NUREG/CR-6889

BNL-NUREG-74978-2005

Seismic Analysis of Simplified Piping

Systems for the NUPEC Ultimate

Strength Piping Test Program

Brookhaven National Laboratory

U.S. Nuclear Regulatory Commission

Office of Nuclear Regulatory Research

Washington, DC 20555-0001
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NUREG/CR-6889

BNL-NURiEG-74978-2005

Seismic Analysis of Simplified Piping

Systems for the NUPEC Ultimate

Strength Piping Test Program

Manuscript Completed: September 2005

Date Published: December 2005

Prepared by

G. DeGrassi and C. Hofmayer

Brookhaven National Laboratory

Upton, NY 11973-5000

S. Ali, NRC Project Manager

Prepared for

Division of Engineering Technology

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555-0001

NRC Job Code N6076
ABSTRACT

The Nuclear Power Engineering Corporation (NUPEC) of Japan has conducted a multi-year test program

for the Ministry of Economy, Trade and Industry (METI) of Japan to investigate the behavior of typical

Nuclear Power Plant (NPP) piping systems under large seismic loads. The objectives of this program are

to develop a better understanding of the elasto-plastic response and ultimate strength of nuclear piping

systems, to ascertain the seismic safety margins in current piping design codes, and to assess new code

allowable stress rules. The test program consisted of a series of static and dynamic material tests, piping

component tests, simplified piping system tests, and large scale piping tests. As part of collaborative

efforts between the United States and Japan on seismic issues, the U.S. Nuclear Regulatory Commission

(NRC) and the Brookhaven National Laboratory (BNL) participated in this program by performing both

pre-test and post-test analysis for selected tests, and by evaluation of program results. This report

presents a summary of the NUPEC tests, describes the BNL post-test analyses of selected simplified

piping system tests, and discusses the insights gained from this program.

iii
FOREWORD

This report documents the collaborative efforts between the U.S. Nuclear Regulatory Commission (NRC)

and the Nuclear Power Engineering Corporation of Japan (NUPEC) under the NRC-NUPEC agreement

in the area of seismic engineering research. Under contract to the NRC, Brookhaven National Laboratory

(BNL) participated in this program and performed this study. This piping test program is one of a series

of large-scale seismic testing programs of nuclear power plant (NPP) structures, systems and components

conducted by NUPEC. The research conducted and information exchanged under this collaborative agreement

have allowed the NRC to obtain valuable seismic data from large-scale test programs that are not available

anywhere else in the world. In addition, this program has enhanced interactions with various Japanese

organizations to promote the exchange of information and awareness of ongoing seismic research in Japan.

The NUPEC Ultimate Strength Piping Test Program was a multi-year program to investigate the behavior

of NPP piping systems under large seismic loads. The objectives were to develop a better understanding

of the elasto-plastic response and ultimate strength of nuclear piping systems, ascertain the seismic safety margins

in current piping design codes, and assess new code-allowable stress rules. The program included static

and dynamic loading tests of piping material specimens and piping components, and seismic shaking table tests

of piping systems. The seismic shaking tests involved small "simplified" piping system test specimen configurations,

as well as large-scale representative piping system configurations. In both series of seismic tests, the specimens

were shaken to levels well above typical design earthquake excitation levels. A summary description of these tests

and their results is included in this report.

The NRC's primary contribution to this collaborative effort was the analysis of selected simplified piping system

test runs. Specifically, the analyses included (1) linear analyses and stress evaluations in accordance with

the piping design rules of the American Society of Mechanical Engineers (ASME) Boiler and Pressure Vessel Code,

and (2) nonlinear analyses to predict the response of simplified piping systems under beyond-design-basis levels

of earthquake motion. The results of the linear analyses provided comparative information on the minimum

safety margins in various types of piping components provided by different versions of the ASME Code

including the latest NRC-endorsed version (1993) and the most recent version (2004) which is still undergoing

NRC review. The nonlinear analyses focused on the capability of current theoretical methods and computer programs

for predicting the response of piping systems under large earthquake loads. Investigations using different

material plasticity models demonstrated the difficulty of accurately predicting the piping response when it is

subjected to large earthquake loads, and the importance of validating the analytical model against test data.

From a regulatory standpoint, it is essential to understand the limitations and sensitivities of these nonlinear

analysis methods and the importance of validating these methods against test data. The NUPEC tests have

provided valuable test data, which can be used to benchmark these analytical models and methods to provide

a better understanding of their limitations. The NRC is continuing the collaborative efforts on the Ultimate

Strength Piping Test Program and will analyze the large-scale piping system test runs and compare the results

with test data. Further insights from these efforts will be provided in a followup report.

Carl . Paperie Directo

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

V
CONTENTS

Page

Abstract .................

Foreword ............. v

Executive Summary ............ xi

A cknow ledgm entsv....................................................................................................................................... xv

1.0 INTRODUCTION ..

2.0 DESCRIPTION OF NUPEC TEST PROGRAM .. 3

2.1 Introduction .3

2.2 Material Tests .3

2.3 Piping Component Tests .3

2.4 Simplified Piping System Tests .4

2.5 Large-Scale Piping System Tests .6

3.0 BNL LINEAR ANALYSES AND CODE EVALUATIONS . .27

3.1 Introduction .27

3.2 ASME Code Criteria .27

3.3 Linear Analysis and Evaluation of Simplified Piping Systems .29

4.0 BNL COMPONENT TEST NONLINEAR ANALYSES . .39

4.1 Introduction .39

4.2 Development of Piping Elbow Model .39

4.3 Development of Plasticity Models .40

4.4 Determination of Material Parameters .41

4.4.1 Bilinear Kinematic Hardening Model .41

4.4.2 Multilinear Kinematic Hardening Model .42

4.4.3 Chaboche Nonlinear Kinematic Hardening Model .42

4.5 Elbow Static Cycling Test Nonlinear Analysis .43

5.0 BNL SIMPLIFIED PIPING TEST NONLINEAR ANALYSES . .47

5.1 Introduction .47

5.2 Development and Pre-Test Analysis of Two-Dimensional Simplified Piping System

Model .47

5.3 Additional A-2 Simplified Piping System Test Investigations .48

5.4 Final A-2 Simplified Piping System Test Analyses and Comparisons .50

5.5 Final Two-Dimensional and Three-Dimensional Simplified Piping System Test

Analyses. 51

vii
6.0 CONCLUSIONS AND RECOMMENDATIONS ...................................................... 77

7.0 REFERENCES ...................................................... 79

Figures

2-1 Test Specimens from the NUPEC Material Tests ...................................................... 10

2-2 Monotonic Loading Stress-Strain Curve for STS410 Carbon Steel ............................................. 11

2-3 Uniaxial Cycling Test Hysteresis Curves for STS410 Carbon Steel ............................................. 11

2-4 Cyclic Stress-Strain Curve for STS410 Carbon Steel ...................................................... 12

2-5 Static Elbow Cycling Test Setup ...................................................... 12

2-6 Static Elbow Cycling Test Strain Gage Locations ...................................................... 13

2-7 Elbow Static Cycling Test SE4 Hoop Strain vs. Cycle at Elbow Flank ....................................... 13

2-8 Dynamic Elbow Test Setup ...................................................... 14

2-9 Model A Simplified Piping System Test Setup ...................................................... 15

2-10 Model A Test Instrumentation ...................................................... 16

2-11 Model A Test Strain Gage Locations ...................................................... 16

2-12 Model B Simplified Piping System Test Setup ...................................................... 17

2-13 Model B Test Instrumentation ...................................................... 18

2-14 Model B Test Strain Gage Locations ...................................................... 18

2-15 A-I Test Acceleration Time History Input Motion ...................................................... 19

2-16 A-2 Test Acceleration Time History Input Motion ...................................................... 19

2-17 B-I Test Acceleration Time History Input Motion (Horizontal) ................................................... 20

2-18 B-2 Test Acceleration Time History Input Motion (Vertical) ...................................................... 20

2-19 B-3 Test Acceleration Time History Input Motion (Horizontal and Vertical) .............................. 21

2-20 A-2 Test - Pipe Center Displacement - DI (X) ...................................................... 22

2-21 A-2 Test - Elbow A Opening Displacement - D2(L) ...................................................... 22

2-22 A-2 Test - Hoop Strain at Elbow A Flank - SAD-3H ...................................................... 23

2-23 A-2 Test - Axial Strain a Elbow A Flank - SAD-3A ...................................................... 23

2-24 B-3 Test - Pipe Center Horizontal Displacement - DI(X) ...................................................... 24

2-25 B-3 Test - Pipe Center Vertical Displacement - DI (Z) ...................................................... 24

2-26 B-3 Test - Hoop Strain at Nozzle - SA16-2H ...................................................... 25

2-27 B-3 Test - Axial Strain at Nozzle - SA17-2A ...................................................... 25

2-28 Large Scale Piping System Test Setup ...................................................... 26

3-1 Simplified Piping System Test - Models A and B Linear Elastic ANSYS Piping Element

Models.34

viii
3-2 A-I Test Horizontal Response Spectrum (2% and 5% Damping) .............................................. 35

3-3 A-2 Test Horizontal Response Spectrum (2% and 5% Damping) .............................................. 35

3-4 B-I Test Horizontal Response Spectrum (2% and 5% Damping) .............................................. 36

3-5 B-2 Test Vertical Response Spectrum (2% and 5% Damping) .............................................. 36

3-6 B-3 Test Horizontal Response Spectrum (2% and 5% Damping) .............................................. 37

3-7 B-3 Test Vertical Response Spectrum (2% and 5% Damping) .............................................. 37

4-1 SEA Component Test Elbow Specimen .............................................. 44

4-2 SEA Test ANSYS Shell Element Model .............................................. 44

4-3 BKIN and MKIN Model Stress-Strain Curves .............................................. 45

4-4 CHAB Model Stress-Strain Curve .............................................. 45

4-5 Elbow Static Cycling Test SE4 - Hoop Strain vs. Number of Cycles at Elbow Flank ................ 46

5-1 ANSYS Model A Elasto-Plastic Shell and Beam Element Model ................................................ 55

5-2 A-2 Nonlinear Simulation Analysis Results - Displacement at Pipe Center - DI (X) .................. 56

5-3 A-2 Nonlinear Simulation Analysis Results - Hoop Strain at Elbow Flank - SAD-3H ............... 57

5.4 ANSYS Model B Elasto-Plastic Shell and Beam Element Model ................................................ 58

5.5 A-I Test Pipe Center Displacement Comparison - DI(X) ............................................................ 59

5-6 A-I Test Hoop Strain at Elbow A Flank Comparison - SAD-3H ................................................. 60

5-7 A-I Test Hoop Strain at Elbow A Flank Inside Surface Compared to Test Hoop Strain at

Outside Surface - SAD-3H ................................................................. 61

5-8 A-2 Test Pipe Center Displacement Comparison - DI (X) ............................................................ 62

5-9 A-2 Test Elbow A Opening Displacement Comparison - D2(L) .................................................. 63

5-10 A-2 Test Hoop Strain at Elbow A Flank Comparison - SAD-3H ................................................. 64

5-11 A-2 Test Axial Strain at Elbow A Flank Comparison - SAD-3A ................................................. 65

5-12 B-I Test Pipe Center Horizontal Displacement Comparison - DI(X) ............... ........................... 66

5-13 B-i Test Hoop Strain at Nozzle Comparison - SA16-2H ............................................................. 67

5-14 B-I Test Axial Strain at Nozzle Comparison - SA17-2A ............................................................. 68

5-15 B-2 Test Pipe Center Vertical Displacement Comparison - DI (Z) ............................................... 69

5-16 B-3 Test Pipe Center Horizontal Displacement Comparison - DI (X) .......................................... 70

5-17 B-3 Test Pipe Center Vertical Displacement Comparison - DI(Z) ............................................... 71

5-18 B-3 Test Hoop Strain at Nozzle Comparison - SA16-2H ............................................................. 72

5-19 B-3 Test Axial Strain at Nozzle Comparison - SA 17-2A ............................................................. 73

5-20 B-3 Test Hoop Strain at Elbow A Flank Comparison - SAD-7H ................................................. 74

5-21 B-3 Test Hoop Strain at Elbow A Flank Inside Surface Compared to Test Hoop Strain at

Outside Surface - SAD-7H ................................................................. 75

ix
Tables

2-1 NUPEC Static Piping Component Test Load Cases ............................................ 7

2-2 NUPEC Dynamic Piping Component Test Load Cases ............... .............................. 7

2-3 NUPEC Simplified Piping System Test Load Cases ............................................ 8

2-4 NUPEC Large-Scale Piping System Test Load Cases ............................................. 9

3-1 A-I Test Maximum Code Stresses and Margins ............................................ 32

3-2 A-2 Test Maximum Code Stresses and Margins ............................................ 32

3-3 B-I Test Maximum Code Stresses and Margins ............................................ 33

3-4 B-2 Test Maximum Code Stresses and Margins ............................................ 33

3-5 B-3 Test Maximum Code Stresses and Margins ............................................ 33

5-1 Summary of Post-Test Analysis Results and Test Comparisons ............................................ 54

x
EXECUTIVE SUMMARY

This report presents the results of BNL collaboration efforts on the NUPEC Ultimate Strength Piping Test

Program. NUPEC conducted a multi-year test program to investigate the behavior of typical NPP piping

systems under large seismic loads. NUPEC's objectives were to develop a better understanding of the

elasto-plastic response and ultimate strength of nuclear piping systems, to ascertain the seismic safety

margins in current piping design codes, and to assess new code allowable stress rules. This test program

included monotonic and cyclic loading tests of piping material specimens, static and dynamic tests of

piping components such as elbows and tees, seismic shaking table tests of two simple piping systems and

of a representative large-scale piping system.

NRC and BNL participated in this NUPEC program under a collaboration agreement between the United

States and Japan on seismic issues. As part of the collaboration, BNL performed pre-test and post-test

analyses for selected tests. NUPEC presented their test results to NRC and BNL and provided selected

test data in electronic format. Periodic technical meetings were held in Japan and the U.S. during the

course of the program to discuss and exchange ideas on the test program plans, to review and evaluate the

test results, and to review and evaluate both pre-test and post-test analyses. This report provides a

summary of the NUPEC tests, describes the BNL post-test analyses, and discusses the insights gained

from this program.

A major goal of the NUPEC test program was to perform a seismic proving test of a representative NPP

large-scale piping system. In order to gain a better understanding of the behavior of piping under severe

seismic loads, NUPEC first carried out a series of material and piping component tests and shaking table

tests of small (simplified) piping systems. The material specimen tensile tests provided stress-strain

curves and basic material strength properties including yield and ultimate stresses, percent elongation and

ultimate strain. Strain-controlled cycling tests were also performed which provided cyclic stress-strain

curves and stress-strain hysteresis curves for incremental cycling amplitudes. The piping component tests

included static and dynamic cycling tests of pressurized elbows, tees, nozzles and reducers. The tests

were carried out at high strain levels and illustrated the accumulation of ratcheting strain. The specimens

were cycled until a through-wall crack occurred.

The simplified piping system tests were designed to excite the systems to seismic levels well above the

typical design earthquake levels in order to induce elasto-plastic response in the system's components.

The test systems were simple two-dimensional (Model A) and three-dimensional (Model B) geometric

configurations of 65 mm (2 12 inch) nominal diameter pipe. The Model A pipe routing configuration was

in the horizontal plane and was excited on a shaking table in the horizontal direction. The Model B

routing'was in the horizontal and vertical planes and was subjected to horizontal, vertical, or simultaneous

horizontal and vertical seismic motions. The piping systems were pressurized and instrumented with

accelerometers, displacement sensors and strain gages at critical locations which were monitored during

each test. A number of tests to high strain levels were carried out to investigate different loadings and

design variations. There were no failures reported during the application of any single time history.

However, NUPEC repeated selected high input level tests for both a Model A and a Model B specimen to

induce failure. In these tests, cracking at an elbow was observed in the Model A configuration after five

tests at a high seismic level. For the Model B, 18 applications of a high level seismic input resulted in a

crack at a nozzle weld. In both cases, the failures were attributed to fatigue ratcheting.

In the final phase of the test program, NUPEC carried out a series of shaking table tests on a

representative large-scale piping system. The test specimen was a 200 mm (8 inch) nominal diameter

Schedule 40 carbon steel pipe. It included straight pipe, nine elbows, a tee, and a 1000 kg (2200 lb)

added mass representing a valve. The system was pressurized and supported by nozzles, an anchor, five

two-directional restraints, and a spring hanger. The three-dimensional routing of the piping was

xi
representative of nuclear power plant piping systems. The system was instrumented with accelerometers,

displacement sensors, and strain gages at critical locations.

