NUREG/CR-6804

ANL-02/28

Second U.S. Nuclear

Regulatory Commission

International Steam Generator

Tube Integrity Research

Program

Final Project Summary Report

Argonne National Laboratory

U.S. Nuclear Regulatory Commission 4

Office of Nuclear Regulatory Research

Washington, DC 20555-0001
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NUREG/CR-6804

ANL-02/28

Second U.S. Nuclear

Regulatory Commission

International Steam Generator

Tube Integrity Research

Program

Final Project Summary Report

Manuscript Completed: August 2002

Date Published: September 2003

Prepared by

S. Bakhtiari, K. E. Kasza, D. S. Kupperman,

S. Majumdar, J. Y. Park, W. J. Shack,

D. R. Diercks

Argonne National Laboratory

9700 South Cass Avenue

Argonne, IL 60439

J. Muscara, NRC Project Manager

Prepared for

Division of Engineering

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 2055540001

NRC Job Code W6487
NUREG/CR-6804, has been reproduced

from the best available copy.
Abstract

This report summarizes work performed under the USNRC Steam Generator Integrity

Program at Argonne National Laboratory. The areas addressed by the program

include assessment of procedures and equipment used for in-service (ISI) inspection of steam

generator (SG) tubes; recommendations for criteria and requirements to improve the reliability

and accuracy of ISI: validation and improvement of correlations and models that are used to

evaluate integrity and leakage of degraded SG tubes: and validation and improvement of

correlations and models that predict the generation and progression of degradation in SG tubes

as a function of aging.

iii
Contents

Abstract.iii

Executive Sunnary ........................................................................ xviii

Acknowledgments.................................................................................................................... .xxiv

Acronyms and Abbreviations ........................................................................ xxv

1 Introduction .................................................................... 1

2. Assessment of Inspection Reliability ................................................................... 2

2.1 Steam Generator Mock-Up Round Robin ........................................................... 2

2.1.1 Introduction ....................................................... 2

2.1.2 Steam Generator Mock-Up Facility ..................................................... 3

2.1.3 Design and Organization of Round-Robin .......................................... 14

2.1.4 Comparison of Round-Robin Data Acquisition and Analysis to

Field ISI ...................................................... 17

2.1.5 Strategy for Evaluation of Results ...................................................... 18

2.1.6 Statistical Analysis ...................................................... 19

2.1.7 Results of the Analysis Round-Robin ................................................. 22

2.1.8 Nature of Missed Flaws ..................... ................................. 32

2.1.9 Nature of Ovemalls.............................................................................. 32

2.1.10 Implications of the Round-Robin Results ............................. .............. 33

2.2 Tube Removal from McGuire 1 Retired Steam Generator .............. ................... 33

2.2.1 ReviewofDukeISIData .............................................................. 33

2.2.2 Results of Field NDE of Deplugged McGuire Tubes ........... ................ 34

2.2.3 EC Signal Growth .............................................................. 36

2.3 Effect of Surface Oxide Films on Eddy Current Signals from SCC ....... ............. 36

2.4 Crack Sizing .............................................................. 37

2.5 Array Probes .............................................................. 39

2.6 Use of Lamb Waves .............................................................. 43

3 Research on ISI Technology ................................................................... 45

3.1 Analytical Modeling of EC NDE ................................................................... 45

3.2 Correlation of Fallure Pressure with Bobbin Coil Signals .............. ................... 48

3.2.1 Multiple Linear Regression and Neural-Network Model; ARC

Data Base ............................................................. 48

3.3 Computer-aided Analysis of Eddy Current Inspection Data. . . 56

3.3.1 Data Processing Structure ............................... ....................... 56

3.3.2 Conversion and Calibration Routines .................................................. 57

3.3.3 Signal Processing, Data Analysis, and Display Routines ................... 60

3.3.4 Multiparameter Data Analysis ...................................................... 69

3.3.5 RulFe-based Flaw Identification ...................................................... 74

3.3.6 Data Quality ...................................................... 74

3.4 Multiparameter Data Analysis Results ................................................................... 75

3.4.1 Graphical Display of Multiparameter Analysis of NDE Data ........ ..... 75

v
3.4.2 Assessment of Sizing Accuracy in Presence of Artifacts ..................... 76

3.4.4 Analysis of Field-Degraded Tubes ....................................................... 78

4 Research on Degradation Modes and Integrity . ............................................................. 88

4.1 Production of Laboratory-Degraded Tubes and Characterization of Flaws ...... 88

4.1.1 Materials Properties of Alloy 600 ...................................................... 88

4.1.2 Production of Laboratory Cracked Tubes ........................................... 90

4.1.3 Characterization of Laboratory-Cracked Tube Specimens:

Specimen Set for Control and Verification.......................................... 91

4.1.4 Characterization of Laboratory-Cracked Tube Specimens: MockUp Specimens...................................................................................... 91

4.1.5 Characterization of Field-Serviced Tube Specimens: McGuire SG

Tubes................................................................................................... 98

4.1.6 Characterization of Laser Notched Tube Specimens .......................... 98

4.2 Crack Evolution Studies ............................................................. 99

4.3 Tube Integrity Tests under Design Basis Conditions . ........................ 108

4.3.1 High-Temperature Pressure and Leak-Rate Test Facility ................... 108

4.3.2 Room-Temperature High-Pressure Test Facility ................................. 111

4.3.3 Test Results for Machined Flaws .................................... ........... 115

4.3.4 Laboratory-Grown SCC ............................................... 134

4.3.5 SCC from Operating Reactors ............................................... 141

4.4 Jet Erosion Under Design Basis Conditions ................................ 143

4.5 Tube Failure Models under Design-Basis and Severe Accident Conditions ..... 148

4.5.1 Characterization of Alloy 600 Tubing Material ................................... 148

4.5.2 Low-Temperature Failure Models .................................... ; . 149

4.5.3 Leak Rate Models for Single-Phase Flow ................................... 157

4.5.4 Pressure and Leak Rate Tests on Specimens with Stress

Corrosion Cracks ............................................... 163

4.5.5 High-Temperature Failure Models ............................................... 164

4.5.6 High-Temperature Leak Rate and Jet Impingement Erosion ............. 171

5 Integrated Model for Steam Generator Integrity Assessments . ..................................... 176

5.1 Description of Original CANTIA.......................................................................... 176

5.2 New Features Added in ANL/CANTIA ........... .. ............... 177

5.2.1 Initial Flaw Size Distributions .............................. 177

5.2.2 Flaw Growth Model .............................. 177

5.2.3 Growth Rate of Stress Corrosion Crack .............................. 177

5.2.4 Failure Models .............................. 178

5.2.5 Leak Rate Models .............................. 178

5.2.6 Residual stress.................................................................................... 178

5.3 Applications of ANL/CANTILA .......................... 178

5.3.1 Effect of POD ........................... 178

5.3.2 Effect of Crack Initiation .......................... 182

References................................................................................................................................ 185

vi
Figures

2.1 Schematic representation of tube bundle in steam generator mock-up ......... ........... 4

2.2 Photograph of mock-up during acquisition of eddy current data ........... ................... 5

2.3 Dye-penetrant examination of tube specimen SGL865 showing an LODSCC ........... 6

2.4 Cross-sectional optical metallography. branched LODSCC and LODSCC ................ 6

2.5 Differential bobbin coil Lissajous figure at 400 kHz from Crack In McGuire SG ...... 9

2.6 Differential bobbin coil Lissajous figure at 400 kHz from Argonne-grown mockup crack ................................................................... 9

2.7 Fractography of tube specimen SGL-413 .................................................................. 10

2.8. Sizes and shapes of LODSCCs in tube specimen AGL-536 determined by EC

NDE using the multiparameter algorithm and fractography ...................................... 11

2.9. Comparison of maximum depth determined by the multiparameter algorithm

with that determined by fractography .11

2.10. Comparison of depths determined by the multiparameter algorithm with those

by fractography and regression fit as a function of the multiparameter depth

estimate........................................................................................................................ 12

2.11. Comparison of depths determined by the multiparameter algorithm with those

by fractography ...................................... 12

2.12 Standard deviation in percent throughwall as a function of maximum depth .......... 13

2.13 BC POD for TSP data as a function of maximum depth for LODSCC and

LIDSCC using maximum likelihood fit with the one-sided 95% confidence limit ..... 23

2.14 BC POD for TSP data as a function of maximum depth for LODSCC using

maximum likelihood fit .................................................................. 23

2.15 BC POD for freespan and TSP data as a function of maximum depth for

LODSCC and LIDSCC by using maximum likelihood fit with the one-sided 95%

confidence limit ............................................. ....... 242.............................

2.16 Round-robin TSP bobbin coil resolution analysts results as a function of BC

voltage.......................................................................................................................... 24

2.17 Logistic fit curves for TSP BC POD as a function of voltage for LODSCC and

LIDSCC .................................................................. 24

2.18 Cumulative distribution of normalized standard deviations for BC voltages for

LODSCC at tube support plates....................................................................I.............. 25

2.19 BC POD by team for TSP LODSCC .................................................................. 26

2.20 BC POD by team for freespan LODSCC .26

2.21 BC POD by team for TSP LIDSCC. 26

2.22 BC POD for TSP LODSCC as a funmction of mp by using maximum likelihood fit

with an estimate of the one-sided 95% confidence limit .27

2.23 BC POD for TSP LIDSCC as a function of mp by using maximum likelihood fit

with an estimate of the one-sided 95% confidence limit .................... I..............I ......... 27

2.24 BC POD for freespan data for LODSCC as a function of mp by using maximum

likelihood fit with an estimate of the one-sided 95% confidence limit .27

vie
2.25 MRPC POD as a function of maximum depth for LIDSCC and CIDSCC

combined and LODSCC and CODSCC combined in the tube sheet .......................... 29

2.26 MRPC POD as a function of maximum depth for axial and circumferential

IDSCC and ODSCC in the tube sheet......................................................................... 30

2.27 MRPC POD by team as a function of maximum depth In the tube sheet .................. 30

2.28 BC and MRPC POD for axial and circumferential IDSCC and ODSCC at the

tube sheet.................................................................................................................... 31

2.29 BC and MRPC POD for CIDSCC and LIDSCC at the tube sheet ................................ 31

2.30 BC and MRPC POD for LIDSCC at the tube sheet ..................................................... 31

2.31 TSP LODSCC with maximum depth of 99% TW that was missed by teams

analyzing MRPC data .32

2.32 C-scan plot of tube R45C40 using s 2.92-mm-diameter MRPC showing top of

tube sheet in December 1995 .37

2.33 C-scan plot of tube R45C40 using s 2.92-mm-diameter MRPC showing top of

tube sheet in July 1997 .37

2.34 C-scan plot from 2.92-mm-diameter MRPC: arrow points to indication at top of

tube sheet in tube R39C46 .38

2.35 C-scan plot of same region shown in Fig. 2.34 taken with a 2.92-mm-diameter

MRPC in July 1997 after tube was deplugged .38

2.36 Estimated length with range of uncertainty for CODSCC at top of tube sheet in

tube R39C46 from McGuire retired steam generator before and after plugging........ 38

2.37 Mag-blas BC Lissajous figures before and after corrosion products were formed

in a tube with axial ODSCC by exposing the tube to PWR conditions for about

two months .40

2.38 Maximum depths for a variety of crack morphologies estimated by two

recognized EC NDE experts using the same procedure with +Point data at

300 kHz and by Argonne using the multiparameter algorithm with input mainly

from pancake phase and amplitude data .. 41

2.39 X-Probe 3D Image of EC signal amplitude as function of position around a test

section in tube sheet simulation level of NRC steam generator mock-up .41

2.40 +Point 3D image of EC signal amplitude as function of position around same

mock-up test section as shown in Fig. 2.39 .42

2.41 3D plot of depth vs. position around same mock-up test section as shown in

Figs. 2.39 and 2.40 generated by ANL multiparameter alogrthm. 42

2.42 Cross-section proffle showing depth vs. circumferential position for same mockup test sections as shown in Fig. 2.41 generated by ANL multiparameter

algorithm ...................................................................................................................... 42

2.43 Ultrasonic Lamb wave echo from a CODSCC in Alloy 600 tube with laboratorygrown crack in the roll transition................................................................................ 44

3.1 Distribution+ of pulled-tube and model-boiler indications as a function of failure

pressure used for multivariate analysis...................................................................... 51

3.2 Predicted vs. measured failure pressures for regression and neural-network fits

when all model-boiler and pulled-tube indications are included for training .52

vili
3.3 Predicted vs. measured failure pressures for regression and neural-network fits

when =50% of model-boiler and pulled-tube indications are randomly selected

for training .................................................................. 52

3.4 Predicted vs. measured failure pressures for regression and neural-network fits

for model-boiler and pulled-tube Indications, with LOOT training and

prediction method........................................................................................................ 53

3.5 Residual vs. regression-predicted failure pressures when all model-boiler and

pulled-tube specimens are included for training ....................................................... 54

3.6 Cumulative distribution of residual failure pressure............................. 54

3.7 Regression-predicted vs. measured failure pressures for all model-boiler and

pulled-tube indications, using the APC model ................................................. .......... I 55

3.8 Residual vs. regression-predicted failure pressures for all model-boiler and

pulled-tube specimens, using the APC model ............................................................ 55

3.9 Cumulative distribution of APC model residual failure pressures, based on an

assumed normal distribution function ........................................................ ............... 55

3.10. Schematic diagram showing basic structure of computer-aided data analysis

algorithm...................................................................................................................... 57

3.11 Display of GUIs for main ECTool menu with data calibration submenus, raw

data and trigger channel, and calibrated data in strip chart and Lissajous

display format.............................................................................................................. 59

3.12 Series of MAIIAB-based GUI tools for computer-aided analysis of EC

inspection results acquired with standard commercial instruments ......................... 61

3.13 Original and restored signal for Type-3 laser-cut specimen #5528-3-3 with

nominal OD flaw depth of 40% IW .65

3.14 Original and restored signals with 2-D Lorentzlan kernels for Type-4 laser-cut

specimen #5469-2-1 with nominal OD flaw depth of 80% TW.................................. 66

3.15 Data analysis results for laboratory-grown specimen with longitudinal ODSCC

showing terrain plot of relative OD depth for cracked zone and restored data by

- inverse filtering .67

3.16 Data analysis results for laboratory-grown specimen with longitudinal ODSCC

showing terrain plot of relative OD depth for cracked zone and restored data by

inverse filtering .......... 68

3.17 Demonstration of 2-D signal suppression on RPC traces of simulated raw data

at 400 kHz and 200 kHz frequencies composed of a 40% 1W axial OD notch at

an arbitrarily selected location within a 270° tube support ring and processed

data using 400 kHz primary and 200 kHz auxiliaxy frequency and 300 kHz

primary and 100 kHz auxiliary frequency .................................................................. 73

3.18 Demonstration of 2-D signal suppression on RPC traces of simulated raw data

at 400 kHz and 200 kHz frequencies composed of a 40%'1W axial OD and

40% 'IW circumferential ID notch at arbitrarily selected locations within a 2700

tube support ring and processed data using 400 kHz primary and 200 kHz

auxiliary frequency and 300 kHz primary and 100 kHz auxiliary frequency .73

ix
3.19 Conventional Isometric plot of +PotntTM probe amplitude response, image

display, and multiparameter sizing terrain plot for laboratory-grown

circumferential ODSCC .................................................................. 76

3.20 Conventional Isometric plot of +PointTM probe amplitude response, image

display, and multiparameter sizing terrain plot of same specimen with

laboratory-grown axial ODSCC .................................................................. 77

3.21 Representative display of data analysis results for specimen with

laboratory-grown ODSCC degradation showing image display, terrain plot, axial

projection, and circumferential projection .................................................................. 78

3.22 Representative display of data analysis results for specimen with

laboratory-grown axial ODSCC showing pre- and post-processed images of

normalized data, terrain plot, and axial and circumferential projections .................. 79

3.23 Representative display of data analysis results for specimen with

laboratory-grown axial ODSCC showing terrain, and cross-sectional views of

estimated flaw profile .................................................................. 80

3.24 Data analysis results for McGuire pulled tube section with axial ODSCC

degradation at dented TSP intersection .................................................................. 82

3.25 Data analysis results for McGuire pulled tube section with axial ODSCC

degradation at dented TSP intersection .................................................................. 83

3.26 Data analysis results for McGuire pulled tube section with axial ODSCC

degradation at dented TSP intersection .................................................................. 84

3.27 Data analysis results for McGuire pulled tube section with axial ODSCC

degradation at dented TSP intersection .................................................................. 85

3.28 Data analysis results for McGuire pulled tube section with axial ODSCC

degradation at dented TSP Intersection .................................................................. 86

3.29 Data analysis results for McGuire pulled tube section with axial ODSCC

degradation at dented TSP intersection .................................................................. 87

4.1 Fractography profile and NDE for circumferential IDSCC 1 in specimen SGL101 .... 92

4.2 Fractography proffle and NDE for circumferential IDSCC 2 in specimen SGL1O1 .... 92

4.3 Fractography profile and NDE for circumferential ODSCC in specimen SGL271 ..... 93

4.4 Fractography profile and NDE for axial ODSCC in specimen SGL410 ...................... 93

4.5 Fractography proflle and NDE for axial ODSCC In specimen SGL412 ...................... 94

4.6 Fractography profile and NDE for axial ODSCC In specimen SGL413 ...................... 94

4.7 Fractography profile and NDE for axial ODSCC In specimen SGL434 ...................... 95

4.8 Fractography profile and NDE for axial ODSCC in specimen SGL487 ...................... 95

4.9 Fractography profile and NDE for axial ODSCC in specimen SGIA88 ...................... 96

4.10 Fractography profile and NDE for axial ODSCC In specimen SGL490 ...................... 96

4.11 Fractography profile and NDE for circumferential ODSCC in specimen

PNNL/W 4-02 .................................................................. 97

4.12 Fractography profile and NDE for circumferential ODSCC in specimen

PNNL/W 4-03 .................................................................. 97

x
4.13 Fractography profile and NDE for circumferential ODSCC in specimen

PNNL/W 4-05 .................................................................. 98

4.14 Fractography proffle and NDE for axial ODSCC in McGuire tube specimen

4-43-2 .................................................................. 99

4.15 Fractography proffle and NDE for axial ODSCC In McGuire tube specimen

5-51-2 .................................................................. 100

4.16 Fractography profile and NDE for axial ODSCC in McGuire tube specimen

7-24-3 ................................................................... 1o

4.17 Fractography profile and NDE for axial ODSCC in McGuire tube specimen

14-55-3 .................................................................. 101

4.18 Fractography of axial laser-cut 5528-1-1 and profiles determined by EC/NDE

and fractography ............ 101

4.19 Fractography of axial laser-cut 5528-1-2 and profiles determined by EC/NDE

and fractography ............ 102

4.20 Fractography of axial laser-cut 5528-2-1 and profiles determined by EC/NDE

and fractography ............ 102

4.21 Fractography of axial laser-cut 5528-3-2 and profiles determined by EC/NDE

and fractography .103

4.22 Elght-L autoclaves used for c-ring type static SCC and high temperature

electrochemistry tests.................................................................................................. 103

4.23 Small-diameter Hastelloy pipe autoclave system for CER....................................... 104

4.24 Accuracy and stability of Ag/AgCl reference electrodes that were manufactured

with different oxidation current .................................................................................. 105

4.25 Electrochemical polarization curve for Alloy 600 specimen In 0.4 M sodium

sulfate aqueous solution at 25, 100, 2500C ................................................................ 105

4.26 Anodic polarization of Alloy 600 in 0.4 M Na2SO4 at pH 3.0 and pH 6.8 ................. 106

4.27 Electrochemical impedance and current at 0.02 Hz for Alloy 600 in pH 6.8 and

0.4 M Na2SO4 aqueous solution at room temperature...................................... 106

4.28 Electrochemical impedance and current at 100 Hz for Alloy 600 in pH 6.8 and

0.4 M Na2SO4 aqueous solution at room temperature...................................... 107

4.29 ANL Pressure and Leak-Rate Test Facility ............................................................... 109

4.30 High-Temperature Pressure and Leak-Rate Test Facility ........................................... 110

4.31 Tube test module for Pressure and Leak-Rate Test Facility ....................................... 111

4.32 Schematic diagram of High-Pressure Test Facility pressurizer and associated

components.................................................................................................................. 112

