem because B-indices for pad-reinforced branch connections are not given in the Code. However, the staff believes that B•b for Test 18 can be bounded by using

Equation (12) with T = 0.322 in. (Sch. 40 run pipe) as an

upper bound and T = 0.322 + pad thickness = 0.644 in.

as a lower bound. For T = 0.322 in., Equation (12) gives

Bb = 1.5(4.1515/0.322)2/3(2.1315/4.1515)1/2

(0.237/0.322)(2.1315/2.25) = 4.12

Changing only the T of 0.322 to 0.644 in. gives

Bab = 1.30. In its evaluation of Test 18, the staff used an

average B2b of 2.7.

For Test 18, Run 6, Appendix B of Reference 11 gives

M = 915 in.-kip, 2% damping

M = 542 in.-kip, 5% damping

These moments are used for Mb in Equation (13) to give

S = 2.7x915/3.21 = 770 ksi for 2% damping

S = 2.7x542/3.21 = 456 ksi for 5% damping

For Test 18, B2b = 4.12 was used in Reference 11 without

an explanation of its basis (perhaps coincidentally,

B2 , = 4.12 can be obtained from Equation (12) for an

unreinforced fabricated tee). In Reference 11, the Codeprescribed Zb = 3.21 in. 3 was used.

Test 20 (see Figure 13)

As applied to Test 20, Equation (12) gives

B2 , = 1.5(6.1875/0.375)2/3(2.1315/6.1875)1/2

(0.237/0.375)(2.1315/2.25) = 3.416

Test No.

X/D

S/2Sy

18

20

7.2

20

16

7.9

36 38

15 20

9.9 15

39

21

16

As its evaluations in Tables 6 and 7, it is apparent that the

staff's evaluations in Table 8 are more conservative than

those in Reference 11. Even so, the staff's evaluations

suggest that, for functional capability evaluations, the

present limit on Code Equation (9) can be increased significantly. The lowest S/ 2 Sy in Table 8, 5% damping, is

4.3. This suggests that the 2Sy limit might be increased to

8.6Sy, even when 5% damping is used in calculating the

applied moments.

4.4 Weight Stresses

Tables 6, 7, and 8, column headed "Sw/Sy," show weight

stresses as ratios to yield strength, Sy. These ratios, except

for Tests 30 and 37 (elbows), are not more than 0.08.

Thus, they are of limited usefulness with respect to establishing a reasonable bound on weight stresses combined

with reversing dynamic stresses.

Tests 30 and 37 are discussed in Section 4.6.

4.5 Pressure Stresses

Tables 6, 7, and 8, column headed "Sp/Sy," show the

* nominal pressure stresses, PDo/2t, as ratios to yield

strength, Sy.

Although internal pressure is significant with respect to

pressure boundary evaluation, the data do not suggest any

decrease in functional capability for Sp/Sy ratios up to

0.48. Indeed, as discussed in Section 4.6, internal pressure

in elbows appears to increase their moment capacity. This

"inverse" pressure effect is also apparent in static tests on

elbows; see Reference 3.

4.6 Tests 30 and 37

Tests 30 and 37 were in-plane moment tests on nominally

identical elbows. The test arrangement is shown in Figure

13 NUREG-1367
14. The assemblies were "tuned" (height of vertical arm,

magnitude and location of weights, etc.) so that the first

mode response frequency was about 1.4 Hz. The timehistory input was adjusted so that the peak of the input

response spectrum was at about 1.3 Hz. The adjustment

was made by expanding the time of the time-history input;

a run for Tests 30 and 37 lasted about 110 sec, rather than

the about 20 sec for other simulated earthquake inputs.

The elbows were from the same heat of stainless steel

material with Sy = 34 ksi. They were 6 nominal pipe size

(NPS), Sch. 10, 9-in. bend radius.

The weight stress, at the mid-arc of the elbows, was 10.74

ksi for both elbows with SwISy = 0.32.

The only apparent testing differences were the following:

Test/Run Pressure S, 2% Damping

30/4 400 psi 620 ksi

37/5 0 651 ksi

Test 30 was ended when a fatigue failure occurred. Some

permanent displacements occurred (not quantified in

Reference 11) during the test runs, but it is believed that

none of these deformations were sufficient to reduce

functional capability.

Test 37 consisted of low-level Runs 1 and 2 and then

elbow cross section would have significantly decreased;

that is, functional capability would have been lost.

Although about 5% higher loadings were used in Test 37,

Run 5, than in Test 30, Run 4, the staff believes the major

difference is that Test 37 was run at zero pressure, while

Test 30 was run at 400-psi pressure. The measured moments in Test 37, Run 5, and Test 30, Run 4, were 57 and

112 in.-kip, respectively. This rather directly indicates the

increase in moment capacity due to an internal pressure

of 400 psi with Sp/Sy = 0.24.

Test 37 is a direct indication that a weight load (10.74-ksi

weight stress) that would not cause collapse by itself, in

combination with high reversing dynamic loads (S, 2%

damping = 651 ksi), does cause collapse.

However, Test 37 must be looked at in light of the following:

(1) The test was meant to be an extreme evaluation of

the concept that reversing dynamic loads do not

cause collapse. The pipe parameters selected for

this extreme case included in-plane moments (weakest direction), zero pressure (worst case for pressure), and thin-walled Sch. 10 pipe, for which

Do It = 49 (high Do/ t and pronounced elbow effects

with h = 0.11). The weight stress was such that

SwISy was 0.32 (significantly higher than the usual

weight stresses in piping systems). The very low test

frequency of about 1.3 Hz may have contributed to

the collapse in that 1.3 Hz gives the assembly more

time to displace before reversal of dynamic load

occurs.

(2) The dynamic loads were very high; that is, S/2Sy, 2%

damping, was 9.6, which is 9.6 times the present

Code Level D limit.

(3) Test 37 was a component test. In a piping system,

additional plastic hinges would have to develop before large plastic displacements could occur.

Nevertheless, Test 37 constitutes a "red flag" to indicate

that appropriate control must be placed on steady-state

loadings to avoid the possibility of loss of functional capability during application of high reversing dynamic loads.

I

Run

S, 2% damping, ksi

3. 4 5

79 324 651

During Runs 1 through 4, displacements were relatively

small. During Run 5, after about 45 sec into the run, the

assembly began to ratchet-displace in the elbow closing

direction. The test was terminated at about 72 sec into

Run 5 because the displacements were becoming large

and increasing rapidly with time.

At the termination of Test 37, the upper end of the inertia

arm had displaced several feet and, if the test had been

continued for a few more seconds, the displacements

would probably have increased to the extent that the

NUREG-1367 14
5 ELECTRIC POWER RESEARCH INSTITUTE, NRC, AND GENERAL

ELECTRIC COMPANY TESTS OF PIPING SYSTEMS

Two piping systems, identified as System 1 and System 2,

were tested. The system configurations, testing, and resuits are given in Reference 13. Reference 13 results are

supplemented by additional data provided in a letter from

H. Hwang (General Electric Company) to E. C.

Rodabaugh dated November 14, 1991 (available in the

author's file).

5.1 Piping System Configurations and

Materials

Figures 15 and 16 show the configurations of Systems 1

and 2, respectively.

System 1 was made of carbon steel (A106-B). It is characterized in Reference 13 as follows: "[System 1] was relatively balanced with regard to dynamic strain such that

several different locations had cyclic plastic strains of

about the same magnitude."

System 2 was made of stainless steel (Type 316). It is

characterized in Reference 13 as follows: "[System 2] had

unbalanced stresses with a single high-stress location

where failure was predicted to occur while the remainder

of the piping system was at a relatively lower stress."

5.2 Loadings

Both Systems 1 and 2 were tested with an internal pressure of 1000 psi.

Reference 13 describes the various time-history inputs

used in the system tests. From the standpoint of functional capability, the highest input is of primary significance for both systems; the highest input was associated

with "Time History B," with all sleds acting in unison.

5.3 Comparisons with Theoretical

Limit Moments

In principle, dynamic moments at any location in a piping

system cannot exceed the moment capacity at that location. This aspect of dynamic loading tests is discussed in

Sections 3.2 and 4.2. Comparisons of test measured moments with calculated moments are shown in Figures 1

and 8 through 11 and Tables 3, 4, and 5.

Reference 13 gives measured and calculated moments at

three locations as shown below:

Ref. 13

Table System Location

5-5 1 Node 72, 6 NPS short-radius

elbow

6-9 2 Node 6, 12x4 NPS nozzle, see

Figure 13

6-11 2 Node 52, 6 NPS, Sch. 40 pipe

Table 5-5 Table 6-9 Table 6-11

Mm Mc Mm Mc Mm Mc

66.1 92.5 39 53 110 150

330 634.7 78 96 255 293

725 3228.4 119 334 576 915

717 4994.1 156 572 681 1551

235 1091 837 2658

where Mm = measured moment, in.-kip

Mc = calculated moment, in.-kip, using 2%

damping, + 1-15% peak broadening

The data were sufficient so that plots of Mm versus Mc,

analogous to Figures 8 through 11, could be made. However, such plots are of little value because of the major

uncertainties discussed below.

Reference 13 does not describe how the moments were

measured. However, in a letter from H. Hwang (General

Electric Company) to E. C. Rodabaugh dated November

14, 1991 (available in the author's file), Hwang stated that

the measurement devices shown in Figure 17 were used in

both Systems 1 and 2. To the extent that applied moments

do not exceed the yield moment of the 6 NPS, Sch. 160

pipe on which strain gages were mounted, the strain

measurements can be used to calculate measured moments (e.g., at Node 72 in System 1).

The yield strength of the Sch. 160 pipe used in System 2,

according to Table 6-1 of Reference 13, is 31.3 ksi. The

yield moment of the load measurement device (Figure 17)

is then

My = SyZ = 31.3x17.81 = 557 in.-kip

Thus, the values of Mm from Table 6-11 greater than 557

in.-kip may reflect yielding of the measurement device

and not be an accurate indication of the measured moment. It is this uncertainty that makes comparison with

the limit moment of straight pipe questionable.

15 NUREG-1367
The moments cited from Reference 13, according to the

letter from H. Hwang to E. C. Rodabaugh dated November 14, 1991, are resultant moments; that is,

M = (Mx2 + My2 + Mz2)1/2

The moment capacity of elbows (Table 5-5) and nozzles

(Table 6-9) are significantly dependent on the orientation

of the applied moments. It is this uncertainty that makes

comparisons with limit moments of elbows or nozzles

questionable.