Two series of tests were performed: a design confirmation test and an ultimate strength test. For the

design confirmation test, the first input motion was selected to induce a maximum stress equal to the

Japanese Code primary stress limit of 3Sm. Additional design confirmation tests were then performed

using the same test specimen with the seismic input incrementally increased up to a maximum elasticallycalculated stress of 13.5 Sm in the system. The tests were conducted at room temperature and

simultaneous horizontal and vertical seismic input motions were applied in all test runs. No evidence of

pipe failure was observed.

The ultimate strength test used a similar but modified piping test specimen. An additional mass was

added and a support was removed. The intent of these modifications was to induce failure in the system.

The pipe was internally pressurized and the tests were performed at room temperature. Horizontal

seismic input motion corresponding to a maximum elastically-calculated stress level of 24 Sm was applied.

The test was repeated until failure occurred. During the fifth test run, a through-wall crack developed in

an elbow. An examination confirmed that the failure was the result of fatigue ratcheting.

In accordance with the collaboration agreement, BNL performed finite element analyses for selected

simplified piping test runs. The BNL analyses had two major objectives. First, linear analyses were

performed in accordance with ASME Code requirements. Since the ASME seismic evaluation

requirements have been undergoing significant changes in recent years, a comparative study was carried

out in which the piping systems were analyzed and evaluated according to three different Code versions.

They include the latest NRC-endorsed version (1993 Addenda), the 1994 Addenda which incorporated an

alternative higher seismic stress allowable which was not endorsed by NRC, and the most recent 2004

Edition which reinstated the original (1993) stress allowable but introduced an alternative method for

calculating seismic stresses in certain piping components. Since the NUPEC simplified piping tests

loaded the piping systems to stress levels above ASME Code allowables, the analyses were used to

identify the degree to which the calculated stresses exceeded the Code allowable stresses. The ratios of

calculated to allowable stress define the minimum margins of safety provided by different versions of the

ASME Code rules. The BNL study includes a comparison of the ratios calculated in accordance with the

different rules which provides a measure of the relative safety margins between different code versions.

Additional analyses and Code evaluations are recommended for other tests in the NUPEC program,

including those in which the simplified and the large-scale piping systems were tested to failure, to

provide further substantiation of Code margins.

The second major objective of the BNL analysis was to perform nonlinear analyses to investigate and

evaluate the adequacy of the latest commercially available methods and computer programs for predicting

the elasto-plastic response of piping systems subjected to large earthquake loads. BNL used the ANSYS

computer program for performing these analyses. ANSYS is a computer program commonly used by the

nuclear industry which has been enhanced in recent years to include a number of improvements in its

nonlinear analysis capability. Among these improvements was the inclusion of the Chaboche nonlinear

kinematic hardening model. This plasticity model was expected to provide better simulation of the strain

ratcheting behavior that a pressurized pipe is expected to exhibit under seismic loading than is possible

with classical linear hardening models. BNL utilized results of the NUPEC material tests to develop

input parameters for the piping models. As part of the development process, BNL first created a finite

element model of a piping elbow that was subjected to static cycling tests in the NUPEC component test

program. Three sets of simulation analyses were carried out using three different plastic hardening

models: bilinear kinematic hardening, multilinear kinematic hardening, and Chaboche nonlinear

hardening. The results of the analyses confirmed that the first two linear hardening models did not

xii
adequately predict the strain ratcheting behavior observed in the test. The Chaboche nonlinear model

provided the most accurate simulation of strain ratcheting behavior.

ANSYS models were developed for the two-dimensional (Model A) and the three-dimensional (Model B)

simplified piping systems. Preliminary seismic time history analyses were first performed for the Model

A high seismic level (A-2) test run. A number of studies were performed to investigate the sensitivity of

modeling and analysis parameters on the results. The studies demonstrated the importance of accurately

modeling the piping system to predict the correct system dynamic characteristics in an elasto-plastic

analysis. Small changes in system frequencies resulted in significant differences in response. Analyses

were also conducted using the two linear kinematic hardening models and the Chaboche nonlinear

hardening model. Comparisons of time history results to test data clearly showed that the Chaboche

model predicted displacement and strain ratcheting with the greatest accuracy.

The Chaboche model was used in all subsequent nonlinear analyses of the simplified piping tests.

A total of five analyses were performed: two levels of horizontal seismic time history input motion for the

two-dimensional system (test runs A-I and A-2), and a horizontal, a vertical, and a combined horizontal

and vertical seismic input motion for the three-dimensional system (test runs B-i, B-2 and B-3).

Displacements and strains at critical locations were compared to test measurements. The application of

the Chaboche nonlinear kinematic hardening model to the simulation analyses of the other four simplified

piping system tests demonstrated that the displacement time history responses were predicted with

reasonably good accuracy. Accurate simulation of strain and strain ratcheting was achieved in some but

not all cases which indicates that even the Chaboche nonlinear model has limitations which need to be

better understood. From a regulatory standpoint, these analyses demonstrated the importance of

understanding the limitations and sensitivities of nonlinear analysis methods and of commonly applied

plasticity models. Further studies to investigate the limitations of plasticity models on representative

large-scale piping systems may provide additional insights.

The NRC and BNL are continuing the collaboration efforts with Japan and plan to analyze the large-scale

piping system test models and compare the results with the test data. Further insights from this

collaboration effort will be provided in a follow-up report.

xiii
ACKNOWLEDGMENTS

The research program described in this report was sponsored by the Office of Nuclear Regulatory

Research of the U.S. Nuclear Regulatory Commission. The authors would like to express their gratitude

to Dr. Syed Ali and Dr. Andrew Murphy, NRC Project Managers, for the technical and administrative

support they have provided in performing this study.

This research program was performed as part of the Implementing Agreement between the U.S. Nuclear

Regulatory Commission and the Nuclear Power Engineering Corporation of Japan in the Area of Seismic

Engineering Research. This agreement is an item of the Implementing Arrangement between the United

States Nuclear Regulatory Commission and the Nuclear and Industrial Safety Administration of Japan for

Cooperation in the Field of Nuclear Regulatory Matters and Nuclear Safety Research and Development.

All of the test results and information about the test models included in this report were provided by

NUPEC and are greatly appreciated. The authors especially thank Dr. Kenichi Suzuki, Program Manager,

Seismic Engineering Center of NUPEC, for his support and technical guidance throughout the

collaborative study. The authors also wish to acknowledge the helpfulness and cooperation provided by

Mr. Hiroshi Abe and Dr. Yoshio Namita of NUPEC, Mr. Koichi Tai of Mitsubishi Heavy Industries, and

Mr. Tatsuya Fujiwaka of IHI Heavy Industries.

The authors also express special thanks to Ms. Susan Signorelli for her secretarial help throughout this

program and in the preparation of this report.

xv
1.0 INTRODUCTION

The Nuclear Power Engineering Corporation (NUPEC) of Japan conducted a multi-year test program for

the Ministry of Economy, Trade and Industry (METI) of Japan to investigate the behavior of typical

Nuclear Power Plant (NPP) piping systems under large seismic loads. The objectives of this program

were to develop a better understanding of the elasto-plastic response and ultimate strength of nuclear

piping systems, to ascertain the seismic safety margins in current piping design codes, and to assess new

code allowable stress rules. The test program consisted of a series of static and dynamic material tests,

piping component tests, simplified piping system tests, and large scale piping tests.

As part of collaborative efforts between the United States and Japan on seismic issues, the U.S. Nuclear

Regulatory Commission (NRC) and the Brookhaven National Laboratory (BNL) participated in this

program. The collaboration program included a series of periodic technical meetings held in Japan and

the U.S. to discuss and exchange ideas on the test program plans, to review and evaluate the test results,

and to review and evaluate both pre-test and post-test analyses. In accordance with the collaboration

agreement, NUPEC provided their test results to NRC and BNL in the form of presentations and also

provided selected test data for test runs analyzed by BNL in electronic format. NRC agreed to have BNL

perform pre-test and post-test analyses for selected tests and provide the results in a series of presentations

and reports.

The purpose of this report is to present a summary of the NUPEC test program, to describe the BNL posttest analyses performed to date, and to discuss the insights gained from the program. Section 2.0 of this

report provides a description and summary results of the NUPEC test program which consisted of four

phases of testing: material tests, piping component tests, simplified piping system tests, and large-scale

piping system tests. Section 3.0 presents the BNL linear seismic piping analysis and ASME Code

evaluations for selected simplified piping system tests. These analyses identify the highest stressed

components and the degree to which the maximum stresses exceed Code allowables. By performing the

evaluations in accordance with different versions of the Code, the minimum margins to failure associated

with the NRC endorsed Code version (1993 Addenda) are calculated and compared with the margins

based on the 1994 Addenda and the latest 2004 Edition.

The results of the BNL nonlinear analyses are presented in Sections 4.0 and 5.0. Section 4.0 focuses on

the development of an accurate nonlinear plasticity model. Analyses of an elbow component test were

performed using two linear hardening models and a nonlinear hardening model. Comparison of the

analysis results to test data demonstrated the nonlinear model's superior capability in predicting strain

ratcheting compared to the linear models. Section 5.0 presents the results of the post-test analyses for the

two-dimensional (Model A) and the three-dimensional (Model B) simplified piping system tests and

compares key displacements and strains to test measurements. A summary of conclusions including

regulatory insights and lessons learned from the program, and recommendations for further research are

presented in Section 6.0.

I
2.0 DESCRIPTION OF NUPEC TEST PROGRAM

2.1 Introduction

A major goal of the NUPEC test program was the performance of a seismic proving test of a

representative large-scale piping system. In order to gain a better understanding of the piping behavior

prior to performing the proving test, NUPEC conducted a series of material and piping component static

and dynamic tests, and simplified piping system seismic tests. This chapter provides a summary

description of each of these series of tests.

2.2 Material Tests

NUPEC carried out a series of static monotonic loading and cyclic loading tests to develop stress-strain

curves and properties for typical piping materials. The materials included carbon steel (grades STS410

and SCV410), stainless steel (grade SUS304TP), and low-alloy steel (STPA24). Typical test specimens

are shown in Figures 2-1(a) and 2-1(b). The tests were performed at room temperature and at 300'C

(5721F). In the monotonic loading tests, the specimens were tensile tested to failure. The results of the

tests provided the stress-strain curves and basic material properties including yield strength, ultimate

strength, percent elongation, percent area reduction, and ultimate strain. A typical stress-strain curve for

STS410 carbon steel is shown in Figure 2-2. In the cyclic loading tests, the specimens were subjected to

strain-controlled incremental cycling in consecutive blocks of twenty increasing amplitude cycles

followed by twenty decreasing amplitude cycles. Within each block of cycles, the strain amplitude was

incrementally increased linearly from 0% to a maximum of 2.5% strain (5% strain range) at 20 cycles,

and then incrementally decreased back to 0% strain. These tests provided stress-strain hysteresis curves

for the four piping materials at room temperature and at 300'C (5721F). They also provided cyclic stressstrain curves for the materials for strains up to 2.5%. A typical series of hysteresis curves for STS410

carbon steel are shown in Figure 2-3. A typical cyclic stress-strain curve is shown in Figure 2-4.

NUPEC also carried out a series of low-cycle strain-controlled ratcheting fatigue tests for typical piping

materials at room temperature and at 300'C (5720F). Both uniaxial and biaxial loading tests were

performed. The uniaxial tests also used the test specimen shown in Figure 2-l(b), as well as the same

four materials identified above. A unique aspect of the uniaxial fatigue tests is that in some of the test

runs, an accumulated mean strain was artificially induced at a given rate per cycle to simulate the

ratcheting strain that may develop in components. Test run variations included different cyclic strain

ranges, mean strain accumulation rates, and temperatures. Test results provided data on strain versus

number of cycles for different strain ranges and mean strain accumulation rates. They demonstrated that

higher mean strain accumulation rates result in shorter fatigue lives for all materials and temperatures.

The biaxial tests used a miniature pipe specimen shown in Figure 2-1(c) to simulate the stress field of

piping components subjected to internal pressure and seismic loading. These tests were performed for

STS410 carbon steel and SUS304TP stainless steel specimens. Test run variations included internal

pressure, displacement amplitude, and temperature. The test results demonstrated that the test specimens

experienced circumferential swelling due to ratcheting and developed through-wall circumferential cracks

in all but two cases in which longitudinal through-wall cracks developed. The test results were plotted on

S-N diagrams (strain range versus number of cycles to failure), and on Dd-Df diagrams (where Df =

(failure cycles)/(failure cycles without ratcheting) and Dd = (accumulated strain)/(fracture strain)) [1, 2].

2.3 Piping Component Tests

In the second phase of the program, NUPEC conducted static and dynamic tests on typical piping

components which included elbows, tees, nozzles and reducers. A total of sixteen static tests were

performed. A typical test setup for a static elbow test is shown in Figure 2-5. A summary of the test

3
conditions for each load case is presented in Table 2-1. Test specimen variations included component

type, material, diameter, and pipe schedule. All tests were conducted at room temperature except for one

elbow test which was conducted at 300 C (5720F). The test specimens were subjected to either cyclic or

monotonic loads. Nearly all of the tests were deflection controlled with the exception of one elbow test

which was load controlled. In the monotonic loading tests, the specimens were unpressurized and loaded

to failure to determine the lower bound collapse moment. In the cyclic loading tests, the test specimens

were pressurized to induce an internal pressure stress equal to Sm, and then subjected to quasi-static

sinusoidal displacements until a fatigue crack developed. The applied displacement amplitudes were

selected to induce a maximum strain range of 3% in all cases but one, in which a larger displacement

amplitude corresponding to a 6% strain range was applied. The specimens were cycled until through-wall

cracks were detected. The test specimens were instrumented with strain gages at critical locations as

illustrated in Figure 2-6. Load cells and displacement transducers were used to monitor force-deflection

data as shown in Figure 2-5. NUPEC reported the monotonic test results as plots of load versus

deflection and strain versus deflection up to the collapse load. The cyclic test results were reported as

plots of load-deflection hysteresis curves, strain distributions at critical component sections, and

maximum strain versus number of cycles. The strain versus cycle plots illustrated the accumulation of

ratcheting strain during the tests. A typical plot of strain versus number of cycles up to failure for the SE4 test is shown in Figure 2-7.

The component dynamic tests were performed on a seismic shaking table. The test specimens were

subjected to either sinusoidal or seismic wave acceleration input motions. A large mass on rollers was

attached to one end of the test specimen. A typical test setup for a dynamic elbow test is shown in Figure

2-8. A total of twelve dynamic load tests were carried out. A summary of the test conditions is presented

in Table 2-2. Test specimen variations included pipe component type, material (STS410 carbon steel or

SUS304TP stainless steel), and diameter. All tests were conducted at room temperature. All specimens

were pressurized to induce a stress level of either Sm or 0.5 Sm. The acceleration input levels were

selected to induce a maximum strain range of 3% in ten load cases and 1.5% in two load cases. The

specimens were tested until though-wall cracking was observed. The test specimens were instrumented

with strain gages at critical locations are shown in Figure 2-6. Accelerometers and displacement

transducers were used to monitor acceleration input and response and displacement response as shown in

Figure 2-8. The test results were reported as plots of load-deflection hysteresis curves, displacement time

histories, and strain time histories. The strain versus time plots illustrated the accumulation of ratcheting

strain during the tests.

2.4 Simplified Piping System Tests

In the third phase of the program, NUPEC conducted a series of simplified piping system shaking table

tests to investigate the elasto-plastic response of small systems with typical piping components under

large seismic excitations. These tests included two-dimensional test model configurations (Model A) and

three-dimensional configurations (Model B). In the two-dimensional configuration, the pipe routing was

in the horizontal plane and was subjected to single-direction horizontal seismic input motion. A typical

Model A simplified piping system test setup is shown in Figure 2-9. The system consists of three straight

pipe runs connected by two long radius elbows. The pipe was 65 mm (2-1/2 inch) Schedule 40 carbon

steel grade JIS STS410. The system was supported at both ends by specially designed ball bearing

pinned supports. The system was pressurized to 19.8 MPa (2870 psi) to induce a pressure stress equal to

Sm. Two 200 kg (440 lb) concentrated masses and two 30 kg (66 lb) masses were clamped to the pipe to

provide the desired dynamic characteristics for the system. The larger weights were on rollers which

provided vertical gravity support. The system was instrumented with accelerometers, displacement

sensors, and strain gages at critical locations as shown in Figures 2-10 and 2-11.