4.33 Overall view of major components of High-Pressure Test Facility ............................. 113

4.34 Containment module for testing radioactively-contaminated field-pulled tubes ....... 114

4.35 Side view of tube with bladder and bored plug seal being installed ........................... 114

4.36 End view of tube with bladder and plug installed ...................................................... 115

4.37 Variation of the ratio of flow path length to the crack opening width with crack

length for normal operating and main steam line break conditions ........... ............... 116

XI
4.38 Posttest appearance of specimen OM121, tested with bladder at a

pressurization rate of 13.8 MPa/s............................................................................... 119

4.39 Top and side views of tube OM101 with a 12.7-mm-long 100% 'W EDM axial

OD notch, tested with a 3.2-mm-thick bladder at a pressurization rate of

13.8 MPa/s.1.................................................................................................................. 121

4.40 Specimen OM133 with 12.7-mm-long 100% '1W OD axial EDM notch tested

with 3.2-mm-thick hard bladder and small backup foil at a pressurization rate

of 13.8 M Pa/s .............................................................................................................. 121

4.41 Specimen OM149 with two aligned axial notches after Phase 1 testing .124

4.42 Specimen OM150 with two shifted axial notches after Phase 1 testing .................... 124

4.43 Specimen OM151 with two shifted axial notches after Phase 1 testing..................... 125

4.44 Specimen OM152 with two shifted axial notches after Phase 1 testing .125

4.45 Specimen OM 160 with two aligned axial notches after Phase 1 testing .125

4.46 Specimen OM161 with two shifted axial notches after Phase 1 testing .125

4.47 Specimen OM162 with two shifted axial notches after Phase 1 testing .125

4.48 Specimen OM152 with two shifted axial notches after Phase 2 burst testing........... 125

4.49 Specimen OM159 with two shifted axial notches after Phase 2 burst testing .126

4.50 Specimen OM162 with two shifted axial notches after Phase 2 burst testing .126

4.51 Configurations of notches in various types 1-5 of laser-cut tube specimens 128

4.52 Configurations of notches in various types 6-10 of laser-cut tube specimens.......... 129

4.53 Appearance of Specimen 5516-4-3 with Type 2 flaw before and after Stage 1

testing.......................................................................................................................... 131

4.54 Appearance of Specimen 5528-3-3 with Type 3 flaw before and after Stage 1

testing.......................................................................................................................... 131

4.55 Appearance of Specimen 5469-2-2 with Type 4 flaw before and after Stage 1

testing.......................................................................................................................... 131

4.56 Appearance of Specimen 5469-2-4 with Type 5 flaw before and after Stage 1

testing........................................................................................................................... 133

4.57 Calculated variations of crack opening displacements with pressure .135

4.58 Pre-test image of Specimen SGL- 177 with flaw highlighted by dye penetrant

and digital image processing....................................................................................... 137

4.59 Pre-test image of Specimen SGI- 195 with flaw highlighted by dye penetrant

and digital image processing....................................................................................... 137

4.60 Pre-test image of Specimen SGb.219 with flaw highlighted by dye penetrant

and digital image processing....................................................................................... 137

4.61 Pre-test dye-penetrant digital image of Westinghouse ODSCC tube produced

using doped steam...................................................................................................... 138

4.62 Photograph of failed region of Specimen SGL- 195 after testing................................. 139

4.63 Photograph of failed region of Specimen SGL- 177 after testing .139

4.64 Post-test appearance of complex ODSCC flaw in Specimen SGL-226 exhibiting

both axial and circumferential cracking..................................................................... 141

xit
4.65 Post-test appearance of complex ODSCC flaw in Specimen SGL-363 exhibiting

a large opening resulting from both axial and circumferential cracking ................... 141

4.66 Set-up for jet-impact erosion test showing EDM hole in Alloy 600 tube and

Alloy 600 impact specimen fastened by spot-welded straps to the tube ................... 144

4.67 Close-up view of 0.79-mm hole in jet-impact specimen after 265 min of testing

at 2820C and 17.2 MPa with 3.7 MPa back pressure ................................................. 144

4.68 Depth of the erosion zone shown in FIg. 4.67, as determined using

interferometric techniques........................................................................................... 145

4.69 Jet erosion penetration depth as a function of water temperature for 0.40-,

0.79-, and 1.59-mm-diameter EDM holes .................................................................. 147

4.70 Flow stress curves for various product forms of Alloy 600 ......................................... 149

4.71 Bilinear fit to existing creep data for Alloy 600 ........................................................... 150

4.72 Observed vs. predicted ligament rupture pressures for tubes with EDM notches

using k = 0.55 and =1. ........................................................... 153

4.73 Observed vs. predicted ligament rupture pressures for tubes with =80% deep

laser-cut notches using k = 0.55 and 1 ................................................................. 153

4.74 Determination of equivalent rectangular crack .................................... 157

4.75 Predicted and observed rupture pressures for through-thickness ligament as a

function of axial or circumferential ligament widths for two part-throughwall

EDM notches 6-mm and 13-mm long .158

4.76 Variation of average ligament thickness with pressure for two 6-mm-long, 100%

throughwall notches separated by 1.3-mm-wide ligament and variation of

ligament rupture pressure with ligament width for two 13-mm-long and two

6-mm-long, 100% throughwall notches...................................................................... 158

4.77 Geometries for multiple axial notch specimens.......................................................... 159

4.78 Observed vs. predicted ligament rupture pressures for multiple laser-cut

notches......................................................................................................................... 160

4.79 Calculated and experimentally measured leak rates at 201C for as-received and

heat-treated 22-mm-diameter tubes with 25.4 mm and 12.7 mm throughwall

axial EDM notches ....................................................................................................... 162

4.80 Calculated and measured pressure vs. leak rate plots for tests OM169 and

OM171 both conducted on as-received tubes at 200C .162

4.81 Calculated vs. experimentally measured leak rates in as-received 22-mmdiameter tubes with a 12.7 mm flaw at 2820C .163

4.82 Crack depth profiles and measured and predicted pressure vs. leak rate plots

for Specimen SGL-480 .165

4.83 Crack depth profiles and measured and predicted pressure vs. leak rate plots

for specimen SGL-493 .165

4.84 Crack depth profiles and predicted and measured pressure vs. leak rate plots of

Specimen SG731. .166

4.85 Crack depth profiles and predicted and measured pressure vs. leak rate plots of

Specimen SGL-822 .166

xiii
4.86 INEEL-calculated severe-accident temperature ramp and simulated ramp used

in ANL severe-accident tube-rupture tests ................................................................ 169

4.87 EPRI-calculated severe-accident temperature ramp and simulation used in ANM

severe-accident tube-rupture tests ................................. ............................... 169

4.88 Creep-rupture-model predicted vs. observed failure temperatures for hightemperature failure tests conducted with the INEEL and EPRI temperature

ramps........................................................................................................................... 169

4.89 Flow-stress-model predicted vs. observed failure temperatures for hightemperature failure tests conducted with the INEEL and EPRI temperature

ramps........................................................................................................................... 170

4.90 Flow stresses of Alloy 600 and Electrosleeve material as a function of

temperature..........................................................................0 ....................................... 170

4.91 Reference geometry for Electrosleeved steam-generator tube with axial crack ......... 170

4.92 Time-temperature history used for testing Electrosleeved tubes with notches in

the parent tubes at ANL .................................................................. 172

4.93 Observed vs. predicted failure temperatures of ANL and FIT tests using actual

notch and Electrosleeve geometry and, for the FF1 tests, actual test temperature

ramp ................................................................ 172

4.94 Comparison of velocity of small particles entrained in flow with fluid velocity ......... 173

4.95 Predicted and experimentally measured notch opening displacement for two

6.4-mm circumferential notches loaded axially and subjected to the Case 6RU

temperature transient and predicted crack-opening area with crack length at

final temperatures 700 and 7501C for a tube subjected to Case 6RU transient . 174

5.1 Variation of probability of burst of one or more tubes with time for initial

residual stress = 30 ksi, POD = 0.9, and inspection interval of 10,000 h .179

5.2 Variation of the distribution of nondimensional crack depth and crack length

with time for initial residual stress of 30 ksi and POD = 0.9 .179

5.3 Variation of probability of burst of one or more tubes with time for initial

residual stress = 30 ksi, POD = 0.6, and inspection interval of 10,000 h .179

5.4 Variation of the distribution of crack depth and crack length with time for

initial residual stress of 30 ksi and POD = 0.6 .180

5.5 Variation of probability of burst of one or more tubes with time for initial

residual stress = 30 ksl, POD = 0.01, and inspection interval of 10,000 h .180

5.6 Variation of the distribution of crack depth and crack length with time for

initial residual stress of 30 kst and POD = 0.01 .180

5.7 Variation of the distribution of crack depth and crack length with time using

POD curve for TSP ID flaws and a constant POD = 0.6 . 182

5.8 Variation of probability of rupture of one or more tubes with time in effective

full power years for a constant POD = 0.6 and a POD curve for TSP/ID flaws . 182

5.9 Two cases of cumulative Welbull distributions assumed for crack initiation

times............................................................................................................................. 183

5.10 Evolution of the distribution for the number of flaws with time for Weibull

initiation Case 1 and Case 2 .................... 183

xiv
5.11 Distributions of crack depth calculated without initiation and with Weibull

initiation Case 1 and-Case 2 at 30,000 and 100,000 h .............................................. 184

5.12 Distributions of crack length calculated without initiation and with Weibull

initiation Case 1 and Case 2 at 30,000 and 80,000 h ................................................ 184

xv
Tables

2.1 Distribution of flaw types............................................................................................. 7

2.2 Flaw types and quantity .................................................................. 7

2.3 Bobbin coil calls of primary, secondary and resolution analysts for SCC at the

TSP .29

2.4 Bobbin coil voltages before plugging and EC voltages after deplugging for

selected tubes from hot-leg side of McGuire steam generator D ................................ 35

2.5. Rotating pancake coil voltages before plugging and EC voltages after deplugging

for selected tubes from hot-leg side of McGuire steam generator D ........................... 35

2.6 Comparison of BC data before and after tube pull ..................................................... 39

2.7 BC voltages and phase angles for two tubes before and after exposure to PWR

water conditions, resulting in formation of corrosion products .40

3.1 Pulled-tube specimens from ARC database incorporated for failure pressure

analysis .......................................................................................................................... 49

3.2 Destructive examination and failure pressure measurements for pulled .tubes

from ARC model-boiler database that contained ODSCC degradation at TSP

intersections................................................................................................................. 49

4.1 Elemental analysis of Alloy 600 tubing material ........................................................ 89

4.2 Room-temperature tensile properties, hardness and grain size of Alloy 600

tubing........................................................................................................................... 89

4.3 Results from pressure and leak-rate tests on tubes with 38.1-mm-long 80% TW

axial EDM flaws ............................................................... 118

4.4 Summary of room-temperature pressurization tests conducted using HighPressure Test Facility on tubes with axial EDM notches to determine effect of

internal Tygon tube bladder and 0.13-mm-thick brass backup foil on failure

pressure....................................................................................................................... 119

4.5 Test results for flawed tubes containing 6.35-mm-long EDM OD axial notches ....... 122

4.6 Pressures at first flaw opening and unstable burst for a nine-tube set of 22.2-

mm-diameter Alloy 600 tubes containing pairs of 6.35- or 12.7-mm-long axial

EDM notches of 70 and 80% 'W separated by various lengths of

full-wall-thickness ligaments...................................................................................... 124

4.7 Sustained pressure and leak rate during Phase 1 testing after flaw opening of

two paired 6.35-mm-long axial notches...................................................................... 126

4.8 Opening pressures of six flaws as a function of separating ligament size for

paired 6.35- and 12.7-mm-long interacting circumferentially shifted axial

notches, all 80% TW ............................................................... 127

4.9 Test results from Stage 1 quasi-steady-state testing of flawed tubes containing

OD laser-cut notches................................................................................................... 130

4.10 Test results from Stage 2 testing of flawed tubes containing OD laser-cut

notches......................................................................................................................... 135

4.11 Flaw opening widths and internal pressures for PSI tubes . 136

xvi
4.12 Summaxy of results from tests on four Argonne tubes and Westinghouse tube

with axial ODSCC cracks............................................................................................. 138

4.13 Results from pressure tests on 14 SCC flaws ............................................................. 140

4.14 Results from pressure and leak-rate tests on tubes from McGuire SG ......... ............ 142

4.15 Erosion data from jet-impact erosion tests utilizing jets from various sizes of

EDM holes Impacting Alloy 600 erosion coupons positioned 6.35-mm from the

hole exit at 2820C and room temperature ................................................................... 145

4.16 Erosion data from jet-impact erosion tests utilizing jets from OAO-mm-diameter

EDM holes Impacting Alloy 600 erosion coupons positioned 6.35-mm from the

hole exit for various water temperatures ................................................................... 146

4.17 Erosion data from jet-impact erosion tests utilizing jets from 0.79-mm-diameter

circular EDM holes impacting Alloy 600 erosion coupons positioned 6.35-mm

from the hole exit for various water temperatures. ...................................... 146

4.18. Room-temperature failure pressures and axial tensile properties of unflawed

Alloy 600 tubing tested by ANL and PNNL .................................................................. 149

XVII
Executive Summary

Assessment of Inspection Reliability

The development and implementation of two key concepts, condition monitoring and

operational assessment, is a major outcome of regulatory activity over the past ten years. That

effort was intended to develop guidance for tube integrity assessments. Condition monitoring

involves an assessment of the current state of the steam generator (SG) relative to the

performance criteria for structural integrity. An operational assessment is an attempt to

assess the state of the generator relative to the structural integrity performance criteria at the

end of the next inspection cycle. Predictions of the operational assessment from the previous

cycle can be compared with the results of the condition monitoring assessment to verify the

adequacy of the methods and data used to perform the operational assessment.

The reliability of nondestructive evaluation (NDE) techniques used to determine the flaw

distribution in terms of detection and characterization of flaws and the capability to assess

their impacts on the structural integrity (i.e., structural and leakage integrity) of SG tubes is a

key factor in establishing the reliability of operational assessments and condition monitoring.

An NDE round-robin exercise has been used to independently assess SG inspection reliability.

The NDE round-robin exerecise was conducted using a steam generator mock-up at Argonne

National Laboratory (ANL).

The steam generator mock-up consists of 400 Alloy 600 tubes made up of nine test

sections, each 0.3 m (1 ftJ long. The test sections are arranged in nine levels, each having

400 tube sections. The lowest level simulates the tube sheet, while three other levels simulate

tube support plate (TSP) intersections. The remaining five levels are free-span regions. Tubes

rolled into ferritic steel collars simulate the tube sheet geometry. Thus, both the roll transition

geometry and the effect of the ferritic tube sheet are simulated. In the TSP crevice, the

presence of magnetite was simulated by filling the crevice with magnetic tape or a

ferromagnetic fluid. A mixture of magnetite and copper powder in an epoxy binder simulated

sludge deposits.

The mock-up contains hundreds of cracks and simulations of artifacts, such as corrosion

deposits, support structures, and tube geometry variations that generally make the detection

and characterization of cracks more difficult. An expert NDE Task Group from ISI vendors,

utilities, ANL, the Electric Power Research Institute (EPRi), and the U.S. Nuclear Regulatory

Commission (NRC) has reviewed the eddy current signals from laboratory-grown cracks used in

the mock-up to ensure that they provide a realistic simulation of cracks obtained in the field.

The number of tubes inspected and the number of teams participating in the round-were

selected to provide better statistical data on the probability of detection (POD) and

characterization accuracy than is currently available from industry performance demonstration

programs.

The reference state for each flaw in the mock-up, i.e., crack geometry and size, was

established by using an ANL-developed algorithm for multiparameter eddy current (EC) data

analysis. Both pre- and post-assembly inspection results were used for this purpose.

Throughout the development stage of the algorithm, the NDE predictions were compared with

results obtained by destructive analyses for dozens of flaws. A final validation was performed

xHill
by comparing the NDE results to destructive analyses in a blind test on a set of 23 flawed

specimens. The results from this comparison were used to estimate the uncertainties

associated with the depth estimates from the multiparameter algorithm. Further validation

was carried out by destructive examination of selected tubes removed from the mock-up and

tubes removed from steam generators from McGuire Unit One.

Eleven teams participated In the analysis round-robin. Each team provided nine reports:

a primary, a secondary, and a resolution analysts' report for each of the three optical data

disks containing the inspection results Ibobbin coil (BC) data for all tubes, rotating pancake

coil (RPC) data for all tube-sheet test sections, and RPC data for a set of selected test sections].

Results were analyzed for all teams. Team-to-team variation in the POD presented, along with

the population average. The detection results for the 11 teams were used to develop POD

curves as a function of maximum depth and the parameter mp, which can be interpreted as a

stress multiplier that relates the stress in the ligament ahead of the crack to the stress in an

unflawed tube under the same loading. Because mp incorporates the effect of both crack depth

and length, it better characterizes the effect of a flaw on the strength (i.e., structural and

leakage integrity) of a tube than do traditional indicators such as maximum depth. The effect

of both statistical uncertainties inherent in sampling from distributions and the uncertainties

due to errors in the estimates of maximum depth and mp were estimated. The 95% one-sided

confidence limits, which include errors in maximum depth estimates, are presented along with

the POD curves.

Deplugged tubes and tube sheet sections removed from retired steam generators at

McGuire Unit One were inspected before removal using +Point, '3-coir pancake, and bobbin

coil probes. The results were compared with results from previous inspections of the tubes. In

all cases the BC voltages increased over time. The most likely cause of the voltage change is a

change in the character of the corrosion deposits on the cracks, but other factors that could

result in voltage increases include crack growth and a change in crack width.

Several flawed tubes from McGuire have been reexamined at Argonne in a glovebox

facility that can handle contaminated tubes. In general, the voltages measured at Argonne are

slightly lower than those measured before removal, with the phase essentially unchanged,

except for one flaw with a particularly large voltage. The observed differences could be the

result of variations in the mix and changes in conduction paths across the crack faces

resulting from the tube pulling procedure (i.e., crack faces now touching.

Research on Inservice Inspection Technology

Research on improved ISI technology of steam generator tubes focused on four areas:

evaluation of analytical methods for the prediction of EC responses as a function of probe

design, flaw characteristics, and material properties; assessment of effective signal analysis

procedures; assessment of flaw imaging and display methods for simple and accurate flaw

characterization; and evaluation of improved probe designs that use directional arrays so that

axial and circumferential defects can be examined simultaneously. The -reliability and

effectiveness of improved Inspection techniques were evaluated through laboratory testing of

SG tubes that contain various flaws and the SG tube bundle mock-up. Some validation efforts

utilized the In-service-degraded SG tubing from McGuire.

xix
Several analytical and computational approaches to electromagnetic modeling for

simulating EC probe response in the NDE of SG tubing have been examined. Most of the

efforts focused on the applications of a general-purpose 3-D FEM-based code for steady-state

and transient-field calculations.

In conventional EC data analysis of SG tubing, human analysts are trained to

differentiate flaw-induced signals from artifacts by analysis of the behavior of signals, typically

extracted from their Impedance plane trajectory and signal amplitude, through application of a

series of rules. For example, clockwise (CW) and counter-clockwise (CCW) rotation of the

impedance plane signal trajectory is commonly used to determine the flaw origin. When the

phase angles at all frequencies from a TW flaw are set to a fixed value, flaw indications of OD

origin exhibit CCW rotations, and those of ID origin have smaller CW rotations.

Manual analysis of multiple-frequency EC data is a tedious and challenging process. No

qualified technique. manual or automated, currently exists that could provide reliable

estimation of flaw size over a wide range of SG tubing damage. Conventional data analysis

methods become rather subjective when dealing with complex forms of degradation such as

SCC. Signal distortion by interference from internal/external artifacts in the vicinity of a flaw

further complicates discrimination of flaw signals from noise. Analogous to the manual

analysis of EC inspection data, computer-aided data analysis algorithms can be developed

from analysis of the characteristic behavior of EC signals as a function of frequency. Because

all available information can be examined rather than a selected subset in the case of a manual

analysis, such algorithms allow more effective identification of subtle forms of degradation.

Computer-based algorithms that imitate some form of human decision making are generally

labeled as expert systems.

Multiparameter algorithms for the computer-aided analysis of EC NDE data have been

developed and implemented in a set of data analysis and signal processing algorithms that

have been written using the PC-based software MATLAB. a high-level scripting language that

provides an efficient environment for data manipulation and computation, together with

convenient graphical user interfaces (GUls) and graphical displays of the results. Phase angle

information from multiple channels is used to construct a model that correlates the NDE

results to flaw size and origin. The flaw size and origin are assumed to be related to the phase

angle information at the different frequencies. Eddy current measurements on a calibration

standard tube are used to determine the regression coefficients in the relation. The

effectiveness of the multiparameter algorithms for the computer-aided analysis of EC NDE data

has been evaluated in studies on electrodischarge machined (EDM) notches, laboratory-grown

flaws, and in-service flaws.