Accordingly, no meaningful comparisons can be made

between Reference 13 measured moments and either

limit load theory or tests or the component tests of Reference 11 discussed herein in Section 4.

5.4 Comparisons with Elastic Analysis

Reference 13 gives the following calculated moments:

Mc, in.-kip, Ref. 13 Sy, for Damping of

Table System Location ksi 2% 5%

5-5 1 Node 72, 43.8 4994

6 NPS shortradius elbow

6-9 2 Node 6, 12x4 35.7 1091 775

NPS nozzle

6-11 2 Node 52, 35.0 2658 1860

6 NPS Sch.

40 pipe

where Mc is the moment calculated by an elastic response

spectrum analysis, 15% peak broadening, damping of 2%

or 5% as indicated. These values are taken from the

indicated Reference 13 tables under the columns "Full

Sled-4 ARS" for System 1 and "FULL UNIF" for System

2. Sy is the material yield strength according to Reference

13, Tables 5-1 and 6-1. For the 12x4 nozzle, Sy is for the

12 NPS pipe.

The staff evaluated these calculated moments in a manner analogous to that for Tables 2, 6, 7, and 8, that is,

develop ratios of S/ 2 Sy, where S = B2MIZ and Sy = yield

strength of the material. The results are summarized as

follows:

Ref. 13 Sy, Z, 2% Damping 5% Damping

Table ksi B2 in. 3 S, ksi S/2Sy S, ksi S/2Sy

5-5 43.8 4.29 8.50 2520 29 - (16)

6-9 35.7 3.42 3.21 1162 16 826 12

6-11 35.0 1.00 8.50 313 4.5 219 3.1

For Table 5-5, a significant uncertainty existed concerning the actual wall thickness of the short-radius elbow.

This uncertainty goes back to the measured wall thicknesses that were available for Component Test 13, which.

indicated that the average actual wall thickness was 0.425

in. rather than the nominal wall thickness of 0.280 in.

However, in a letter from W. P. Chen (Energy Technology Engineering Center) to E. C. Rodabaugh dated

December 19, 1991 (available in the author's file), Chen

provided the measured wall thicknesses of elbows in Systems 1 and 2. For the short-radius elbow at Node 72, the

average wall thickness is 0.310 in., which is only 11% more

than the nominal wall thickness.

Table 5-5 of Reference 13 does not give a calculated

moment for 5% damping. However, Table 2-1 in the

Executive Summary of Reference 13 gives INPUT X/

LEVEL D = 24.0 for 5% damping. The S/2Sy, shown in

parentheses, was obained using S/2Sy = 24x60/(2x43.8)

= 16.4.

Table 6-9 of Reference 13 pertains to the evaluation of a

nozzle that, according to H. Hwang (General Electric

Company) in a letter to E. C. Rodabaugh dated November 14, 1991 (available in the author's file), was dimensionally the same as that used for Component Test 20; see

Figure 13 herein. The calculation of B2 = 3.42 is discussed herein in Section 4.3.3.2.

Table 6-11 of Reference 13 pertains to the evaluation of

straight pipe, with no complications. Staff ratios of S/2Sy

and Reference 13 ratios of S/3Sm are as follows:

Ref. 13 S/2Sy S/3Sm

Table 2% 5% 2% 5%

5-5 29 16 42 24

6-9 16 12 21 15

6-11 4.5 3.1 5.2 3.7

As its evaluations in Tables 6, 7, and 8, it is apparent that

the staff's evaluations are more conservative than those in

Reference 13, largely because the staff incorporated the

material property, Sy, rather than using Sm = 20 ksi.

Even so, the staff's evaluations indicate that, for functional capability, the present Code Equation (9) limit of

2Sy could be increased to 9Sy for 2% damping or to 6Sy

for 5% damping.

NUREG-1367 16
5.5 Weight Stresses

Figures 18 and 19 show weight stresses, Swr, for Systems

1 and 2. Reference 13 gives no further information on

weight stresses. However, in a letter from H. Hwang

(General Electric Company) to E. C. Rodabaugh dated

November 14, 1991 (available in the author's file), Hwang

provided the following information:

(1) SWT = B2MwIZ, or Bzb Mwb/Z, or Br Mwr/Z

where B2, Bah, B2, are Code indices

Mw = resultant moment due to weight at

elbows

Mwb = resultant moment due to weight on

branch of tees and at Node 8, System 1,and Node 6, System 2

Mwr = moment due to weight on runs of

tees

Z = pipe section modulus

(2) Mw, Mwb, and Mwr were calculated by analyses of

the piping systems with weight loading, including the

weight of water in the systems.

(3) SWr is in units of ksi.

For the highest weight stress in each system:

Swr, Sy, System Fig. Node Component Ba, ksi ksi SWTr/Sy

1 18 8 Vesselet 3.48 8.5 45.4 0.19

2 19 6 Nozzle 3.61 10 35.7 0.28

Thus, the staff concludes from the two piping system tests

that weight stresses of up to about 0.25Sy in combination

with the high reversing dynamic loads such as those applied in the tests will not impair the functional capability

of piping systems.

5.6 Pressure Stresses

Systems 1 and 2 were tested with an internal pressure of

1000 psi. The pressure stresses, PDo/2t, of particular relevance are the following:

System/ Sy, Sp,

Node ksi Dimensions ksi Sp/Sy

1/72 43.8 6 NPS, 0.280-in. wall 11.3 0.26

2/52 35.0 6 NPS, 0.280-in. wall 11.3 0.32

2/6 35.7 4 NPS, 0.237-in. wall 8.99 0.25

From the standpoint of pressure boundary integrity, using

an internal pressure of 1000 psi was appropriate, although using a higher internal pressure would have been

even more appropriate. However, from the standpoint of

functional capability, using zero internal pressure might

have been more bounding; that is, the elbows in the systems would have had lesser moment capacity.

17 NUREG-1367

6 OTHER PIPING SYSTEM TESTS

Piping systems, other than the two systems discusssed in

Section 5, were tested using dynamic loadings. Table 9

(Refs. 14-20) identifies the tests the staff reviewed for

this report.

The major purpose of reviewing other piping system tests

was to see if any of those tests might invalidate conclusions drawn from the evaluations in Sections 3, 4, and 5.

The Hanford Engineering Development Laboratory

(HEDL) tests (Ref. 14) are the most significant because

large displacements occurred to the extent that functional

capability was threatened. The HEDL tests are described

in Section 6.1. Except for the HEDL tests, the other

piping system tests, like the two system tests discussed in

Section 5, did not result in any threat to functional capability and, like most tests in Sections 3, 4, and 5, provide

lower bounds on combinations of steady-state (weight)

and dynamic loadings that will not cause loss of functional

capability.

For direct comparison with the tests discussed in Sections

3, 4, and 5, it would be ideal to have elastic response

spectrum analyses using 2% or 5% damping and + /- 15%

peak broadening. If the yield strength of the material at

the highest stress location were available, calculation of

S/2Sy in direct analogy to those calculated in Sections 3, 4,

and 5 would be possible. Of the cited references (Refs.

14-20), none provide exactly what would be needed.

However, each of the references does provide elastic

analyses, which are sufficient for the major purpose of

determining if any of these tests invalidate conclusions

drawn from the evaluations in Sections 3, 4, and 5.

6.1 Hanford Engineering Development

Laboratory Tests (Reference 14)

The HEDL piping system has undergone numerous tests,

starting in 1979 with design verification tests and continuing to 1985 (see Table 1-1 of Reference 14). The tests of

primary interest in this report are identified in Reference

14 as "modified four-support configuration" tests and are

discussed below.

The configuration is shown in Figure 20a. The system was

subjected to sinusoidal input at a frequency of 2 Hz, with

step increases of maximum acceleration levels up to 2.8g.

The 2-Hz input frequency coincides with the measured

first mode natural frequency of 2 Hz. The calculated first

mode frequency is 2.14 Hz. The system was not pressurized.

Figure 20b shows the displacement of point A as a function of input acceleration. At the end of the 2.8g test,

point A was permanently displaced 18 in. in the positive Z

direction. Plastic hinges developed at points B and C, but

there was no significant reduction in cross-sectional flow

area.

Directional changes in the piping system were made by

cold-bending the pipe to a bend radius of 3 in. The B2

index for the bends is 1.3h 2/3 = 1.3/1.142 2/3 = 1.19. The

ratio of the in-plane, closing, limit moment, using Equation (9), to the straight pipe limit moment with the same

Sy, using Equation (2), is 1.17. Thus, even ignoring the

cold-bending effect on yield strength, the bends had only

a little less moment capacity than the straight pipe. Coldbending significantly increases the yield strength of an

austenitic stainless steel material. Thus, the bends would

be expected to be stronger than the straight pipe and,

indeed, the hinge at point B formed in the straight pipe.

The significance of this lies in generalization of the "no

significant loss in flow area." If, for example, ANSI B 16.9

elbows with a bend radius of 1.5 in. had been used in the

HEDL system, and if the elbow material yield strength

were not greater than that of the pipe, the hinge at point

B might form in the elbow with incipient loss of flow area.

Reference 14 gives the results of an elastic response spectrum analysis using 10% damping, no peak broadening.

The response spectrum used is shown in Figure 5-1 of

Reference 14, that is, a response spectrum for a

sinusoidal input at 2 Hz. The highest stress in the system

was calculated to occur at the bend at point B in Figure

20b. At an input of 0.36g, the highest stress was calculated

to be 60 ksi. The yield strength of the pipe material is

stated to be 40 ksi. For a maximum input of 2.8g,

S/ 2 Sy = 60x2.8/(0.36x2x40) = 5.83 10% damping

An estimate of S/2Sy for 5% damping is needed. There

are indications in Reference 14 that the system essentially

responded as a single-degree-of-freedom system to the

2-Hz sinsusoidal input. Then, the calculated stress for 5 %

damping would be two times that for 10% damping, leading to

S/2Sy = 5.83x2 = 11.7 5% damping

This S/2Sy, rounded off to 12, is shown in Table 9. Using

the hypothesis of single-degree-of-freedom response,

peak broadening is meaningless and the SI2Sy ratios for

the first three entries in Table 9 are deemed to be reasonably comparable.