4
Pipe routing for the three-dimensional configuration (Model B) was identical to the Model A routing in

the horizontal plane with the addition of a branch line in the vertical plane as shown in Figure 2-12. The

branch line was connected to the horizontal run pipe with a tee, and consisted of a riser pipe, a long radius

elbow, and a horizontal straight pipe which terminated at a nozzle. An additional 30 kg (66 lb)

concentrated mass was attached to the horizontal portion of the branch line. The system was supported by

ball bearing pinned supports at the ends of the lower horizontal run and by the fixed nozzle at the end of

the upper branch pipe. Instead of using roller supports at the large masses, variable spring hangers were

used to provide gravity support to the system without constraining vertical motion. The system was

pressurized to 19.8 MPa (2870 psi). In this three-dimensional configuration, the Model B piping could be

subjected to horizontal, vertical, or simultaneous horizontal and vertical seismic input motion. The

system was instrumented with accelerometers, displacement sensors, and strain gages at critical locations

as shown in Figures 2-13 and 2-14.

The basic seismic input excitations for the tests were selected to represent typical horizontal and vertical

building floor response spectra for Japanese PWR and BWR plants. However, the time scales were

adjusted to provide near resonant responses and the amplitudes of the seismic waves were modified to

obtain the desired large elasto-plastic responses. The tests were performed at two excitation levels. The

basic level identified as "medium" was intended to produce a maximum strain range of 1.5% with

equivalent pipe stress of 6 Sm to 9 Sm and an anticipated fatigue life of 1000 cycles. The higher input

level identified as "large" was selected to produce a maximum strain range of 3% corresponding to

equivalent pipe design stress of 10 Sm to 12 Sm and a fatigue life of 60 to 300 cycles.

NUPEC conducted a number of tests of both Model A and Model B specimens to investigate different

loading and design variations. Table 2-3 provides a summary of the load cases. In all cases, the test

specimen was a 65 mm (2-V2 inch) Schedule 40 pipe pressurized to 19.8 MPa (2870 psi). The basic twodimensional test (A-I) subjected the pipe to a single horizontal direction medium level seismic load. In

test A-2, a large level seismic load was applied. In test A-3, one of the pinned supports was changed to a

fixed support. In test A-4, the frequency of excitation was reduced to about one-half of the resonant

frequency. Test A-5 used a stainless steel instead of carbon steel test specimen. In test A-6 an actuator

was used at one of the pinned supports to provide independent support motion. Seismic acceleration time

history input motions for the A-I and A-2 load cases are shown in Figures 2-15 and 2-16.

The three-dimensional test (Model B) load cases subjected the test specimens to horizontal, vertical and

simultaneous horizontal and vertical motions as summarized in Table 2-3. Test B-i subjected the pipe to

a single horizontal direction medium level seismic load. Test B-2 subjected the pipe to a vertical

direction medium level seismic load. In test B-3, simultaneous horizontal and vertical medium level

seismic motions were applied. In test B-4, a large level seismic motion was applied. Finally in test B-5,

lower frequency horizontal and vertical seismic input motions were applied simultaneously. Seismic

acceleration time history input motions for the B-I, B-2 and B-3 load cases are shown in Figures 2-17, 2-

18, and 2-19.

The piping systems were monitored during each test to measure accelerations, displacements, and strains

at selected locations. Representative time history response plots for the A-2 test including displacements

at the pipe center, elbow A opening and closing displacements, and hoop and axial strains at the elbow A

flank are shown in Figures 2-20 though 2-23. Representative B-3 test time history responses including

horizontal and vertical displacements at the pipe center, and hoop and axial strains at the nozzle are

shown in Figures 2-24 through 2-27. The plots illustrate the accumulation of permanent deformations

and strains that occurred during the tests. There were no failures reported during the application of a

single seismic time history. However, NUPEC repeated selected high input level tests to induce failure.

NUPEC reported that cracking at the elbow A location was observed after five tests of the A-2 high level

seismic time history. Cracking at the nozzle occurred after 18 applications of the B-4 high level seismic

5
time history. For both the two-dimensional and three-dimensional tests, the failures were attributed to

fatigue ratcheting.

2.5 Large-Scale Piping System Tests

In the final phase of the test program, NUPEC performed a series of seismic shaking table tests on a

representative large-scale piping system. The tests were performed using the large high performance

shaking table at the Tadotsu Engineering Laboratory. The test specimen was a 200 mm (8 inch) nominal

diameter) Schedule 40 carbon steel (STS410) pipe. It included straight pipe, nine elbows, a tee, and a

1000 kg (2200 lb) added mass representing a valve. The system was supported by nozzles, an anchor,

five two-directional restraints, and a spring hanger. The three-dimensional routing of the piping system

represented typical configuration characteristics of safety-related Japanese nuclear power plant piping

systems as illustrated in Figure 2-28. The system was instrumented with accelerometers, displacement

sensors, and strain gages at critical locations.

Two series of tests were performed using two different test specimens. The first was a design

confirmation test and the second was an ultimate strength test. The seismic input motion was selected

based on the Japanese "S2" seismic wave for a PVYTR with modified time pitch and excitation level to

achieve the desired test responses. For the design confirmation test, the first input motion was selected to

induce a maximum stress equal to the Japanese Code primary stress limit of 3Sm. Additional design

confirmation tests were then performed using the same test specimen in which the seismic input was

incrementally increased until a maximum elastically-calculated stress of 13.5 Sm was reached. In all tests,

the piping was internally pressurized to induce a hoop stress equal to the design stress intensity Sm. The

tests were conducted at room temperature and simultaneous horizontal and vertical seismic input motions

were applied in all test runs. No evidence of pipe failure was observed. A summary of the test cases is

presented in Table 2-4.

The ultimate strength test used a similar but slightly modified test specimen. An additional 1000 kg

(2200 lb) mass was added and a two-directional support was removed. These modifications were needed

in order to induce failure at the highest stressed elbow within the shaking capability of the Tadotsu

shaking table. The pipe was internally pressurized and the tests were performed at room temperature. As

indicated in Table 2-4, horizontal seismic input motion corresponding to a maximum elasticallycalculated stress level of 24Sm was applied. The test was repeated until failure occurred. During the fifth

test run, a through-wall crack developed in the elbow. An examination confirmed that the failure was the

result of fatigue ratcheting.

6
Table 2-1 NUPEC Static Piping Component Test Load Cases

Test Component Matenal O.D. Thickness Internal Ternp. Load Load Load Control Viewpoint

No. Type Pressure Wave Level Direction Type

SE-1 Cyclic Elbow Basic Case

Sm Room Deflection (Carbon Steet)

SE-2 200A Sch40 Temp. Cyclic Load Load Control

SE-3 (A) None 3005 Large Temperature

SE-4 65A Pipe Daimeter

SE-5 200A Srh20 Cyclic DOameler/Tuckness

SE-6 SGV41 IO 400A sTD In-plane Deflection Ratio

SE-7 S 20 Elbow Basic Case

SE-t TP Extra Load Level

SE-9 STD None Room S Monotonic Lower Collapse Load

Temp___Deflection ________

ST-I STS41 0 200A4 Sch40TO Tee Basic Case

(A) 200A Sm L arge Cyclic

ST-2 Tee Deflection Diameterl Ttickness

SGV410 400AI STD Ratio

400A Pore -T-3 Monotonic Lower Collapse Load

SN-1 Bend Cyclic Nozzle Basic Case

Nozzle 200AN Sch4O Sni Large Tron Deflection St-2 N te STS4 I10 20500A/ Sc4 Sr ag Tors o elefon Loaa Directlion

-3 (A) Sct2 None M onotonic Lower Collapse Load

__ __S_ end Deflection _ _ _ _ _ _ _ _

St-I Reducer 200 Sctt40 Sm Large Cyclic Reducer Basic Case

1_ SOA_ .Deflection

Note) Load Level-Extra Large (around 6% In Max. Strain Range). Large (around 3% In Max. Strain Range), Small (Current Design Allowable Level)

O.D.--O OutsideDiameter(65A.2.SB. 1sOA-68. 200A885 250A=t08. 40OA.168)

Table 2-2 NUPEC Dynamic Piping Component Test Load Cases

Test Componenl Material O.0. Thickness Internal Temp. Load Load Load Control V owporit

No. Type Pressure Wave Level Direction Type

_ _tSress -I

DE-1 200A Sine Elbow Basic Case

Large (Carbon Steel)

DE-2 Sm Seismic Load Wave STS410 DE-3 (A) Sine Mdi.LaLei

DEt- 0.5Sm Internal Pressure

DE-5 Elbo In-laneLevel DE-5 Elbow 65A seismic Large ln-ptane Diameter

oE-6 200A Sm Room Sine Acceleration Elbow Basic Case

Temp. (Stainless Sleet)

DE-7 SUS304 - Sch40 Seismic Load Wave

DETt Middle load Level

Sine DE-9 .0.5m Internal Pressure Level

or- T ee ZQOA/ Tee B~asic Case * ~~200A Lrg

ON-1 None STS41O 200A Sm Seismic Nozzle Basic Case

(A) 250A Bend

DR-i Reducer gOON Reducer Basic case

Note) Load Level--- Large (around 3% in Max. Strain Range). Middte (around 1.5% In Max Strain Range). Small (Curnent Design Afowable Level)

7
Table 2-3 NUPEC Simplified Piping System Test Load Cases

Test No A-1 A-2 A-3 A-4 A-S A-6

o Pipe size:2 I/t'Xsch4O

(07S.3mmX5.2em) Yt X

Conceptual drawing o Stress due to internal P ji rtA -ttesi

of piping system pressure ( A- nanelG0,,

2D-PIPlng o Stress due to dead weight /I'"' e 1

2D-Pptn S 0.55. rh af1nnm' sys tem .r fed1.nl

Investigulion Purpose of the test BDse data Loading condition Urn-symmetric piP'g roul Lower freouency Input Uaterial Forced displacement

on sieato- (Test parameter) (symmetric pipire route) [refer A-t1 (reafr A-l. A-b] (refer A-l1 [refer A-t] [refer A-3]

plastic

behtlvior ot wave type artificiel seismic wane artificial seismic wave artificial seismic wove artificial seismic wove artificial seismic wave artificial seismic wane siodi nt including resonance Including resonance including resonance including resonance including resonance

COnyonents pipingfirlqeno q inludn Lwfeec condition ncy frequency frequnency freuncy I _din Low _frequency frequency frequency

direction V (fonced disalacement * Y)

Loading Ievel Medium Large Medium Medium Medium medium

P1p5 material JIS:SIS4i0 JIS:STS4I0 JIS:SYS4I0 JIS:SS1410 JIS:SUS304TP JIS:SIS4I0

(Carbon steil) (Carbon steel) (Carbon steel) (Carbon strut) (equivalent 51312 TP304) (Carbon steil)

Test No B-1 B-2 B-3 8-4 B-S

aPipe sfzt :-l/Z'Xsch40 z

(461. 3emXS. 2mm) Y x

Conceptual drnwing a Stress due to Internal

of piping syetetn tSmu s

o Stress due to dead weight

3D-Piplng :S 0.ssm

system * !aidII~a~aI *na

Base date Loading direction Loeding direction Loading condition Lower frequency Input

Invest(gatlon Purpose of the test (horizontal) (vertical) (vertical 4 horizontal) (vertical t horizontal) (vertical t horloontal)

en dynailc (Test Plraneter) (rofer A-li [refer B-li (refer B-l. B-2i [refer A-2. B-31 (refer A-t. B-3]

behavior of piping wane type artificial seismic wave artificiul seismic wave artificial seismic wane urtificial seismic wave artificial sesmice wave

spite loadiny g n inducing resonance Inducing resonance Inducing resonance inducing resonance inducing resonance

condition frI uec f1 requency frequency frequency frequency frequency

direction Y Z Y. Z Y. Z Y. Z

Loading level tedloum Medium medium Large nedium

Pipe materiaI JIS:STS410 JIS:STS410 JIS:STS110 JIS:STS410 JIS:STS410 _ _ *S (Curbon steel) (Carbon steel) (Carbon steel) (Carbon steel) (Carbon steel)

8
Table 2-4 NUPEC Large-Scale Piping System Test Load Cases

Test Case Excitation Excitation Design Stress Dominant

Wave Direction Level Frequency

DM1-_ Sweep Horizontal (Elastic)

Preliminary Sweep Vertical (Elastic)

Test Seismic Horizontal (-DM2- 1,2) Ofresonance

DM1-2Ofreoac

Design Method Seismic Vertical (=DM2-1,2)

Confirmation Allowable DM2-1 Seismic H + V 3Sm(-S2 limit) Off-resonance

Test Stress Test DM2-2 Seismic H + V 4.5Sm

Elasto-Plastic DM4-1 Seismic H+ V 6Sm

Response DM4-2(1) Seismic H + V 1.5Sm On-resonance

Test DM 4-2(2) Seismic H + V 1 3.5Sm

Preliminary US1 Sweep Horizontal (Elastic)

Ultimate Test Sweep Vertical (Elastic)

Strength Test Ultimate US2 Seismic Horizontal 24Sm On-resonance

Strength Test

9
(a) Uniaxial Static Monotonic Loading Test Specimen

(b) Uniaxial Cyclic Loading Test Specimen

0o. < Detail A

(c) Biaxial Ratcheting Fatigue Test Specimen

Dimensions shown in mm; I mm = 0.03937 in

Figure 2-1 Test Specimens from the NUPEC Material Tests

10
---

tEE

to'

5UU

450

400 -

350-

300 -

150 - Io

100

50

0.00 0.50 1.00 1.50

Strain (%)

2.00 2.50 3.00

Figure 2-2 Monotonic Loading Stress-Strain Curve for STS410 Carbon Steel

(I N/mm 2= 145 psi)

600

400

200

E

z

so

0

-200

-400

-600 _-

-3.00 -2.00 -1.00 0.00 1.00 2.00

Strain (%)

3.00

Figure 2-3 Uniaxial Cycling Test Hysteresis Curves for STS410 Carbon Steel

(I N/mm2 = 145 psi)

I1I
500

450

a-

_______ ____I_

350 -

'300 1 l

200 l

150 ___

0.00 0.50 1.00 1.50

Strain (%)

2.00 2.50 3.00

Figure 2-4 Cyclic Stress-Strain Curve for STS410 Carbon Steel

(I N/mM2 =145 psi)

Fatigue Testing Machine

Dimension shown in mm;

I mm = 0.03937 in

Figure 2-5 Static Elbow Cycling Test Setup

12
D

C 20

25-

B

Section D

SD-2A liol* SD-IA

SD-21- SD-IH

120* 6 SD3A SDOA

SD3H 85 - *-85 SD-BH

SDA 600

SD-4H ' SD-SA

; SD-5H

Section C

SC-3H- SC-6H

Sections B, F

, SB.1A

* SF-1A

, SB.SA

* SF-SA

* Strain Gage Location

SD-lA

Strain 7/ Direction

Section

Numrber

A: Axial

H: Hoop

Figure 2-6 Static Elbow Cycling Test Strain Gage Locations

18

16

14

12

.5 10

m'U

C 8

0

6

4

2

0

0 20 40 60 80 100 120 140 160 180 200

Cycle (N) /

Figure 2-7 Elbow Static Cycling Test SE-4 Hoop Strain vs. Cycle at Elbow Flank

13
*-' Acceleration

-* *Displacement ~ 

200A x S40

1117

Length dimensions shown in mm; 1

I mm= 0.03937 in

Figure 2-8 Dynamic Elbow Test Setup

14
(30kg)

Length dimensions shown in mm;

1 mm = 0.03937 in; I kg = 2.2 Ibm

Figure 2-9 Model A Simplified Piping System Test Setup

is
A2(X.Y)

A4(X) ; Vibration Table

P1 ; Internal Pressure

Al

Acceleration: 7ch lL D

Displacement: 4ch k

Temperature: 1ch

Pressure : Ich

A3 C. Y)

Figure 2-10 Model A Test Instrumentation

j 1

Scc.AI

A. tc.A

II.... .4 II, -- - 1-d , , ElbowA 1,4w. V

11 r1 e IT _ =,(

5 ec.A5

;7S cl

. _in

Sc.C Scc.D

Sec.B a 2r Sec.E

2* Sec.F

Sec.D

F o. M~IA

MaAO~ o..%

S-MA r-.SD-& SO-4A

; 30-M

T . . -- , , , w ,III

Elbow B Sec.AS

\ 'S Sec.A7 LIMV

Sec.C,E X

SC4H SC

SE41

Sec.B,F ' * Strain gange location

(j) Strain 1 floop

, " SecID \Direction Vlf4A Number

Strain gauge channel:

Elbow: 22chx 2elbows-44ch

Straight Pipe: 2chx 8sections-16ch

Figure 2-1 1 Model A Test Strain Gage Locations

16
Nozzle

A'

Variable

Small Mass

Large M~ass )Length dimensions shown in mm;

As (200kg) I mm = 0.03937 in; I kg = 2.2 Ibm

Figure 2-12 Model B Simplified Piping System Test Setup

17
A2(X.Y)

fD4)

Ti: Themlo-Couple

A5(XY): Vibration Table '

P1: Internal Pressure

A-A'

Figure 2-13 Model B Test Instrumentation

0 Strain gauge location

60' !