Research on Degradation Modes and Integrity

Because degraded tubing from operating or retired SGs is difficult and expensive to

obtain and the availability of such tubing is limited, it was necessary to produce prototypical

degradation in tubes as a part of the program. The degraded tubes were used in the SG

mock-up and for pressure and leak-rate testing. Techniques have been developed to produce

SCC degradation of various sizes, depths, and orientations that are prototypical of

fleld-degraded SG tubes. During the course of the program, a total of about nine-hundred

22.2-mm (7 /87in.)-diameter Alloy 600 tubes were processed to produce axial and

circumferential outer-diameter stress corrosion cracking (ODSCC). axial and circumferential

'DC
inner-diameter stress corrosion cracking (lDSCC), and SCC in skewed orientations. Some of

the specimens were produced with multiple cracks or combination of multiple cracks of

different orientations. Most of the specimens were straight tubes, but some were

roll-expanded, dented, or ovalized.

One of the phenomena considered in NUREG-1570 was the potential for the propagation

of tube failures during severe accidents. In the event of a tube failure, the escaping jet of

superheated steam could entrain particles from the core debris and impinge on a neighboring

tube. The particle-laden jet could then erode through the neighboring tube, leading to another

escaping jet and the possibility of a cascading failure event in the steam generator. Some

analyses were reported in NUREG-1570 to estimate the time required for a jet from a failed

tube to induce failure in an adjacent tube. It was noted that the analyses were at best scoping

estimates. An investigation was carried out in the present program to analyze the potential for

such a failure in more detail. The effort included high-temperature erosion tests conducted at

the University of Cincinnati, a computational fluid dynamics analysis of the escaping jet, and

development of a model for calculating the crack opening area as a function of time. The

results showed that, even after some conservative assumptions are made, failure of an adjacent

tube would take >10 h. Thus, jet impingement is very unlikely to play any significant role In

the failures.

Under design-basis conditions, two basic erosion damage mechanisms, droplet impact

and cavitation, are believed possible from fluid jets associated with flawed SG tubes. In the

case of jets consisting of water droplets, the impact of the droplet gives rise to fluctuating

stresses in the target, which produce fatigue damage. In the case of cavitation, bubbles form

in the liquid stream and then collapse on impact, again producing fluctuating stresses and

fatigue in the target.

Tests has been conduced to determine the susceptibility of steam generator tubes to

erosive damage from impacting jets of superheated water or steam leaking from adjacent tubes.

The tests show that the impact erosion depth varies with water temperature, and the erosion

rate is substantially reduced with decreasing and increasing subcooling on either side of the

peak damage temperature range. The peak erosion rates occur for degrees of subcooling

typical of the cold leg, though cracking is much less likely to occur here. Erosion rates are

much lower for conditions more typical of the hot leg. Even in the case of the cold leg, where

the erosion rates are higher, the likelihood of failure propagation by jet erosion appears to be

very low.

Analysis and testing have been performed to develop models for predicting failure

pressures and leak rates in steam generator tubes with axial and circumferential cracks under

normal operation and design-basis and severe accident conditions. The models can account

for effects of the constraint exerted by support structures such as the tube support plates. In

addition to models for the unstable failure (burst) of the tube, extensive tests on rectangular as

well as trapezoidal and triangular part-throughwall flaws were conducted to derive a

correlation for predicting the failure of radial ligaments in tubes with axial notches, either

shallow or deep.

A procedure has been developed for defining an equivalent rectangular crack for an

arbitrarily shaped notch or crack. Structural and leak rate analyses of the flaws and

comparison with experimental data have shown that the radial ligament failure pressure in

xod
nonrectangular notches or SCCs can be predicted reasonably well by using a flow stress model

that Is strictly applicable to rectangular part-throughwall cracks and replacing the actual crack

profile by an equivalent rectangular crack. An empirical approach for calculating crack growth

by stable ligament rupture has been developed by an extension of the equivalent rectangular

crack method. The approach has been used to calculate the incremental ligament rupture of

non-rectangular notches as well as stress corrosion cracks.

Interaction effects between two or more throughwall notches as influenced by ligaments

separating the notches are analyzed by finite element analyses. Interaction effects (with

respect to radial ligament rupture) between two or more part-throughwall notches can be

estimated by the equivalent rectangular crack approach.

The calculated crack opening displacements (areas) for rectangular and trapezoidal

throughwall notches were in reasonable agreement with measured values. Leak rates for tests

on specimens with EDM notches or open cracks can be predicted reasonably well by using the

well-known formula for single-phase flow through smooth circular orifices and by replacing the

orifice area by the crack opening area. In tests conducted to date, the equations can predict

the ligament rupture pressures and leak rates for rectangular and trapezoidal EDM notches at

both room temperature and 2820C if the appropriate densities of water are used. The

correlation will break down at lower levels of leak rates for very tight stress corrosion cracks,

particularly at elevated temperatures where flashing may occur inside the cracks; but such a

limit has not yet been established.

A creep rupture-based criterion was developed for predicting radial ligament rupture

during severe accident conditions. The ligament rupture criteria for part-throughwall cracks

were validated with SG tubes having rectangular EDM notches in tests conducted at room

temperature, 2820C, and >6000C.

The behavior of flawed tubes with ElectrosleeveTM repairs during severe accidents was

examined. Although the nanocrystalline ElectrosleeveTM has superior mechanical properties

compared to Alloy 600 at reactor operating temperatures, It loses strength at high

temperatures because of grain growth. A flow-stress based ligament rupture criterion was

developed for predicting failure of Electrosleeved tubes with 80-100% deep throughwall cracks

in the parent tubes and was validated with high-temperature tests on EDM-notched

Electrosleeved tubes supplied by Framtome Technologies, Inc.

Integrated Model for Steam Generator Integrity Assessments

The models for predicting radial ligament rupture, unstable burst, and leak rate of flawed

Alloy 600 tubes developed in the program have been incorporated into the integrity assessment

code for SG tubes, CANI1A. which was originally developed by Dominion Engineering for the

Canadian Nuclear Safety Commission. Several other modifications were also made in

ANL/CANTIA. The source language of the code was updated to Visual BASIC 6. The treatment

of the basic flaw was changed from one- to two-dimensional, and the calculation routines for

crack growth were made more rigorous. Two well-known SCC growth models from the

literature were Included as options, and a simple residual stress model was added.

ANL/CANTIA can handle either axial or circumferential cracks, but not both simultaneously.

Ddi
Preliminary trial runs demonstrated that it is very difficult to grow stress corrosion cracks

in SG tubes by pressure stress alone. Several trial runs with residual stress were conducted to

demonstrate the importance of POD and initiation of new cracks on rupture and leak rate

probabilities.

oDiii
Acknowledgments

The authors thank C. Vulyak and L. Knoblich for their contributions to the experimental

effort and P. Heasler and R. Kurtz for discussions related to the statistical analysis. The

authors also thank the NDE Task Group Members, G. Henry and J. Benson (EPRI), T. Richards

and R Miranda (Framatome Technology), D. Adamonis and R. Maurer (Westinghouse).

D. Mayes (Duke Engineering and Services), S. Redner (Northern States Power), and B. Vollmer

and N. Farenbaugh (Zetec, Inc.). They also thank H. Houserman and H. Smith for their input

in this effort. The authors acknowledge the contributions of C. Gortemliler, C. Smith,

S. Taylor, and the staff from Zetec Inc. to the data acquisition and analysis effort. The authors

also thank proctors S. Gopalsami, K. Uherka, and M. Petri, as well as ABB-CE, Anatec, DE&S,

FTI, KAITEC, Ontario Power Generation, Westinghouse, and Zetec for providing round-robin

analysis teams. This work is sponsored by the Office of Nuclear Regulatory Research of the

U.S. Nuclear Regulatory Commission under Job Code W6487. The Program Manager is Dr. J.

Muscara, who provided helpful guidance in the performance of this work.

xxiv
Acronyms and Abbreviations

ABB-CE ASEA Brown-Boveri-Combustion Engineering

AECL Atomic Energy of Canada, Ltd.

ANL Argonne National Laboratory

APC alternative plugging criterion

ARC alternative repair criterion

ASME American Society of Mechanical Engineers

BC bobbin coil

CCW counter clockwise

CIDSCC circumferential Inner-diameter stress corrosion crack/cracking

CODSCC circumferential outer-diameter stress corrosion crack/cracking

COD crack opening displacement

CW closckwlse

DE&S Duke Engineering and Services

EC eddy current

ECT eddy current testing

EDM electrodlscharge machining

EFPY equivalent full power years

EM electromagnetic

EPRI Electric Power Research Institute

FEA fintie element analysis

FEM finite element method

FIR finite-impulse response

FS freespan

FTI Framatome Technology

ID inner diameter

IGA intergranular attack

INEEL Idaho National Engineering and Environmental Laboratory

ISI inservice inspection

LIDSCC longitudinal inner-diameter stress corrosion crack/cracking

LODSCC longitudinal outer-diameter stress corrosion crack/cracking

LS least squares

MLR multiple linear regression

MRPC motorized rotating pancake coil

MSLB main steam line break

NDE nondestructive evaluation

NDD nondetectable degradation

NRC U.S. Nuclear Regulatory Commission

OD outer diameter

OPG Ontario Power Generation

OSL One-sided 95% confidence limits

PCR principal component regression

PLS partial least squares

PNNL Pacific Northwest National Laboratory

POD probability of detection

PSI Paul Scherrer Institute

PWR pressurized water reactor

ocv
PWSCC prlmary-water stress corrosion crack/cracking

RPC rotating pancake coil

RR round-robin

RTZ roll transition zone

SCC stress corrosion crack/cracking

SG steam generator

S/N signal to noise ratio

SS stainless steel

SVD singular value decomposition

TS tube sheet

TSP tube support plate

ITS top of tube sheet

'1W throughwall

UT ultrasonic testing

Vpp peak-to-peak voltage

W Westinghouse
1 Introduction

The objective of this program is to provide the experimental data and predictive

correlations and models needed to permit the U.S. Nuclear Regulatory Commission (NRC) to

independently evaluate the integrity of steam generator (SG) tubes as plants age and

degradation proceeds, new forms of degradation appear, and new defect-specific management

schemes are implemented. The areas addressed by the program include assessment of

procedures and equipment used for inservice inspection (ISI) of SG tubes; and

recommendations for criteria and requirements to improve the reliability and accuracy of

ISI; validation and improvement of correlations and models that are used to evaluate integrity

and leakage of degraded SG tubes; and validation and improvement of correlations and models

that predict the generation and progression of degradation in SG tubes as a function of aging.

This program is divided into four technical tasks and one management task: Assessment

of Inspection Reliability; Research on ISI Technology; Research on Degradation Modes and

Integrity: Integration of Results. Methodology, and Technical Assessments for Current and

Emerging Regulatozy Issues; and Program Management.

This report summarizes the results and accomplishments over the duration of this program

from its inception in August 1995 through its conclusion in December 2001.

1
2. Assessment of Inspection Reliability (D. S. Kupperman and

S. Bakhtiari)

2.1 Steam Generator Mock-Up Round Robin

2.1.1 Introduction

One major outcome of regulatory activity over the past 10 years intended to develop

guidance for tube integrity assessments Is the development and implementation of two key

concepts: condition monitoring and operational assessment. Condition monitoring involves an

assessment of the current state of the SG relative to the performance criteria of structural

integrity. An operational assessment is an attempt to assess what the SG state will be relative

to the structural integrity performance criteria at the end of the next inspection cycle. The

predictions of the operational assessment from the previous cycle can be compared with the

results of the condition monitoring assessment to verify the adequacy of the methods and data

used to perform the operational assessment. The reliability of the in-service inspection (ISI) is

critical to the effectiveness of the assessment processes. Quantitative information on

probability of detection (POD) and sizing accuracy of current-day flaws for techniques used for

SG tubes is needed to determine if tube-integrity performance criteria were met during the last

operating cycle, and if performance criteria for SG tube integrity will continue to be met until

the next scheduled ISL. Information on Inspection reliability will permit estimation of the true

state of SG tubes after an ISI by including the flaws that were missed because of imperfect

POD. Similarly, knowledge of sizing accuracy will permit corrections to be made to flaw sizes

obtained from ISL.

Eddy-current (EC) inspection techniques are the primary means of ISI for assessing the

condition of SG tubes in current use. Detection of flaws by EC depends on detecting the

changes in impedance produced by the flaw. Although the impedance changes are small

(=10 6), they are readily detected by modem electronic instrumentation. However, many other

variables, including tube material properties, tube geometry, and degradation morphology, can

produce impedance changes, and the accuracy of distinguishing between the changes

produced by such artifacts and those produced by flaws is strongly influenced by EC data

analysis and acquisition practices (including human factors). Similarly, although it can be

shown that the depth of a defect into the tube wall is related to the EC signal phase response,

in practice, those factors that affect detection also affect sizing capability.

The most desirable approach to establishing the reliability of current ISI methods is to

carry out round-robin (RR) exercises in the field on either operating SGs or those removed from

service. However, access to such facilities and validation of the results would difficult. Such

work would also be prohibitively expensive. In addition, obtaining data on all morphologles of

interest would require tubes from many different plants.

The approach chosen for this program was to develop an SG tube bundle mock-up that

simulates the key features of an operating SG so that the inspection results from the mock-up

would be representative of those for operating SGs. Considerable effort was expended in

preparing realistic flaws and verifying that their EC signals and morphologies are

representative of those from operating SGs. The mock-up includes stress corrosion cracks of

different orientations and morphologles at various locations in the mock-up and simulates the

2
artifacts and support structures that may affect the EC signals. Factors that influence

detection of flaws include probe wear, eddy current signal noise, signal-to-noise ratio, analyst

fatigue, and the subjective nature of interpreting complex eddy current signals. In this

exercise, all analysts examine the same data provided on optical disks that contain copies of

the data to be analyzed. The team-to-team variation in detection capability is the result of

analyst variability in interpretation of eddy current signals. The mock-up will also be used as a

test bed for evaluating emerging technologies for the ISI of SG tubes.

2.1.2 Steam Generator Mock-Up Facility

The mock-up tube bundle consists of four-hundred Alloy 600 tubes (22.2-mm j0.875-in.1-

diameter) consisting of 9 test sections, each 0.3 m (1 ft) long. Test sections are arranged in

nine levels, with 400 test sections at any elevation. The centers of the tubes are separated by

3.25 cm (1.28 in.). Me rods hold the test sections together. The ends of each test section are

pressed into 19-mm (0.75-in.)-thick high-density polyethylene plates that hold the test sections

in alignment. One end of each tube is spring-loaded. The lowest level (A) has a roll transition

zone (RIZ) and simulates the tube sheet, while the 4th, 7th, and 9th levels simulate drilled hole

tube support plate (TSP) intersections. The other five levels are freespan regions. Above

the 9th level is a 0.91-m (3-ft)-long probe run-out section (see Fig. 2.1 for the tube bundle

diagram and Fig. 2.2 for photograph of the mock-up).

Most of the degraded test sections were produced at ANL, although some were produced

by Pacific Northwest National Laboratory (PNNL); Westinghouse; Equipos Nucleares, SA

(ENSA); and the Program for the Inspection of Steel Components (PISC).

Alloy 600 test sections at ANL were cracked by using a 1 M aqueous solution of sodium

tetrathionate at room temperature and atmospheric pressure. Localized environmental

exposure, low applied load, and electrochemical potential were utilized to produce various

crack geometries. Masking by coating areas of the tubes with lacquer was used to limit or

localize the cracking area. The tubes were internally pressurized to generate hoop stresses to

produce axial cracks and axially loaded to produce circumferential cracks. Various OD and I)

crack geometries were produced: axial, circumferential, skewed, or combinations of these.

Many of the specimens contained multiple cracks separated by short axial or circumferential

ligaments. Prior to exposure to the sodium tetrathionate solution, specimens were sensitized

by heat-treating at 600'C (11 12°F) for 48 h to produce a microstructure that is susceptible to

cracking. The mock-up was seeded with sensitized flaw-free test sections with and without

artifacts so that the possibility of distinguishing sensitized from unsensitized test sections

would not be an indicator that a flaw was present in that test section. In addition, many

cracks from Westinghouse were grown without sensitizing the test sections.

Dye-penetrant examination (see Fig. 2.3 for example), as well as destructive

metallography, were performed on many samples to provide documentation of the mock-up

crack morphology. Figure 2.4 shows examples of LODSCC. Cracks in the mock-up provided

by PNNL (about 50) were produced by Westinghouse using a proprietary doped steam method.

Axial and circumferential cracks, both ID and OD, were produced for the freespan, TSP, and

roll transitions. Several IGA specimens, as well as fatigue and wastage samples, were also

provided by PNNL.

3
Figure 2.1. Schematic representation of tube bundle in steam generator mock-up.

Table 2.1 shows the distribution of flaw types in the mock-up. The flaw depths are

distributed into three ranges, 0-40%TW, 41%-80%TIW, and 81-100%1W. The distribution is

skewed toward deeper cracks. This skewing is necessary to obtain high confidence in the high

POD for the deeper cracks. Draft Regulatory Guide 1074 ("Steam Generator Tube Integrity)

describes criteria for performance demonstrations to quantify defect detection performance

(POD for a given defect). While the distribution of flaw sizes for the round-robin is not as

uniform as required in Draft Regulatory Guide 1074, other requirements involving extraneous

signals, signals from fabricated defects, and detection and false calls have, for the most part.

been met.

4
Ii;

Figure 2.2. Photograph of mock-up during acquisition of eddy current data.

5
Figure 2.3.

Dye-penetrant examination of tube

specimen SGL865 showing an

LODSCC.

Figure 2.4 Cross-sectional optical metallography: (a) branched LODSCC and (b) LODSCC.

In the tube support plate ([SP) regions, filling the crevice with magnetic tape or a

ferromagnetic fluid simulated magnetite in the crevices. A mixture of magnetite and copper

bonded with epoxy was used to simulate sludge deposits. Simulated sludge was placed above

the RTZ and at some TSP intersections. Many test sections had sludge or magnetite but no

flaws. LODSCC and LIDSCC, both planar and segmented, and are present at TSP locations

with and without denting. Some flaw-free test sections were dented. Cracks in the freespan

are primarily LODSCC, both planar and segmented. Axial and circumferential ID and OD

cracks are found in the RIZ. A small number of other flaw types such as IGA and wear were

placed in the tube bundle. The mock-up also contains test sections with electric-dischargemachined (EDM) notches and laser-cut slots. Table 2.2 summarizes the degradation types and

their locations in the mock-up.

6
Table 2.1. Distrbution of flaw types

Max. EDM & ODSCC PWSCC Wear/

Depth Laser- ODSCC ODSCC Free- PWSCC PWSCC Free- WastRange Cut Slots IGA 7S TSP span TDS TSP span age Fatigue

0-40%TW 7 2 3 14 15 4 8 1 6 0

41-8096 W 13 9 2 14 26 8 16 0 6 0

81-100%TW 1 2 16 41 49 35 7 3 0 3

Table 2.2. Flaw types and quantity

EDM &

Laser-Cut Wear/

Location Slots IGA ODSCC PWSCC Wastage Fatigue

TopofTubeSheet - 21 47 - -

Freespan 14 8 90 4 3 -

TSPs 7 5 69 31 9 3

Magnetite-filled epoxy markers placed at the ends of all test sections provide a reference

for the angular location of flaws when collecting data with a rotating or array probe. The

analysts were instructed to Ignore the region 25 mm (1 In.) from each test section end.

Prior to assembly, flawed test sections in the tube bundle were examined with both

bobbin coil (BC) and a three-coil rotating probe that Incorporates a +Point coil, a 2.9-mm

(0.115-in.)-diameter pancake coil, and a 2-mm (0.080-inj-dlameter shielded pancake coil. In

addition to a full EC examination, many cracked test sections were examined by dye-penetrant

before being Incorporated Into the mock-up tube bundle. If EC data, dye penetrant results, or

crack growth parameters indicated that a crack must be present, the test section was included

in the mock-up. Because primary interest is with deep flaws, the majority of cracks selected

for the mock-up had a +Point phase angle consistent with deep (> 60% 1W) cracks. Because of

the large fraction of deep flaws, high voltage signals are more common in the mock-up than in

an actual steam generator.