Reference 14, Figure 5-2, indicates weight stresses did

not exceed about 10 ksi, SwISy < 0.25. Thus, the HEDL

tests cannot be used to defend high reversing dynamic

loads in combination with steady-state stresses higher

than 0. 25Sy. However, the HEDL tests obviously support

a 4Sy limit on Code Equation (9). Indeed, from only these

tests, a limit of about 20Sy is defensible.

19 NUREG-1367
6.2 References 15-20 Tests

The piping system tests in References 15 through 20 did

not result in any threat to functional capability. In the

following sections, the staff briefly describes how the

S/2Sy ratios shown in Table 9 were derived from the cited

references and what information on weight stresses can

be gleaned from the references.

6.2.1 Reference 15

Figure 4 of Reference 15 is a response spectrum for 2%

damping. It indicates that the maximum test input was

about four times that required to produce a maximum

calculated stress of 2.4Sh. For the A106-B material, Sh =

15 ksi, 2.4Sh = 36 ksi. The yield strength, Sy, of the pipe

material is not given. Using a typical Syof 45 ksi leads to

S/2Sy = 4x36/(2x45) = 1.6 2% damping

The effect of peak broadening is not discussed in Reference 15. Weight stresses are not given, but SwISy was

probably less than 0.1.

6.2.2 Reference 16

Reference 16 states that the piping system without

branches withstood seismic inputs that were approximately four times the input required to produce a calculated stress equal to the Level D stress limit for Class 2

piping. It appears that the calculated stresses are from an

elastic response spectrum analysis using 3% damping

and, probably, no peak broadening. Although Reference

16 does not give important details, it appears that this

would translate to an S/2Sy between 2 and 4; a value of 3 is

shown in Table 9. This S/2Sy is very approximate.

Reference 16 does not give analogous information for the

piping system with branches. Weight stresses are not

given, but SwISy was probably less than 0.1.

It appears that for both systems, Reference 16 tests were

low-level tests relative to References 13, 14, and 17 tests.

S/2Sy = 21x60/(2x45) = 14 5% damping, 3 NPS

system

The yield strength of the A106-B material for the 6 NPS

system, according to Table 6-2 of Reference 17, is 54 ksi.

Thus,

S/ 2 Sy = 15x60/(2x54) = 8.3 5% damping, 6 NPS

system

Reference 17 indicates the weight stresses are not more

than 10 ksi; SwISy was less than about 0.2.

6.2.4 Reference 18

Reference 18 gives the results of an elastic response spectrum analysis using 2% damping, +1-15% peak broadening. The maximum calculated S/3Sm was 2.5; 3Sm = 60

ksi. The yield strength, Sy, of the austenitic stainless steel

piping material is not given. Using a typical Sy of 35 ksi

leads to

S/2 Sy = 2.5x60/(2x35) = 2.1 2% damping

Weight stresses are not given, but SwISy was probably less

than 0.2.

6.2.5 Reference 19

Reference 19 gives the results of elastic response spectrum analyses using (probably) 3% damping, no peak

broadening. Calculated stresses are summarized in Table

7 of Reference 19. The highest calculated stress is 25.5 ksi

at location "QA100" for Test T41.21.2. This test is for the

"KWU support configuration" at an input excitation of

"300%SSE." The maximum input was 800%SSE; thus,

S(nom) = 25.5x8/3 = 68 ksi. Stresses in Reference 19

were calculated using M/Z, not B2M/Z. Location

"QA100" is at an 8 NPS, 0.535-in. wall, 12-in. bend radius

elbow for which B 2 = 2.42. Thus, S = B2MIZ = 2.42x

68 = 165 ksi at 800%SSE input. The yield strength, Sy, of

the austenitic stainless steel pipe material is not given.

Using a typical Sy of 35 ksi leads to

S12Sy = 165/(2x35) = 2.4 3% damping

Weight stresses are not given, but SwiSy was probably less

6.2.3 Reference 17 than 0.2.

Reference 17 gives the results of elastic response spectrum analyses using 5% damping, no peak broadening.

Table 2-1 of Reference 17 shows the following:

S13Sm = 30/1.4 = 21 for 3 NPS system

S/3Sm = 30/2.0 = 15 for 6 NPS system

where 3Sm = 60 ksi

The yield strength, Sy, for the A106-B pipe material is not

given for the 3 NPS system. Using a typical Sy of 45 ksi

leads to

6.2.6 Reference 20

Reference 20 gives the results of a linear elastic time

history analysis using (probably) damping equivalent to

about 1% in a response spectrum analysis.

The calculated results are summarized in Reference 20 as

follows:

Maximum test input

Maximum allowable input

for Code Equation (9) = 60 ksi

1895 gal (1.93g)

240 gal (0.24g)

NUREG-1367 20
Figures 6 through 10 of Reference 24 indicate that the

yield strength of the pipe material was about 25 kg/mm =

36 ksi. Thus,

S/2Sy = (1895/240)60/(2x36) = 6.6

The S/2Sy of 6.6 in Table 9 is shown in parentheses because a response spectrum analysis was not available. It

may be that, on the basis of a response spectrum analysis,

Reference 20 tests were low-level tests relative to References 13, 14, and 17 tests.

Reference 20 cites a weight stress of 0.1 kg/mm = 0.14

ksi, presumably at the location at which S is maximum. No

other weight stresses are given, but SwISy probably did

not exceed 0.2 at any location.

6.3 Summary of Other Piping System

Tests

Results of the other piping system tests do not invalidate

the conclusions drawn from the evaluations in Sections 3,

4, and 5 of this report. In particular, the HEDL tests (Ref.

14) support a limit of 4Sy on Code Equation (9). None of

the other piping system tests provide a defense of steadystate stresses greater than about 0.25Sy when combined

with reversing dynamic stresses of 4Sy.

21 NUREG-1367

7 OTHER DYNAMIC LOADS

Dynamic loads applied in the tests discussed in Sections 3

through 6 were rapidly reversing in nature. The "rapidly"

is quantified as dominant reponses of 2 Hz or more. Provided the dominant response is not less than about 2 Hz,

these tests support, in regard to asssurance of functional

capability, an increase in the Code Equation (9) Level D

limit from 2Sy to 4Sy, with steady-state stresses up to

about 0.25Sy.

Other dynamic loads are the result of the following:

" fluid hammer

- fluid pressure waves

- slug flow

* relief-valve actuation

- steady-state forces

- short-time effects

* postulated pressure boundary breaks

* vibrations (e.g., piping connected to a reciprocating

pump)

The question discussed in the following sections is: Can

the Code Equation (9) Level D limit be increased when

other dynamic loads are applied to piping systems, either

alone or in combination with rapidly reversing dynamic

loads?

7.1 Fluid Hammer

A part of the EPRI, NRC, and General Electric Company

program consisted of water-hammer tests. These tests are

described in Section 7, "Pipe System Water Hammer

Tests," of Reference 13.

Five piping systems were tested; the tests were identified

as Test 28, Test 29, MS-l, MS-2 Runs 1-5, and MX-2

Runs 6 and 7. Water-hammer tests consisted of

(1) piping systems filled with water, sudden pressure

increase at one end of system: "solid water-hammer

load"

(2) piping systems partially filled with water, sudden

pressure increase at one end of system: "slug-type

loading"

The conclusions quoted from Reference 13 are the following:

7.3.4 Water Hammer Test Conclusions

In general solid water wave load, because of

quick load reversal, does not cause pipe collapse,

even when the calculated moment exceeds the

limit moment.

Strut failure due to water hammer can occur, but

in the test the failure load exceeded 10 times of

its rated load.

Slug type loading of long duration (simulating

static loads) can cause pipe "collapse."

A fluid (e.g., steam or water) pressure wave load could be

caused by closing of a valve or slamming of a check valve.

Time-history analyses are used to evaluate such loads,

and damping is not very significant. The staff agrees with

the conclusion in Reference 13 that pressure wave loads

are appropriately included with other rapidly reversing

dynamic loads.

As indicated by the third conclusion in Reference 13, slug

flow may produce collapse and thus constitutes a threat to

functional capability. No increase in Code Equation (9)

can be defended.

Slug flow is, of course, difficult to anticipate in the design

stage. Designs should include drains and vents, and operating procedures should be implemented so that the possibility of slug flow is minimized.

7.2 Relief-Valve Actuation

The steady-state thrust (e.g., acting for one or more seconds) should be evaluated as equivalent to a weight stress.

The time-variable effects would depend on whether there

is any slug flow. However, whether there is slug flow or

not, the information is insufficient to defend any increase

in Code Equation (9) Level D limits for those portions of

piping systems on which the relief valve is mounted.

In boiling-water reactors, relief-valve actuation may

cause building vibration. The effect of this buildingfiltered vibration on piping systems is appropriately included with other rapidly reversing dynamic loads.

7.3 Postulated Pressure Boundary

Breaks

A concern is whether the postulated pressure boundary

break might cause loss of functional capability of piping

23 NUREG-1367
systems other than the system in which the break is postulated.

Because the break effects will be "filtered" at other piping systems, the staff believes that the effects of postulated breaks can be considered to be rapidly reversing for

the purpose of evaluating piping systems other than the

system in which a break is postulated.

7.4 Vibrations

Vibrations, such as those induced by attached equipment

or fluid flow, are difficult to anticipate in the design stage.

The staff believes that such vibrations are best evaluated

during preoperational testing.

NUREG-1367 24
8 SUMMARY AND LIMITATIONS

The objective of this report is to examine present Code

rules and potential changes in Code rules to see if they are

sufficient to ensure maintenance of functional capability.

Stresses calculated by using Code Equation (9) [Equation

(1) herein] are limited as indicated in Section 2.1 of this

report.

As indicated in Section 2.2 of this report, the staff believes

that for static loadings, meeting Code Equation (9) with

Level D limits does not adequately demonstrate functional capability. However, as discussed in the previous

sections of this report, the results of many dynamic tests

show that the functionality of piping systems has been

maintained at equivalent stress levels significantly higher

than Level D limits. The following sections summarize

the findings and the limitations for ensuring piping functionality.

8.1 Reversing Dynamic Loads

Reversing dynamic loads are those due to earthquakes

and building-filtered loads such as those due to vibration

of buildings caused by relief-valve actuation in boilingwater reactors.

The test data evaluated in Sections 3, 4, 5, and 6 of this

report are relevant to this type of dynamic loading.

A significant aspect of the test data is that, with one

exception discussed in Section 4.6, none of the tests resulted in loss of functional capability. Thus, the staff's

evaluations are based on the premise that the test data

provide lower bounds on combinations of steady-state

(e.g., weight) and dynamic loadings that will not cause loss

of functional capability. This lower bound premise may

introduce conservatisms in the staff's recommendations.