Sec.C

X A: Axial

I \/N H:Hoop

120- | W'.

SDF -S 4 = -5

Bs, 60 I.

Sec.D

1,x, .=Jl At .- U

Strain gauge channel:

Elbow:9ch. lelbow+3chs 2elbows=lSch

Nozzle:8ch- Isection+4chb 2sections=16ch

Straight Pipe :2ch. 16sections=32ch

A -A'

Figure 2-14 Model B Test Strain Gage Locations

18
1SO0O

10000 + ±

c

E

I

e0

5000

0

4- 4 -4-

I ,11, I

] 1 "I "r'

-b(UU - I4 _ _ _ _

-10000-~

0 50 100 ISO 200 250

1 minIsec 2 = .03937 in/sec2 Time (sec)

Figure 2-15 A-1 Test Acceleration Time History Input Motion

0

-C

0 50 100 150 200

1 mm/sec2 = .03937 in/sec2 Time (sec)

Figure 2-16 A-2 Test Acceleration Time History Input Motion

250

19
8000

6000

4000

i 20001

1 E01- 0

.0 -2000

_ I I -I _- l _

I.1111~ I

0 20 40 60 80 100

I mrnnec2 = .03937 Wnlsec2 Time (sec)

Figure 2-17 B-I Test Acceleration Time History Input Motion (Horizontal)

120

4000

-4 2000

E

E

.2

4

< .2000

0 20 40 60 80 100

1 mm/sec2 =.03937 in/sec 2 Time (sec)

Figure 2-18 B-2 Test Acceleration Time History Input Motion (Vertical)

120

20
2000

E

200

-4000

-6000

-8000

0 20 40 60 80 100 120

1 mmtsec2 = .03937 in/sec2 ime (sec)

8000

6000 - === _ -_

|Vertical|

4000-

200

0C 0 R~ TT

-2000

-4000

0 20 40 60 80 100 120

1 mm/sec2 = .03937 in/sec2 Time (sec)

Figure 2-19 B-3 Test Acceleration Time History Input Motion (Horizontal and Vertical)

21
E,

.50

.150

0 50 100 150 200 250

1 mm =.03937 in Time (sec)

Figure 2-20 A-2 Test - Pipe Center Displacement - DI (X)

30

-15

-20

15a LLL __

0 50 too ?50 200 250

1 mm = .03937 in Time (sEc)

Figure 2-21 A-2 Test - Elbow A Opening Displacement - D2(L)

22
50 100 ISO 200

Time (sec)

250

Figure 2-22 A-2 Test - Hoop Strain at Elbow A Flank - SAD-3H

4. a

UU .

2.50-

2.00

1.50

1.00

0.50 A L___._ ___- ______ _____,______ __-_I_

________-_______

0 50 100 150 200 250

Time (sec)

Figure 2-23 A-2 Test - Axial Strain at Elbow A Flank - SAD-3A

23
1100.

Jl I 80 4 

l

40

? 20

.EA

00 F3 -20 -

-40 -

-60

-80

-100

~II

Ik I

I . I

I !

0 20 40 60 80 100

1 mm - .03937 in Time (sec)

Figure 2-24 B-3 Test - Pipe Center Horizontal Displacement - DI(X)

120

8

v

E

a,

0 20 40 60 80 100

1 mm =.03937 in Time (sec)

Figure 2-25 B-3 Test - Pipe Center Vertical Displacement - DI(Z)

120

24
2.00

. _,

I.75 - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _

.50 - I _ _ __ _ _

.2 _ _ 5 _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ __ _ _ _ _ _ _ _ _ __I

02

0.75

0.50 -&-V

0.25

0.00

-0.25 -~_ _ _ _ _ _ _ _ _ _ _ _ _ _ _

0 20 40 60

Time (sec)

80 100 120

Figure 2-26 B-3 Test - Hoop Strain at Nozzle - SA I 6-2H

2.00

20 40 60 80 100

Time (tec)

120

Figure 2-27 B-3 Test - Axial Strain at Nozzle - SA I 7-2A

25
I 3-D routeE

u *[1'1',

I:

High pressure core High pressure core

spray system

I Nozzle

Han

Residual heat removal system

flooder system

I Heavy weight

(Valve) removal system

Figure 2-28 Large Scale Piping System Test Setup

26
3.0 BNL LINEAR ANALYSES AND CODE EVALUATIONS

3.1 Introduction

BNL conducted a series of linear elastic seismic analyses and ASME Code evaluations for the NUPEC

simplified piping system tests. The primary objective of these studies was to calculate and compare the

minimum factors of safety to failure in piping systems designed in accordance with different versions of

the Code. In 1994, the ASME Code rules for seismic design of piping underwent a major revision based

on a number of studies. Due to NRC staff concerns regarding the technical basis for the revisions, this

Code version was not endorsed by the NRC. As a result of efforts to resolve the staff's concerns, the

Code Committee proposed new rules in 2001. The new rules were approved with some additional

changes and were incorporated into the 2004 Edition of the ASME Code.

Since the simplified piping system tests were expected to load the piping systems to stress levels beyond

Code allowables, the test results would provide data on minimum safety margins in critical piping

components designed in accordance with different Code rules. During the pre-test analysis phase of this

program, BNL performed linear response spectrum analyses using the defined seismic wave input

motions and carried out the Code evaluations in accordance with the NRC-endorsed (1993) Code rules,

the 1994 revised rules and the 2001 proposed rules. During the test program, NUPEC determined that the

damping values for the simplified piping systems were lower than anticipated and subsequently revised

the table input motions in order to achieve the desired output. Therefore, during the post-test analysis

phase of the program, BNL repeated these calculations using the revised seismic table motions and the

different versions of the ASME Code rules. However, for these evaluations, the final 2004 rules were

used in place of the proposed 2001 rules. The post-test analysis results are presented below.

3.2 ASME Code Criteria

The latest version of the ASME Code Piping Design Criteria endorsed by NRC is the 1992 Edition

including the 1993 Addenda. The piping design rules are defined in Section III, Division 1, Subsections

NB, NC, and ND for Class 1, 2 and 3 components, respectively. Stress limits for piping subjected to

seismic and other Level D loads are given in paragraphs NB/NC/ND-3656. For Class I piping, the Code

stress equation for Level D primary stresses is as follows:

Bi PD"+ B 2 DoS Mj•<3. Os (3-1) 2t 21(3l)

where:

P = pressure

Do = outside diameter

t = wall thickness

B1, B2 = primary stress indices

I = moment of inertia

M; = resultant moment

S = allowable stress intensity

The resultant moment term in the above equation includes deadweight, seismic and other Level D

moments. The revised 1994 Addenda to the Code provided an alternate stress limit which was 50%

higher than the old limit which could be applied when certain additional conditions were met. The new

equation was:

27
B PDO +B 2 Do ME < 45S (3-2) 2t 21

The resultant moment term, ME, in the above equation includes deadweight, seismic inertia and other

"reversing" Level D moments. Additional conditions for application of this equation included the

following:

(1) The analysis must be based on the linear elastic response spectrum method with 5% damping and

15% peak broadening.

(2) The ratio of outside pipe diameter to wall thickness shall not exceed 50.

(3) Stress due to weight moments is limited to 0.5 Sm.

(4) Stress due to seismic anchor motion moments is limited to 6 Sm.

(5) Stress due to seismic anchor motion longitudinal forces is limited to Sm.

(6) Maximum pressure during the SSE must not exceed the design pressure.

In 2001, the following changes were proposed to the alternate Level D primary stress limit equation:

BL P2 °+B2 DO ME <3.OSm (3-3) 2t 21

where:

B2' = B2 for all piping components except:

B 2 ' = 3/4 B2 for elbows, bends, and tees

B2'= 1.33 for girth butt welds between items which do not have nominally identical

thicknesses.

When this equation is used, the same additional conditions must be met, except that the ratio of outside

pipe diameter to wall thickness shall not exceed 40. This equation reinstated the original 3.0 Sm stress

limit but modified the stress indices for critical components. Compared to the 1993 Code, for straight

pipes, the equation is identical to equation (3-1). For elbows, bends and tees, the change is approximately

equivalent to using a stress limit of 4.0 Sm in equation (3-1). For girth butt welds at equipment nozzles,

the change is approximately equal to a stress limit of 2.25 Sm in equation (3-1).

Equation (3-3) and the new B2' stress indices for seismic loads listed above were acceptable to the NRC

staff and would have been formally endorsed. However, ASME comments submitted March 22, 2004, on

the proposed 10 CFR 50.55a rulemaking (Federal Register Volume 69, Number 4, January 7, 2004)

indicated that these B2' stress indices were a typographical error that will be corrected by an errata.

According to the ASME position, the criteria should be B2' = 2/3 B2 for elbows, bends, and tees.

Subsequently, the following alternate Code stress equation for seismic and other "reversing" stresses was

published in the 2004 Edition of the ASME Code as follows:

B, 2t ° B2'10 ME 3OS,,m (3-4)

where:

28
B2'= B2 for all piping components except:

B2' = 2/3 B2 for elbows, bends, and tees

B2' = 1.33 for girth butt welds between items which do not have nominally identical

thicknesses.

As in Equations (3-2) and (3-3) above, the resultant moment term, ME, includes deadweight, seismic

inertia and other "reversing" Level D moments. Additional conditions for application of this equation are

the same as those for Equation (3-3):

(1) The analysis must be based on the linear elastic response spectrum method with 5% damping and

15% peak broadening.

(2) The ratio of outside pipe diameter to wall thickness shall not exceed 40.

(3) Stress due to weight moments is limited to 0.5 Sm.

(4) Stress due to seismic anchor motion moments is limited to 6 Sm.

(5) Stress due to seismic anchor motion longitudinal forces is limited to Sm.

(6) Maximum pressure during the SSE must not exceed the design pressure.

Compared to the 1993 Code, for straight pipes, this equation is identical to equation (3-1). For elbows,

bends and tees, the change is approximately equivalent to using a stress limit of 4.5 Sm in equation (3-1).

For girth butt welds at equipment nozzles, the change is approximately equivalent to a stress limit of 2.25

Sm in equation (3-1). Although this alternate stress criterion has been published in the 2004 Edition of the

ASME Code, to date, it has not been endorsed by the NRC.

3.3 Linear Analysis and Evaluation of Simplified Piping Systems

The two-dimensional (Model A) and three-dimensional (Model B) simplified piping systems were

analyzed and evaluated using ASME Code seismic criteria by applying linear analysis methods and

assumptions in accordance with nuclear industry practice. For consistency with the nonlinear simulation

analyses performed under this program, the ANSYS general purpose finite element analysis code [3] was

used. Both the Model A and Model B piping specimens were modeled using ANSYS three-dimensional

linear elastic straight pipe (PIPEI 6) and curved pipe (PIPE18) elements as shown in Figures 3-1 (a) and 3-

I (b). MASS21 elements were used to represent the added masses attached to the pipes. For the pre-test

analyses, it was assumed that the added masses would also provide additional stiffness to the pipes.

Therefore, over the length of the added masses, the pipe's cross-sectional properties were modified to

reflect the rigid stiffness of the added masses. However, frequency test results for the actual piping

specimens demonstrated that this assumption resulted in excessively stiff piping models with higher

frequencies than were observed in the tests. Therefore, for the post-test analyses, the stiffnesses at the

locations of the added masses were not included. This modeling assumption is consistent with the final

nonlinear models as discussed in Section 5.3 of this report. In Model B, the nozzle was represented by a

series of straight pipe elements with cross-sectional properties of the nozzle. The spring hangers in

Model B were represented by linear spring elements (COMBIN14). Model A boundary conditions

included vertical translational restraint at the large mass rollers and restraint of all translations and

horizontal axis rotations at the pinned supports. In Model B, the nozzle end nodes and spring hanger

ground attachment point nodes were fully restrained, and pinned support point nodes were restrained from

all translations and from rotations about the horizontal axes.

Linear response spectrum analyses were performed in accordance with industry practice and regulatory

guidelines [4] for piping design with only two exceptions: (1) since actual piping dimensions had been

provided by NUPEC, as-built rather than design dimensions were used in the models; and (2) since the

input motion was defined and the objective was to identify minimum margins to failure, unbroadened

29
instead of broadened response spectra were applied in the analyses. The response spectra for twodimensional simplified piping tests A-1 and A-2 are shown in Figures 3-2 and 3-3. The response spectra

for the three-dimensional piping tests B-i, B-2 and B-3 are shown in Figures 3-4 through 3-7. The modal

damping values used in the analyses were consistent with the practices and requirements of the respective

Code version. The 1993 Code version did not specify a damping value for analysis. For this case, two

different analyses were performed with damping values based on industry practice. In the first analysis,

two percent damping was used based on NRC Regulatory Guide 1.61 for pipe diameters less than 12

inches (300 mm). In the second analysis, frequency dependent damping values based on ASME Code

Case N-41 I damping were used. These values are 5 percent for modes with frequencies less than 10 Hz

and 2 percent for modes with frequencies greater than 20 Hz. For frequencies between 10 Hz and 20 Hz,

the damping values are linearly interpolated between 2 and 5 percent. For the 1994 and 2004 Code

versions, both analyses used the Code-specified value of five percent damping.

Linear analyses were carried out for pressure, deadweight and seismic loading. For each test case, the

results were post-processed to calculate stresses in accordance with the three different code versions. The

highest stresses in the critical components (elbows for Model A; elbows, tee, and nozzle for Model B)

were determined in accordance with the different equations, stress indices and stress allowables of each

code version. The component with the largest ratio of stress to allowable stress was identified as the

critical component. The results for each load case are summarized in Tables 3-1 though 3-5.

In order to properly interpret the results of these analyses, it should be noted that the "margins" presented

in Tables 3-1 through 3-5 represent the ratios of the highest stress in a piping system (calculated in

accordance with the specific code version rules which include the stress equation, stress indices, and

damping) divided by the allowable stress (defined by the specific code version). Since none of the

simplified piping systems failed during a single seismic load application, they should not be interpreted as

margins to failure. Instead they should be considered minimum margins to failure or minimum Code

stress margins. For example, Table 3-1 shows that the piping system in test A-I was able to sustain a

maximum stress equal to 2.99 times the Code allowable without failure according to the rules of 1993

ASME Code and R.G .1.61 damping. The same test shows that the piping system was able to sustain a

maximum stress of 1.20 times the Code allowable according to the 2004 ASME Code rules. Had the test

been conducted to higher stress levels resulting in failure the calculated margins to failure would have

been proportionally higher.

In addition to providing information on minimum margins to failure, the results provide a measure of the

relative safety margins between the different code versions. The margins in Table 3-1 indicate that the

1994 and 2004 Code rules would provide essentially equal margins to failure while the 1993 rules with N411 damping would provide 50% greater margin. The 1993 rules with R.G. 1.61 damping would provide

about 150% greater margin. The margin comparisons can be used to evaluate the relative conservatisms

between the different Code versions.

For both the A- I and A-2 tests, the highest stresses were at Elbow A. Since the level of input excitation

was greater in the A-2 test, the margins were also greater for all code versions. The 1994 and 2004 Code

calculations provided nearly identical margins of 1.83 and 1.84, respectively. The 1993 Code calculations

provided margins of 2.74 with N-411 damping and 3.56 with R.G. 1.61 damping.