The flaw types selected for the mock-up are those currently found in operating steam

generators. Since about 1980, SG tube degradation has been dominated by SCC, which can

occur on either the primary or secondary side, unlike the wastage and denting that occur

exclusively on the secondary side (OD) of the tubes. Primary water stress corrosion cracking

(PWSCC) Is most likely to occur at regions of high residual stress, such as at the tube

expansion transition and immediately above the tube sheet, at U-bends (particularly the smallradius U-bends), and in tube regions deformed by secondary-side denting. As a result, the

mock-up consists primarily of ID and OD SCC at the TSP (with and without dents), at and

above the roll transitions, as well as in the freespan. Outer-diameter intergranular attack (IGA)

commonly occurs in crevices or under corrosion product scales. Such locations include the

TSP crevice, the region near the top of the tube sheet, freespan areas under corrosion products

or deposits, and regions under sludge buildup. As a result, some outer-diameter IGA is

7
present in the mock-up. In addition, there are some fatigue cracks, some test sections with

wastage, and some with wear.

The distribution of voltages in the BC data from the mock-up was analyzed. The

histograms show a reasonable distribution of BC voltages (up to 20 V) for cracks and

conditions, and cracks alone. The voltages and phases for mock-up cracks are similar to those

observed in field data such as from the McGuire Nuclear Plant.

To provide additional validation of the similarity of the mock up specimens to cracks in

operating SGs, pulled tubes from a retired McGuire steam generator have been inspected at

Argonne in a large glovebox modified for NDE studies. The NDE glovebox allows for the EC

inspection of radioactive test sections, including samples with loose radioactive contamination.

Figure 2.5 shows the BC UIssajous figure for an LODSCC from a dented TSP region of the

McGuire steam generator. Two standards (one an ASME standard, the other an 18 EDM notch

standard) were in line with the McGuire test section during the inspection. The BC data from

the standards are seen in the linear traces on the left side of Fig. 2.5. Figure 2.6 shows a

comparable BC Lissajous flgure from an LODSCC in the mock-up. The similarity of the BC

signals for the two LODSCC provides evidence of the ability to grow, under laboratory

conditions, SCC that mimics the EC signals of field flaws.

As part of the development of the multiparameter algorithm, results have been compared

to fractographic results on a wide variety of SCC cracks and EDM and laser-cut notches. To

provide an objective benchmark, however, an additional set of 29 SCC cracks was produced

and used in a blind test of the predictions of the algorithm against fractographic measurements

of the crack geometry. Six of the benchmark samples have not yet been destructively analyzed

because they will also be used for leak and ligament rupture tests that have not yet been

performed.

The stress corrosion cracks for the blind test were produced by the same technique in

1 M aqueous solutions of sodium tetrathionate as described above. A variety of OD and ID

crack geometries was produced: axial, circumferential, skewed, and combinations of these.

Many of the specimens contained multiple cracks separated by short axial or circumferential

ligaments. Cracked tubes were examined by dye-penetrant techniques. conventional eddy

current NDE, the multiparameter algorithm, and destructive methods.

Data collection procedures for multifrequency inspection of the mock-up tubes are

described in Ref. 3. These guidelines define the instrumentation setup (coil excitation

frequencies, gain setting, cable length, sampling rate, probe speed, etc.) and calibration

procedures for a given probe (e.g., bobbin, rotating, and array probes). Although the quality of

data affects both detection and sizing, this Issue is of particular concern when quantitative

estimates of flaw size are to be determined. The multiparameter algorithm used to obtain flaw

size estimates for the mock-up requires data at three frequencies and the minimum sampling

rate recommended in the Examination Technique Specification Sheet (ETSS) for MRPC probes

131. The multiple-frequency EC data were acquired with a standard three-coil rotating probe

that incorporates a 2.92-mm (0.1 15-in.)-diameter mid-range primary pancake coil, a mid-range

+PointTm, and 2.03-mm (0.080-in.)-diameter high-frequency pancake coil. Initial amplitude

proffles are obtained from the +Point coil at a single channel. The final estimated depth profiles

are obtained by using multichannel information from the mid-range primary pancake coil for

8
Figure 2.5. L _ W Differential bobbin coil Lissajous figure at 400 kHz from crack in McGuire SG

: _ _ . 9 . (LODSCC7243). EC data were taken

from a McGuire pulled tube using

Argonne's NDE glove-box facility.

. '-- l 

_____ X - - _ ._ ,~ ___

Figure 2.6.

: / - . R . . Differential bobbin coil Lissajous figure

- - _ l at 400 kHz from Argonne-grown mock-

_ . _ _ . - up crack (LODSCC300). The BC

signal shape, amplitude, and phase of

the McGuire and Argonne SCC are

similar.

multiparameter data analysis. The algorithm and the data quality issue are discussed in detail

in Ref. 2, which also describes the conversion of Eddynet formatted data to a standard format

for off-line analysis.

For the destructive examination, the samples were heat-tinted before fracture to permit

differentiation of the SCC and fracture opening surfaces. The specimens were then chilled In

liquid nitrogen, and cracks were opened by fracture. The fracture surfaces were examined

macroscopically and with optical and scanning electron microscopy. The fractography and

NDE data were digitized to obtain tabular and graphical comparisons of the depths as a

function of axial or circumferential position. Well-defined markers on the test sections

provided a means to accurately overlap the profiles. The fractured surfaces were recorded by

digital photography at magnifications of 0.2-IOX (Fig. 2.7).

9
Figure 2.7. Fractography of tube specimen SGL-413.

Crack profiles were obtained from digitized photographs of the fracture surfaces. The

sampling distance depends on the complexity of the crack geometry. Short sampling distances

were used for complicated geometries over a small scale, while longer distances were used for

simpler geometries, e.g., straightline or smooth contours. Fractography and NDE results were

plotted in a same figure for comparison (e.g., Fig. 2.8). Drawing lines through the EC data

points generates the NDE profiles (nominally 12 points per centimeter 130 per inch] around the

circumference and 12 points per centimeter [30 per inch] axially). The NDE and fractography

profiles were then compared at many axial and circumferential positions, and the differences

were used to establish the NDE uncertainty as a function of depth. The NDE uncertainties

were then used to generate the lower 95% confidence limits for the POD curves presented in

this report.

In the development of the multiparameter algorithm. the results from the algorithm have

been compared to fractographic results on a wide variety of SCCs and EDM and laser-cut

notches. To provide an objective benchmark, however, additional SCCs were produced and

used for a blind test of the predictions of the algorithm against fractographic measurements of

the crack geometry. Crack profiles from the destructive analyses are compared with those

obtained from the multiparameter algorithm in Refs. 1-3. Figure 2.9 the maximum depths as

determined by both fractography and by the multiparameter algorithm. A linear regression fit

and 95% confidence bounds for the observed data as a function of the multiparameter

estimates are shown in the figure. The overall root-mean-square error (RMSE) in the predicted

maximum depths is 13.7%. If the comparison is limited to deeper cracks, the RMSE is smaller,

namely 9.7% for depths 30-100% TW and 8.2% for ODSCC of depths 50-100% lW. The data

are too few, however, to determine whether the apparent variation of the RMSE with depth Is

statistically significantBecause the field of view of the rotating pancake probe is limited, the depth

measurements at points 25 mm (?0.2 in.) apart along the crack profile are essentially

independent, and additional comparisons of the estimated depths with those determined by

fractography were made at various points along the crack profile. To avoid observer bias in the

selection of the data for comparison, the intersections of the crack profiles with the major

10
120.00

100.00 _

0.00 Figure 2.8. F l: E- Sizes and shapes of LODSCCs In tube

6B.00 specimen AGL-536 determined by EC

NDE using the multiparameter

40oo.00f / \|algorithm (dotted curve) and

J / ; fractography (smooth curve).

20.00 20.00 - # I . ,

-2 0 2 4 6 8 10 12

Crack Length (mm)

100 ,,I|I|I'.- ' e

so a

*0 I

0.00~~

R 60 - 0' 6 °" Figure 2.9.

S~~- - . - -o .Comparison of maximum depth f ° F / - - ~~~~~~determined by the multiparameter

40 - algorithm with that determined by

20 -k -, - \ epth m21.8+0.79.depthfractography.

20 _ , dePt b 21.8 + 0.79-depth.

O _ I I .I i I_

0 20 40 60 80 100

gridlines in the graphs of the superimposed profiles were chosen as the points for comparison.

This corresponds. in most cases, to a spacing of 5-10 mm (0.2-0.4 in.) between points.

Figures 2.10 and 2.11 show the results for 89 points from 20 different cracks, axial and

circumferential, ID and OD. A linear regression curve and 95% confidence bounds for the

observed data as a function of the multiparameter estimates are shown. The intercept in

Fig. 2.10 Is 13.a somewhat less than that generated from the maximum depth data (Fig. 2.9),

but the slope of 0.78 is almost identical to the linear regression line slope for the maximum

depth data.

2.1.2.1 Accuracy of Maximum Depth for Mock-Up Cracks

Accuracy In estimating the maximum depth of cracks in the mock-up was determined by

a comparison between crack profiles generated by ANLs multiparameter algorithm and profiles

determined from fractography. The overall RMSE for all cracks of all depths Is 15.1%, but the

RMSE varies significantly with depth. The RMSE value Is significantly better for 80-100% TW

cracks than for cracks with other depths.

11
100

0 00 Figure 2.10.

- 00 Comparison of depths (% TW)

60 _ OX * '°o , . determined by the multiparameter

I -/ - S. algorithm with those by fractography

40 , 8 and regression fit as a function of the

0O a to.,.' multiparameter depth estimate.

*O \ f .-Estimated 95% bounds for the

20 p 0tO ° at- observed depth are also given.

depth Om =13.8+0.78 depth

0 _ 1 ,i. I I I I I I7 0 20 40 60 so 100

Deph,,,t, (%)

100 _ , ' o 1

- 0 0

0 - 0 0

80 . 0 O O >d°° _00 Figure2.11. -o .- /O zComparison of depths (% TW)

_,60 0 0 0 0/ o b determined by the multiparameter

O- 0 0 algorithm with those by fractography.

I K 00 0/O ' O In this case, the observed and

predicted results are shown In terms of

40d 0 O , S the uperfect agreemenr 1:1 line and

estimated 95% bounds that account for

20 P Athe variation in the uncertainty with

0 / / 1: Perfect Agreement depth.

001 L 0I 1 I I I I II II

0 20 40 60 80 100

Depth,,,,, (%)

Two sets of RMSE values are given. One set is based on the values obtained directly from

the multiparameter algorithm, and the other on 'correctedu values obtained from the regression

fit shown in Fig. 2.10. For the shallowest cracks, the 'corrected' values give a significantly

lower RMSE value, but when all the data are considered, the differences in the RMSE for

corrected and uncorrected predictions are small. This finding indicates little systematic bias in

the predictions of the multiparameter algorithm, i.e.. the errors are random.

These sizing-accuracy results can be used to estimate the uncertainty in POD curves if

the multiparameter algorithm is used to determine the 'true' state of the mock-up for the NDE

round-robin. Instead of characterizing the error in the depths in terms of the overall average

for all depths (-15%), the error was taken as a function of depth. The values of RMSE given in

FMg. 2.12 are assumed to apply at the midpoint of the depth range for each bin. The error at

other depths is then estimated by linear interpolation of these values.

12
0.2

0.16

ci5 FIgure 2.12.

X 0.14 Standard deviation In percent

throughwall as a function of

X 0.12 maxdmum depth.

co0.1

0.08 , , I , , I , , , I , , , I , , 0 0.2 0.4 0.6 0.8 1

Depth

A set of 20 test sections analyzed with the multiparameter algorithm Was destructively

analyzed by PNNL In an exercise carried out before the 23 test-section set was evaluated.

While these 20 test sections, which were prepared using doped-steam techniques, represented

a small subset of the mock-up test sections, the NDE results from this set provided guidance

for selecting the multiparameter approach. Many NDE techniques were evaluated before

selecting the multiparameter algorithm for establishing the reference state of the mock-up

flaws. The evaluated techniques include phase analysis of EC +Point data, multivariate

regression analysis of EC data, multiparameter analysis of EC data with neural networks, highfrequency ultrasonics (tI) from the OD, Lamb waves, acoustic microscopy, and a combination

of UT and EC data (from the ID). The multiparameter algorithm provided the best accuracy for

sizing the cracks.

Following the completion of the analysis round robin, several test sections in the mock-up

were removed and destructively analyzed to further help validate the multiparameter algorithm

for profiling and providing maximum depth. The agreement between the results of the

destructive analysis and the multiparameter algorithm was good.

Although the probability of detection is normally expressed in terms of the maximum

depth of the crack, it Is also useful to express POD in terms of a parameter that better

characterizes the structural integrity (i.e., ligament rupture) of the tube. A useful parameter

for this purpose Is mp, which is defined as

mp = b,(1)

PSC

where Pb is the burst pressure of an unflawed tube, and prc is the ligament failure pressure of

a part-throughwall crack. The parameter mp can be interpreted as a stress multiplier that

relates the stress In the ligament ahead of the crack to the stress in an unflawed tube under

the same loading. Because this parameter incorporates the effect of both crack depth and

13
length, it better characterizes the effect of a flaw on the integrity of a tube than does maximum

depth. For short cracks, the ligament may not correspond to tube bursting pressure. The

crack may just 'pop' through without increasing in length.

In Fig. 2.12, the standard deviation in depth (expressed as % TW) is plotted against the

maximum depth. The largest uncertainty is in the 50-70% TW range. While the deviation is

small for the shallow cracks, it represents a relatively large error. For example, the standard

deviation for 20% TW is 12% 'W.

The overall RMSE for all cracks of all depths is 15.1%, but it is somewhat misleading to

look at this overall value because there is a significant variation in the RMSE with depth. The

RMSE value is significantly better for the 80-100% TW bin than for the other depth bins.

The reference-state table for the mock-up provides all the relevant information for

analyzing the round-robin analysts' results. Reference 3 provides the primary information for a

flaw in the table. The flaw type, BC volts, BC phase, whether ID or OD, three-letter code for

the flaw, maximum depth as determined by multiparameter algorithm, flaw length, average

depth, and mp are all included in the flaw table.

2.1.3 Design and Organization of Round-Robin

As noted previously, an expert Task Group was formed to provide input to help ensure

that the RR simulated an ISI inspection. The Task Group included members from utilities,

vendors, EPRI, NRC, and ANL. The industry members are G. Henry and J. Benson (EPRI),

T. Richards and R. Miranda (FI), D. Adamonis and R. Maurer (Westinghouse). D. Mayes

(Duke), S. Redner (Northern States Power), and B. Vollmer and N. Farenbaugh (Zetec).

The Task Group provided input related to the make-up of the mock-up, the quality of the

data collected for the RR. the nature of the flaws, procedures for analyzing data, and

documentation. The Task Group provided input on the analysis guidelines, data acquisition,

degradation assessment, the training manual, and examination technique specification sheets

(ETMSs). The RR began only after the Task Group approved the documentation used for the RR

and concluded that the flaws in the mock-up had EC signals similar to those observed under

field conditions.

Consistent with current practice, the analysis teams had five members, including a

primary, a secondary, and two equally qualified resolution analysts to analyze the EC data.

The fifth member, the independent qualified data analyst (QDA), came from a utility. The

primary and secondary analysts reported their observations independently. The resolution

analysts reviewed calls when the primary and secondary analysts' calls differed. The

independent QDA monitored the effort, looking for possible excess overcalls and sampled 40

test sections to ascertain If flaws might have been missed.

The mock-up was treated as a steam generator owned by a utility. The role of the utility

in this case was taken by ANL. The ISI followed the process and procedures used by industry.

ANL was responsible for preparing documentation. monitoring data collection, monitoring data

analysis, and carrying out statistical analysis.

14
Four documents were prepared for the mock-up testing and for the RR data analysis.

They are ANLOOI Rev. 2 "Argonne Analysis Guideline,' ANLOO2 Rev. 3 "Multifrequency EC

Examination of Tubing within the ANL SG Mock-up,' ANLOO3 Rev. 3 "SG Mock-up Tubing

Degradation Assessment and Technique Qualification," and ANLOO4 Rev. 3 "Training Manual."

These documents are discussed in some detail in Ref. 3.

The 'SG Mock-up Tubing Degradation Assessment and Technique Qualification" (ANLOO3

Rev. 3) was prepared per the requirements specified in NEI-97-06 [51 and Revision 5 of the

EPRI PWR Steam Generator Examination Guidelines. In accordance with Rev. 5 of the EPRI

PWR Steam Generator Examination Guidelines 161, EPRI-quallfled techniques were reviewed to

ensure that application of these techniques was pertinent to site-specific conditions of the

mock-up.

The document "Multifrequency EC Examination of Tubing within the ANL SG Mock-up'

(ANLOO2 Rev. 3) provides all information necessary to collect the RR data. The procedures

mimic those of an actual ISI. The basis of the data acquisition is the EISS. The document

defines the frequencies, axial and rotational speeds, and calibration procedures. The two

ETSSs developed for the ANL mock-up were based on reviews of EPRI ETSSs for the various

degradation mechanisms in the mock-up and combining them into the two used.

The 'ANL Analysis Guideline (ANLOO1 Rev. 3) provided the technical direction for the

performance of EC examinations of the ANL SG mock-up. This procedure was applicable to all

examination personnel and generally mimics industry ISI guidelines.

A "Training Manual' (ANLOO4 Rev. 3) was developed for review by all analysts before the

RR exercise was begun. The manual provided information on the mock-up design, including a

schematic diagram, information on type of artifacts present, a discussion regarding the

presence of the circumferential markers, discussion of how the data for the RR were acquired,

and a table showing the format for entering data.

Before development of the mock-up ETSS, the essential variables for all EPRI-qualifled

techniques were reviewed. This ensured that the applications of the EPRI techniques are

pertinent to the site-specific mock-up steam generator conditions.

During June and August of 1999, a qualified three-man team from Zetec collected data

from the mock-up. The data acquisition team included a QDA Level Ha and a QDA Level Hla.

Data were acquired with a 10-D pusher-puller, MIZ3O with 36-pin cables, and Eddynet

software. BC data from a mag-biased probe were collected from all 3600 test sections of the

mock-up. The BC data were calibrated before and after the 4-h interval required to collect the

data. No change in voltage from the standard was detectable during this time period. A magbiased rotating three-coil probe that includes a +Point. 2.9-mm (0.1 15-in.)-diameter pancake,

and high-frequency shielded coil was used to collect data from all 400 tube sheet test sections

and all special-interest (spin call) test sections. A comparison of magnetically biased and

unbiased coils showed that biasing eliminates the voltage shift and noise in the +Point EC

signal resulting from tube sensitization.

The BC data were taken at 0.53 m/s (21 in./s), maintaining a digitization rate of

37 samples per inch. Bobbin coil data were taken at 400, 200, 100, and 20 kHz (differential

and absolute). MRPC data were taken from all degraded test sections in addition to hundreds

15
of clean test sections and test sections with artifacts. An ASME standard and a standard with

18 ID and OD axial and circumferential EDM notches (20, 40. 60, 80, and 100% TW) were used

for calibration. MRPC data were taken at 900 rpm and an axial speed of 12.7 mm/s (0.5 ln./s)

to maintain a digitization rate of 12 samples/cm (30 samples per in.) in the circumferential

direction and 12 samples/cm (30 samples per in.) in the axial direction. Data were taken at

600, 400, 300, 200, and 100 kHz. Nine +Point probes were used during this exercise because

of probe wear. MRPC probes were replaced when one of the channels could not be nulled.

This condition appears on the computer screen as an alert "flag."

The mock-up data collected by Zetec were analyzed at ANL (by ANL personnel) with

Eddynet98 software. The locations of the flaw signals were checked against the location data of

the flaw map.

Companies participating in the RR provided a list of analysts who would be available to

participate. For those companies who supplied more names than would be needed, ANL

randomly selected the team members from the list provided. The team members were expected

to be available during the entire exercise, generally seven to eight working days. Analysts were

generally QDA Level l~a or m.. The resolution analysts and independent 9DAs were Level m or

I1a. During the RR exercise, the primary and secondary analysts did not communicate with

each other or the resolution analysts. Upon submitting their reports, the primary and

secondary analysts could discuss their reports with the resolution analysts, but their reports

were not changed as a result. The resolution analysts provided the final report used for

establishing POD.