But, as will become apparent in the following discussions,

this premise leads to a significant relaxation of the present NRC position on functionality; that is, present Code

Level D limits ensure piping functionality provided

steady-state stresses do not exceed 0.2 5Sy and the dynamic loadings are similar to those induced by earthquake

internal loadings.

8.1.1 Method of Calculating Mi in Code

Equation (9)

The moment, Mi, represents both steady-state (e.g.,

weight) loads and dynamic loads. Values of Mi are obtained by analyses of piping systems. In the past, the

dynamic portion of Mi has been obtained by an elastic

response spectrum analysis with +/-15% peak broadening and as low as 0.5% damping. The present trend is to

use 2% and up to 5% damping. Thus, the staff's evaluations are focused on 2% or 5% damping.

However, it is within the state-of-the-art to more accurately calculate the dynamic portion of Mi using an

elastic-plastic analysis. The approach used by the staff in

Sections 3.2, 4.2, and 5.3, "Comparisons with Theoretical

Limit Moments," was to look at Code Equation (9) with

the thought that Mi might be more accurately calculated.

The staff concludes that, if Mi is accurately calculated,

Code Equation (9), with a 2Sy limit, is not conservative.

If elastic-plastic analyses of piping systems in nuclear

power plants become routine, the staff believes that, for

ensuring piping functionality, a revised set of guidelines

might be needed for NRC's acceptance of such analyses.

Thus, its recommendations discussed herein apply only to

elastic response spectrum analyses.

8.1.2 Summary of SI2Sy Evaluations

S/2Sy

2% Damping 5% Damping Sw/Sy

Table Min. Max. Avg. Min. Max. Avg. Max.

2 0.62 5.5 3.5 0.25 2.2 1.4 0.15

6 3.4 20 10 2.2 10 5.6 0.06

7 9.1 24 13 5.2 15 7.4 0.32

8 7.2 16 11 4.3 8.4 6.4 0.08

N/A* 4.5 29 16 3.1 12 7.6 0.28

In the staff's judgment, the averages of S/2Sy are reasonable indicators of lower bounds on functional capability,

since functional capability was not lost in any tests other

than Test 37 of Reference 11. Also, in Table 2 the values

of S/2Sy < 1 do not mean that the pipe could not withstand higher dynamic loads; rather, no attempt was made

to apply higher dynamic loads.

Thus, the staff finds that the dynamic test results clearly

demonstrate that with certain limitations discussed in

Sections 8.1.3, 8.1.5, and 8.1.7, Code Equation (9) with a

stress limit of 2Sy, using 5% damping, provides assurance

that piping functional capability will be maintained.

8.1.3 Steady-State Stresses

Weight stresses should be considered as design conditions. The Code limit on Equation (9) for design

conditions is 1.5Sx, where Sx =Sm for Class 1 piping,

Sx = SA for Class 2 piping. In the bounding case in which

P = 0, Sm = Sh = 0.9Sy (austenitic steel at 650'F), the

*N/A = not applicable; results were obtained from Section 5 of this

report.

25 NUREG-1367
allowable moment due to weight using Equation (9), for

1.5Sx = 1.35Sy, and for straight pipe (B2 = 1.0), is

M = 1.35 ZSy

Equation (2) gives

ML = (4/hr)ZSy = 1.27 ZSy

Thus, in this bounding case, the moment due to weight is

about equal to the theoretical limit moment.

At Level D, the stress due to combinations of weight plus

dynamic loads is limited to 2Sy. Thus, at Level D, there is

a spectrum of allowable combinations ranging from

Sw = 0, Sd = 2.0Sy, to Sw = 1.35Sy, Sd = 0. 6 5Sy, where

Sw = weight stress, Sd = dynamic stress.

The test evaluations clearly indicate that the combination

of Sw = 0, Sd = 2.OSy maintains functional capability.

Unfortunately, no tests are available that show that the

combination of SW = 1.35Sy, Sd = 0.65Sy maintains

functional capability. That is, if a straight pipe were

loaded to its limit load by weight and then subjected to a

dynamic stress of + /-0.65Sy, would functional capability

be maintained?

In the absence of relevant test data, the staff recommends

that steady-state stresses be limited to 0.25Sy. Its judgment is based mainly on Reference 11 Tests 30 and 37. In

these tests, the combinations were

Test 30 S/2Sy = 5.2 (5% damping), SwISy = 0.32,

no collapse

Test 37/4 S/2Sy = 2.8 (5% damping), SwISy = 0.32,

no collapse

Test 37/5 S/2Sy = 5.5 (5% damping), SwISy = 0.32,

collapse

Using Code Equation (9) Level D limit of 2Sy,

S/ 2 Sy = (2-0.32)/2 = 0.84

Comparing this value with the corresponding value from

Test 37, Run 4, indicates that permitting SwISy up to 0.25

is adequate to ensure maintenance of functional capability. The value of 0.25 was deliberately chosen to be a bit

less than that in Test 37.

A conceptually more direct method of controlling steadystate stresses might be to introduce an Equation (9a),

which would directly, and independently of Code Equation (9), limit steady-state stresses to 0.25Sy. However,

the staff's present goal is to be able to say that meeting

Code Equation (9) Level D limits also ensures functional

capability.

The staff's recommended limit on steady-state stresses of

0.25Sy is not deemed to be onerous if the steady-state

stresses are due to weight. Typically, weight stresses do

not exceed about 3 ksi. Some examples of the 0.25Sy limit

are

Material Temp., 0F Sy, ksi 0.25Sy, ksi

A106-B 100 35 8.75

A106-B 650 25.4 6.35

Type 304 100 30 7.5

Type 304 650 17.9 4.48

However, if the steady-state stress is due to the steadystate thrust of a relief-valve discharge, then the staff's Sw

limit of 0.25Sy might be restrictive.

8.1.4 Pressure Stresses

Code Equation (9) includes the term [see Equation (1)

herein] B 1

PDo/2t. For most components, P > 0, this term

reduces the allowable combinations of Sw and Sd.

The staff's test data evaluations did not indicate any adverse effect of P > 0 on functional capability. Indeed, for

elbows, P > 0 tended to increase the moment capacity.

This aspect is partially recognized in the Code by

B1 = -0.1+0.4h but not < 0 nor > 0.5

Thus, for elbows with h < 1/4, the pressure term becomes

zero. But for other components, and elbows with h > 1/4,

the pressure term, P > 0, tends to add to the margin for

assurance of functional capability.

However, there is a potential for external pressure to

jeopardize functional capability. An external pressure

might arise for piping inside the containment when the

containment is pressurized under accident conditions.

nThe staff's Do/t limit (see Section 8.1.5) partially addresses this concern. However, the staff's recommendations include a restriction that external pressure must not

exceed internal pressure, as a reminder that this special

condition might need to be considered.

8.1.5 Dolt Limit

The available test data are mostly for components with

Do/t < 50, for example, 6 NPS, Sch. 10, Do/t =

6.625/0.134 = 49.4. Three Beaney tests (Ref. 10) were on

straight pipe with Do/t = 103/1.5 = 69, but incipient

buckling occurred in one of these three tests.

Thus, the staff deems it prudent to limit its recommendations for functional capability evaluation to components

NUREG-1367 26
with Do/t < 50. The Code also applies this limit to applicability of B-indices.

8.1.6 Future Changes in B-indices

The staff's recommendations are based on B-indices as

given in the present Code (Ref. 1). Code committees

constantly review newly developed data relevant to stress

indices and, sometimes, these reviews lead to reducing

the magnitude of stress indices. However, the Code committees are interested in pressure boundary integrity, not

necessarily functional capability. Thus, it becomes incumbent on the NRC staff to review any future Code changes

in B-indices from the standpoint of their effect on functional capability.

8.1.7 Future Changes to Code Equation (9)

Stress Limits

Code committees have been reviewing from the standpoint of pressure boundary integrity the same sets of test

data reviewed in this report from the standpoint of functional capability. It is possible that the Code Equation (9)

Level D limit of 2Sy might be increased to 4Sy.

It would be highly desirable that, if the Level D limit were

increased to 4Sy, it could be demonstrated that meeting

the Code would also ensure functional capability.

As stated in Section 8.1.3,

Using the Code Equation (9) Level D limit of 4Sy,

S/2Sy = (4-0.32)/2 = 1.84

Comparing this allowable value (1.84) with Test 37, Run

4, S/2Sy = 2.8, no collapse, indicates that permitting

SwISy up to 0.25 is adequate to ensure maintenance of

functional capability, even if the Level D limit on Code

Equation (9) is increased to 4Sy.

However, the boundary between static loading and dynamic loading is not well-defined. Use of a Code Equation (9) limit of 4 Sy can only be defended by the available

test data for rapidly reversing dynamic loads. For Component Test 37, discussed in Section 4.6, and the HEDL test,

discussed in Section 6.1, the dominant response frequencies were about 2 Hz. Both of these tests resulted in an

incipient threat to functional capability. Thus, the staff

believes that it is prudent to restrict a Code Equation (9)

limit of 4Sy to piping systems for which the elastic response spectrum analysis indicates that the response

stress contribution at 2 Hz and less is not more than Sy.

8.2 Other Dynamic Loads

Section 7 contains a brief discussion of other dynamic

loads. The staff concludes that it is appropriate to include

fluid-hammer pressure wave loads in the category of reversing dynamic loads. Those dynamic loads that are not

clearly in the category of reversing dynamic loads, and

combinations of reversing with nonreversing dynamic

loads, will require special consideration. Some suggestions are included in Section 7.

Test 30 S/2Sy

Test 37/4 SI2 Sy

Test 37/5 S/2Sy

= 5.2 (5% damping),

no collapse

= 2.8 (5% damping),

no collapse

= 5.5 (5% damping),

collapse

SwISy = 0.32,

SwISy = 0.32,

SwISy = 0.32,

27 NUREG-1367

9 CONCLUSIONS

9.1 Functional Capability Assurance,

Present Code Requirements

The staff concludes that piping functional capability is

ensured by meeting the present Code (Ref. 1) requirements, provided

(1) Dynamic loads are reversing. This includes loads due

to earthquakes, building-filtered loads such as those

due to vibration of buildings caused by relief-valve

actuation in boiling-water reactors, and pressure

wave loads (not slug-flow fluid hammer).