For the Model B tests, the stresses and margins were generally lower. In the B-I and B-3 tests, the

highest stresses were either at the tee or at the nozzle weld. Stresses at these locations were very close in

value in nearly all cases except for the 2004 code stress calculation which has lower stress indices for tees

and higher indices for nozzle welds. In the B-2 vertical motion test, the stresses were always highest at

the nozzle weld but the magnitudes were lower and the margins were less than one in all cases. The

highest margins were in the B-3 (horizontal + vertical) tests. The 2004 Code calculations provided a

30
margin of 1.76 at the nozzle weld. The other code versions also provided highest stresses at the nozzle

welds. The margin based on the 1994 Code was less than one (0.91). The margins based on the 1993

version were 1.37 with N-411 damping and 1.85 with R.G. 1.61 damping.

In summary, the linear analyses described above demonstrated that the NUPEC simplified piping system

high level seismic tests subjected the piping components to stresses well above Code allowables. By

performing analyses and evaluations in accordance with different versions of the Code rules, minimum

margins to failure were determined, and the relative differences in safety margins between the different

versions were compared for the tested components. Based on the 2004 Code rules, these evaluations

demonstrated minimum margins to failure of 1.84 for elbows and 1.76 for nozzle welds. The most

conservative margins were provided by the 1993 Code version with R.G. 1.61 damping. Those

evaluations demonstrated minimum margins to failure of 3.56 for elbows and 1.85 for nozzle welds. It

should be clearly noted, however, that these margins are simply minimum safety margins based on the

limited simplified piping system test data. Since no pipe failures were observed during a single test, the

actual margins to failure are higher. As discussed in Section 2.4 of this report, NUPEC conducted tests to

failure for a Model A and a Model B piping system. For the Model A test, a crack at the Elbow A

location occurred after five high level seismic tests. For Model B, a nozzle crack was observed after 18

applications of a high level seismic time history. The failures in both cases were attributed to fatigue

cracking.

31
Table 3-1 A-i Test Maximum Code Stresses and Margins

I MPa = 145 psi

Table 3-2 A-2 Test Maximum Code Stresses and Margins

Code Edition Damping Maximum Stress Critical Location Stress/Allowable Coe(MPa) Cnia oain(Margin)

1993 R.G. 1.61 1461 Elbow A 3.56

1993 N-411 1126 Elbow A 2.74

1994 5% 1126 Elbow A 1.83

2004 5% 755 Elbow A 1.84

I MPa = 145 psi

32
Table 3-3 B-I Test Maximum Code Stresses and Margins

Code Edition Damping Maximum Stress (MPa) Critical Stress/Allowable

Code _______ Elbow Tee Nozzle Location (Margin)

1993 R.G. 1.61 564 733 728 Tee 1.78

1993 N-411 405 536 539 Nozzle Weld 1.31

1994 5% 405 536 539 Nozzle Weld 0.87

2004 5% 274 424 692 Nozzle Weld 1.68

I MPa = 145 psi

Table 3-4 B-2 Test Maximum Code Stresses and Margins

Code Edition Damping Maximum Stress (MPa) Critical Stress/Allowable

Elbow Tee Nozzle Location (Margin)

1993 R.G. 1.61 117 124 262 Nozzle Weld 0.64

1993 N-411 98 109 226 Nozzle Weld 0.55

1994 5% 98 109 226 Nozzle Weld 0.37

2004 5% 70 93 276 Nozzle Weld 0.67

1 MPa = 145 psi

Table 3-5 B-3 Test Maximum Code Stresses and Margins

Maximum Stress (MPa) Critical Stress/Allowable

Cod eEdition Damping Elbow Tee Nozzle Location (Margin)

1993 R.G. 1.61 577 747 759 Nozzle Weld 1.85

1993 N-411 413 544 564 Nozzle Weld 1.37

1994 5% 413 544 564 Nozzle Weld 0.91

2004 5% 280 430 725 Nozzle Weld 1.76

I MPa = 145 psi

33
Elbow C

Nozzle Elbow A

Elbow A

Elbow B

Elbow B

(a) Model A (b) Model B

Figure 3-1 Simplified Piping System Test- Models A and B

Linear Elastic ANSYS Pipe Element Models

34
4.5

F

Ca

0

F

I2 2.5 _____a_

0.5 J - -=

0 _ l l l I 

0 2 4 6 8 10

Frequency (Hz)

12 14 16 18 20

Figure 3-2 A-i Test Horizontal Response Spectrum (2% and 5% Damping)

.r

a

0f

I

2 4 6 8 10 t2 14 16 18

Frequency (Hz)

Figure 3-3 A-2 Test Horizontal Response Spectrum (2% and 5% Damping)

35
T

02

0

7 A __ _ I 

6 _ _ _ _ _ _ _ _ _ _ _ _ _ _ _

T 11 - apigA__

3- 1 apn

0I

0 2 4 6 8 10

Frequency (Hz)

12 14 16 18 20

Figure 3-4 B-i Test Horizontal Response Spectrum (2% and 5% Damping)

2.5

T

0

1.5

0.5

0 2 4 6 8 10

Frequency (Hz)

12 14 16 18 20

Figure 3-5 B-2 Test Vertical Response Spectrum (2% and 5% Damping)

36
T

20

aI

00

.-C

7 I

4 A__

3 _________

3

I0

0 2

Figure 3-6

4 6 8 10 12 14 16 18

Frequency (Ha)

B-3 Test Horizontal Response Spectrum (2% and 5% Damping)

20

22 .

00

j .

0 2 4 6 8 10 12 14 16 18

Frequency (Hz)

Figure 3-7 B-3 Test Vertical Response Spectrum (2% and 5% Damping)

37

20
4.0 BNL COMPONENT TEST NONLINEAR ANALYSES

4.1 Introduction

A major goal of the NRC/BNL collaboration effort was the performance of pre-test and post-test analyses

for selected piping component and piping system tests. The objective of these analyses was to investigate

and evaluate the adequacy of current commercially available methods and computer programs for

predicting the elasto-plastic response of piping systems subjected to large earthquake loads. Upon

initiation of the collaborative effort, NUPEC indicated that the ABAQUS computer program would be

used by NUPEC to perform their analyses. In order to complement the NUPEC analytical effort, BNL

selected the ANSYS computer program [3] for performing these analyses. ANSYS is a widely used,

general-purpose finite element analysis program. In recent years, the program was enhanced to include a

number of improvements in its nonlinear analysis capability. Among these improvements was the

inclusion of the Chaboche nonlinear kinematic hardening model [5]. This plasticity model was expected

to provide better simulation of the strain ratcheting behavior that a pressurized pipe is expected to exhibit

under seismic loading than is possible with classical linear hardening models.

The major focus of the BNL simulation analyses was on the prediction of selected simplified piping

system test responses. These included the two-dimensional Model A test load cases A-I and A-2, and the

three-dimensional Model B test load cases B-1, B-2, and B-3. However, in order to develop, calibrate and

test elasto-plastic elbow components for the simplified piping system nonlinear models, analyses were

performed to simulate the elbow component static cycling test SEA. Results of the NUPEC material tests

were used to develop the plastic material properties needed in the component and piping analyses. The

following sections describe the modeling and analysis approach, and the analysis results for the elbow

component test.

4.2 Development of Piping Elbow Model

The overall flexibility and dynamic properties of a piping system are typically dependent on the flexibility

of the elbows in the system. Furthermore, under large seismic deformations, the largest plastic strains are

expected to be concentrated in the elbows. Therefore, the development of accurate detailed finite

element models of the piping system elbows was a critical step in the simulation analysis. A shell

element model of the elbow used in static component test SEA was first developed and compared with

the test data. The test specimen was a 65 mm (2 1/2 inch) Schedule 40 carbon steel long radius elbow. At

each end, the elbow was welded to a short section of straight pipe which terminated at a welded flange.

Details of the specimen configuration are shown in Figure 4-1 and the test setup is shown in Figure 2-5.

As-built measurements of the specimen outer diameter and wall thickness at various cross-sections were

provided by NUPEC. The pipe was filled with water and pressurized to 19.8 MPa (2870 psi). The test

was conducted at room temperature. Cyclic displacement of -15.5 mm (LO.6 in) was applied at the ends

of the specimen at a frequency of one cycle/min.

The ANSYS finite element model of the SEA test elbow is shown in Figure 4-2. The model utilized

ANSYS SHELLI81 large strain shell elements to represent the pipe wall. This element is appropriate for

analyzing thin to moderately-thick shell structures. It is a four-noded element with six degrees of

freedom at each node. This type of element is well suited for linear, large rotation, and large strain

nonlinear applications. The element can be used with all ANSYS kinematic hardening plasticity models

including Chaboche nonlinear kinematic hardening. The triangular option of this element was used to

represent the flanges at the ends of the pipe. ANSYS elastic pipe elements (PIPEI6) with rigid stiffness

properties were used to represent the end links that connect the pipe to the test fixture. Since the test

specimen was symmetric about two perpendicular planes, a quarter model with appropriate symmetry

39
boundary conditions was used. The pipe diameter and thickness of the shell element model was based on

an average of the as-built values provided by NUPEC.

The elastic and plastic material properties for the elbow models were based on NUPEC material test data

for carbon steel (STS410) specimens. The elastic modulus used for all models was 203,000 MPa (29.4 x

106 psi) which was based on a monotonic tension test (MTT-1-1). The stress-strain curve for this test is

shown in Figure 2-2. The development of plasticity parameters was dependent on the plasticity model.

Analyses were performed using three different plastic hardening models. The development of the

plasticity parameters for the pipe is described in the next section.

4.3 Development of Plasticity Models

Plasticity theory provides a mathematical relationship that characterizes the elasto-plastic response of

materials. The theory considers three basic components: the yield criterion, the flow rule, and the

hardening rule. The yield criterion determines the stress level at which yielding is initiated. The flow

rule determines the direction of plastic straining. The hardening rule describes the changes in the yield

surface with progressive yielding, so that the stress states for subsequent yielding can be established. The

ANSYS program provides several options for characterizing elasto-plastic material behavior. In

performing the simulation analyses under this collaboration program, three different plasticity options

were investigated: classical bilinear kinematic hardening, multilinear kinematic hardening, and nonlinear

kinematic hardening. These options are based on the von Mises yield criterion, the associated flow rule,

and different variations of kinematic hardening models as described below.

The kinematic hardening rule assumes that the difference between yield stresses under tension and

compression loadings remains constant. Under plastic deformation, the elastic domain will retain a

constant size but will move about in the stress space by translation. This translation can be characterized

in terms of the location of the center of its elastic domain or yield surface, which is also referred to as a

backstress. The ANSYS classical bilinear kinematic hardening option is based on the linear kinematic

hardening rule introduced by Prager [6]. This theory assumes that the center of the yield surface moves

linearly with plastic strain. For the simple case of uniaxial loading, this rule is expressed by the following

incremental equation:

da = Codep (4-1)

where a is the center of the yield surface (backstress), ep is the plastic axial strain, and CO is a material

constant equal to the plastic modulus. The total stress during plastic flow is expressed as follows:

a = a ± CO (4-2)

where ais the total axial stress and o0 is the yield stress, and the sign is dependent on the direction of

plastic flow. In applying the bilinear kinematic hardening (BKIN) option of ANSYS, the material stressstrain curve is represented as a bilinear curve in which the first slope represents the elastic modulus up to

the yield stress, and the second slope represents a linear plastic modulus. When a material is loaded into

the plastic range and subsequently unloaded, the reverse path follows the elastic slope back to a range of

twice the yield stress and then follows the plastic slope.

The multilinear kinematic hardening (MKIN) option in ANSYS is based on the Besseling multilayer

model [7]. When this rule is applied, the stress-strain curve is represented by several linear segments.

This allows for a more accurate representation of a stress-strain curve. Under reversed loading, the load

path again follows the elastic slope back to twice the yield stress and then follows the path of the multiple

40
plastic slopes. However, using either rule, the plastic modulus is always the same for loading and

unloading and is unaffected by the presence of a mean stress. As a result, for a prescribed uniaxial stress

cycle with a mean stress, the loading and reversed loading hysteresis curves will always produce a closed

loop with no ratcheting.

The nonlinear kinematic hardening option (CHAB) in ANSYS is based on the work of Chaboche [5] who

proposed the superposition of several "decomposed" Armstrong-Frederick hardening rules. Armstrong

and Frederick [8] had introduced a nonlinear kinematic hardening rule with a "recall" term which

introduced a fading memory effect of the strain path. For the case of uniaxial tension-compression

loading, the translation of the yield surface is expressed by the following incremental equation:

da = Cdcp -yadep I(4-3)

where C and y are material dependent constants. In applying this rule, the plastic modulus, H, is

expressed by the following equation:

H=C:Fra (4-4)

In this equation, the negative sign is used for the forward loading curve and the positive sign is used for

the reverse loading curve. This rule provides an exponential plastic modulus. For a uniaxial stress cycle

with mean stress, this rule produces changes in shape between forward and reverse loading paths.

Therefore the loop does not close and ratcheting occurs. A disadvantage of this model, however, is that

an experimental stress-strain curve is not necessarily exponential in nature and an attempt to simulate it

by a single exponential equation does not yield a good fit. In addition, for a high strain range, the model

does not produce a constant plastic modulus as exhibited in experiments, but will instead always stabilize

to a zero plastic modulus.

In order to improve the simulation of the hysteresis loop, Chaboche proposed the superposition of three

hardening rules. The first rule (a,) should simulate the initial high plastic modulus at the onset of

yielding; the second rule (a2) should simulate the transient nonlinear portion of the curve; and the third

rule (a3) should be a linear hardening rule (y3 = 0) to represent the subsequent linear part of the hysteresis

curve at a high strain range. As is the case for the single Armstrong-Frederick rule, for a uniaxial stress

cycle with mean stress, the rule will also produce a change in shape between the forward and reverse

loading paths which results in ratcheting. However, because of the incorporation of the linear kinematic

hardening rule (a3), this model will eventually approach the complete shakedown of ratcheting which is

not in agreement with experimental results. The ratcheting simulation can be improved by assigning a

small value to y3, which introduces a slight nonlinearity which will improve the ratcheting simulation

without producing a noticeable change in the hysteresis loop. The procedure for defining the parameters

is described in the next section.

4.4 Determination of Material Parameters

The NUPEC material tests provided the basic material parameters needed for each of the three plasticity

options investigated. The methodology used to define the parameters is described below.

4.4.1 Bilinear Kinematic Hardening Model

The stress-strain curve from a monotonic uniaxial tensile test of an STS41O carbon steel test specimen

was used to define the parameters for the bilinear kinematic hardening (BKIN) model. The parameters

included modulus of elasticity, yield stress, and plastic modulus. Consistent with the theory, the plastic

41
modulus was assumed to be constant and was approximated from the slope of the stress-strain curve from

the yield strain to 3 percent strain as shown in Figure 4-3. The values used for this model were 203000

MPa (29.4 x 106 psi) for the elastic modulus, 1220 MPa (177000 psi) for the plastic modulus, and 270

MPa (39000 psi) for the yield stress.

4.4.2 Multilinear Kinematic Hardening Model

The cyclic stress-strain curve developed from a uniaxial strain cycling test of an STS410 carbon steel test

specimen was used to define the parameters for the multilinear kinematic hardening (MKIN) model. The

stress-strain curve was approximated by nine linear segments from zero to 2.5 percent strain as shown in

Figure 4-3.