Before the RR exercise was started, a training manual, supplemental schematics, and

final reports for the training data were sent to the teams for review. The training optical disk

was either sent for review before the exercise was started or was provided by the proctor upon

arrival. The ANL proctor arrived at the analysts' site with exams, documentation (analysts'

guidelines, etc.), and optical disks containing all the data to be analyzed. The proctor provided

nondisclosure agreements signed by all analysts participating in the RR and collected all

analyst certifications. After the analysts finished studying the training manual, analyst

guidelines, the training disk, and supplemental schematics, the ANL proctor gave and graded

the written and practical site-speciflc exams. The passing grade for the written exam was 80%.

For the BC practical exam, the analysts had to correctly call all 'I' codes without excessive

overcalling. For the MRPC data, the analysts had to correctly indicate the presence of all

cracks and their orientations (circumferential vs. axial). About 10%3 of the analysts had to take

the second practical exam, which they passed. The ANL proctor retrieved the exam disk after

testing was completed. The process of evaluating the analysts closely followed standard

Industry practice.

After the analysts completed the site-specific exam, the proctor provided a third disk

containing all BC data. The primary and secondary analysts analyzed the BC data, and their

reports were recorded on the disk. The resolution analysts resolved the primary/secondary

discrepancies. A resolution analyst's report was provided along with primary and secondary

analyst reports. The Argonne BC disk contained primary, secondary, and resolution analysts

reports for BC data at the conclusion of the BC analysis. The proctor collected hard copies of

these reports and the data disk.

16
The ANL proctor then provided a fourth disk containing MRPC special-interest data. The

primary and secondary analysts analyzed the MRPC data, and their reports were recorded on

the disk except for tube sheet data, which were analyzed later. Analysts reported the depth at

maximum amplitude and location information, following the instructions in the training

manual. The resolution analysts resolved the prlmary/secondary discrepancies and provided a

report. Upon completion of the special-interest data, the special interest disk contained

primary, secondary, and resolution analysts' reports for MRPC data at other than the tube

sheet level. The proctor collected hard copies of these reports and the data disk.

The ANL proctor then provided a fifth disk containing only tube sheet (Level A) MRPC

data from all 400 tubes. The primary and secondary analysts analyzed the data for the tube

sheet level and provided their report. The resolution analysts resolved any discrepancies and

provided their report. The completed tube sheet disk contains primary, secondary, and

resolution analysts' reports for the tube sheet. The report was printed, and the hard copy

given to the ANL proctor, who also collected the data disk.

After testing, the proctor returned to ANL with all the optical disks containing the

analysis reports for the team.

2.1.4 Comparison of Round-Robin Data Acquisition and Analysis to Field ISI

The RR exercise was intended to closely mimic a field ISI. The procedures, equipment

used, and documentation are based on industry practices for the inspection of steam

generators. Similar to field inspections, a Zetec MIZ3O instrument, along with a 10-D pusherpuller and Eddynet 98 software, was used to collect the data. A standard mag-blased bobbin

coil and an MRPC with 2,9-mm (0.115-in.)-diameter pancake, +Point, and 2.0-mm (0.080-in.)-

diameter shielded high-frequency pancake coils were used. Round-robin teams used

Eddynet98 software to analyze the data. While flaws and flaw responses have been shown to

be representative of field flaws, the mock-up is mechanically different from a steam generator.

The mock-up had no U-bends. Test sections are in contact with each other, resulting in strong

EC signals similar to a 360° 100% 1W circumferential crack at the test section ends. The

analysts, through training and practice, easily adjusted to these signals. and there is no

indication that the PODs reported are compromised by this mechanical arrangement. Another

physical difference is the presence of a circumferential marker at the bottom of each test

section. Again, through review of training examples, the analysts quickly became adjusted'to

the marker signals, and their presence appears to have had no effect on the POD results.

Noise levels in the mock-up are generally lower than those in field data. Although many

of the test sections had sludge and magnetite on the OD, many test sections with flaws did not

Noise as severe as that in the U-bends of plants such as Indian Point 2 was not present in the

mock-up freespan and TSP levels. A review of BC field data from seven plants provided a

general idea of the noise from a BC field inspection. Baseline noise in the BC voltage trace of

field data was generally -0.7 V (excluding U-bends and tube sheet). The mock-up BC baseline

noise level was less, about 0.3 to 0.4 V. This low noise suggests that the results from the

mock-up are an upper limit on POD for the TSP and freespan levels for flaws with low-voltage

BC signals. While it is possible to have a deep crack with a low BC voltage, the difference in

baseline noise levels between the field and the mock-up would affect the POD only somewhat

for shallow cracks and is not expected to have a significant effect on the logistic fit. Noise in

the mock-up tube sheet level, however, was significant and did play a role in the ability of

17
analysts to detect and correctly characterize the flaws in and around the roll transition zone

(RTZ. The tube sheet noise is present In the mechanically expanded portion and in the roll

transition. Variations in the geometry of the RTZ contribute to the difficulty of analyzing data

from the tube sheet.

As in field inspections, the analyst involved with the RR decides whether the quality of the

data is sufficient to analyze for flaws. In one example from the field (Callaway), the bobbin coil

is replaced when the peak-to-peak voltage (Vpp) exceeds twice the initial control level (from a

reference tube). In the RR the quality of the bobbin coil data did not vary during the time BC

data were collected. At Callaway, for example, the signals from notches in the standard must

be clearly discernible from background noise when MRPC data are collected; otherwise, the

probe is replaced. A similar protocol was followed for the RR. except that if the MPRC probe

could not be nulled, it was replaced. This procedure led to high-quality MRPC data from the

mock-up test sections.

EC examination parameters are typical of current ISI practice. The 100, 200, 300, and

400 kHz frequencies used for the BC allow use of the conventional 100-400 kHz mix to

suppress the TSP indication. The range of frequencies used for the MPRC data covers the

requirements of the EPPI ETSSs for flaws present in the mock-up. The mock-up may have a

greater variety of flaws than might be present in any given steam generator. Nevertheless, the

analysts were familiar with the EC responses to all types of flaws in the mock-up, as

demonstrated by their passing the EPRI personnel qualification exams. The variety of flaws in

the mock-up should not impact the POD results.

The reporting requirements are slightly different than for a field ISI. In the analyst

reports for the mock-up, an extra column shows whether the flaw indication is OD or ID. In

addition, the location of an indication is given by data point, not number of inches from a

physical reference. Another variation from conventional reporting is the requirement to report

no more than four flaws (two axial and two circumferential) at any given axial location. These

variations were necessary to provide an exercise that could be completed in a reasonable time

(7-8 days) and provide as much information as possible while not negatively affecting the work

of the analysts. These variations from standard practice are carefully described In the training

manual and analyst guidelines. The analysts made the adjustment to the mock-up reporting

requirements quickly.

A primary objective of the RR is to establish the POD for deep flaws. While some deep

flaws may result in relatively low EC signal amplitude, the deep flaws generally have high

signal amplitudes. As a result, although the voltage histogram for the mock-up flaws looks

reasonable, there are more high-voltage signals than expected from a field inspection. A review

of field data, such as from the McGuire Nuclear Plant (a better than average plant), shows that

while BC signals from TIS can be many volts in amplitude (>10 V), the signals from the TSP

regions are primarily less than 3 V, with most less than 1 V in the 100-400 kHz mix channel.

Stronger TSP flaw signals can be found in the mock-up because of the emphasis on deep flaws.

2.1.5 Strategy for Evaluation of Results

POD has been determined for the flaws in the mock-up as a function of flaw type and flaw

location (i.e., freespan, TSP, and tube sheet). The PODs have been plotted against maximum

depth, mp, average depth, and, for the case of circumferential cracking, area. Logistic fits have

18
been calculated and Include errors in depth sizing and false call rates. Upper and lower 95%

confidence limits are included in the logistic fit curves. An analyst is given credit for detecting

a flaw if the call is an 'Ib code (e.g., NQI, DNI, DTI for BC calls, MAI, SAI, SCI, MCI, and MMI for

MRPC calls) and the location is within 25 mm (1 in.) of the ends of the flaw.

The location error allowed for calls made from bobbin coil data Is 25 mm (1 in.) from

either end of the flaw along the tube axis. This allowed error converts to 30 data points for

bobbin coil data. For MRPC data, the error allowed in the axial direction is also 25 mm (1 in.)

from the ends of the flaw along the tube axis. This allowed error converts to 3000 data points

for the MRPC data.

Analysts' reports were used to determine the false call rate. The rate was determined

from a review of randomly selected flaw-free test sections in the mock-up and the number of "I'

codes called in those test sections. An or code call (NQI, DTI, DNI) signifies a flaw indication in

the section even though no flaw was actually present. A total of 522 test sections was

analyzed. No known stress corrosion cracks were in any of the test sections. With 11 teams,

there were 5742 chances to make a false call. The result was a 6% false call rate for the tube

sheet level using MRPC data. 1.7% for the TSP with BC data, and 0.1% for the freespan with

BC data. These rates are low enough to avoid any consideration of penalizing the analysts for

false calls. The false call rates for the TSP and freespan are lower than in field inspections

because of the lower noise levels in the mock-up. The false call rate for the tube sheet is of the

same order found in the field. Since higher false call rates would lead to higher POD curves,

the results presented in this report could be considered conservative. However, even doubling

the false call rate would have no discernible effect on the POD curves presented in this report.

Data from the eleven teams participating in the RR exercise were first handled by using

the EPRI "Shell' software. The optical disks used by the analysts contained the analysts'

reports and were read by the 'Shellr software. The 'Shells program sorts the data. Calls from

primary, secondary, and resolution analysts can be compared to the results of expert opinion.

Note that the result of the comparison to expert opinion Is not the result sought because expert

opinion does not always provide the true state of the flaws. The reference state of the flaws was

provided by the ANL flaw characterization algorithm, which uses a multiparameter approach to

analyze the EC data taken at multiple frequencies. Three reports are analyzed for each team

for each of the three parts of the RR the BC data, the MRPC tube sheet data, and the MRPC

special-interest (spin call) data. The "Shell' program sorts the data by degradation and, for

LODSCC at the tube support plate, by voltage. One team was not able to complete the specialinterest MRPC disk.

The Eddynet software provides a series of files that contain the reports of results from

each analyst who participated In the RR. These data are saved under an Eddynet environment

and are Identified by extensions that refer to primary, secondary, and resolution analysts'

reports. These files are then read by a text editor and converted into a format usable for offline manipulation. The text files were then imported into Excel. Excel macros were written to

sort the results and carry out the grading.

2.1.6 Statistical Analysis

The probability of detection as a function of depth was assumed to be expressed as a

linear logistic function of x:

19
P (x l +ea+bx (2.2)

or

px = logistic(a + bx), (2.3)

where a and b are parameters that are determined by comparison with the observed results.

Other forms for the POD curve can be chosen. but the linear logistic curve has been widely

used for this purpose and is widely used in other fields to describe binomial responses

(detected or not detected) 171.

The Method of Maximum Likelihood 181 is widely used to estimate statistical parameters

such as a and b. For quantities that are normally distributed, it can be shown to be equivalent

to the familiar method of least squares 18.91. It is more generally applicable, however, and can

be applied to events such as detection of cracks that are not normally distributed. If p(x) is the

probability that a crack of depth x will be detected by an inspection team, the probability that

the crack will not be detected is l-p(x). The probability that n out of N teams of inspectors will

detect a crack of depth x is

(N) pn (I )N.n , (2.4)

where (N)= is the combinatorial symbol. Equation (2.4) assumes that the teams

are equally capable and are independent of each other.

The probability L that a collection of K cracks of depth xl, X2, ... , xK will be detected

successfully ni, n2 , ..., nK times is just the product of the probabilities for the individual

cracks:

L= r(nk)pk (1pk)Nk k (2.5)

where Pk = p(a, b, xk) and a and b are the parameters of the logistic fit. The Method of

Maximum Likelihood seeks to determine a, b such that the probability of the observed

outcome, L, is maximized. It is more convenient to deal with the log of Eq. (2.5)

In(L) = , In((Nk})+ I [nkln(Pk)+ (Nk -nk)In(l-pk)]. (2.6a)

K

D= X [nkln(Pk)+(Nk-nk)ln(l-Pk)]. (2.6b)

k-l

The first summation in Eq. (2.6a) is a constant that is independent of the choice of a and

b. Defining D as the second summation in Eq. (2.6a), we can determine the choice of a and b

that maximizes D or L by solving

20
aD= O

aa

aD O'

fib

or

, nk -NkPk apt o

k=lPkQ-Pk) aa (2.7)

K nk-NkPk 4 = o

k=lPkQl-Pk) ab

Differentiating Eq. (2.7), we find that

4~k =-P p(- Pa

2.k.Op (1-pl)xk | (2.8)

ab

Using Eqs. (2.5) and (2.6). Eqs. (2.7) and (2.8) reduce to

K

I (nk -Nkpk) =O

K =1 (2.9)

X (nk - Nkpk)xk = 0

k=1

Equations (2.8) and (2.9) are a pair of simultaneous nonlinear equations for a and b. For

computation, It Is generally more convenient to determine a and b by algorithms that directly

maximize D rather than to attempt to solve these equations. Excel spreadsheets were

developed for this purpose and benchmarked against the commercial statistical software

package STATA.

Equations (2.9) can be solved for a and b. These values depend on the round robin

results, i.e. on nj, n2 , etc. If the round robin was repeated with a different set of teams or a

different set of cracks, different values would be obtained for a and b, i.e., there will be a

distribution of values for a and b. Similarly, the depths of the cracks, xk, are not known

exactly; instead we have a measured value ik = xk +Ek, where ek is the error in the measured

value of xk. The errors will be random variables. The distributions for a and b can be

characterized by mean values and variances. The mean values can be found by solving

Eq. (2.9), although it is generally easier to obtain a and b by direct maximization of D

(Eq (2.6b)I However, Eq. (2.9) involve the unknown quantities Xk*, where, in reality, only the

measured values, xlk are known. The Impact of the uncertainties in x can be estimated. The

details are given in Ref. 3.

In comparing POD results, one must be sure that differences are statistically significant.

There are several ways to test whether two POD curves are the same. The test described below

21
is the easiest to carry out, because it requires only the logistic regression results. A logistic

regression is run on two sets of data. Each regression fit has as a result a set of parameter

estimates u = (a,b) and an associated covariance matrix C. The two data sets are designated by

Greek letters a and A, and the two regression fits are described by

paL = logistic(a. I + ba.2 xlJ (2.10)

pp., = logistic(ap,I + bO,2 X(3. (2.11)

The regression fits produce the estimates ua and Ub, along with covariance matrices Ca

and Cb. To test whether ua = ub, one forms a chi-squared statistic:

X2 = (ua - uiF [CCa + CpD-1 (ua - u, (2.12)

and compares x2 to a critical value obtained from a chi-squared table. The degree of freedom

associated with the critical value equals the number of model parameters, in this case, two.

The two sets of parameters are equal when x2 is less than the critical value. For example, to

conduct the test at a 10% level of significance, the critical value would be 4.61.

A chi-squared table can also be used to assign a p-value to the statistic x. When

performing this test, a less stringent level of significance than typical can be used, such as 10%

or 20%/6 instead of the typical 5%. This approach has been used to determine if POD curves by

different teams using the same data are different by chance or if the difference is significant.

2.1.7 Results of the Analysis Round-Robin

2.1.7.1 Bobbin Coil Results

- Figure 2.13 shows the maximum likelihood logistic fit for the POD with LODSCC and

LIIJSCC at the TSP as a function of maximum depth (as determined from the multiparameter

algorithm). The NDE uncertainty in depth is included in the one-sided 95% lower confidence

limit (OSL). In general, the curves are consistent with the behavior typically expected for POD

curves. The POD for ID cracks is higher than for OD cracks (99%A with 98% OSL at 60% 'W vs.

75% with 65% OSL at 60% TW). Figure 2.14 shows raw data and the logistic fit curve for the

BC POD for one of the RR teams. As can be seen, the probability of a hit is very high for the

deepest flaws, which are dominant in this set of cracks.

Figure 2.15 compares the POD results (with OSL) for the TSP cracks with the results for

freespan cracks. While as expected the POD for freespan LODSCC (95% at 600% TW) is higher

than the POD for TSP LODSCC (75% at 60% TW), it is lower than the POD for TSP LIDSCC

(99% at 60% TW).

In addition to examining the RR data as a function of flaw depth, the POD has also been

evaluated as a function of BC voltage for TSP SCC. The results are shown in Fig. 2.16. A

pattern similar to that found for POD versus depth is observed for the POD vs. TSP bobbin coil

voltage. In this figure, the percentage of correctly calling a flaw is plotted against binned data

as indicated in the graph. Figure 2.17 shows the logistic fits to the POD vs. voltage data for

LODSCC and LIDSCC, along with the 95% one-sided confidence limits. In contrast to the case

22
1

0.8

a

0

IL

0.6

0.4

--i I I I I I

IDI

I

/ Tube Support Plate

LOD & LIDSCC

~~POD - - - --5OSL OD - - --POD - ID ------- 95 OSL ID

I I .*.,, I, *I 1

1.

I

Figure 2.13.

BC POD for TSP data as a function

of maximum depth (as fraction

throughwall) for LODSCC and

LIDSCC using maximum likelihood fit

with one-sided 95% confidence limit

(OSL). Maximum depth uncertainty

is Included In the OSL. Depths are

determined with the multiparameter

algorithm.

0.2

Avl

0 0.2 0.4 0.6 0.8 1

Depth

1

0.8 Figure 2.14.

BC POD for TSP data as a function

of maximum depth (as fraction

0.6 / - throughwall) for LODSCC using

maximum likelihood fit (solid line).

0.4 _ ZThe logistic fit Is to data from one - / team only. The circles show the raw

Tube Support Plate - data from which the curve Is

0.2 Data ODSCC generated. Depths are determined

atare fOr one team with the multiparameter algorithm.

0 0.2 0.4 0.6 0.8 1

Depth

of POD as a function of depth, the POD as a function of voltage for LEDSCC at the TSP is lower

than that for LODSCC. The lower POD vs. voltage curve for LIDSCC is possibly the result of

missing shallow cracks that are in dents with high voltages.

Because BC voltages are directly used to characterize tube Integrity for plants operating

under Generic Letter 95-05 1101. the BC voltages reported for LODSCC at tube support plates

have been analyzed to determine the team-to-team variability in the reported voltages. For

each LODSCC, an average BC voltage and a corresponding standard deviation were computed.

The cumulative distribution of the normalized standard deviations (i.e., the standard deviation

divided by the corresponding value of the average voltage) can be fit well by a Weibull

distribution (the RMS difference between the observed distribution and the Weibull fit is <0.03).

The fitted distribution is shown in Fig. 2.18. For almost 90% of the indications, the normalized

23
1

0.8 _/ Figure 2.15.

, o 1 t BC POD for freespan and TSP data

S, / /' as a function of maximum depth (as

0.6 1/ 1 fraction throughwall) for LODSCC

O / and LIDSCC by using maximum

CL ~~~~~~~~~~~~~~~likelihood fit with the one-sided 95% 0.4

. -- -' PODFSOD / @ , confidence limit. Maximum depth

' - | - --' 95OSLFSOD uncertainty is included. Depths are

0.2 j .. .. 95 OSL TSP OD- determined with the multiparameter

-f-/*- POD TSP ID algorithm. -'-------95OSLTSPID ,- ,' { -

0

0 0.2 0.4 0.6 0.8 1

Depth

100 1''1''''1''' '''

8 o ; I Na

M 1 Figure 2.16.

80 B Round-robin TSP bobbin coil resolution analysts' results as a

| I .. °function of BC voltage. The BC POD

3 40 _ was evaluated for LODSCC at the

Upp eand Lower 95% TSP. Depths are determined with

20 - Shown as Error Ba the multiparameter algorithm.

0 0.5 1 1.5 2 2.5 3

Voltage Range for Correct TSP BC Calls (V)

TSP LODSCC / , 0.8- and LIDSCC

0.8 7*

0 o.8 - Figure 2.17.

0 , / Logistic fit curves for TSP BC POD

0. , as a function of voltage for LODSCC

0.4 /, ,@ _ / PODOD - and LIDSCC.