(2) Dynamic moments are calculated using an elastic

response spectrum analysis with +/-15% peak

broadening and with not more than 5% damping.

(3) Steady-state (e.g., weight) stresses do not exceed

0.25Sy.

(4) DoIt does not exceed 50.

(5) External pressure does not exceed internal pressure.

9.2 Functional Capability Assurance,

Future Code Requirements

Until such time as Code changes are made, the staff can

make no specific conclusions concerning such changes.

If the Code Equation (9) Level D limit is increased to, for

example, 4Sy, the staff concludes that in addition to restrictions (1) through (5) in the previous section, an additional restriction would be needed; that is, the elastic

response spectrum analysis must show that the response

stress contribution at 2 Hz and less is not more than Sy.

(See Section 8.1.7.)

Any changes in B-indices in the present Code should be

reviewed to determine whether such changes would adversely affect the assurance of functional capability.

With the use of a limit greater than 2Sy, increased vigilance would be needed to provide assurance that such

components as piping supports, anchors, restraints,

guides, and anchors have sufficient load capacity.

29 NUREG-1367

10 REFERENCES

1. American Society of Mechanical Engineers, Boiler

and Pressure Vessel Code, Section III, Division 1,

"Nuclear Power Plant Components," New York,

1989 Edition.

2. United States of America Standard B31.1-1967,

"USA Standard Code for Pressure Power Piping,"

American Society of Mechanical Engineers, New

York.

3. U.S. Nuclear Regulatory Commission, NUREG/

CR-0261, "Evaluation of the Plastic Characteristics

of Piping Products in Relation to ASME Code Criteria," Rodabaugh and Moore, July 1978.

4. --- , NUREG-0800, "Standard Review Plan for the

Review of Safety Analysis Reports for Nuclear

Power Plants," Section 3.9.3, "ASME Code Class 1,

2, and 3 Components, Component Supports, and

Core Support Structures," Rev. 1, July 1981.

5. General Electric Company, Nuclear Energy Engineering Division, "Functional Capability Criteria

for Essential Mark II Piping," NEDO-21985, San

Jose, California, September 1978.

6. U. S. Nuclear Regulatory Commission,

NUREG-1061, "Report of the U. S. Nuclear Regulatory Commission Piping Review Committee" (5

Volumes), Volume 2, "Evaluation of Seismic Designs-A Review of Seismic Design Requirements

for Nuclear Power Plant Piping," April 1985.

7. E. M. Beaney, "Response of Tubes to Seismic Loading," TPRD/B/0605/N85, Central Electricity Generating Board, Berkeley Nuclear Laboratories,

Berkeley, Gloucestershire, United Kingdom, January 1985.

8. --- , "Response of Pipes to Seismic Excitation-Effect of Pipe Diameter/ Wall Thickness Ratio and

Material Properties," TPRD/B/0637/N85, Central

Electricity Generating Board, Berkeley Nuclear

Laboratories, Berkeley, Gloucestershire, United

Kingdom, July 1985.

9. --- , "Response of Pressurized Straight Pipe to Seismic Excitation," TPRD/B/0826/R86, Central Electricity Generating Board, Berkeley Nuclear Laboratories, Berkeley, Gloucestershire, United Kingdom,

February 1986.

10. --- , "Response of Stainless Steel Pipes to Seismic

Excitation," TPRD/B/ 105 1/R85, Central Electricity

Generating Board, Berkeley Nuclear Laboratories,

Berkeley, Gloucestershire, United Kingdom, April

1988.

11. General Electric Company, Nuclear Energy Engineering Division, "Piping and Fitting Dynamic Reliability Program," Volume 2, "Component Test Report," EPRI Contract RP 1543-15, Draft, San Jose,

California, December 1989.

12. J. Spence and G. E. Findlay, "Limit Loads for Pipe

Bends Under In-Plane Bending," Paper No. 1-28,

Proceedings of the 2nd International Conference on

Pressure Vessel Technology, American Society of Mechanical Engineers, New York, October 1973.

13. General Electric Company, Nuclear Energy Engineering Division, "Piping and Fitting Dynamic Reliability Program," Volume 3, "System Test Report,"

Draft, San Jose, California, February 1990.

14. Hanford Engineering Development Laboratory,

"High-Level Dynamic Testing and Analytical Correlations for a One-Inch Diameter Piping System,"

M. R. Lindquist, M. J. Anderson, L. K. Severud, and

E. 0. Weiner, HEDL-TME 85-24, Richland, Washington, February 1986.

15. G. E. Howard, B. A. Johnson, W. B. Walton, H. T.

Tang, and Y. K. Tang, "Piping Extreme Dynamic

Response Studies," Proceedings of the 7th Structural

Mechanics in Reactor Technology Conference, Vol. F,

August 1983.

16. U. S. Nuclear Regulatory Commission, NUREG/

CR-3893, "Laboratory Studies: Dynamic Response

of Prototypical Piping Systems," ANCO Engineers,

Inc., August 1984.

17. --- , NUREG/CR-5023, "High-Level Seismic Response and Failure Prediction Methods for Piping,"

Westinghouse Hanford Co., January 1988.

18. B. Charalambus, E. Haas, and R. Mihatsch, "Comparisons of Dynamic Test Data with Results of Various Analytical Methods," Nuclear Engineering and

Design, Vol. 96, pp. 447-462, 1986.

19. U. S. Nuclear Regulatory Commission, NUREG/

CR-5757, "Verification of Piping Response Calculation of SMACS Code with Data from Seismic Testing of an In-Plant Piping System," Argonne National

Laboratory, September 1991.

20. --- , NUREG/CR-5585, "The High Level Vibration

Test Program," Brookhaven National Laboratory,

May 1991.

31 NUREG-1367

Load

Frame

Displacement

Transducer

Actuator

Servo

Straight Pipe

Pivot

Pivoted Strain

Link Gauge

Pivot Test Pipe Accelerometer

L

Figure 1 Test arrangement

Source: References 7, 8, 9, and 10.

z

C1

Actuator

Piston
1/2% 1% 2% 5%

gr 1

g 2i

22

20 ,

18 -

16 3

o 14

(D

12 12

OL 4

10

Pipe Symbol

611 -0-<-

2 ...

4 3

4

2~ damping factor

0 - 1

0 1 2 3

Input Acceleration, g

Figure 2 Response versus input acceleration

Source: Reference 8.

1

2

21.3

15.6

16.9

11.5

18.8

NUREG-1367 34
-11 *1I Input Acceleration, g

X...__ LUnc4 Positive~ Strain

N

C

6-.5

-- g • x-×Ne-gative Strain

Figure 3 Strain at pipe midspan versus input acceleration

Source: Reference 7.

------_x

35 NUREG-1367
E C6

Shape

300 200 100 .0 100 200 300

Distance from Center Line of Tube, mm

Figure 4 Deformed shape and permanent strain after tests

Source: Reference 7.

400 500

NUREG-1367 36
Input Acceleration

0 2 3

-0-21

I...

, •. \1

-0-4-*

-0... ......... ..

- 2 .... X ... 0.11.

0 V

-1. \

-1-2

Central Offset v

Pipe Gage Gage Sw/Sy -'6 1 -o-- -- 0.076 :

-1-8 3 ,- -- -&- 0.11\'-

4. -.- '+... •-.*-. 0.15 "

5 .---•- 0.13 4

- 2~0

0

Figure 5 Mean strain versus input acceleration

Source: Reference 8.

37 NUREG-1367
_2

150 125 100 7S 50 25 0 25 50 75 100 125

t Distance from Center of Span

Figure 6 Deformed shape of upper surface of 103-mm pipe, Test 16

Source: Reference 10.

150

NUREG-1367 38
Elbow

Test

Fixture

•~~~ .T:h•o.- in ~ 9 in. Direction

"L ] : _ of Motion

Clamp

Figure 7 In-plane elbow test arrangements, Tests 1, 3-8, 13, 19, and 31

Source: Reference 11.

39 NUREG-1367
250 r200

o6

10

e9 011

e7

E

0

(D

CO

ca

150

ML = 148 in.-kip

o5

100 H

50h All Data Points:

Simulated Earthquake Input;

Run Numbers Are Shown.

I 0 I I

0 500 1000 1500 2000 2500 3000

Calculated Moment, in.-kip, 2% Damping

Figure 8 4 NPS, Sch. 40 stainless steel pipe, Test 15

Source: Reference 11.

NUREG-1367 40
600 r- 07 912

ell

500 l-

@10

e9 ML = 490 in.-kip

Cý

(D

E

0

400

300

200

100

0 Sinesweep Input

0 Simulated Earthquake Input

Run Numbers Are Shown.

-0 5

08

1000 2000 3000 4000

Calculated Moment, in.-kip, 2% Damping

5000 6000

Figure 9 6 NPS, Sch. 40 carbon steel pipe, Test 34

Source: Reference 11.

41 NUREG-1367
200

.2-

C:

a)

E

0

(D

=)

150

All Data Points:

Simulated Earthquake Input;

Run Numbers Are Shown.

.ll

10

09

e8

@7

06

ML = 52.3 in.-kip

.5

100

50

A

0 250 500 750 1000

Calculated Moment, in.-kip, 2% Damping

Figure 10 6 NPS, 9-in. bend radius, Sch. 10 stainless steel elbow, Test 3

Source: Reference 11.

NUREG-1367 42
400

E

0

ca

(D

Cu

ML = 380 in.-kip

Average Wall Thickness

7-8 9 .10

o6

ML = 189 in.-kip

Nominal Wall Thickness

5

300

200

100 H All Data Points:

Simulated Earthquake Input;

Run Numbers Are Shown.

0

0 500 1000 1500 2000 2500

Calculated Moment, in.-kip, 2% Damping

Note:

This elbow had a measured average wall thickness

of 1.52 times nominal wall thickness.

Calculated moments are based on nominal wall thickness.

Figure 11 6 NPS, 6-in. bend radius, Sch. 40 carbon steel elbow, Test 13

Source: Reference 11.

43 NUREG-1367
1.0

0.95 - Range:

0.8- 3 7Tests i @

0.80 _ (4,5,6,7,| 0

19, 35, 41)j

0.75 - 3 t0.156 0.15

0.70 - 431

0.65

0.60 30

0.55 - 8E

0.50 - 813, t -0.425h DP 0.450 - 0. 02t Sy0.45-- /k, J /•/ •ML =0.8hO6 tsP=0

0.40

0.35.