4.4.3 Chaboche Nonlinear Hardening Model

The Chaboche (CHAB) model required the definition of elastic modulus, yield stress, and the C1 and Yi

constants for each of the three Armstrong-Frederick hardening rules. A procedure for determination of

the C and y parameters was described by Bari and Hassan [9]. All parameters, except for y3, were

determined from the stress-strain hysteresis curves generated from a strain-controlled uniaxial strain

cycling test of an STS410 carbon steel test specimen shown in Figure 2-3. The procedure specifies that

the decomposed backstress terms (a, or a2) should start from -C7/yi at the starting plastic strain -so, and

reach a value of C/y, at the final plastic strain cp,. The third linear backstress term should go through the

origin. The equations used for the loading part of the hysteresis curve are as follows:

3

a = 0oe + Eaz (4-5)

; = Ci [I - 2 expl- Yi (up - (- c, for i = 1 and 2 (4-6)

y i

a 3 = C3 cp (4-7)

where £pI is the strain limit of the stable hysteresis loop. The procedure first required the selection of a

hysteresis loading curve with a reasonably large strain range. The yield stress and yield strain were then

estimated from the curve and the curve was adjusted for elastic strain to produce a stress versus plastic

strain plot. Using the above equations, the C, and yi parameters were selected to match this test curve. A

spreadsheet program was used to compare the loading part of the hysteresis test curve to the a curve

from Eq. (4-5). The C3 parameter was first selected to match the slope of the linear segment of the

hysteresis curve at a high strain range. The C, parameter was next selected as a large value such that the

slope of a, nearly matched the plastic modulus at initial yielding, and the corresponding Ty value was also

a large value to stabilize the hardening of a, immediately. The C2 and y2 parameters were selected by trial

and error to match the overall shape of the curve. Additional adjustments to C, and y, were then made to

match the initial slope of the test curve as well as the stress at both the start and end points. Figure 4-4

illustrates the contribution of each rule to the simulation of the stress-strain curve, and compares the final

analytical stress-stress curve to the test curve. As suggested in [9], the final parameter, y3, may be

determined from a uniaxial ratcheting test to provide a best fit. However, for this analysis effort, this

parameter was determined from the results of an elbow component test strain cycling test as described

below. The parameters determined for the simulation analyses using the Chaboche nonlinear hardening

model were as follows: ao= 270 MPa (39000 psi), E 203000 MPa (29.4 x 106 psi), C, = 70000 MPa

(10.2 x 106 psi), C2 = 30000 MPa (4.35 x 106 psi), C 3 = 1200 MPa (174000 psi), y7 = 3500, y2 = 210, y3

=1.

42
4.5 Elbow Static Cycling Test Nonlinear Analysis

As a preliminary test of the ANSYS shell element elbow model, a nonlinear static cycling analysis was

performed to simulate the elbow component test SE-4. The quarter elbow model shown in Figure 4-2

was subjected to a constant internal pressure of 19.8 MPa (2870 psi) and equivalent cyclic end

displacements of 1 5.5 mm (±0.6 in) were imposed statically for 185 cycles. A separate analysis was

performed for each of the three plastic hardening rules as described below. Additional analyses

investigated the significance of geometric deformation effects by applying the large deflection analysis

option in ANSYS and comparing results against those obtained using the small deflection analysis option.

These studies demonstrated that the changes in geometric configuration due to elbow ovalization had a

very small effect on elbow load deflection characteristics and on strain ratcheting. Therefore these effects

were judged to be insignificant within the range of this application, and the final analyses presented below

were based on small deflection theory.

For in-plane loading of elbows, the maximum strain is expected to occur at the elbow flank. The outer

surface of the elbow flank was instrumented with strain gages which were monitored during the test. In

the SEA test, the elbow was cycled until crack penetration occurred at the flank at 185 cycles. The hoop

strain versus cycle history measured during the test is shown in Figure 4-5(a). As expected, significant

strain ratcheting was observed during the test. As shown in the Figure 4-5(a), the rate of strain ratcheting

was highest during initial cycling and then leveled off to a nearly constant rate until failure. Shakedown

did not occur. At failure, the accumulated hoop strain was nearly 15 percent.

The corresponding strain histories predicted by the ANSYS nonlinear analyses using the three plastic

hardening models are shown in Figures 4-5(b), 4-5(c), and 4-5(d). Using the bilinear kinematic hardening

model (BKIN), initial strain ratcheting was overpredicted during the first ten cycles. The rate of strain

then leveled off to zero (shakedown) with no additional strain accumulation. At 185 cycles, the

accumulated strain was about 7 percent. The multilinear kinematic hardening model (MKIN) analysis

predicted shakedown after only five cycles with a total strain accumulation of only 2 percent at 185

cycles.

As described in Section 4.4.3 above, the parameters needed for the Chaboche nonlinear kinematic

hardening model (CHAB) were determined from a strain-controlled uniaxial strain cycling test of an

STS410 carbon steel test specimen with the exception of the y3 parameter. This parameter was chosen by

performing parametric studies which varied the ys parameter and selecting the value that provided the best

match to the strain ratcheting results. The first analysis used a value of y3 equal to zero. The hoop strain

versus cycle history is shown in Figure 4-5(d). This curve provided a better simulation of the initial strain

ratcheting behavior but also gradually leveled off and achieved shakedown with an accumulated strain of

8 percent. Results of analyses using yv values of I and 2 are also shown in Figure 4-5(d). These strain

histories more accurately simulated the test results by leveling off to a nearly constant rate of ratcheting.

The analysis with a yj value of I provided the best match to test results and was used in subsequent

analytical models.

43
Dimensions shown in mm; 1 mm = 0.03937 in

Figure 4-1 SEA Component Test Elbow Specimen

Figure 4-2 SEA Test ANSYS Shell Element Model

44
4500 Soo _-AL400 _2 _ _ 7 7 _ _ _ _ _ _ _ _

E2 300

-250 a,

0

- 200

150

100

50

__ - __ __ __ __ _ __ __ _ __ __ __ _ M notonr c Test -

BKIN Model

-_ *. Cydic Test

lMKIN Model I

o a0.00 0.50

1 N/mm2 = 145 psi

1.00 1.50 2.00 2.50

Strain (%)

3.00

Figure 4-3 BKIN and MKIN Model Stress-Strain Curves

500

400

300

"E 200

z

E 100

0

-100

Test - I

I I_

/ O3 I

-200 I I

-3.0% -2.0% -1.0% 0.0%

Plastic Strain (%)

1.0% 2.0% 3.0%

1 N/mm 2= 145 psi

Figure 4-4 CHAB Model Stress-Strain Curve

45
16E

1 4

- 1 1

n

U~c

00

I

-ilI4

4 _________

2 -- = -- L ==

a

0 20 40 60 80 100 120 140 160 180 200

Cycle (N)

0 20 40 60 80 100 120 140 160 180 200

Cycle (N)

F. 0UC

0

M

18~

16 - -

14 (c)MN Model j

1 2 l2- III 1 -= -4-f -

10 I __.I

aj8 ___ __-.. _.. _ ._- -

8 -- .I- --- - -

4 -

2 WAM UIIMu

o 1111I1117IlHmrmNmmDUU INi~l INiiimlll IlD1111111 l

C

00Ct

0 20 40 60 80 100 120 140 160 180 200

Cycle (N)

0 20 40 60 80 100 120 140 160 180 200

Cycle (N)

Figure 4-5 Elbow Static Cycling Test SE-4 - Hoop Strain vs. Number of Cycles at Elbow Flank

46
5.0 BNL SIMPLIFIED PIPING SYSTEM TEST NONLINEAR ANALYSES

5.1 Introduction

Initial analytical models of the piping systems and components utilized the ANSYS straight and curved

plastic pipe line elements. Preliminary analyses, however, demonstrated that the curved plastic pipe

elements were not adequate for simulating the elbow strain ratcheting behavior observed in the tests. As a

result, the final models incorporated detailed shell element representations of elbows and other piping

components expected to experience plastic deformation. The elbow models were based on the shell

element elbow models developed for the SE4 component test. The following sections describe the

modeling and analysis approach, analysis results, and comparisons to test results.

5.2 Development and Pre-Test Analysis of Two-Dimensional Simplified Piping System Model

The ANSYS model of the two-dimensional Model A Simplified Piping System test specimen shown in

Figure 2-9 consisted of a combination of elements. Because of symmetry about the horizontal plane, only

the top half of the system was modeled and appropriate symmetry boundary conditions were defined at

the horizontal mid-plane. The ANSYS model is shown in Figure 5-1. The elbows were modeled in the

same manner as the SEA4 test elbow model described in Section 4.2 above using ANSYS SHELL181

large strain shell elements to represent the pipe wall. These elements were used to represent the elbows

plus a 100 mm (4 in) length of straight pipe at each end of the elbows. Since it was anticipated that the

plasticity would be localized to the elbows, the remaining straight pipes were represented by twodimensional linear elastic beam elements (BEAM3) with dimensional properties equivalent to the crosssectional properties of half the pipe. The transition from the straight pipe shell model to the straight pipe

beam model was achieved through a series of rigid beam elements arranged in a wheel spoke

configuration as shown in the elbow detail in Figure 5-1. The shell element representation of the

additional length of straight pipe at the ends of the elbow was needed to ensure that the constraints at the

transition from shell to beam elements did not inhibit elbow ovalization required for appropriate

characterization of elbow flexibility. The selection of the 100 mm (4 in) length considered the guidance

provided in the ASME Code, Section 111, Subsection NB-3686.2 which states that the code elbow

flexibility equations are applicable only if there are no flanges or stiffeners from the end of the elbow to

within a distance equal to the mean radius of the pipe. For this case, a 100 mm (4 in) length is

approximately three times the mean radius. Additional static load deflection analyses comparing a full

shell element model to this combination shell element/beam element model verified the adequacy of the

simplified shell/beam element model. ANSYS MASS2 1 elements with half the actual masses represented

the added masses attached to the piping system. Symmetry boundary conditions were specified to

constrain vertical translation and horizontal axis rotations for all nodes in the horizontal (x-y) plane.

Horizontal translational restraints were defined at the two pinned support points.

I

During the pre-test analysis phase of the program, seismic analyses were performed to predict the

responses for the A-l and A-2 tests. A primary objective of these analyses was to investigate and

compare responses using the three different plasticity models described in Section 4.3 and to select the

best model for application in the post-test analysis. The analyses applied the ANSYS nonlinear full

transient analysis option which utilizes the Newton-Raphson iteration procedure and the Newmark time

integration method to solve the nonlinear equations of motion. An internal pressure of 19.8 MPa (2870

psi) was applied to the system. System damping was defined in terms of Rayleigh damping coefficients

based on the expected modal damping value of 3 percent provided by NUPEC. The expected input

motions were defined as x-direction digitized acceleration time histories of the shaking table divided into

4000 equal time increments. Duration times were 191.6 seconds for the A-1 test and 215.6 seconds for

the A-2 test. Since the ANSYS program does not accept base acceleration input motion, the input was

applied as an equivalent force time history to a fictitious large mass of 10" kg (2.2 x 10" Ibm) connected

47
to the support points. The ANSYS automatic time stepping option was applied with an integration time

step upper limit of one quarter of the input motion load step (.012 sec. for A-i and .0135 sec. for A-2).

Both the A-i and A-2 analyses were performed for each of the three plastic hardening models. Key

responses including displacements, rotations, stresses and strains at critical locations were saved at every

load step for comparison with test results. Differences in both displacement and strain ratcheting results

were observed between the different plasticity models.

During preliminary testing of the two-dimensional Model A simplified piping system, NUPEC

determined that the system modal damping was lower than expected (1.8 percent versus 3 percent). As a

result of this finding, NUPEC revised the final table acceleration input motions to achieve their target

strain levels. Because of this change, a comparison between the BNL pre-test analysis results and test

results was not meaningful. NUPEC made adjustments to both the time and acceleration level scales of

the input motions and provided the new inputs, as well as the test results, to BNL for post-test analysis.

Using the same methodology and the analytical model with Chaboche nonlinear kinematic hardening,

BNL repeated the A-2 test analysis with the new seismic input motion and lower damping values. A

comparison of these post-test analysis results to test results showed that the analysis significantly

underpredicted displacements and strains. Peak displacement at the center of the pipe predicted by

analysis was 50 mm (2 in) versus 178 mm (7 in) recorded in the test. The analysis predicted no strain

ratcheting compared to 3.4 percent accumulated strain observed in the test. A series of additional

analytical studies to investigate and identify the source of the underprediction and improve the accuracy

of the simulation were initiated.

5.3 Additional A-2 Simplified Piping System Test Analytical Investigations

A series of additional A-2 test simulation analyses were carried out to examine the sensitivity of results to

changes in various modeling and analysis parameters. The initial series of analyses investigated changes

in analysis parameters. The following changes were considered:

* Reduction in integration time step from .015 sec. to .003 sec.

* Reduction in output time step from .09 sec. to .03 sec.

The analyses with the above reduced time steps showed no significant differences in results.

The load application method was investigated next. The original analysis had applied the base

acceleration input time history as an equivalent force time history to a large fictitious base mass. Two

alternate methods of load application were investigated:

* A displacement time history was calculated by performing a double integration of the

acceleration input time history. This was applied directly to the model as a displacement input

time history at the support points.

* The acceleration time history was converted to a set of equivalent inertia force time histories and

applied directly to the concentrated masses in the piping system model.

Both the displacement time history and the equivalent inertia force time history analyses showed no

significant differences in results. It was noted, however, that the inertia force time history analysis was

more efficient and this method was applied in subsequent analyses.

Variations in material parameters were investigated next as follows:

* A 10% reduction in modulus of elasticity to better match the Model A test frequency

48
* A 20% and a 40% reduction in yield stress to induce more plasticity

* A modal damping ratio of 0% (vs. 1.8%) to induce greater amplified response

* Application of bilinear kinematic hardening model vs. Chaboche model

The results of these analyses showed small differences in results. The case with no damping provided

slightly higher displacement but no strain ratcheting. The case with 40% reduction in yield stress

demonstrated a small amount of strain ratcheting but no increase in displacement. The other cases

provided nearly identical results to the original analysis.

All of the analyses were underpredicting the global seismic response of the system. Further investigations

were conducted to establish additional reference information that may assist in understanding the cause

for the underprediction. Additional analyses were performed with increased input motion. These

analyses demonstrated that a close match to test results can be achieved by increasing the amplitude of the

input acceleration time history by a factor of two to three times the A-2 input motion. A linear elastic

time history analysis was also performed to establish an elastic baseline analysis. The elastic analysis

considered no material (plasticity) or geometric (large deflection) nonlinearities. The analysis provided

unexpected results. The maximum deflection was nearly twice the deflection predicted by the nonlinear

analysis. Since the linear analysis had removed the geometric as well as the material nonlinearities,

additional analyses were performed to identify the source of the difference. The elasto-plastic analysis

was repeated with no geometric nonlinearity (small deflection theory) and the elastic analysis was

repeated with geometric nonlinearity (large deflection theory). A comparison of deflection results

demonstrated that the differences were due to the application of small deflection versus large deflection

theory. Within the same geometric assumption cases, the differences in peak deflections between elastic

and elasto-plastic analyses were relatively small. However, within the elastic and elasto-plastic cases,

small deflection theory predicted almost twice the deflection than predicted by large deflection theory.

Although these results demonstrated that application of small deflection theory would provide increased

deflections closer to test, they were still below test results and additional investigations were needed.

Investigations on the sensitivity of results to dimensional variations were conducted next. The A-2 model

used as-built dimensions for straight pipes and elbows based on average as-built diameter and wall

thickness measurements provided by NUPEC. A comparison with nominal dimensions showed that the

as-built thickness used in the model was 30% higher than nominal. This difference may be significant

since it results in a stiffer, higher frequency system and lower strain in the elbow. To investigate the

sensitivity of response to dimensional variations, the system was reanalyzed using nominal pipe

dimensions. A frequency analysis demonstrated that by using nominal dimensions, the model provides a

much closer match to the test frequency. The original as-built dimension model had a fundamental

frequency of 2.34 Hz compared to the test frequency of 2.2 Hz. The nominal dimension model had a

frequency of 2.18 Hz. The seismic analysis of the nominal dimension model with small deflection theory

provided the closest match to both displacement and strain ratcheting test results than was observed in any

previous analysis. Based on these results, further discussions were held with NUPEC to verify the

parameters used in the BNL model.

BNL compared key modeling parameters between the ANSYS model and the NUPEC ABAQUS model.

The as-built dimensions were found to be nearly identical, so there was no justification for using nominal

dimensions. Other parameters and modeling assumptions were also compared. A difference in the

modeling of added masses was identified. Both models included concentrated masses at the pipe nodes

located at the centers of gravity of the added masses. The BNL model assumed that the added masses

stiffened the pipe along their length and modified the cross-sectional properties of the pipe to correspond

to the rigid stiffness of the added mass. The NUPEC model did not increase the stiffness of the pipe

along the length of the added mass. Therefore, additional analyses were performed to investigate the

effect of this difference. Frequency analysis of the A-2 model without the added mass stiffness showed a

49
reduction in piping system frequency which more closely matched the test specimen frequency. A

nonlinear seismic analysis of this modified model using small deflection theory showed excellent

comparisons to test displacements and strain ratcheting. Subsequent analyses for the simplified piping

systems discussed in the following sections used this modeling assumption and were based on small

deflection analysis.