/ , -* - 95OSL OD

0.2 , - - POD ID-_

>< - , - - 95 S 0SL ID

0 0 1 I 1 5 I 2

0 0.5 I 1.5 2

Voltage (V)

24
1

C

0.6 Figure 2.18.

* Cumulative distribution of normalized

Z 0.4 L standard deviations for BC voltages

E - for LODSCC at tube support plates. E

0.2

0

0 0.2 0.4 0.6 0.8 1

Standard Deviation

standard deviation in the reported voltages is <0.15. This result is consistent with Generic

letter 95-05 [101, which assumes a 15% cutoff for the voltage-response variability distribution

is acceptable. The indications with larger variations are not associated with particularly high

or low voltages (i.e., approximately half the signals with standard deviations >0. 1 have voltages

of >2 V). but rather with the complexity of the signal and the difficulty in identifying the peak

voltage.

The results were analyzed by the teams to determine whether there were strong

team-to-team variations in the POD. For this exercise, all teams were given optical disks

containing the same data sets to analyze. All analysts were given the same instructions and

documents related to analyzing the data. Team-to-team variations thus result from variations

in analyst interpretation of the same signals. The results as a function of team for freespan

and TSP LODSCC combined are shown in Fig. 2.19. The performance of most of the teams

cluster rather tightly, although significant variation exists between best and worst. Figure 2.20

shows team-by-team variation for the freespan LODSCC alone. Figure 2.21 shows team-byteam variation for the TSP LIDSCC alone.

From these findings, the probability that team-to-team variations in logistic fits to data

are due to chance can be estimated. For LIDSCC at the TSP, the variation from best to worst

(Fig. 2.21) is very significant. The probability is <0.1% that the difference is due to chance

(DTC). For FS OD, the variation from best to worst (Fig. 2.20) is probably sigificant (DTC Is

<20%). For TSP OD, the variation from best to worst DTC is >60% and is thus probably not

significant.

Figure 2.22 shows the POD logistic fits for LODSCC at the TSP as a function of mp.

Figure 2.23 shows the corresponding results for LIDSCC at the TSP. Figure 2.24 shows the

logistic fits for POD for axial SCC In freespan test sections as a function of mp. The errors in

calculating mp by using the NDE characterization of the crack geometry compared to using

fractography data have been determined with the 23-tube set. Because only one value of mp

25
0.8 - Figure 2.19.

1 U / BC POD by team for TSP LODSCC.

0.6 / X ' , - The maximum crack depth (as o hi °fraction of wall) was determined by 23/ A the multiparameter algorithm. The

0.4 - highest solid line represents the best

team, the lowest dashed line

- fig a TSP represents the worst team, and the

0.2- ,LODSCC other symbols represent the

remaining teams.

0

0 0.2 0.4 0.6 0.8 I

Depth

1 r C nG d- i

D ,/ g v x

08 - / g~~~4 v X- 0.8 - a x Figure 2.20.

* 3 BC POD by team for freespan

0.6_ | LODSCC. The maximum crack

o / t ° x depth (as fraction of wall) was X / determined by the multiparameter

0.4 - algorithm. The solid line represents

-v x / / the best team, while the symbols and

0.2 - / # xFreespan dashed line represent the remaining 0.2 -/ X LODSCC teams.

/ Ox(

0

0 0.2 0.4 0.8 0.8 1

Depth

1~~~~~~~

VA - -/o

0.8 _ v + 

/ . Figure2.21.

A . BC POD by team for TSP LIDSCC.

0.8 0 + The maximum crack depth (as

X /fraction of wall) was determined by

V the multiparameter algorithm. The

0.4 0 / - solid line represents the best team,

AT the dashed line represents the worst

0.2 A/: 8 + / TSP UDSCC _ team, and the symbols represent the

+ v remaining teams.

0 0.2 0.4 0.8 0.8

Depth

26
1

0.8 _

/ , Figure 2.22.

0.6 / BC POD for TSP LODSCC as a

0.6 - / / TSP LODSCC - function of mp by using maximum

0 - . likelihood fit with an estimate of the

/ 4 one-sided 95% confidence limit. The

0.4 / values of mp are derived by using

- /, POD depths from the multiparameter

0.2 _ -_ - 95OSL algorithm.

'I0 , I IA, I,, I

1 1.5 2 2.5 3

m~p

-' /1

0.8 /

Figure 2.23.

BC POD for TSP LIDSCC as a

0.6 - function of mp by using maximum

o -TS UICC likelihood fit with an estimate of the

OA _ X one-sided 95% confidence limit. The

0.4-

-1; values of mp are derived by using

POD - depths from the multiparameter

0.2 -- 95OSL algorithm.

0 , ,, I, I''

1 1.5 2 2.5 3

0.8 i

* { Figure 2.24.

BC POD for freespan data for

0.6 Freespan LODSCC as a function of mp by 0 [ using maximum likelihood fit with an

0.4 _ [ Lestimate of the one-sided 95%

I confidence limit. The values of mp

- ; POD are derived by using depths from the

0.2 -/ -95OSL multiparameter algorithm.

0 1

1 1.5 2 2.5 3

. M P

27
per crack is obtained, fewer data are available than In the case of depth (multiple points per

crack); hence, estimates of mp have greater uncertainty. In all three graphs, the 95% one-sided

lower confidence limit includes the error due to the use of NDE data to calculate mp, as well as

the statistical uncertainties associated with finite samples. In the TSP and FS regions, the

POD for cracks that would fail or leak under 3Ap internal pressure (corresponding to mnp 2.3)

is >95%, even when accounting for depth uncertainties.

The analysis presented in this section is based on the results of the resolution analysts'

reports. In some cases, the bobbin coil signal was difficult to analyze, and significant

disagreement occurred between the calls of the resolution analysts and the primary and

secondary analysts. Three examples for the TSP are presented here. The first is for a 23% IW

LIDSCC, the second is for a 67% TW LODSCC, and the third is for a very short 99% TW

LODSCC. Table 2.3 shows that, while for some teams all analysts were in agreement with

respect to making I-code calls with the bobbin coil data (13 out of 33), for other teams there

was disagreement. In some cases (also 13 out of 33), the resolution analyst incorrectly

dismissed a correct call by the primary and/or secondary analyst. In four cases, the resolution

analysts made a correct call, while the primary and/or secondary analysts did not. These

examples show that significant team-to-team variations exist for difficult-to-analyze signals,

and that It might be more prudent to limit the impact that resolution analysts have in making

the final call for these types of SCC.

2.1.7.2 MRPC Tube Sheet Results

Figure 2.25 shows the POD for SCC in the tube sheet region as a function of maximum

depth for LIDSCC and CIDSCC combined and LODSCC and CODSCC combined. (Note that the

false call rate for the tube sheet Is fairly high, =6%). The POD for IDSCC is higher than for

ODSCC, as expected. At 60% TW, the POD for IDSCC is =90% with an OSL of -75%.

Figure 2.26 shows the POD as a function of maximum depth for axial and circumferential

SCCs in the tube separated into a POD curve for LIDSCC only and for LIDSCC combined with

CIDSCC. The highest POD curve is for LIDSCC, where the POD at 60% TW is 95%.

Figure 2.27 shows MRPC POD by team as a function of maximum depth (as estimated by the

multiparameter algorithm) for axial and circumferential IDSCC and ODSCC in the tube sheet.

The POD at 60% TW ranges from 90% to 70%. Based on the discussion in Section 2.5.1, the

probability that team-to-team variations in logistic fits to data are due to chance can be

estimated.

Figures 2.28-2.30 compare the tube sheet BC POD and tube sheet MRPC POD.

Figure 2.28 shows, for all SCC, a higher POD curve for the MRPC (80% vs. 40% at 60% T"V.

Figure 2.29 compares BC and MRPC PODs for tube sheet LIDSCC and CIDSCC, with the MRPC

curve substantially higher. Figure 2.30 compares BC and MRPC POD for tube sheet LIDSCC

only. The MRPC POD at 60% 'lW is 95%, while the BC POD is only 40%.

The MRPC results were reviewed for TSP LODSCC BC voltages between 2.0 and 5.6 V was

carried out. Such calls are normally reviewed to confirm or dismiss the BC call. There were

17 TSP LODSCC flaws with BC voltages in the range 2.0 to 5.6 V. and maximum depths were

estimated to be >70% 'W (by the multiparameter algorithm). The average correct call using the

MRPC data for this set of cracks was 98% (with a 96% lower 95% confidence limit). One other

LODSCC in the TSP with BC voltage in the 2.0-5.6 V range had an estimated maximum depth

of 28% TW. None of the teams correctly called this flaw with the MRPC data.

28
Table 2.3 Bobbin coil calls of primary, secondary and resolution analysts for SCC at the TSP. Note

that the deep (990%TW) LODSCC is very short.

LEDSCC LODSCC LODSCC

(23% IW) (67967W) (99% MW

Number of teams where the resolution, 4 6 3

primary, and secondary analysts all made a

correct bobbin coil I-code call.

Number of teams where the resolution 1 3 0

analysts made a correct bobbin coil I-code

call, but the primary and/or secondary

analyst did not.

Number of teams where the resolution 5 2 6

analysts did not make a correct bobbin coil

I-code call, but the primary and/or

secondary analyst did.

Number of teams where the resolution, 1 0 2

primary, and secondary analysts all failed to

make a correct bobbin coil I-code call.

1

0.8

0

0.

0.6

0.4

Figure 2.25.

MRPC POD as a function of

maximum depth (as fraction of wall)

for LIDSCC and CIDSCC combined

and LODSCC and CODSCC

combined at the tube sheet. Curves

include maximum likelihood fit and

an estimate of the one-sided 95%

confidence limit that includes

uncertainty In maximum depth.

Depths are determined with the

multiparameter algorithm.

0.2

0

0 0.2 0.4 0.6 0.8

Depth

1

29
1

Tube Sheet - / Figure 2.26.

0.8 Axial & CIrc SCC/ MRPC POD as a function of

0.8 maximum depth (as fraction of wall)

7.' for axial and circumferential IDSCC

0.6 - and ODSCC in the tube sheet. The

a -/ upper dashed line is for all-LIDSCC

2 / * . only, the lower dashed line is for

0.4 - LIDSCC and CIDSCC. Depths are

determined with the multiparameter

0.2 /' a,/PODCID&O0 algorithm. Combined ID and OD

0 -- / POD LID should not be used except for

comparison of relative effectiveness

I I I I of probes.

0 0.2 0.4 0.6 0.8 1

Depth

1 '''I' ''l

Tub. Sheet ID &OD

0.8 _ Axial & Clro SCC * I Figure 2.27.

/ !,8' MRPC POD by team as a function of

u& / g maximum depth (as fraction of wall)

0.6 U / in the tube sheet. The solid line

0 - i; , - represents the best team, the

0.4 _ /t/ _ dashed line represents the worst

0.4 / ' team,. and the symbols represent the

remaining teams. ID and OD are

0.2 - combined for the purpose of

. J comparing teams.

0 I I l l I I

0 0.2 0.4 0.6 0.8 1

Depth

The possibility of a crack with a high BC voltage being missed by the subsequent MRPC

data analysis could arise when a flaw is shallow and long. shallow and volumetric, or deep and

short. An example is shown In Fig. 2.31. The crack profile In this case is generated from

mock-up data with the multiparameter algorithm. This axial TSP LODSCC with maximum

depth of 99% TW was missed by all teams analyzing MRPC data. In this case, the MRPC

+Point voltage at 300 kHz was only =0.2 V. The largest part of the segmented crack has a

length of about 10 mm. Figure 2.31b shows the crack along the test section axis. The mp for

this flaw is -4.5, Indicating that the tube would leak at pressures well below 3Ap. The BC

voltage for this crack can, depending on analyst, vary from 4.5 to 8 V. The dye penetrant

image of the crack intersection with the tube OD is consistent with the Isometric image

generated by the multiparameter algorithm.

30
1 .. ... . . . . - . . . .

_ ' I I ' I I ' ' I I I ' I .

- Tube Sheet ,. - *_Im S f0B1= Lfr

0.8 _ unu Figure 2.28.

MRPC . BC and MRPC POD for axial and ~ - - - 95OSLMRPC / circumferential IDSCC and ODSCC

0.6- ---- -- 95 OSL BC / * at the tube sheet. The lower 95%

o one-sided confidence limits (OSL)

ff 0.4 /.. - / # - are shown (includes uncertainties In . .* Xdepths). Combined ID and OD

/ - .. should not be used except for

0.2 - , - . comparison of relative effectiveness

of probes.

o

o 0.2 0.4 0.6 0.8 1

Depth

Tube Sheet

-UDSCC and CIDSCC

0.8 -

-0.6 / - Figure 2.29.

o , - 2 r BC and MRPC POD for CIDSCC and

0.~ ~ ~ ~ ~ ~ ~ ~ ~ ~~~~IDC ttetb he.Dph 0.4 - are determined with the

/RPG . 'D multiparameter algorithm. - / , - ~~~~~MRPC L&C ID

0.2 - - -BCL&C ID

0O I I I I I I I 1 1 , I , I.1 

o 0.2 0.4 0.6 0.8 1

Depth

Tube Sheet

0.8 ~ LIDSCC

0.6 FIgure 2.30.

o / , - : BC and MRPC POD for LIDSCC at 0 , the tube sheet. Depths (as fraction

0.4 - of wall) are determined with the

/ -. multiparameter algorithm.

/ * , MRPC UD

0.2 - / _ - ---.- BCLID

0 0.2 0.4 0.6 0.8 1

Depth

31
10MM""MVIR I 11~~~~~~~~~~~~~=4-

R 118"NIM " I . I . . It

.Ikll 0. " P. ' 4 7-14'.E ti4 1-1 ", lz.i

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

------- ---- -- 4 .

so ------ -----

20 .. .... -- - ---.

0 10 M 30 All so

(a) (b)

Figure 2.31. (a) TSP LODSCC with maximum depth of 99% TW that was missed by teams analyzing

MRPC data. The largest piece of the segmented crack has a length of about 10 mm

(0.4 in.). (b) Projection of the crack along the test section axis.

These results suggest that flaws detected correctly by the bobbin coil could subsequently

be dismissed upon further examination of MRPC data even when flaws are relatively deep. The

MRPC probes are very effective In characterizing defects, compared to bobbin coils, but may be

less effective than bobbin coils in recognizing that a crack is present.

2.1.8 Nature of Missed Flaws

The primary cause of missing a flaw is that the flaw signal is distorted by geometry or

deposits and is no longer recognized. A tight crack that does not generate a significant EC

signal is another cause for missing a flaw. Another possibility is that the signal from an EC

coil does not conform to what is expected (i.e., the signal could be out of the flaw plane or could

be generated by multiple cracks). Very long flaws may be missed because the analysts may

concentrate on a small portion of the flaw, thereby missing the overall response. Confusion

could also arise from conflicting behavior of two or more coils. For example, there could be a

clear BC signal but nothing reportable from an MRPC. Analysts have a preconceived idea of

what flaw responses at various locations should be like and might not pursue anomalous

indications that are actually from a flaw. A few cracks in the mock-up have been called by the

bobbin coil and dismissed following an MRPC analysis. Some cracks detectable with an MRPC

are not detected by the bobbin coil. The reasons vary, as described above, but in the case of a

crack and a BC indication, the crack is not called because the indication has a very high phase

angle or is out of the flaw plane.

2.1.9 Nature of Overcalls

Overcalls are the result of signals from certain coils that tend to generate flaw-like signals

from geometrical distortions and deposits. Overcalls could also be the result of confusion from

conflicting behavior of two or more coils. In a round-robin exercise, participants tend to make

calls that might not be made under field conditions because there is no penalty for overcalling

32
as long as the overcalling Is not abused. In fact, the reports from resolution analysts show a

reasonably low overcalling rate for the freespan (0.1%) and TSP (1.7%). The overcalling at the

tube sheet level is significantly higher (6%). The complex nature of the roll transition is

probably the root cause of the tube sheet overcalling, although further review and destructive

analysis suggests that unintentional flaws may have been introduced to the tube sheet level

during tube expansion, flaw fabrication, and assembly.

2.1.10 Implications of the Round-Robin Results

In summary, the detection capability of current ISI technology and procedures has been

assessed by carrying out an eddy current RR exercise with a steam generator tube bundle

mock-up. Inspection of the mock-up and analysis of the data mimicked industry ISI practices

conducted on operating steam generators. AU documentation for conducting the inspection

was prepared with Input from an industry-based NDE Task Group, and the realism of the

mock-up was established. Data were acquired in June and August 1999, and these data were

analyzed by I 1 commercial teams in December 2000. Each team consisted of five qualified

analysts. The exercise took seven to eight working days per team.

The conclusion from the analysis of round-robin results is that good POD can be achieved

for deep flaws when commercial techniques are used in a similar manner to the RR exercise.

The level of success in detection of SCC did vary with flaw location. The maximum depth from

eddy current crack profiles and false call rates was estimated to establish POD as a function of

depth and mP. Logistic fits to the data were generated. The BC POD for TSP ID cracks is

higher than for OD cracks (99% with 98% OSL at 60% TW vs. 75% with 65% OSL at 60% 'W).

The BC POD for freespan LODSCC is -95% at 60% 1W. The POD for IDSCC is -90% with an

OSL of =75%. The highest POD curve Is for LIDSCC, where the POD at 60% TW is 95%. The

MRPC results for BC voltage between 2.0 and 5.6 V were reviewed. Such calls are normally

reviewed to confirm or dismiss the BC flaw call. The result for LODSCC >74% 'W is an

average correct call of 98%. All teams missed an LODSCC at the TSP with an estimated

maximum depth of 28% 'W. Another example illustrates the possibility that a deep crack with

a weak MRPC signal might not be called a crack by analysts. The example presented had an

estimated maximum depth of 99% TW with only a few tenths of a volt generated by the +Point

coil at 300 kHz.

No useful correlation was found between signal amplitude or phase and the maximum

depth of the mock-up flaws. When the PODs are considered as a function of mP, in the TSP

and FS regions, the POD for cracks that would fail or leak under 3SAp internal pressure

(corresponding to mp - 2.3) is >95% even when uncertainties are accounted for.

These analyses also established the adequacy of the multiparameter algorithm for

obtaining profiles and maximum depth. The curves for POD as a function of depth or mp were

based on the profiles generated with the multiparameter algorithm.

2.2 Tube Removal from McGuire 1 Retired Steam Generator

2.2.1 Review of Duke ISI Data

Steam-generator EC data were obtained from the Duke Power Company's McGuire plant

on a magneto-optical disk in the form of the raw inspection data. These data were analyzed

33
using Zetec Eddynet95 software in an effort to optimize the selection of tubes for the tube

removal effort from two retired steam generators. Tubes with single and multiple axial and

circumferential indications in the TS and the first TSPwere the main priority, and several

tubes were selected as candidates for removal. Areas where defects are somewhat clustered

were also identified. These areas are of greatest interest because selecting tubes in a relatively

small area minimizes the total cost of procuring tube samples.

2.2.2 Results of Field NDE of Deplugged McGulre Tubes

Twenty tubes from the McGuire 1 steam generator D (SGD) were selected for plug

removal. Nineteen of these 20 tubes were successfully deplugged in June 1997. Eighteen of

these 19 tubes were subsequently examined by NDE, and the results were used to select tubes

for tube pulling.

Deplugged tubes and TS sections removed from the McGuire Station SGD were inspected

before removal using +Point, "3-coil pancake, and bobbin coil probes. This inspection and the

subsequent analyses were carried out by staff from ANL, Duke, and the EPRI NDE Center. An

ASME standard and a standard with 18 axial and circumferential OD and ID EDM notches

were used in the data collection.

To determine any changes in EC voltages between current and previous inspections of

deplugged tubes, the analysis team followed the procedure specified in the 'Eddy Current

Analysis Guidelines-McGuire Nuclear Station Unit 1.' Unprocessed (raw) bobbin coil data

taken in previous ISIs were used to reproduce, for the tubes selected, the amplitude and phase

results reported in the Duke EOC1O (December 1995) and EOC9 (August 1994) summary

sheets. The same procedure was applied to the data taken July 1997, using the same

reference standards and normalizations employed in the previous ISI. The BC voltages are

presented in Table 2.4; in all cases, the voltage increased after deplugging. Factors that could

result in voltage increases include crack growth, widening of cracks, or a change in deposits.

Note that for the one tube plugged in August 1994 (R7 C24), the voltage increase was the most

dramatic.

In Table 2.5. results for a standard 2.92-mm (0.1 15-in.)-diameter pancake coil are

compared. Reproducing the results in the Duke summary sheets from raw EC data of previous

ISIs was more difficult for the RPC data than for the bobbin coil because the analysis is

somewhat more subjective. After deplugging, the RPC voltages, except for R33 C33 and

R5 C17, were about the same or increased with time. Again, for the tubes plugged in 1994

(R7 C24 and R39 C46), the voltage increase was most dramatic.