Q 0.30 -37

0.25

o P>0

0.20 Reference 3 Dynamic Tests

(see Table 5)

V 24 (4) and 24 (5), P>0

Reference 11 Dynamic Tests

0.15 (see. Table 4)

* P=O

* P>o

P =internal pressure

0.10 I I

0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0

h = tR/r 2

Note: Number next to symbol is test number.

Figure 12 Elbows: Static in-plane closing moment capacity and dynamic in-plane

moment capacity tests

Source: References 3 and 11.

NUREG-1367 44
Cap

M

Detail

12 NPS

I

Detail A

Nominal Dimensions

Figure 13 Test 20 configuration: 4 NPS nozzle in 12 NPS vessel

Source: Reference 11.

45 NUREG-1367
I -

LII~~ I F

Inertia Arm

48 in.

f12, - 6) - I

V1 IF

Inertia Arm

Extension 36 in.

-1-

27 in.

Elbow

Test

Fixture

10 in. NRC Sled

in. Direction

of Motion

Clamp

Figure 14 Test arrangement, Tests 30 and 37

Source: Reference 11.

NUREG-1367 46
Hanger:

EL= 1823 ft 4-1/2 in.

EL11f 72i. 66 ~3-in. PIPE88 18 in. 30 3 8

N.j fl " 92 90 86 1901I24in

B25 50b 6Ib) 26•3

.- I 6 6in.

6-n Sch. 16 36 28inLJE 33in~ 

60 2 504b 46 6in.

6-in. Sch. 40Z P 6i Sh•c. 18 688 •4 18),z

7 7i k---96-in Sch. 40

Input

Note: Valves 18, 38, and 66 are

simulated by lump weight.

Valve 86 is a motor-operated valve.

Figure 15 Piping System 1 configuration [material: carbon steel (AO16-B)]

Source: Reference 13.

z

~11

0
Y

z x

W8x15 (or

6.625-in. Sch. 160 Pipe,

'or W8x21)

150 Ib

400 lb

50 -,< I'-

42 in. q

C14

3

6.375 in.

A-SLED 4

6-in. PIPE

Sch. 40

Sch. 40

SLED[ 1

28

ASLSLED 2

Residual Heat Removal Near Containment

Note: Valves 42 and 14 are simulated by weight.

Figure 16 Piping System 2 configuration [material: stainless steel (Type 316)]

Source: Reference 13.

NUREG-1367 48
Attach to Sled

Outer Diameter of Hub

6-in. 900 Class Slipon Flange

6-in. Sch. 160

0.718 in, A

'A

Connection to Sled Y

Designed by Others

Uniaxial

SG1 Through SG6

SG13 Through SG18

(6 Channels Each Place) A-A

Typical for Two Places

Note: SG - strain gauge

Figure 17 Load measurement device at Sleds 2 and 4

Source: Reference 13.

49 NUREG-1367
z

I

0h

34

72\

SLED 4

SWT = weight stress

Sb = bending stress due to g ARS

(plastic, assumed sled capacity)

R = Sb/ 3 Sm

SLED 1

z

Indicates the High-Stress Locations

EI~ K~Z

SLED 2

Figure 18 Piping System 1, weight stresses, SWT

Source: Reference 13.
(1) Input: 25 times safe shutdown earthquake

TH-B 35% ARS unbroadened

(2) High-weight stress system

(3) fn = 4.1(x), 5.3(z), 6.0(Y), 6.1(Z), 8.3(z) 36

ST= 5.6

Sb = 1.26

R =2.1 ID7= > ,,, • I I

SwT = !. D.x = 2x2.8 38DN 5

-Isb = 1221 D0 = 2x3._ _ _

W= 5.6 2

M= 1350 (inkip) Sb 1 SWT = 0.8

R =2.= 5 06T 

2 sw -r= 1.8 

R

b = 1.6

Sb = Bending stress, B2 M/Z, at ETEC sled capacity

SWT = weight stress

R = Sb/3Sm

6-in. SCH 40 elbow MLIM = 232 in.-kip

4-in. SCH 40 elbow MLIM = 102 in.-kip

4-in. x 6-in. tee MLIM = 96 in.-kip

X

B2 = 2.02

1~/~

I-.

Ay = 4.b IL

Az= 19

R= =

R =1.1

28

2, weight stresses, SWTi

ce 13.

MT = 724

(Torsion)

z

Figure 19 Piping System ;

Q Source: Referen
Rigid Strut

-E) Snubber

1.315" 0.0., 0.133" Wall

Stainless Steel Pipe

40"

~•fl24"1

18"

A

z

(a) 1-in.-Diameter Pipe Loop, Modified Four-Support Configuration

20 1 1 1 1 1

A 

C

E

(D 10-

C.) A2

aB

lo0 .5 1.0 1.5 2.0 2.5 3.0

Sinusoidal Strongback Input Acceleration (G) (ACCEL El1-11)

HEDL 8511-041.3

(b) High-Level Sinusoidal Test, Permanent Displacement at Upper Elbow

Figure 20 Hanford Engineering Development Laboratory piping system

Source: Reference 14.

NUREG-1367 52
Table 1 Beaney (Refs. 7, 8, 9, and 10) Straight Pipe Tests: Materials, Yield Strengths,

Dimensions, Sinusoidal Input Test Frequencies, Pressures, and Test Planes

Pipe

Sy, Do, t, L, fP, Test

Ref. Test Mtl. MN/m 2 mm mm mm Hz MN/m 2 Plane

(a) (b) (c) (d) (e) (M) (g) (h) (i)

7 (j) CS 209 25.4 2.64 3530 5 0.00 V

8 1 CS 298 25.4 0.91 3739 5 0.00 V

2 CS 219 25.4 2.64 3617 5 0.00 V

3 CS 248 25.4 6.35 3386 5 0.00 V

4 CS 161 34.1 4.06 4165 5 0.00 V

5 CS 223 51.8 4.47 5664 5 0.00 V

9 1 CS 162 34.14 4.06 4166 5 29.2 V

2 CS 162 34.14 4.06 4166 5 29.2 V

3 CS 162 34.14 4.06 4166 5 29.2 V

4 CS 298 25.4 0.91 3734 5 14.3 V

5 CS 162 34.14 4.06 2946 9 29.2 V

6 CS 298 25.4 0.91 2642 9 14.3 V

7 CS 298 25.4 0.91 2642 9 14.3 V

8 CS 162 34.14 4.06 2946 9 29.2 V

10 1 CS 162 34.1 4.06 2946 9.7 0.00 H

2 CS 162 34.1 4.06 2946 9.4 29.5 H

3 CS 162 34.1 4.06 4166 4.8 0.00 H

4 CS 162 34.1 4.06 4166 4.7 29.5 H

5 CS 195 25.4 2.64 3556 4.8 0.00 H

6 CS 195 25.4 2.64 3556 4.6 31.0 H

7 SS 247 25.4 2.64 3556 4.7 0.00 H

8 SS 247 25.4 2.64 3556 4.6 38.2 H

9 SS 247 25.4 2.64 2891 6.7 38.2 H

10 SS 247 25.4 2.64 2946 6.5 57.4 H

11 SS 261 78 1.5 6121 5.2 0.00 H

12 SS 261 78 1.5 3988 7.5 0.00 H

13 SS 261 78 1.5 3988 7.5 10.3 H

14 SS 335 103 1.5 6039 7.3 0.00 V

15 SS 335 103 1.5 5490 5.1 0.00 V

16 SS 335 103 1.5 5490 5.0 7.2 V

(a)

(b)

(c)

(d)

(e)

(f)

(g)

(h) (i)

(Ij)

Test identification according to references.

CS carbon steel; SS = Type 316 stainless steel.

Sy = yield strength of pipe material (from references).

Do= pipe outside diameter.

t = pipe wall thickness.

L = pipe span length.

f = sinusoidal input test frequency.

P = internal pressure in pipe while being tested.

V = dynamic loading in vertical plane; H = dynamic loading in horizontal plane.

Straight pipe test of Reference 7.

53 NUREG-1367
Table 2 Beaney (Refs. 7, 8, 9, and 10) Straight Pipe Test Results Evaluated

in Relation to Elastic Analysis, 2% or 5% Damping

S/2Sy f Sy,

Ref. Test Hz MIgr gmi ksi = 0.02 • = 0.05 SwISy Sp/Sy

(a) (b) (c) (d) (e) (M) (f) (g) (h)

7 (i) 5 178.4 4.2 30.3 5.19 2.08 0.11 0.00

8 1 5 67.5 3.6 43.2 2.78 1.11 0.076 0.00

2 5 170.0 3.5 31.8 3.93 1.57 0.11 0.00

3 5 299.9 3.0 36.0 3.40 1.36 0.11 0.00

4 5 455.4 2.7 23.4 4.18 1.67 0.15 0.00

5 5 1057 2.7 32.3 2.48 0.99 0.13 0.00

9 1 5 456.9 1.8 23.5 2.77 1.11 0.15 0.67

2 5 456.9 (j) 23.5 --- 0.15 0.67

3 5 456.9 3.6 23.5 5.54 2.22 0.15 0.67

4 5 67.7 0.8 43.2 0.62 0.25 0.076 0.65

5 9 282.0 5.0 23.5 4.75 1.90 0.076 0.67

6 9 41.7 2.1 43.2 1.00 0.40 0.038 0.65

7 9 41.7 2.6 43.2 1.24 0.50 0.038 0.65

8 9 282.0 4.8 23.5 4.56 1.82 0.076 0.67

10 1 9.7 241.8 5.7 23.5 4.66 1.86 0.076 0.00

2 9.4 257.5 5.3 23.5 4.61 1.84 0.076 0.67

3 4.8 493.9 2.8 23.5 4.67 1.87 0.153 0.00

4 4.7 515.1 2.6 23.5 4.52 1.81 0.153 0.67

5 4.8 190.8 3.3 28.3 4.67 1.87 0.121 0.00

6 4.6 207.7 2.7 28.3 4.16 1.67 0.121 0.69

7 4.7 199.0 2.8 35.8 3.27 1.31 0.095 0.00

8 4.6 207.7 1.4 35.8 1.70 0.68 0.095 0.67

9 6.7 148.2 4.4 35.8 3.82 1.53 0.063 0.67

10 6.5 151.6 5.5 35.8 4.89 1.95 0.065 1.00

11 5.2 1168 3.7 37.9 3.46 1.38 0.073 0.00

12 7.5 1323 3.6 37.9 3.81 1.52 0.080(k) 0.00

13 7.5 1323 2.3 37.9 2.43 0.97 0.080(k) 1.01

14 7.3 1422 5.3 48.6 2.66 1.06 0.042 0.00

15 5.1 3526 2.3 48.6 2.86 1.14 0.107(k) 0.00

16 5.0 3669 1.9 48.6 2.46 0.98 0. 107(k) 0.73

NUREG-1367 54
Table 2 (Continued)

Table Notes:

(a) Test identification according to references.