Some additional analyses were performed to investigate why small deflection analysis was providing

better results than large deflection analysis. Large deflection analysis was originally used to account for

the effects of elbow ovalization on elbow stiffness. This change in cross-sectional geometry results in

decreased elbow stiffness on closing and increased elbow stiffness on opening. In theory, a large

deflection analysis should account for this change. A model of the SE4 elbow was statically loaded by

applying displacements in opening and closing directions using small deflection and large deflection

analyses. A plot of force versus deflection showed that for small deflections, both analyses predicted the

same elbow stiffness in both the opening and closing directions, as expected. At larger deflections, the

large deflection analysis predicted higher elbow stiffness upon opening and lower stiffness upon closing.

The same effect was observed with an elastic model and an elasto-plastic model. These analyses

demonstrated that the ANSYS elbow model behaved as expected for large and small deflection analysis.

A similar static analysis was performed for the A-2 piping system model. In these analyses, the center of

the pipe was displaced in the positive and in the negative x-directions. In both directions, the large

deflection analysis predicted about 40% higher stiffness at all values of displacement. This was observed

for both elastic and elasto-plastic models. The higher initial stiffness seen in the large deflection analysis

is believed to be an anomaly. At small deflections, the geometric effect is insignificant, and a large

deflection analysis should provide the same results as a small deflection analysis. It appears that the large

deflection analysis of the A-2 model introduces artificial stiffening which does not represent the actual

behavior of the system and provides erroneous results. The higher stiffness would explain the reduced

dynamic global response of the piping system seen in the original analysis. The cause of this anomaly is

unknown. There may be incompatibilities between the different elements used in the model, but further

investigations are needed to ascertain the root cause of this problem.

As noted above, subsequent simplified piping system analyses used small deflection analysis. Although

this type of analysis does not account for the geometric effects of elbow ovalization, the effect was

believed to have a small impact on the results. Some additional analyses were performed to support this

assumption. Since large deflection analysis provided reasonable results for the SEA4 elbow, this model

was further analyzed to compare large deflection analysis with small deflection analysis results. The load

versus deflection curves from the monotonic loading analyses of the elbow describe above had already

demonstrated that the difference in elbow stiffness between large and small deflection analysis was small.

Additional static cycling analyses simulating the SEA test were carried out. A comparison of results

demonstrated that the large deflection cycling analysis predicted only a slight increase in elbow

equivalent stiffness along with a slight increase in strain ratcheting in the elbow. The differences in

results were less than 10 percent. Since the strain levels in the elbow tests were comparable to the strain

levels in the simplified piping system test, it was therefore concluded that within the range of interest, a

small deflection analysis of the simplified piping system model would provide a reasonable level of

accuracy in the results.

5.4 Final A-2 Simplified Piping System Test Analyses and Comparisons

Final nonlinear seismic analyses of the A-2 test were carried out based on the methodology described

above using each of the three plastic hardening models. Comparisons of displacements and elbow strain

time history results between the test and the different models are shown in Figures 5-2 and 5-3. As

shown in Figure 5-2(a), test measurements indicated peak displacements of +170 mm (+6.7 in) and -178

mm (-7 in) at the pipe center with a permanent deformation of-12 mm (-0.47 in) at the end of the test.

so
The BKIN model (Figure 5-2(b)) predicted displacements of +1 46 mm (+5.75 in) and -166 mm (-6.54 in)

with +2 mm (+0.08 in) permanent deformation. The MKIN model (Figure 5-2(c)) predicted +165/-161

mm (+6.5/-6.34 in) and -6 mm (-0.24 in) permanent deformation. The CHAB model (Figure 5-2(d))

predicted +162/-167 mm (+6.38/-6.57 in) and -6 mm (-0.24 in) permanent deformation. A close

examination of the displacement time history plots shows that the Chaboche nonlinear kinematic

hardening model predicts the closest match to the test displacement time history, although the multilinear

kinematic hardening model also produces a reasonably good match.

A comparison of the Elbow A flank strain time history results shows more significant differences. As

shown in Figure 5-3(a), the test results indicated hoop strain ratcheting in the elbow for the first 125

seconds of motion. By the end of the test, the accumulated strain was 3.35%. The BKIN model (Figure

5-3(b)) overpredicted the strain ratcheting with shakedown occurring at 75 seconds with a final permanent

strain of 5.5%. The MKIN model (Figure 5-3(c)) also predicted strain ratcheting during the first 75

seconds but the final permanent strain only reached 1.3%. Finally, the CHAB model (Figure 5-3(d))

initially underpredicted the test strain ratcheting response during the first 50 seconds but continued

ratcheting until 125 seconds with a final accumulated strain of 3.25% by the end of the test. An

examination of the strain time history plots clearly shows that the Chaboche nonlinear kinematic

hardening model provided the best simulation of the strain ratcheting observed in the test. These findings

demonstrated the superior performance of the Chaboche plasticity model in predicting both global

displacement and local strain ratcheting response. Therefore, the remaining post-test analyses were based

on the Chaboche model.

5.5 Final Two-Dimensional and Three-Dimensional Simplified Piping System Test Analyses

The same analytical methodology, plastic hardening model, material properties, and modeling

assumptions used in the final A-2 analysis were applied in the post-test simulation analyses for the twodimensional A-I test and for the three-dimensional B-I, B-2 and B-3 simplified piping system tests. The

A-1 analysis used the same half-symmetry, combination shell and beam element model shown in Figure

5-1. Revised equivalent inertia force time histories at the concentrated masses corresponding to the lower

level acceleration time history input for the A-I test were calculated and applied to the model. The

analysis results are discussed below.

For the three-dimensional Model B tests, a new ANSYS finite element model based on the test specimen

configuration shown in Figure 2-12 was developed. The Model B specimen consisted of a similar run of

pipe in the horizontal plane as Model A with an additional branch line in the vertical plane which

terminated at a nozzle. It included an additional concentrated mass and relied on spring hangers rather

than rollers for gravity support. The system was also pressurized to 19.8 MPa (2870 psi). Because of the

three-dimensional configuration, a full model rather than a half-symmetry model was needed. The

ANSYS model is shown in Figure 5-4. The three elbows were modeled using ANSYS SHELLI81 large

strain shell elements around the entire circumference of the pipe wall. A 100 mm (4 in) length of straight

pipe was included at each end of the elbows. The same types of shell elements were used to model the

nozzle plus 300 mm (12 in) of straight pipe. It was assumed that the plasticity would be localized to the

elbow and nozzle, so the remainder of the piping system was modeled using three-dimensional elastic

straight pipe elements (PIPE 16). The straight pipe elements were connected to the elbow and nozzle shell

element models by a series rigid beam elements arranged in a wheel spoke configuration as shown in

Figure 5-4. ANSYS MASS21 elements were used to represent the added masses in the system.

Boundary conditions included full constraint of all degrees of freedom at the end of the nozzle and

restraint of translations and horizontal axes rotations at the pinned supports. Constant vertical forces were

applied at the spring hanger attachment points to provide gravity support. Loads due to gravity and an

internal pressure of 19.8 MPa (2870 psi) were applied to the system. System damping was defined in

terms of Rayleigh damping coefficients corresponding to the measured modal damping of 0.7 percent

51
from the test. The analyses were performed in the same manner as the Model A analyses using the

ANSYS nonlinear full transient analysis option. The actual horizontal and vertical acceleration time

history input motions from the NUPEC tests were used to develop equivalent inertial force time histories

at the concentrated masses. In the B-1 analysis, the forces were applied in the horizontal x-direction. In

the B-2 analysis, they were applied in the vertical z-direction. In the B-3 analysis, the horizontal and

vertical forces were applied simultaneously. The results of these analyses are discussed below.

NUPEC provided simplified piping system digitized test data at selected critical locations for comparison

to BNL analysis results. This included displacement and strain time history response data. Table 5-1

summarizes the selected types of data and measurement locations. Plots of analysis and test time history

results are shown in Figures 5-5 through 5-21 with analysis results shown on top for comparison to the

test results shown on the bottom. A summary of the comparisons is presented below.

A-I Test:

The comparison of test vs. analysis displacements at the pipe center shown in Figure 5-5 shows excellent

agreement. The test indicated a slight negative permanent deformation which was not predicted by the

analysis. Figure 5-6 compares hoop strains at the outside surface of the Elbow A flank. It shows that the

analysis significantly underpredicts strain ratcheting in this case. Figure 5-7 shows the inside surface

strain time history from analysis compared to the outside surface test strain time history (all test

measurements were on the outside surfaces). This is not a one-to-one comparison, but demonstrates that

the analytical model predicts a comparable amount of strain ratcheting at this elbow location.

A-2 Test:

A comparison of test vs. analysis displacements at the pipe center shown in Figure 5-8 shows excellent

agreement in terms of overall shape of the plots, peak displacements and permanent deformation. The

comparison of Elbow A opening displacements shown in Figure 5-9 also shows similar agreement. The

comparison of Elbow A hoop strains shown in Figure 5-10 shows very good overall agreement. The

comparison of axial strains at the same elbow location (Figure 5-1 1) shows that the analysis

underpredicted the strain cyclic amplitudes and strain ratcheting. The magnitude of strain ratcheting in

the axial direction, however, was much lower than in the hoop direction.

B-I Test:

A comparison of test vs. analysis horizontal displacements at the pipe center shown in Figure 5-12 shows

good agreement in terms of overall shape of the plots, peak displacements and permanent deformation.

The comparison of hoop strains at the nozzle shown in Figure 5-13 shows that the analysis significantly

overpredicted the ratcheting strain. The strain cyclic amplitudes, however, were observed to be small and

appeared comparable between analysis and test. The comparison of axial strains at the nozzle location

(Figure 5-14) shows comparable strain cyclic amplitudes and some overprediction of strain ratcheting.

B-2 Test:

Figure 5-15 shows that the peak vertical displacements at the pipe center were somewhat overpredicted

by the analysis. The vertical displacements, however, were of smaller magnitude than the displacements

in the horizontal motion tests. Both analysis and test showed no permanent deformation at the end of the

test.

52
B-3 Test:

A comparison of test vs. analysis horizontal displacements at the pipe center presented in Figure 5-16

shows good agreement in terms of overall shape of the plots, peak displacements and permanent

deformation. The comparison of vertical displacements presented in Figure 5-17 shows that the analysis

overpredicts the test results. The vertical displacements, however, were of smaller magnitude than the

horizontal displacements and the permanent deformation at the end of the test predicted by analysis was

comparable to test. The comparison of hoop strains at the nozzle shown in Figure 5-18 was consistent

with the B-1 test comparisons, which showed that the analysis significantly overpredicted ratcheting

strain. As in the B-i comparisons, the strain cyclic amplitudes were observed to be small and appeared

comparable between analysis and test. The comparison of axial strains at the nozzle location presented in

Figure 5-19 shows comparable strain cyclic amplitudes but overprediction of strain ratcheting. In this

case, it was noted that in comparison with the B-i test axial strains shown in Figure 5-14, the strain

ratcheting observed in the B-3 test was less than half of the strain ratcheting observed in the B-i test,

despite the similar input excitation level in the horizontal direction. Furthermore, the analyses predicted

nearly identical axial strain ratcheting for the B- 1 and B-3 tests as expected. NUPEC, however,

investigated and confirmed that the B-3 axial strain data was correct. The comparison of hoop strains at

the Elbow A flank (Figure 5-20) shows comparable cyclic strain amplitudes, but the analysis

underpredicted strain ratcheting. In Figure 5-21, hoop strain at the inside surface of the Elbow A flank is

compared to the hoop strain at the outside surface of the Elbow A flank. Although this is not a one-to-one

comparison, it demonstrates that the analytical model predicts a comparable amount of strain ratcheting at

this elbow location.

In summary, the comparisons of simplified piping system post-test simulation analysis results

demonstrated that displacement time history response was predicted with good accuracy. Accurate

prediction of seismic strain and strain ratcheting was achieved in some cases, but in other cases, both

overpredictions and underpredictions were observed.

53
Table 5-1 Summary of Post-Test Analysis Results and Test Comparisons

Test Data Location Figure

Displacement at Center of Pipe DI(X) 5-5

A-1 Hoop Strain at Elbow A SAD-3H 5-6

Hoop Strain at Elbow A (Inside Surface) SAD-3H 5-7

Displacement at Center of Pipe DI(X) 5-8

Opening Displacement of Elbow A D2(L) 5-9

A-2

Hoop Strain at Elbow A SAD-3H 5-10

Axial Strain at Elbow A SAD-3A 5-11

Horizontal Displacement at Center of Pipe DI(X) 5-12

B-i-I Hoop Strain at Nozzle Section A16 +90° SA16-2H 5-13

Axial Strain at Nozzle Section A17 +90° SA17-2A 5-14

B-2-1 Vertical Displacement at Center of Pipe DI(Z) 5-15

Horizontal Displacement at Center of Pipe DI(X) 5-16

Vertical Displacement at Center of Pipe DI(Z) 5-17

Hoop Strain at Nozzle Section A16 +90° SA16-2H 5-18

B-3

Axial Strain at Nozzle Section A17 +900 SA17-2A 5-19

Hoop Strain at Elbow A Section D -85° SAD-7H 5-20

Hoop Strain at Elbow A Section D -85° (Inside Surface) SAD-7H 5-21

54
Elbow A

Detail

Elbow B

Figure 5-1 ANSYS Model A Elasto-Plastic Shell and Beam Element Model

55
200

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*100 -

CL

.1a. -'150==^L 

j 5s

mm = .03937 in F e me (s..) N O 200 2e0 

i200 

s:A 

a i2 o 100 Ti 150(sec)

Figure5-2 A-2NonlinearSimulationAnalysisResults 

- Displacement at Pipe Center- DI(X)

200 

250

56
4.00 7.00

6.50

6.00

ZMS

'

0

0

Z

OC

V0. 8.

00

Z

C

a.

00

0

x

U.UU

4.50

4.00

3.50

3.00

2.50

2.00

1.50

1.00

0.50

0.00

-0.50

-II

-,-I

_, I 

. ! I

0 50 100 150 200 250 0 50 100 150 200 250

Time (sec) Time (sec)

4.00

7

a. 00C

0 50 100 150 200

Time (sec)

250 0 . 50 100 150 200

Time (sec)

250

Figure 5-3 A-2 Nonlinear Simulation Analysis Results - Hoop Strain at Elbow Flank - SAD-3H

57
Elbow C

Detail

Nozzle

Detail --

, Elbow A

vB

ail

..i.

Figure 5-4 ANSYS Model B Elasto-Plastic Shell and Beam Element Model

ElboM

Det;

58
ANSYS A-1 Analysis Results (a1d2 Model)

Pipe Center Displacement

E

0.

a

0 50 100 150 200 250

Time (sec) 1 mm = .03937 in

A-1 Test Results

Pipe Center Displacement

V

E

2

a

0 50 100 150 200

1 mm = .03937 in Time (sec)

250

Figure 5-5 A-I Test Pipe Center Displacement Comparison - DI(X)

59
ANSYS A-1 Analysis Results (al d2 Model)

Hoop Strain at Elbow A Flank

2

en

-0.5

0 so 100 150 200

Time (sec)

250

A-1 Test Results

Hoop Strain at Elbow A Flank

SS

T

ra

0 50 100 150 200 250

rime (sec)

Figure 5-6 A-I Test Hoop Strain at Elbow A Flank Comparison - SAD-3H

60
ANSYS A-1 Analysis Results (a1d2 Model)

Hoop Strain at Elbow A Flank Inside Surface

F

.5

2I

50 100 150 200

Time (sec)

250

A-1 Test Results

Hoop Strain at Elbow A Flank

F

r

2W

(A

0 50 100 150 200

Time (sec)

250

Figure 5-7 A-I Test Hoop Strain at Elbow A Flank Inside Surface Compared to Test Hoop Strain at

Outside Surface - SAD-3H

61
ANSYS A-2 Analysis Results (a2d41 Model)

Pipe Center Displacement

0r

,E

0

-100

-200

so 100 IS1 200 250

1 mm = .03937 in Time (sec)

A-2 Test Results

Pipe Center Displacement

E

E0.