In addition. several new indications were identified from relatively weak EC signals.

These resulted primarily from new signals from probes similar to those used before and from

the availability of new +Point data (not taken during previous ISI). A review of some earlier EC

data indicated that these new calls would not have been made with the data available during

previous inspections because either (1) the EC signals initiated after the December 1995 ISI, or

(2) the EC signals were too weak to be called flaw indications with probes used at that time but

had since grown sufficiently to be detectable. An example is shown in Figs. 2.32 and 2.33.

These figures present standard pancake data in the form of an isometric plot (c-scan) around

34
Table 2A. Bobbin coil voltages before plugging and EC voltages after deplugging for selected tubes

from hot-leg side of McGuire steam generator D.

Bobbin Coil Volts Difference

Tube Before After Call/

Position Pluggonga Deplugginga Locationb Volts Percent

R4 C43 0.56 1.56 MAI/1stTSP 1.00 179

R5 C17 0.32 0.70 MAI/2nd TSP 0.38 119

R5 C44 8.0 14.00 MCI/TO S 6.00 75

R5 C51 0.45 1.30 MAI/Ist TSP 0.85 189

R7 C24 0.56 6.66 MAI/2nd TSP 6.10 1089

R11 C96 10.54 12.50 sciarsc 1.96 19

R13 C89 3.88 5.22 SCI/TS 1.34 35

R14 C55 0.35 0.91 MAI/2nd TSP 0.56 160

R17 C90 4.07 4.90 MAI-MCI/TIS 0.83 20

R19 C20 0.52 0.54 Vol/5th TSP+ 0.02 4

R33 C33 6.61 10.00 MCI-MAI/TIS 3.39 51

R39 C57 1.14 1.80 SAI/IstTSPC 0.66 58

R45 C40 0.27 0.57 MAI/lstTSPC 0.30 111

STubes plugged 12/95, except R7 C24 plugged 8/94: all tubes deplugged 7/97.

bMAM = multiple axial indication. MCI = multiple circumferential indication, SAI = single axial indication,

SCI = single circumferential indication, Vol volumetric indication.

CNearly throughwall indication from ISI before plugging.

Table 2.5. Rotating pancake coil voltages before plugging and EC voltages after deplugging for

selected tubes from hot-leg side of McGuire steam generator D.

RPC Volts Difference

Tube Before After Call/ b

Position Plugginga Deplugginga Location Volts Percent

R4 C43 0.47 1.20 MAI/1stTSP 0.73 155

R5 C17 0.48 0.37 MAI/2nd TSP -0.11 -23

R5 C44 0.59 0.90 MCI/OIS 0.31 53

R5 C51 0.22 0.93 MAI/1stTSP 0.71 323

R7 C24 0.35 2.75 MAI/2nd TSP 2.40 686

R11 C96 7.64 - SCI/I'-Sc

R13 C89 0.34 0.66 SCI/1TS 0.32 94

-R14 C55 0.34 0.69 bAI/2nd TSP 0.35 103

R17 C90 0.52 2.40 MAI-MCI/TTS 1.88 362

R19 C20 0.93 1.04 Vol/5th TSP+ 0.11 12

R33 C33 2.74 1.02 MCI-MAI/TTS -1.72 -63

R39 C46 0.7 2.80 MAI-SCI/TfIS 2.10 300

R39 C57 0.57 1.17 SAT/1stTSPC 0.60 105

R45 C40 0.27 0.39 MAI/ I St TSPC 0.12 44

aTubes plugged 12/95. except R7C24 plugged 8/94: all tubes deplugged 7/97.

bMAI = multiple axial indication, MCI - multiple circumferential indication, SAI = single axial Indication,

SCI = single circumferential indication, Vol . volumetric indication.

CNearly throughwall indication from ISI before plugging.

35
the top of the tube sheet ('TS). In 1995. there was no clear evidence of a crack. In 1997, the

EC signal pattern had changed, and multiple axial and circumferential crack calls were made.

This example suggests growth of EC signals in plugged tubes.

2.2.3 EC Signal Growth

The growth of an EC signal attributed to an SCC can be seen in the c-scan plots of

Figs. 2.34 and 2.35. which were generated using a 2.92-mm (0.1 15-in.)-diameter MRPC. In

Fig. 2.36, the arrow points to an EC signal at the ITS in tube R39 C46. This indication was

called an SCC in August 1994. Figure 2.35 shows the ITS In July 1997. after the tube was

deplugged. The signal has grown considerably relative to the roll transition (Fig. 2.36) for one

or more of the reasons suggested above.

Several SCCs from McGuire have been examined at Argonne using a modified glovebox. a

4D Pusher Puller, a MIZ30, and Eddynet 98 software. Table 2.6 compares the BC voltage and

phase angle measured at McGuire before pulling and at Argonne after pulling. The Argonne

data include results both with and without a ring over the crack to simulate the TSP.

Measurements made at Argonne followed the same procedure as that used at McGuire. This

procedure entailed use of an ASME standard, setting the 550-130 kHz mix channel for the four

20% 'W holes to 2.75-V and the phase angle setting of the 100% '1W hole to 400. In general,

the voltages measured at Argonne are slightly lower, with the phase essentially unchanged

except for SCC noted as A in the table. The differences noted could be the result of variations

in the mix and changes in conduction paths across the crack faces resulting from the tube

pulling procedure (i.e.. crack faces now touching). All test sections were pressurized at room

temperature to a maximum pressure of 51.7 MPa (7500 psi). Tube B failed at 49.0 MPa

(7100 psi) and leaked at a rate of 32.2 L/min (8.5 gal/min). Tube E failed at 35.2 MPa

(5100 psi). The others did not fall at the maximum system pressure of 51.7 MPa (7500 psi).

2.3 Effect of Surface Oxide Films on Eddy Current Signals from SCC

The effect of corrosion products (thin oxide films) formed on crack faces on the EC signal

from an SCC has been investigated. Alloy 600 tubes with axial ODSCC were exposed under

PWR water chemistry conditions (3000C and 3-8 ppb oxygen) for about two months. The

cracks were examined with both magnetically biased (mag-biased) BCs and a +Point coil

before and after corrosion products were formed. The voltages for the BCs increased

significantly with the creation of the thin oxide film. However, the general shape of the

Lissajous figures remained unchanged. In contrast, the results for the +Point coil were

inconclusive. In two cases, the +Point voltage did not change, while in a third, the voltage

dropped significantly after the oxide film formed. Figure 2.37 shows the BC Lissajous figure,

along with voltage and phase information for one of the tubes. Table 2.7 summarizes the BC

results for two tubes. While voltages increased significantly, the phase angle did not. The

creation of corrosion products in the crack could lead to a reduction in the number of

electrically conducting paths from contacting crack faces. In that case, the EC signal would be

expected to increase, as observed, while the depth remains essentially the same.

36
-_ RPC: SGDjH-CL00211 SUN 13:21:12 DEC-24-95 SG D.

Filtwr:FF P uttca xfr.Qs-24 Ytreds xeot45 Vat.3m

A+~~~~~~

20~~~~

Figure 2.32.

C-scan plot of tube R45C40 using a

2.92-mm (0.115-in.)-diameter MRPC

showing top of tube sheet in December

1995. No crack call was made at that

time. The tube axis Is from lower left to

upper right, with full 360° shown. The

roll transition In Indicated by the arrow.

Figure 2.33.

C-scan plot of tube R45C40 using a

2.92-mm (0.115-in.)-diameter MRPC

showing top of tube sheet in July 1997.

Multiple axial and circumferential

cracks were called at that time. The

tube axis Is from lower left to upper

right, with full 3600 shown. The roll

transition is Indicated by the arrow.

The variation in EC signal from 1995

(Fig. 2.32) can easily be seen.

240~ ~ ~ ~ ~ ~~~~~~R4

0 00~~~

11d.7

2.4 Crack Sizing

Mock-up data for several cracks were used to compare the profiling results from industry

experts to the results obtained from application of the Argonne multiparameter algorithm. The

objective of the exercise was to obtain information of the variation possible for maximum depth

estimates when using the same EC data. Figure 2.38 shows the maximum depths for a variety

of crack morphologles estimated by two recognized EC NDE experts using the same procedure

with +Point data at 300 kHz and by Argonne using the multiparameter algorithm with Input

primarily from pancake phase and amplitude data. The deeper the crack, the better the

comparison. For one TSP (No. 4), the varation was only about 10%. In another case (No. 3),

37
Figure 2.34.

C-scan plot from a 2.92-mm (0.15-

in.)-diameter MRPC; arrow points to

indication at top of tube sheet (TTS)

- , In Htube R39 C46. This indication was

called an SCC in August 1994. Tube

axis Is from lower left to upper right,

with 360- of tube shown.

N _,* got k Ad _ _t_.A .

Figure 2.35.

C-scan plot of same region shown in

Fig. 2.34 taken with a 2.92-mm

(0.11 5-in.)-dlameter MRPC in July

1997 after tube was deplugged.

. I Signal g (arrow) has grown

cnsiderably relative to roll transition.

Tube axis is from lower left to upper

right, with 360° of tube shown.

Upper and lower Omits for estimated length of CODSCC at RT In 1994 and 1997

1.50 '3.81

Figure 2.36.

Estimated length with range of I 1.00 _254 uncertainty for CODSCC at top of

2 . . s tube sheet in Tube 39-46 of McGuire .1.. _retired _ steam generator before I T - aE (August 1994) and after (July 1997)

0.50 1.27 plugging.

0.00 - 0.00

38
Table 2.6. Comparison of BC data before and after tube pull

BC Voltage and Phase BC Voltage and Phase

Measured at McGuire Measured at Argonne

Flaw Type (550-130 kHz Mix) before (550-130 kHz Mix with and

and Location Tube was Pulled without TSP Simulation

A. 1ItTSP, LODSCC 1.3V/830 1.1 V/130 °withTSP

1.2 V/1230, no TSP

B. Ist TSP, LODSCC 1.6 V/870 1.3 V/ 870 with TSP

1.0 V/ 690, no TSP

C. 2nd TSP, LODSCC 6.7 V/810 3.9 V/ 790 with TSP

4.3 V/ 820, no TSP

D. 2nd TSP, LODSCC 0.9 V/1200 0.8 V/1330 with TSP

0.9 V/ 1370, no TSP

E. 16t TSP, LODSCC 1.8 V/460 1.8 V/ 400 with TSP

1.8 V/ 400, no TSP

the variation was over 50% C1 CIDSCC). One analyst (B) was not able to make a good

estimate for that case. The variation was greatest for the TS SCC.

2.5 Array Probes

The X-Probe is a transmit-receive EC array consisting of pancake coils in a bracelet

configuration. The transmit coil establishes the EC, while receive coils read flaw-created

disruptions in currents, represented as phase and amplitude variations in the voltage plane.

The advantages of the X-Probe are its sensitivity to flaw orientation and characterization of

noise signals. The array can travel through tubing at speeds comparable to BCs and much

faster than rotating coils, while still providing spatial and orientation information for both axial

and circumferential cracks.

A group of NDE experts brought the X-probe to ANL to Inspect the NRC steam generator

mock-up. The purpose was to test the probe performance with the mock-up and finalize

analyst guidelines for an X-Probe evaluation exercise. The data acquired were analyzed

without previous knowledge of flaw location. The results are currently being reviewed by ANL

personnel. Documentation, including training manuals and schedules, are currently being

prepared.

Figures 3.39-3.42 compare X-Probe and +Point 3D signal amplitude plots with a 3D

image generated by the Argonne multiparameter algorithm that uses amplitude and phase

information to generate depth profiles. The +Point probe is the rotating probe of choice for ISIs

involving the steam generator TS as well as special-interest locations. The multiparameter

results are derived from phase and amplitude data from a standard rotating pancake coil. Note

that the ANL multiparameter plot provides depth vs. position, while the X-Probe and +Point

plots show signal amplitude and do not represent depth very well, if at all. These images

illustrate the differences in spatial resolution and ability to suppress noise for one of the mockup cracks in the TS simulation.

39
PIA-MO1 40 S P- -7L~4T :

0.77 v/d span 40 rotl79 Uj7-1v d spani40- rot

L!j.y43 volts es dg32vls 8ida 1%-0--

+ 5354 7~+-835-:I

(a) (b)

Figure 2.37. Mag-biased BC Lissajous figures before (top) and after (bottom) corrosion products (thin

oxide film) were formed in a tube with axial ODSCC by exposing the tube to PWR

conditions for about two months. Voltage increased from 2.43 to 3.24 V, suggesting less

contact at crack faces. Estimated depth changed from 87 to 81% TW.

Table 2.7. BC voltages and phase angles for two tubes before and after exposure to

PWR water conditions, resulting in formation of corrosion products. Each

tube has an axial ODSCC that is -1 5-mm (0.60-in.) long.

BC BC

Voltage Phase BC

Tube No. (V] Angle % 1W

SGL-197 (before) 2.43 66° 87

SGL-197 (after) 3.24 760 81

SGL-149 (before) 3.27 87° 73

SGL-149 (after) 4.51 89° 72

The mock-up was also examined with an array probe developed by Mitsubishi Heavy

Industries (MMII) of Japan. This probe combines the speed of the BC with the detectability of a

rotating coil. The 24-channel design includes thin-film pick-up coils. Results of their study

were presented at the 20h Annual EPRI Steam Generator NDE Workshop in Orlando, Florida.

July 9-11, 2001.11 For mock-up cracks, the quality of images from the MHI array are

comparable to images generated by the X-Probe, though the MHI probe exhibited higher spatial

resolution in some cases.

40
120

100

iso8 e0%

t 60

0

Z 40

20

0

I

TSP

LODSCC

2 3 4 6 6

TS TS TSP TSP TS

CIDSCC CIDSCC LODSCC LODSCC CODSCC

Figure 2.38. Maximum depths for a variety of crack morphologies estimated by two recognized EC NDE

experts (A and B) using the same procedure with +Point data at 300 kHz and by Argonne

using the multiparameter algorithm with input mainly from pancake phase and amplitude

data.

Figure 2.39.

77 * X-Probe 3D image of EC signal

l; V .] lamplitude as function of position

around a test section in tube sheet

a no simulation level of NRC steam

I_-,+ . generator mock-up. Tube axis is from

I_+x* #slower left to upper right.

41
IMMIMM-1911--

11L' }_.S. 11 i Z. si:5g 6£t-z .. :hsvi- * ..... d vi<v'-; e. ssv t e 4t i) '>

i Aw"FA  ko

eFigure 2.40.

+Point 3D image of EC signal

amplitude as function of position

around same mock-up test section as

shown in Fig. 2.39. Tube axis is from

lower right to upper left.

Figure 2.41.

3D plot of depth vs. position around

same mock-up test section as shown in

Figs. 2.39 and 2.40 generated by ANL

multiparamneter algorithm. Tube axis is

from lower left to upper left. Improved

spatial resolution Is evident, along with

crack depth Information not available in

X-Probe and +Point images. Phase

and amplitude information are used to

generate these images.

Figure 2.42.

Cross-section profile showing depth vs.

circumferential position for same mockup test section as shown In Figure 2.41

generated by ANL multiparameter

algorithm. Improved spatial resolution

Is evident, along with crack depth

information not available from X-Probe

and +Point Images. Phase and

amplitude information are used to

generate these images.

o a

-11- 1-1- -1.1 I - - 1--1----.-111 1. - 1-1

NM #a -  if&!! K' ILijdang,"

.~............ ..

710----------- .... ..-. -.-.

.0+ -- - .--- ------ 1-

0 50 MO 150 2C 250 3OO

42
2.6 Use of Lamb Waves

The possibility of characterizing the mock-up tubes with UT waves to obtain depth

estimates has been Investigated. Lamb waves were explored as a possible method to help

characterize defects in the steam generator mock-up tubes and to help identify attributes that

may be useful for characterizing defects in field tubes. A specially fabricated Lamb wave probe

with a 6-mm (0.24-in.) piezoelectric crystal on a wedge, operating in a pulse echo mode, was

employed for the results described in this report. Lamb wave inspection was carried out from

the inside of the tube with the tube filled with water and the contoured plastic wedge of the

probe in contact with the tube ID surface. One significant advantage of Lamb waves is their

insensitivity to the presence of roll transitions, making It easier to detect cracks at the

transition of rolled tubes. In addition, Lamb waves are less sensitive to the presence of

artifacts such as support plates and deposits and can be used to detect flaws at a distance.

Figure 2.43 shows the echo from a CODSCC in an Alloy 600 tube, 22.2-mm (0.875-in.)-

diameter, with a laboratory-grown crack in the transition of a mechanically rolled tube.

Ultrasonic diffraction and EC evaluation of the crack suggests a maxdmum depth of 80% 1W.

With the probe on the Inside of the tube and 75 mm (3 in.) from the crack, the Lamb wave echo

Is clearly evident. The crack Is about 1400 in circumferential extent. With the probe on the

opposite side from the crack, only a relatively weak signal from the roll transition is evident.

The echo is visible with only moderate variation in signal amplitude as the probe is moved

along the axis. Thus, by moving the probe a measured amount and noting the change in

transit time of the echo, the Lamb wave velocity can be measured. For this tube, a value of

4800 m/s (15,750 ft/s) was measured at a UT frequency of 1.2 MHz. Rose et al,.1 2 reported

for an Alloy 600 tube, a measured velocity of 4790 m/s (15,720 ft/s) and a theoretically derived

L(0,2) mode group velocity of 4670 m/s (15,320 ft/s) at 1.2 MHz and a wall thickness of

1.25 mm (0.047 in.). Consistent with the current evaluation, Rose et al. noted. that the

frequency of 1.2 MHz was especially effective for water-filled tubes. In general, Lamb waves

have relatively low attenuation, and echoes from defects in the tube wall can, in principle, be

obtained for long distances between probe and defect.

43
Figure 2.43. Ultrasonic Lamb wave echo (top half of figure) from a CODSCC in Alloy 600 tube with

laboratory-grown crack in the roll transition. UT and EC evaluation of crack suggests

a maximum depth of 80% TW.

44
3 Research on ISI Technology (S. Bakhtlari and D. S. Kupperman)

The objective of this task is to evaluate advanced NDE and signal analysis techniques for

the reliable ISI of original and repaired SG tubes. Improved correlations between EC probe

response and flaw morphology, leak rate, and failure pressure will be developed and validated.

In addition, the reliability of the voltage parameter and other EC parameters and techniques

will be evaluated with respect to their ranges of applicability.

The work In the task focuses on four primary areas: (a) analytical prediction of EC

response as a function of probe design, flaw characteristics, and material properties;

(b) development of effective signal analysis procedures; (c) development of flaw imaging and

display methods for simple and accurate flaw characterization; and (d) evaluation of improved

probe designs that may use directional arrays so that defects of arbitrary orientation can be

examined simultaneously. The reliability and effectiveness of the proposed inspection

techniques and the robustness of potential correlations were substantiated through laboratory

testing on the SG tube bundle mock-up and on SG tubes that contain laboratory-grown flaws.

Some additional validation was performed on service-degraded tubes.

3.1 Analytical Modeling of EC NDE

The interaction of induced fields and currents produced by an EC coil with flawed steam

generator tubes is a complex phenomenon. One must distinguish between real defect signals

and those due to background, such as probe response from tube support plates, probe wobble

(liftoff), material inhomogeneities, and thickness variations, etc. Experiments do not always

provide the level of detail and repeatability needed to understand the interactions. A-priori

knowledge of the probe response is helpful in the interpretation of the measured data.

Several analytical and computational EM modeling approaches to the simulation of EC

probe response for NDE of SG tubing have been examined. Most of the efforts focused on the

application of a general-purpose 3-D FEM-based code for steady-state and transient-field

calculations. In the FEM model, the governing EM field equations are expressed in terms of

magnetic-vector and electric-scalar potentials in conducting media and in terms of reduced or

total scalar potentials in nonconducting regions. The field equations are solved by

finite-element discretization. The signal trajectory in the impedance plane due to probe motion

is determined by calculating the response at discrete points along the tube axis. Various

earlier reports 113-161 provide descriptions of the governing field equations, the lumped

element equivalent circuit of the probe and test piece, and conversion formulas for comparison

of numerical with NDE measurement parameters of interest. Results of representative test

cases that simulate steady-state solutions with both differential and absolute bobbin coils are

discussed here.