(b) f = sinusoidal input test frequency.

(c) M/gr = 386EI/(412 L 2)

where M = moment at center of pipe span, in.-lb

gr = response acceleration

E = modulus of elasticity, 30,000 ksi used

I = section modulus of pipe cross section, in.

f = sinusoidal input test frequency, Hz

L = pipe span length, in.

(d) gmi = maximum input acceleration during each test (from figures in the references).

(e) Sy = yield strength of pipe material (from references).

(f) S = (M/gr)gmi/(2 gZ)

where g = damping factor, 0.02 or 0.05

Z = section modulus of pipe cross section, in.

(g) Sw = stress at center of pipe span due to weight.

(h) Sp = stress due to internal pressure = PD/(2t)

where P = internal pressure

D = pipe mean diameter = Do- t

t = pipe wall thickness

(See Table 1 for values of P, Do, and t.)

(i) Straight pipe test of Reference 7.

(j) No gmi given in Reference 9 for Test 2.

(k) In Reference 10, Tests 12, 13, 15, and 16, the pipe was filled with water.

55 NUREG-1367
Table 3 Reference 11 Pipe Tests: Limit Moments and-Measured Moments

Pipe

Test Run Sy, PDo

No. No. Type NPS Sch. Mtd. ksi 2tSy ML Mm MLIMm

(a) (b) (c) (d) (e) (f)

9 6 T 6 40 SS 40.8 0.472 420 540 1.29

10 7 T 6 40 SS 40.8 0.278 446 491 1.10

11 6 T 6 10 SS 39.7 0.244 219 143 0.65

12 6 T 6 40 SS 40.8 0.472 420 492 1.17

14 6 T 6 40 CS 41.5 0.464 429 564 1.32

15 10 R 4 40 SS 37.0 0.413 148 189 1.27

16 6 R 4 40 CS 49.5 0.309 205 260 1.27

33 -- P 6 40 CS 44.5 0.255 490 532 1.09

34 12 P 6 40 CS 44.5 0.255 490 605 1.23

40 5 R 4 40 SS 37.0 0.000 159 202 1.27

(a) T = 6x6x6 ANSI B16.9 tee, fixed at both run ends, branch loaded.

R = 4 NPS pipe between 8x4 and 6x4 ANSI B16.9 reducers.

P = straight pipe.

Maximum loads are due to earthquake-type dynamic input, except for Test 33, during which sinesweep dynamic

input was used.

(b) SS = stainless steel, SA312 Type 316; CS = carbon steel, SA106-B.

(c) Sy = yield strength of material, ksi (from Appendix D of Reference 11).

For tees (no data for pipe), tee data were used.

For reducers, pipe data were used.

For pipe, Sch. 40 pipe data were used.

(d) P = internal pressure; D = mean diameter of pipe; t = nominal wall thickness of pipe.

(e) ML = calculated limit moment, in.-kip, = D2tSy[1-0.75(PD/2tSy)2]1/2

(f) Mm = maximum measured dynamic moment, in.-kip (from Appendix B of Reference 11).

For Tests 12 and 14, Mm was adjusted by dividing the Reference 11 measured moment by 1.09 to obtain

estimate of measured moment at the failure location.

NUREG-1367 56
Table 4 Reference 11 Elbow (6 NPS, 900) Tests: Limit Moments and Measured Moments

Elbow

Test/ Sy, Test PD,/

Run Sch. Mtd. h ksi Plane 2tSy - ML Mm MLIMm

(a) (b) (c) (d) (e) (f)

1/8 80 CS 0.41 40.0 In 0.269 424 569 1.34

2/8 80 CS 0.41 40.0 Out 0.269 ? 574

3/11 10 SS 0.11 34.0 In 0.285 52.3 163 3.12

4/? 40 CS 0.25 47.8 In 0.237 246 396 1.61

5/8 40 CS 0.25 47.8 In 0.403 246 478 1.94

6/8 40 SS 0.25 54.2 In 0.355 279 457 1.64

7/8 40 SS 0.25 54.2 In 0.209 279 426 1.53

8/8 40 SS 0.25 54.2 In 0.000 232 342 1.47

13/10 40 CS 0.17 47.0 In 0.241 189 400 2.12

17/? 40 CS 0.17 47.0 Tor 0.241 ?- __

19/8 40 SS 0.25 54.0 In 0.525 278 450 1.62

23/4 40 CS 0.25 42.3 In 0.268 217 470 2.16(g)

25/14 10 SS 0.11 34.0 In 0.570 52.3 380? 7.31(h)

26/? 40 CS 0.25 42.3 In 0.455 217

30/4 10 SS 0.11 34.0 In 0.285 52.3 112 2.14

31/7 10 SS 0.11 38.6 In 0.251 59.4 150 2.53

35/? 40 CS 0.25 42.3 In 0.455 217 394 i.82

37/5 10 SS 0.11 34.0 In 0.000 43.6 57 1.31

41/? 40 CS 0.25 44.0 In 0.438 226 398 1.76

(a) All except Tests 13 and 17, 9-in. bend radius; Tests 13 and 17, 6-in. bend radius.

Maximum loads are due to earthquake-type dynamic input, except for Test 25, during which dynamic input in

the middle-range frequency was used, and Test 26, during which sinesweep dynamic input was used.

(b) CS = carbon steel, SA106-B; SS = stainless steel, SA312 Type 316.

(c) h = elbow parameter = tR/r2

where t = elbow nominal wall thickness

R = elbow bend radius

r = mean elbow cross-section radius

(d) Sy = material yield strength (from Appendix D of Reference 11).

(e) ML = limit moment calculated using Equation (9), in.-kip; conceptually, in-plane, closing limit moment.

(f) Mm = maximum measured moment, in.-kip (from Appendix B of Reference 11).

(g) Assembly restrained with a strut. Significance of the measured moment is not clear.

(h) Reference 11, Appendix B, states: "Mid-freq. moment measuring method is still in study, results will be

changed."

57 NUREG-1367
Table 5 Reference 3 Static and Dynamic In-Plane Moment Capacity Tests on Elbows

(See Figure 12 for plot of these data.)

Test Sy, PD, 12t, M, D2 tSy,

Iden. Mtl. ksi ksi in.-kip in.-kip MID2 tSy h

(a) (b) (c) (d) (e) (M)

22(2) CS 50.0 0 261 + 563.6 0.46 + 0.25

22(5) CS 50.0 17.0 347 + 563.6 0.61 + 0.25

22(8) CS 37.8 0 450+ 626.3 0.72+ 0.41

22(11) CS 39.6 0 202+ 446.4 0.45+ 0.17

22(15) SS 37.7 0 206+ 425.0 0.48+ 0.25

22(16) SS 37.7 0 202+ 425.0 0.48+ 0.25

22(17) SS 35.6 0 200+ 401.3 0.50+ 0.17

22(18) SS 35.4 0 381 586.5 0.65 0.41

22(19) CS 46.0 0 202 518.5 0.39 0.17

22(20) CS 34.6 0 369 573.3 0.64 0.27

23(1) SS 36.3 0 3300 6515 0.51 0.15

13(1) CS (45) 15.0 269+ 488.1 0.55+ 0.26

13(5) SS (35) 0 166+ 379.6 0.44+ 0.26

13(6) SS (35) 18.6 313 379.6 0.82 0.26

13(7) SS (35) 0 122 274.3 0.44 0.18

13(8) SS (35) 0 79 221.4 0.36 0.14

13(9) SS (35) 0 78+ 106.4 0.73+ 0.40

24(4) SS (35) 16.5 +/-43.8(g) 47.5 0.92 0.18

24(5) SS (35) 17.3 +/-45.0(g) 47.5 0.95 0.18

(a) Identification according to Table 4 of Reference 3.

(b) CS = carbon steel; SS = stainless steel.

(c) Sy = yield strength as listed in Reference 3. For References 13 and 24 in Reference 3, yield strengths

were not given. Typical values of 45 ksi for carbon steel and 35 ksi for stainless steel were used.

(d) P = internal pressure; D = mean diameter of elbow; t = nominal wall thickness of elbow.

(e) From Table 4 of Reference 3, column headed "Mm". A "+" indicates that the moment capacity was not

reached in the static loading test.

(f) h = elbow parameter = tR/r2

where t = elbow wall thickness

R = elbow bend radius

r = elbow mean cross-section radius

(g) These values were derived from sinusoidal dynamic loading tests.

NUREG-1367 58
Table 6 Reference 11 Pipe Tests: Comparisons with 2Sy Limit

Pipe 2% Damping 5% Damping

Test/ Sy, S, S,

Run Type NPS Sch. MtI. ksi ksi S/2Sy ksi S/2Sy SwlSy SpiSy

(a) (b) (c) (d) (e) (fM (g)

9/6 T 6 40 SS 40.8 589 7.2 330 4.0 0.02 0.47

10/7 T 6 40 SS 40.8 600 7.4 335 4.1 0.02 0.28

11/6 T 6 10 SS 39.7 269 3.4 178 2.2 0.04 0.24

12/6 T 6 40 SS 40.8 737 9.0 401 4.9 0.02 0.47

14/6 T 6 40 CS 41.5 542 6.5 304 3.7 0.02 0.46

15/9 R 4 40 SS 37.0 787 11 428 5.8 0.06 0.41

16/6 R 4 40 CS 49.5 1979 20 1011 10 0.04 0.31

33/? P 6 40 CS 44.5 --,- 0.00 0.25

34/12 P 6 40 CS 44.5 731 8.2 419 4.7 0.01 0.25

40/5 R 4 40 SS 37.0 1345 18 786 11 0.06 0.00

(a) T = 6x6x6 ANSI B16.9 tee, fixed at both run ends, branch loaded..

R = 4 NPS pipe between 8x4 and 6x4 ANSI B16.9 reducers.

P = straight pipe.

Maximum loads are due to earthquake-type dynamic input, except for Test 33, during which sinesweep dynamic

input was used.