E0

U0

-200

0 so 100 150 200

1 mm = .03937 in rime (sec)

250

Figure 5-8 A-2 Test Pipe Center Displacement Comparison - DI(X)

62
ANSYS A-2 Analysis Results (a2d41 Model)

Elbow A Opening Displacement

E 10

E

E 5

Ut

. 0 a

-20

0 50 100 150 200 250

1 mm = .03937 in Time (sec)

A-2 Test Results

Elbow A Opening Displacement

E10 E

E 5

.E 0 a

0 50 100 150 200

1 mm = .03937 in Time (sec)

250

Figure 5-9 A-2 Test Elbow A Opening Displacement Comparison - D2(L)

63
ANSYS A-2 Analysis Results (a2d41 Model)

Hoop Strain at Elbow A Flank

S0

In

0 50 100 150 200 250

Time (sec)

A-2 Test Results

Hoop Strain at Elbow A Flank

Fc

I,>

0 50 100 150 200 250

Time (sec)

Figure 5-10 A-2 Test Hoop Strain at Elbow A Flank Comparison - SAD-3H

64
ANSYS A-2 Analysis Results (a2d41 Model)

Axial Strain at Elbow A Flank

S7

£

0 50 100 150 200

Time (sec)

250

A-2 Test Results

Axial Strain at Elbow A Flank

0 50 10o 150 200

Time (sec)

250

Figure 5-1 1 A-2 Test Axial Strain at Elbow A Flank Comparison - SAD-3A

65
ANSYS B-1-1 Analysis Results (b1d2 Model)

Pipe Center Displacement

E 20

Ec

.n -20 a

0 20 40 60 80 100

1 mm = .03937 in rime (sec)

120

B-1-1 Test Results

Pipe Center Displacement DI (X)

_ 20

E 0

-

.- -20 B

0 20 40 60 80 100

1 mm = .03937 in Time (sec)

120

Figure 5-12 B-I Test Pipe Center Horizontal Displacement Comparison - DI(X)

66
ANSYS B-1-1 Analysis Results (bId2 Model)

Hoop Strain at Nozzle Section SA16-2H

U)

6 1 __

5

4

3

0 20 40 60 80 100 12C

Time (sec)

B-1-1 Test Results

Hoop Strain at Nozzle Section SA16-22H

6.00

5.00

4.00 -

3.00

2.00 - I i I

1 .0 0 _ _ - _ _ _ _ _ _ f n nn _ _ _ _ _ _ _ _ _ _ _ _

cT

0 20 40 60 80 100 120

Time (sec)

Figure 5-13 B-I Test Hoop Strain at Nozzle Comparison - SA16-2H

67
ANSYS B-1-1 Analysis Results (bld2 Model)

Axial Strain at Nozzle Section SA117-2A

so

2.5

2

IT WI 1w PF4.5

0 20 40 60 so 100 120

Time (sec)

B-1-1 Test Results

Axial Strain at Nozzle Section SA117-2A

2

S

2I

-0.50

0 20 40 60 80 100

Time (sec)

120

Figure 5-14 B-I Test Axial Strain at Nozzle Comparison - SA17-2A

68
ANSYS B-2-1 Analysis Results (b2d2 Model)

Pipe Center Displacement

25

I _ I I i I

I

-E

EC

a.

15

5

-5 -- 2

-10 - II

-12 0 I 1

-20 _ _ _

0 20 40 60

Time (sec)

80 100 120

I mm =.03937 in

B-2-1 Test Results

Pipe Center Vertical Displacement

25

E

e.

0.

an

0 20 40 60 80 100

1 mm = .03937 in Time (sec)

120

Figure 5-15 B-2 Test Pipe Center Vertical Displacement Comparison - DI(Z)

69
ANSYS B-3 Analysis Results (b3d2 Model)

Pipe Center Horizontal Displacement

-E

E0

a1

-100

0 20 40 60 s0 100

1 mm = .03937 in Time (sec)

120

B-3 Test Results

Pipe Center Horizontal Displacement

E

Z

E0

v0

0 20 40 60 80 100

1 mm = .03937 in Time (sec)

120

Figure 5-16 B-3 Test Pipe Center Horizontal Displacement Comparison - DI(X)

70
ANSYS 8-3 Analysis Results (b3d2 Model)

Pipe Center Vertical Displacement

25 - _ ]_ _ _ --

15

10*

E

E S

E 0-

00

-10I

-15

- - 

+

v

E

E

E1

20.

a

-20 l } ll}

0 20 40 60 80 100 12

1 mm = .03937 in Time (sec)

B-3 Test Results

Pipe Center Vertical Displacement

25

20

15

-5

.15 I l

*20 _ _ _ L -

25 . ___

0

0

1 mm=.03937 in

20 40 60

Time (sec)

80 100 120

Figure 5-17 B-3 Test Pipe Center Vertical Displacement Comparison - D1(Z)

71
ANSYS B53 Analysis Results (b3d2 Model)

Hoop Strain at Nozzle Section SAI 6-2H

F

.520

ch

5 ~

4 -

3 1

2-

'I _ _.-- Il

0 20 40 60

Time (sec)

80 100 120

B-3 Test Results

Hoop Strain at Nozzle Section SA116-2H

6.00

.S

0 20 40 60 80 100

Time (sec)

Figure 5-18 B-3 Test Hoop Strain at Nozzle Comparison - SA I 6-2H

120

72
ANSYS B-3 Analysis Results (b3d2 Model)

Axial Strain at Nozzle Section SA17-2A

--i

.5

2I

rn

20 40 60 80 100

Time (sec)

B-3 Test Results

Axial Strain at Nozzle Section SA17-2A

120

7

1.50 -

1.00

0.50 i j 1- -

Adnr - -- -

oo Adsr I

0.00s I

0 20 40 60 80 100 120

Time (sec)

Figure 5-19 B-3 Test Axial Strain at Nozzle Comparison - SA17-2A

73
ANSYS B-3 Analysis Results (b3d2 Model)

Hoop Strain at Elbow A Flank

Jo_

0 20 40 60 80 100

Time (sec)

B-3 Test Results

Hoop Strain at Elbow A Section SAD-7H

120

-A

2I

oa

-0.50

20 40 60 80 100

Time (sec)

120

Figure 5-20 B-3 Test Hoop Strain at Elbow A Flank Comparison - SAD-7H

74
ANSYS B-3 Analysis Results (b3d2 Model)

Hoop Strain at Elbow A Flank Inside Surface

S

EI

0 20 40 60 80 100

Time (sec)

B-3 Test Results

Hoop Strain at Elbow A Section SAD-7H

120

-0.50

0 20 40 60 80 100

Time (sec)

120

Figure 5-21 B-3 Test Hoop Strain at Elbow A Flank Inside Surface Compared to Test Hoop Strain at

Outside Surface - SAD-7H

75
6.0 CONCLUSIONS AND RECOMMENDATIONS

The NUPEC Ultimate Strength Piping Test Program has provided valuable new test data on high level

seismic elasto-plastic behavior and failure modes for typical nuclear power plant piping systems. The

NUPEC program was a comprehensive program which included basic piping material tests, piping

component static and dynamic tests, simplified piping system seismic tests, and shaking table seismic

tests of a representative large-scale piping system to levels well above the typical design earthquake. A

selected number of component and piping system tests including the large-scale piping test were carried

out to failure. The test specimens were heavily instrumented to record displacements, accelerations and

strains versus time at critical locations. The component and piping system tests demonstrated the strain

ratcheting behavior that is expected to occur when a pressurized pipe is subjected to cyclic seismic

loading. All pipe failures that were observed in this test program were characterized as through-wall

cracks which occurred as a result of fatigue ratcheting. This report provides a summary of tests

performed under this program and selected test results which were used by BNL for comparison to posttest analysis.

The test data generated under this program augments the database of existing information on full-scale

piping system and component seismic tests performed in the U.S. Although the NUPEC program was

geared toward demonstrating margins in Japanese piping design, there are a number of similarities in

materials and design requirements that make the test results applicable to U.S. piping designs. Technical

justification for recent changes in the ASME Code piping design rules relied primarily on data collected

under the EPRI/NRC Piping and Fitting Dynamic Reliability Program. The NUPEC data may be

evaluated in the same manner to provide additional technical justification from an independent test

program. As reported in Section 3.0 of this report, BNL performed linear analyses and code evaluations

for selected simplified piping system tests. Linear analysis methods were applied in accordance with

current and earlier ASME Code requirements. The results demonstrated that the highest level NUPEC

simplified piping seismic tests subjected the piping components to stress levels above Code allowables

and illustrated the differences in minimum stress margins between recent revisions of the Code. Based on

the latest NRC endorsed ASME Code rules (1993) and Code Case N-41 I damping, the analyses showed

minimum margins of 2.74 for elbows and 1.37 for nozzle welds. For the same test runs analyzed and

evaluated according to the 2004 ASME Code rules, the analyses showed minimum margins of 1.84 for

elbows and 1.76 for nozzle welds. Additional analyses and Code evaluations could be performed for

other tests including the large-scale piping test to provide further substantiation of Code margins.

The NUPEC Ultimate Strength Piping Test Program also provided excellent test data for benchmarking

inelastic analysis methods and computer codes. The BNL post-test nonlinear analyses described in

Sections 4.0 and 5.0 of this report provided useful insights into the application of state-of-the-art elastoplastic analysis methods for predicting the seismic response of a piping system.

The nonlinear analyses demonstrated that the use of classical linear kinematic hardening rules (which

have been applied in NPP nonlinear analyses) may not be adequate for predicting the strain ratcheting

behavior in piping systems. The application of either bilinear or multilinear kinematic hardening rules in

the analysis of static elbow component cycling predicted high rates of initial strain ratcheting with

shakedown occurring after only five to ten cycles. The tests demonstrated lower initial strain ratcheting

eventually stabilizing to a constant rate of ratcheting without shakedown. The application of the

Chaboche nonlinear kinematic hardening model provided an improved strain ratcheting simulation. The

elbow cycling simulation analysis demonstrated that the Chaboche model more closely predicted the

initial ratcheting followed by a constant rate of ratcheting in subsequent cycles as observed in the test.

For the A-2 simplified piping system test, the Chaboche model also provided the most accurate overall

simulation of strain ratcheting behavior. The accurate simulation of ratcheting is particularly important

for Code evaluation of piping based on plastic or equivalent plastic analyses that rely on demonstration of

77
shakedown. For prediction of displacement response, the A-2 simplified piping system analyses showed

that the Chaboche model also provided the best simulation although the multilinear kinematic hardening

model also provided a reasonably good match. The application of the Chaboche nonlinear kinematic

hardening model to the simulation analyses of the other four simplified piping system tests demonstrated

that the displacement time history response was predicted with reasonably good accuracy. Accurate

simulation of strain and strain ratcheting was achieved in some but not all cases which indicates that even

the Chaboche nonlinear model has limitations which need to be better understood. From a regulatory

standpoint, the reviewer must be aware of the limitations of commonly applied plasticity models. To gain

further understanding of the model limitations, the NRC staff may consider additional studies to

investigate the limitations of plasticity models on representative large-scale piping systems.

The difficulty in achieving precise test simulations during the preliminary A-2 simplified piping system

analyses illustrated the importance of accurately modeling the piping system to predict correct system

dynamic characteristics (frequencies and mode shapes) in an elasto-plastic analysis. In a linear response

spectrum analysis of a piping system, modeling uncertainties such as dimensional variations in the system

may be accounted for by peak broadening. This is important when the fundamental frequency of the

system is on a steep slope of the response spectrum and a small shift in frequency can result in a large

increase in response. In a nonlinear analysis, a small change in frequency of the model may have an even

more significant impact on the response since it may lead to plastic strain ratcheting response versus

elastic response. The application of 15% peak broadening may not provide sufficient safety margins

when inelastic analysis is performed. Further studies to investigate the sensitivity of response due to

modeling uncertainties for large scale piping systems may provide additional insights.

78
7.0 REFERENCES

1. K. Suzuki, et al., Seismic Proving Test of Ultimate Piping Strength, ICONE-9, Paper No. 155,

2001.

2. K. Suzuki, et al., Seismic Proving Test of Ultimate Piping Strength, ICONE-IO, Paper No. 22225,

2002.

3. ANSYS, User's Manual, Release 5.6., ANSYS Inc., Canonsburg, PA 15317, U.S.A.

4. U.S. Nuclear Regulatory Commission, Standard Review Plan for the Review of Safety Analysis

Reports, NUREG-0800, Sections 3.7.3 and 3.9.2, Revision 2, 1989.

5. Chaboche, J.L., Time Independent Constitutive Theories for Cyclic Plasticity, International

Journal ofPlasticity, Vol. 2, pp. 149-188, 1986.

6. Prager, W., A New Method of Analyzing Stresses and Strains in Work Hardening Plastic Solids,

Journal of Applied Mechanics, Vol. 23, pp. 493-496, 1956.

7. Besseling, J.F., A Theory of Elastic, Plastic and Creep Deformations of an Initially Isotropic

Material, Journal ofApplied Mechanics, Vol. 25, pp. 529-536, 1958.

8. Armstrong, P.J. and Frederick, C.O., A Mathematical Representation of the Multiaxial Bauschinger

Effect, CEGB Report No. RD/B/N 731, 1966.

9. Bari, S. and Hassan, T., Anatomy of Coupled Constitutive Models for Ratcheting Simulations,

International Journal of Plasticity, Vol. 16, pp. 381-409, 2000.

79
NRC FORM 335 U.S. NUCLEAR REGULATORY COMMISSION 1. REPORT NUMBER

(9.2004) (Assigned by NRC, Add Vol., Supp., Rev.,

NRCMD 3.7 and Addendum Numbers, If any.)

BIBLIOGRAPHIC DATA SHEET NUREGICR-6889

(See instructions on the reverse) BNL-NUREG-74978-2005

2. TITLE AND SUBTITLE 3. DATE REPORT PUBLISHED

Seismic Analysis of Simplified Piping Systems for the NUPEC Ultimate Strength Piping Test MONTH YEAR

Program December 2005

4. FIN OR GRANT NUMBER

N-6076

5. AUTHOR(S) 6. TYPE OF REPORT

DeGrassi, G., Hofmayer, C. Technical

7. PERIOD COVERED (Inclusive Dates)

0. PERFORMING ORGANIZATION - NAME AND ADDRESS (Il NRC, provide Division. Office orRegion. US NuclearRegulatory Commission, and mailing address; if contractor,

provide name and mailing address )

Brookhaven National Laboratory

Energy Sciences and Technology Department

Upton, NY 11973-5000

9. SPONSORING ORGANIZATION - NAME AND ADDRESS (IfNRC, type 'Same aJs above'; icontractor, provide NRC Division, Office or Region, U.S. Nuclear Regulatory Commission,

and mailing address.)

Division of Engineering Technology

U. S. Nuclear Regulatory Commission

Washington, D.C. 20555-0001

10. SUPPLEMENTARY NOTES

Sved . A. .Ali. . . Proaram . . Manaqer ..

11. ABSTRACT (200 words or less)

The Nuclear Power Engineering Corporation (NUPEC) of Japan has conducted a multi-year test program for the Ministry of

Economy, Trade and Industry (METI) of Japan to investigate the behavior of typical Nuclear Power Plant (NPP) piping systems

under large seismic loads. The objectives of this program are to develop a better understanding of the elasto-plastic response

and ultimate strength of nuclear piping systems, to ascertain the seismic safety margins in current piping design codes, and to

assess new code allowable stress rules. The test program consisted of a series of static and dynamic material tests, piping

component tests, simplified piping system tests, and large scale piping tests. As part of collaborative efforts between the United

States and Japan on seismic issues, the U.S. Nuclear Regulatory Commission (NRC) and the Brookhaven National Laboratory

(BNL) participated in this program by performing both pre-test and post-test analysis for selected tests, and by evaluation of

program results. This report presents a summary of the NUPEC tests, describes the BNL post-test analyses of selected

simplified piping system tests, and discusses the insights gained from this program.

12. KEY WORDS/DESCRIPTORS (List words or phrases that will assist researchers in locating the report.) 13. AVAILABILITY STATEMENT

Seismic Analysis, seismic testing, piping design, ASME code evaluation, fatigue ratcheting, plasticity, unlimited

kinematic hardening, fatigue failure. 14. SECURITY CLASSIFICATION

(This Page)

unclassified

(This Report)

unclassified

15. NUMBER OF PAGES

16. PRICE

NRC FORM 335 (9-2004) PRINTED ON RECYCLED PAPER
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on recycleds

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NUREG/CR-6889 SEISMIC ANALYSIS OF SIMPLIFIED PIPING SYSTEMS FOR THE NUPEC ULTIMATE

STRENGTH PIPING TEST PROGRAM

DECEMBER 2005

UNITED STATES

NUCLEAR REGULATORY COMMISSION

WASHINGTON, DC 20555-0001

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