(b) SS = stainless steel, SA312 Type 316; CS = carbon steel, SA106-B.

(c) Sy = material yield strength (from Appendix D of Reference 11).

For tees (no data for pipe), tee data were used.

For reducers, pipe data were used.

For pipe, Sch. 40 pipe data were used.

(d) From Appendix B of Reference 11, Case 2, B2M/Z with B2 = 1.00. Appendix B states: "Case 2 Actual tested

time history used 2% damping amplified response spectrum + /-15% broadening response spectrum analysis."

(e) From Appendix B of Reference 11, Case 3, B2MIZ with B2 = 1.00. Appendix B states: "Case 3 Same as

Case 2 except using 5% damping."

(f) Sw = stress due to weight = Mw/Z where Mw = moment due to weight.

(g) Sp = stress due to internal pressure = PD/2t

where P = internal pressure

D = pipe mean diameter

t = pipe nominal wall thickness

59 NUREG-1367
Table 7 Reference 11 Elbow (6 NPS, 900) Tests: Comparisons with 2Sy Limit

Elbow 2% Damping 5% Damping

Test/ Sy, Test S, S,

Run Sch. Mtl. ksi Plane ksi S/2Sy ksi S/2Sy Sw/Sy Sp/Sy

(a) (b) (c) (d) (e) (M) (g)

1/8 80 CS 40.0 in 890 11 547 6.8 0.01 0.27

2/8 80 CS 40.0 Out 897 11 501 6.3 0.01 0.27

3/10 10 SS 34.0 In 1276 19 752 11 0.04 0.28

4/? 40 CS 47.8 In 1057 11 648 6.8 0.01 0.24

5/8 40 CS 47.8 In 1238 13 674 7.0 0.01 0.40

6/8 40 SS 54.2 In 1158 11 634 5.8 0.01 0.36

7/8 40 SS 54.2 In 1392 13 756 7.0 0.01 0.21

8/8 40 SS 54.2 In 1442 13 776 7.2 0.01 0.00

13/10 40 CS 47.0 In 1255 13 679 7.2 0.02 0.24

17/? 40 CS 47.0 Tor - - -- 0.02 0.24

19/8 40 SS 54.0 In 1331 12 707 6.5 0.01 0.52

25/15 10 SS 34.0 In 1628 24 990 15 0.04 0.57

26/? 40 CS 42.3 In - - - - 0.01 0.46

30/4 10 SS 34.0 In 620 9.1 356 5.2 0.32 0.28

31/11 10 SS 38.6 In 1391 18 738 9.6 0.04 0.25

35/? 40 CS 42.3 in - - - - 0.08 0.46

37/5 10 SS 34.0 In 651 9.6 375 5.5 0.32 0.00

41/? 40 CS 44.0 In ---.--- 0.01 0.44

(a) All except Tests 13 and 17, 9-in. bend radius; Tests 13 and 17, 6-in. bend radius.

Maximum loads are due to earthquake-type dynamic input, except for Test 25, during which dynamic input in

the middle-range frequency was used, and Test 26, during which sinesweep dynamic input was used.

(b) Elbow material: CS = carbon steel, SA106-B; SS = stainless steel, SA312 Type 316.

(c) Sy = material yield strength (from Appendix D of Reference 11).

(d) From Appendix B of Reference 11.

Response spectrum analysis based on 2% damping, + /-15% peak broadening.

(e) Same as (d), except 5% damping was used.

(f) Sw = stress due to weight = B2Mw/Z where M, = moment due to weight at mid-arc of elbow.

(g) Sp = stress due to internal pressure = PD/2t

where P = internal pressure

D = elbow mean diameter

t = elbow nominal wall thickness

NUREG-1367 60
Table 8 Reference 11 Tests on Other Components: Comparisons with 2Sy Limit

Pipe 2% Damping 5% Damping

Test/ Sy, S, S,

Run Type Size Sch. Mtl. ksi ksi S/2Sy ksi S/2Sy SwiSy SpiSy

(a) (b) (c) (d) (e) (M) (g)

18/6 RFT (a) (a) CS 53.4 770 7.2 456 4.3 0.08 0.24

20/7 NZ (a) (a) SS 48.7 770 7.9 436 4.5 0.05 0.34

36/8 TR 6 40 CS 45.5 902 9.9 610 6.7 0.05 0.42

38/6 TB 6 40 SS 40.1 1185 15 654 8.2 0.05 0.48

39/4 TB 6 40 SS 40.1 1248 16 674 8.4 0.04 0.00

(a) RFT = reinforced (with pad) fabricated tee; 4 NPS, Sch. 40 branch; 8 NPS, Sch. 40 run; pad thickness

0.322 in.

NZ = nozzle (see Figure 13).

TR = 6x6x6 ANSI B16.9 tee, loaded through run.

TB = 6x6x6 ANSI B16.9 tee, fixed at one run end, branch loaded.

(b) Pipe material: CS = carbon steel, SA106-B; SS = stainless steel, SA312 Type 316.

(c) Sy = material yield strength (from Appendix D of Reference 11). For Tests 18 and 20, run pipe material yield

strengths.

(d) Moments from Appendix B of Reference 11.

Response spectrum analysis based on 2% damping, + /-15% peak broadening.

See text for conversion of moments to stresses.

(e) Same as (d), except 5% damping was used.

(f) Sw = stress due to weight = B2Mw/Z where Mw = moment due to weight.

(g) Sp = stress due to internal pressure = PD/2t

where P = internal pressure

D = run pipe mean diameter

t = run pipe nominal wall thickness

61 NUREG-1367
Table 9 Reference 13 and Other Piping System Tests: Comparison with Elastic Analyses

Test Damping,

Ref. Location System Description Mtd. S/2Sy %

(a) (b) (c) (c)

13 ETEC System 1, 6 NPS and 3 NPS, Sch. 40 CS 16 5

See Figure 15 and Section 5

13 ETEC System 2, 6 NPS and 4 NPS, Sch. 40 SS 12 5

See Figure 16 and Section 5

14 HEDL 1 NPS, Sch. 40 SS 12 5

See Figure 20

15 ANCO Z bend, 4 NPS, Sch. 40 CS 1.6 2

16 ANCO 8 NPS and 6 NPS, Sch. 40 CS 3 3

No branches

16 ANCO 8 NPS and 6 NPS, Sch. 40 CS ---

Two 3 NPS, Sch. 40 branches

17 ETEC 3 NPS, Sch. 40 CS 14 5

One 3 NPS, Sch. 40 branch

17 ETEC 6 NPS, Sch. 40 CS 8.3 5

One 3 NPS, Sch. 40 branch

18 KWU 4.5-in. outer diameter, 0.165-in. wall thickness SS 2.1 2

(Germany) 2.38-in. outer diameter, 0.114-in, wall branch

19 HDR 18-in.- to 4.5-in.-outer-diameter pipes SS 2.4 3

(Germany) D/t = about 15

20 Tadotsu 1/2.5 scale model of one loop of SS (6.6) (d)

(Japan) PWR primary coolant system

D/t = about 12

(a) ETEC = Energy Technology Engineering Center, Canoga Park, California

HEDL Hanford Engineering Development Laboratory, Richland, Washington

ANCO = ANCO Engineers, Culver City, California

KWU = Kraftwerk Union, Aktiengesellschaft, Federal Republic of Germany

HDR Heissdampfreaktor, Kahl/Main, Federal Republic of Germany

Tadotsu = Tadotsu Engineering Laboratory, Tadotsu-cho, Kagawa Prefecture, Japan

(b) CS = carbon steel pipe material (e.g., A106-B)

SS = austenitic stainless steel pipe material (e.g., A312 Type 304)

(c) S = calculated stress using response spectrum analysis with indicated damping, except for Reference 20.

S for Reference 20 is from a time history analysis.

Sy = yield strength of piping material.

(d) Time history analysis

NUREG-1367 62
NRC FORM 335 U.S. NUCLEAR REGULATORY COMMISSION 1. REPORT NUMBER

(2-89) (Assigned by NRC, Add Vol.,

NRCM 1102, Supp., Rev,, and Addendum Num3201, 3202 BIBLIOGRAPHIC DATA SHEET bers, if any,)

(See Instructions on the reverse) NUREG-1367

2. TITLE AND SUBTITLE

3. DATE REPORT PUBLISHED

Functional Capability of Piping Systems MONTH " I YEAR

November 1992

4. FIN OR GRANT NUMBER

5. AUTHOR(S) 6. TYPE OF REPORT

D. Terao, E. C. Rodabaugh Technical

7. PERIOD COVERED (Inclusive Dates)

8. PERFORMING ORGANIZATION - NAME AND ADDRESS (If NRC, provide Division, Office or Region, U.S. Nuclear Regulatory Commission, and

mailing address; if contractor, provide name and mailing address.)

Division of Engineering

Office of Nuclear Reactor Regulation

U.S. Nuclear Regulatory Commission

Washington, DC 20555

9. SPONSORING ORGANIZATION - NAME AND ADDRESS (If NRC, type "Same as above"; if contractor, provide NRC Division, Office or Region,

U.S. Nuclear Regulatory Commission, and mailing address.)

Same as above

10. SUPPLEMENTARY NOTES

11. ABSTRACT (200 words or less)

General Design Criterion 1 of Appendix A to Part 50 of Title 10 of the Code of Federal Regulations requires, in part,

that structures, systems, and components important to safety be designed to withstand the effects of earthquakes without a loss of capability to perform their safety function. The function of a piping system is to convey fluids from one

location to another. The functional capability of a piping system might be lost if, for example, the cross-sectional flow

area of the pipe were deformed to such an extent that the required flow through the pipe would be restricted.

The objective of this report is to examine the present rules in the American Society of Mechanical Engineers Boiler

and Pressure Vessel Code, Section III, and potential changes to these rules, to determine if they are adequate for ensuring the functional capability of safety-related piping systems in nuclear power plants.

12. KEY WORDS/DESCRIPTORS (List words or phrases that will assist researchers In locating the report,) 13. AVAILABILITY STATEMENT

Unlimited

14. SECURITY CLASSIFICATION

piping' (This Page) operability limits U casfe

functional capability Unclassified

stress limit (This Report)

Unclassified

15. NUMBER OF PAGES

16. PRICE

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NUREG-1367 FUNCTIONAL CAPABILITY OF PIPING SYSTEMS NOVEMBER 1992

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