Test simulations were carried out to verify the accuracy of the FEM approach by

comparing the solutions with detailed experimental measurements. The experimental EC data,

supplied by C. V. Dodd, formerly of Oak Ridge National Laboratory, were obtained from tests on

a large aluminum tube that contained throughwall holes and axial slits. Measurements were

made at three frequencies with a Hewlett-Packard impedance analyzer and a specially

constructed absolute bobbin coil (SN480A). The results were expressed in terms of both

calculated Impedance variations as a function of probe position inside the tube and

45
impedance-plane plots that simulate conventional EC instrument display. Excellent agreement

was found between theory and experiment 1131.

These studies also Investigated the differences in EC probe response to volumetric (e.g.,

hole) and nonvolumetric (e.g., crack) flaw geometries. The calculations demonstrated that BC

signal amplitude cannot be regarded as an absolute indication of the extent of defects when

comparing different flaw geometries. A throughwall hole with a greater volume than an axial

notch results in a smaller perturbation of the coil impedance, which is linearly proportional to

the probe output voltage. Eddy currents always flow through the path of least resistance in a

conducting medium. The discontinuity (infinite resistance) introduced by the thin but long

axial notch forces the currents to take a contour around the defect, which, in turn, gives rise to

a larger impedance mismatch for the probe.

The response of bobbin coil probes to simulated cracks in Alloy 600 tubing was studied.

The effects of the length, width, and ligament size of axial cracks on the EC indications from

conventional ISI bobbin probes were determined. To determine the effect of crack length on

absolute-BC signal amplitude, the absolute-BC probe response to axial slits, 100%

throughwall, on 22.2-mm (0.875-in.)-OD Alloy 600 tubing with a nominal wall thickness of

1.27-mm (0.05-in.) and a conductivity of 100 pa-cm was calculated. The modeled probe was a

conventional 18.3-mm (0.72-in.)-OD absolute bobbin coil. Simulations were carried out at

frequencies off = 100 and 400 kHz. The results were expressed both as current distribution

along the tube and also in terms of change in the magnitude of the probe Impedance, which is

linearly proportional to its output voltage (i.e., related by a constant multiplier that is

proportional to the coil current density). For the test cases considered, the width of the slit

was chosen to be 0.127 mm (0.005 In.), and the length was varied from 0.25 to 50 mm (0.010

to 2 in.). The coil was initially positioned symmetrically in the middle of the slit and the coil's

length and height were taken to be 1.27 and 1.525 mm (0.050 and 0.060 in.). respectively.

The calculations showed that the narrow crack forces circumferentially induced currents

to take a flow contour around the defect. There is a nonlinear relationship between crack

length and EC voltage amplitude that is frequency dependent. For defect lengths < 2 mm

(0.08 in.), the perturbation of the EC absolute-BC signal is small. The slope of the curve

decreases with an increase in defect length. For example, at 100 kHz, the relative variation in

the signal amplitude for an increase in crack length from 5 to 25 mm (0.2 to 1 in.) is more than

five times that due to a change in length from 25 to 50 mm (1 to 2 in.). This saturation occurs

much faster for the differential than the absolute coil configuration. The effect of crack length

on BC probe amplitude is much more significant for deep OD flaws approaching 100% TW.

The effect of length increases with increasing depth of OD flaws. This suggests that a

single-point voltage-based criterion for the determination of OD degradation extent could

result in unreliable estimates for nonvolumetric flaws.

To examine the effect of crack width on absolute BC response, a series of computations

for different defect widths were carried out with the coil located symmetrically under the

discontinuity. The crack length was taken to be 25.0 mm (1 in.) and the width was varied from

0.015 mm (0.0006 in.) to 1.0 mm (0.04 in.). Although the results obtained at higher frequency

show slightly larger variation of the signal amplitude, the variation of the EC voltage amplitude

with defect width over the range of 0.015 to 1 mm (0.0006 to 0.04 in.) is relatively small at both

frequencies (100 and 400 kHz). This finding is in agreement with the concept that conduction

loss is proportional to the path length of EC flow, which is a frequency-dependent parameter.

46
Note that all simulations are based on a nonzero crack width and do not represent a tight

crack with contact areas over the crack faces.

The computations show that the signal amplitude drops exponentially with defect depth.

Also, the response to OD flaws is smaller at the higher frequency in all cases, and the relative

difference in the response between OD and ID flaws Increases as the flaw depth decreases. Due

to higher conduction loss at higher frequencies, the drops In the signal amplitudes as the OD

notch depth decreases are much larger at higher frequencies. Although the resistive and

reactive signal components display different responses as a function of position. the maximum

signal magnitude for all test cases occurs in the center of the defect.

The calculations demonstrate that the calibration of BC amplitude based on 20% 1W

drilled holes on an ASME standard tube does not imply detection of short axial cracks of

comparable depth. For a 6.35-mm (0.25-in.)-long axial notch, the BC probe amplitude begins

to approach that of four 20% 1W drilled holes only for notch depths approaching 80% /1W.

Therefore, under realistic ISI conditions (i.e., in the presence of noise), short, shallow OD

cracks (e.g., <40% TW) could go undetected during BC examinations because of their lower

voltage response. It could be argued that, for nonvolumetric degradations, even the 40% 1W

OD depth may not be a conservative estimate for BC examinations because field-induced

cracks are often much tighter than machined flaws, which could further degrade the

signal-to-noise (S/N) ratio and, in turn, detection capability. The large differences in signal

amplitude between holes and notches can be attributed to both the defect geometry and probe

characteristics. The volume of the four 20% 1W OD holes in the ASME calibration-standard

tube is much larger than that of a notch of comparable length-and depth, and this condition

results in a much greater perturbation of the probe's impedance. The effect of the volume of a

discontinuity on signal amplitude is especially significant for BC probes, where the coil output

at each axial position along the tube represents an integrated reading around the

circumference.

In practice, single isolated cracks are less likely to occur than a series of short cracks

separated by short ligaments. Calculations were performed to examine the effect of such

ligaments on bobbin coil response. The probes modeled were conventional 18.3-mm

(0.72-in.)-OD differential and absolute bobbin coils, and the frequencies considered were 100

and 400 kHz, which are typical for multifrequency ISI of SG tubes. We considered two axial

flaws, 100 and 75% TW, both of which were 12.7-mm (0.5-in.) long and separated by a

ligament of varying length in Alloy 600 tubing the same dimensions and conductivity as before.

Results were obtained in terms of current distribution along the tube and computed resistance

and reactance values as a function of probe position.

We showed that even small ligaments (0.127 mm [0.005 in.J) create a conducting path for

the circumferentlally Induced currents to flow across the discontinuity, which consequently

results in a reduction of the probe signal due to the smaller impedance mismatch. This finding

is more apparent in the impedance plane trajectory (issajous trace) than in the absolute signal

amplitude, where, for example, a 2-mm (0.08-in.) ligament in a 100% 1W axial notch could

significantly reduce the signal amplitude.

47
3.2 Correlation of Failure Pressure with Bobbin Coil Signals

Conventional repair and plugging criteria for SG tube integrity have been based on

estimated degradation depth. Tubes with EC indications that correspond to degradation

greater than a specified limit (typically > 40% of nominal tube wall thickness) are either

repaired (by sleeving) or taken out of service (e.g., plugged). For some types of degradation,

e~g., those characterized by short deep crack segments separated by ligaments, depth

measurements of an Individual crack may not be the best measure of structural integrity. For

ODSCC indications at TSP intersections, a voltage-based alternate plugging criterion (APC) has

been developed. The criterion is based on a correlation between BC voltage amplitude, rather

than estimated defect depth, and failure pressure and leak rate. NRC guidance for the

application and development of such criteria are provided in Generic Letter 95-05 1171. The

correlations between voltage and failure pressure and leak rate were developed empirically from

tests on service-degraded pulled-tube specimens and laboratory-grown cracks.

Multivarlate regression analysis and neural networks were used to develop alternative

empirical models to relate BC probe data with tube failure pressures. This effort was based on

a reanalysis of the data on pulled and model boiler tube specimens that were originally used by

the Westinghouse Owners Group to develop an APC for ODSCC indications at TSP locations in

Westinghouse plants that use drilled-tube support plates. Both multiple linear regression

(MLR) and neural-network models were successful in producing reasonable correlations for the

individual data sets. Signal features, such as maximum and integrated amplitude, were

considered as model parameters. The signal features ultimately selected were the Integrated

amplitudes at multiple original and mixed frequencies. Preliminary studies on pulled-tube

data suggested that an optimal (i.e., minimum number of indications for an acceptable

correlation) calibration/training set should include representative tubes that describe the

range of variation in inspection data that is included from each SG unit. As expected.

prediction confidence intervals were narrower for the model boiler test set. Subsequent

attempts to develop correlation models with a combined set of pulled and model boiler

specimens by using similar EC test parameters failed to produce acceptable prediction

confidence intervals. Eventually, an independent normalization procedure was devised for this

analysis. Note that the robustness of any correlations developed in this study might be limited

by the range of crack morphologles and Inspection artifacts included in the available database.

3.2.1 Multiple Linear Regression and Neural-Network Model; ARC Data Base

Current alternate plugging and repair criteria for ODSCC degradation at TSP

Intersections for 22.2-mm (0.875in.)-dlameter tubing are based on BC voltage amplitude (e.g.,

signal amplitude from 400/100 kHz mixed frequency channel). Multivariate regression

analysis procedures were used to analyze the relationship between BC probe signals for

ODSCC degradation In the TSP area and tube failure pressure. Forty-one indications from the

alternate repair criterion (ARC) database were used in the analysis. The indications were from

22 model-boiler and 19 pulled tubes. All of the examined tubing was 22.2-mm (0.875-in.) In

diameter, with a nominal wall thickness of 1.27-mm (0.05-In.). The pulled-tube indications

were from the first three TSPs on the hot-leg side. The pulled-tube specimens are from three

plants. Table 3.1 lists the number of specimens from each SG unit that were incorporated for

analysis. Table 3.2 lists the crack depth, crack length, and failure pressure from the ARC

model-boiler tube database.

48
Table 3.1. Pulled-tube specimens from ARC

database incorporated for failure

pressure analysis.

Plant SG Unit Indication no.

A 1 1-4

A 2 5-8

B 1 9-11

C 1 12-19

Table 3.2. Destructive examination and failure pressure measurements for

pulled tubes from ARC model-boiler database that contained

ODSCC degradation at TSP Intersections.

Destructive Eaminationa djusted

Specimen Maxdmum Depth LengPressurec

No. (%) (in.) (psi)

528-2 100 0.67 (0.5) 4,668

532-1 100 0.7 (0.52) 4,010

532-2 100 0.75 (0.58) 4.087

535-1 100 0.28 (0.11) 8.265

555-3 100 0.75 (0.42) 4,209

533-4 100 0.34 (0.14) 6,964

536-1 90 0.4 7,959

542-4 N/A N/Ad > 4,867

543-1 N/A N/A > 2,633

543-2 N/A N/A > 3,153

543-4 98 0.52 5,459

547-1 N/A N/A 5,561

576-2 100 0.3 (0.22) 6,628

576-4 100 0.6 (0.43) 5,602

557-1 90 0.18 > 7500

557-2 100 0.52 (0.44) > 4,974

557-4 N/A N/A 11,633

558-1 100 0.4 (0.32) 5,969

568-1 100 0.33 (0.25) 6390

568-2 N/A . N/A > 4,635

571-1 100 0.44 (0.35) > 4,950

574-4 100 0.39 (0.33) 5,506

aMeasurements were carried out by Westinghouse Electric Corporation.

bTotal crack network length, with throughwall degradation length given in

parentheses.

sNormaIzed to 150 ksi flow stress (sum of yield and ultimate stress).

N/A. Data not available.

49
Before a specimen was included in the analysis, the following information about it was

required: an ASME calibration standard data Mfie, a measured failure pressure, and data at a

consistent set of EC inspection test frequencies (400, 200, 100 kHz absolute and differential

channels). Although procedures to correlate EC readings at different test frequencies for

similar probes have been proposed, such procedures were not used to avoid an additional

source of uncertainty.

Figure 3.1 is a histogram of the distribution of measured failure pressure for pulled- and

model-boiler-tube indications. As noted previously, the pulled- and model-boiler tube

populations cluster at high and low failure pressures, respectively, with relatively few data

points at Intermediate pressures. Furthermore, the NDE measurements are affected by such

factors as variations among calibration standards, probe wear, and dissimilarity of

tube/artifact geometry (NDE carried out with or without TSPs in place, presence or absence of

deposits, etc.) as well as actual differences in defect size and morphology.

To reduce the influence of such factors on NDE data, the EC readings were normalized

and frequency mixed. The normalization procedure consisted of automatic phase-angle

adjustment for each channel and amplitude scaling in reference to four 20%, flat-bottom holes

on an ASME calibration standard tube. For 400/100, 400/200, and 200/100 kHz, where the

upper frequency was taken as the primary channel, linear two-frequency least-squares mixes

were used to reduce/suppress the TSP signal. A 400/200/100 kHz mix was also examined

and, based on the regression analysis results, provided a small improvement over the

400/100 kHz mix. The mixing algorithm was carried out for all data files regardless of whether

or not the Mfie contained a TSP indication. The resampled data points for the mixed signals

were chosen to cover the entire support plate region.

A series of computer codes was written to perform all the manipulations of NDE data,

such as searching each Mfie for proper channel configuration, selecting data segments that

contain indications at the TSP region, normalizing, resampling, and constructing

mixed-frequency channels.

The integrated area under each TSP indication was determined by the equation

, = - QH n+ t0VnD, (3-1)

where S represents the integrated signal magnitude at frequency f for each absolute A and

differential D channel, and H and V are the horizontal and vertical signal components at each

frequency. The integrated region was bounded by N /2 points on either side of the TSP center

(n = 0), with the sampling rate of each data segment being adjusted in accordance with the

simulated TSP ring from the corresponding ASME tube standard.

Although the integrated TSP signal of Eq. 3.1 could be correlated reasonably well with the

tube failure pressure when the model-boiler and pulled-tube data sets were considered

separately. the correlation was poor when both types of specimens were included in a single

database. Consequently, additional transformations of the signals were explored. The selected

transformation involved mean-centering and energy-scaling of each data segment

50
4 .. . . . . . . . .. . . . . . .. . . . .. . . . . . . .. . . . . I'.... ...._ .. . . . .

3 . . . ... ... . . . . . . . . . .. . . . . . . .... .

2 .. . .. . . ... . . . . . . . . . . . . ... .. . . . . . . ................

2 _ _ ~~~. _. ......... .............

110 20 30 so 50SO 70 80-

Measured Pressure (MPa)

5.

Figure 3.1.

Distribution of pulled-tube (top) and

model-boiler (bottom) indications

as a function of failure pressure

used for multivariate analysis.

4

12

1

.. HI.H. ..................

.. . . . . .. . . .... .. ''~~ ~~~.......

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

w............ ........... ..............................

..._ ............................. ...............

:n nmnnn ............ . 10 20 so 40 50 60 70 80

Measured Pressure (MPa)

independently. Mean-centering is performed to reduce the effect of signal offsets;

energy-scaling, to render equal weight for each data segment. As in the case of Eq. 3.1, the

integrated area under the transformed signals was determined by

CO-Al D =N + i e (3.2)

n=-N/2

where the transformed H and V components are defined as

.kn = (Xn - 1Xx)/4x, (3-31

where p represents the mean and 4 is the signal energy. For an equal-ength mean-centered

signal component 4x is proportional to the standard deviation ox of each trace.

Transformation in accordance with Eq. 3.3 will retain the shape of the original impedance

plane trajectory (i.e., a Lissajous trace). Training and prediction schemes were subsequently

employed to evaluate the ability of each correlation model to capture all of the useful trends in

the data.

Figure 3.2 displays the multivarlate regression and the neural-network fits when all

samples are incorporated for the model-building stage. Also shown are the upper and lower

bounds of the 95% confidence Interval, calculated through a first-order least-squares fit to the

measured-vs.-predicted pressure. As expected. the neural network (Fig. 3.2b) can better

capture all the variations In the data. However, this test case does not represent the prediction

capability of the model but rather the ability of the model to construct a fit to all of the data.

Figure 3.3 displays the prediction results of the two models when - 50% of the indications from

each tube category are included in the training stage. For this test case, comparable

confidence intervals were achieved by both methods. Finally, Fig. 3.4 displays the prediction

51
70

450

2 C

* .1~~~~~~......

K. -~~~~~~~~~~~~~B

80

70

.~~~ . - -

.1. 0 ..20.....0 . . .... .80 ~~~' 7 --~~~~

.. -.......... ............... ............ ..............- ..............- , ...,:..-x..

.. . . . . ................. r ................. i. ;/;...-;........................!

X.- : ,'.K - ' ' 

~~~~~~~~.......... ...... .. ..... ... ...... ..-. ... .. .... ..

I ~~~~~~~~...... .. :............e ..: ::.........;

........... _.. : . .. . ... . . . . ....... :.'.. .... . . . .: . . . . .

201

10

10 20 .0 .0 .0 . . 10 20 30 40 50 -- O V0

Mewasuede wns

(a)

so 10 20 30 40 50 eo 70 so Mmaud Pressw (MP)

(b)

Figure 3.2. (a) Predicted vs. measured failure pressures for (a) regression and (b) neural-network fits

when all model-boiler (x) and pulled-tube (*) indications are included for training. Also

shown are the upper and lower bounds of the 95% confidence interval, based on a linear

least-squares fit to the data.

80

70- ..........

la, I 50

30

20

10

++ + t

+ 44

.4. ~ ~ ~ ~ ...... ...... ... ..

70

980

1530

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

. . ........ ........ . . ...... ..-:.+. ..+: ....... _'.

. ,.c ~~+ : .+ ,.,

.. .! . ..... .... : . . ...... ...... ....... ...-.- .. .i..

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

.. . . . . . . . . . . . . . . . .. . . . . . . . . ............ '.. .. . .. . . . .

; ; ; ; ; ; ; ;~~~~4

..... :.- ......... o.*'.

. : ."I

....... : .... I ....... ...... I

.1 :

........ I ......... .:.... .

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

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

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

201

10

10 20 30 40 50 60 70 so Meaiarad Praessu (MP)

(a)

Measured Prmsw, (PAW

(b)

70 so

Figure 3.3. (a) Predicted vs. measured failure pressures for (a) regression and (b) neural-network fits

when = 50% of model-boiler and pulled-tube indications are randomly selected for training.

Also shown are the upper and lower bounds of the 95% confidence interval, based on a

linear least-squares fit to the data.

52
-

.... ....... A... : ....... . . . I ,.:':

*70 . 70 .yEj. >, .. ... 70 . ... ........ .... ....... ..... ..., ::. 70 ..x.. .. ................... .. .... . . ... : ........ .. x . . . . .. .. . .. .

} ...,.........oRX.. ..... .x. ,..... ..,,.... ..,....,.....,,... , ..,, ........ 1, ,.._,-

X XX~~;

S°-'' X '''''I ' X 8 X -/XX

________________________________________________________________ .......L.. , " ,X ,.........................x ............ ' 40 .......... ...... . f-,0r--...... ............ ............ 4 . . .. I:

10 20 30 40 8g0 60 70 80 10 20 30 40 60 80 70 80

Measured Pressurel (MWa) Mleasured Pressure (MPa)

(a) (b)

Figure 3.4. Predicted vs. measured failure pressures for (a) regression- and (b) neural-network fits for

model-boiler (x) and pulled-tube (a) indications, with LOOT training and prediction method.

Also shown are the upper and lower bounds of the 95% confidence Interval based on a linear

least-squares fit to the data.

results of the two models based on the leave-one-out training (LOQI) and prediction scheme.

Once more, the confidence intervals for both methods of fitting show comparable prediction

accuracy.

Figure 3.5 is a plot of residual vs. regression-predicted failure pressure when all

model-boiler and pulled-tube specimens are included in the training. The distribution of

residual pressures does not display any distinct pattern, a finding that is consistent with the

assumption that a normal distribution of residual pressures is implicit in the calculation of the

confidence interval. Figure 3.6 shows the changes in residual failure pressures when the

LOOT training and prediction procedure is performed. For multivarlate models, the variation of

residual pressure obtained with the LOOT procedure is an indicator of model stability. For a

stable model, the predictions should be relatively independent of individual data points in the

training set, i.e., the residual pressures in Fig. 3.5 should be small when compared with the

predicted failure pressures. In the present case, the change in the residuals is =10%6. For this

type of data, such a change is reasonably small.

Figure 3.6a shows the cumulative distribution of the residual pressure, assuming that it

is normally distributed when the training set includes all of the available indications.

Figure 3.6b shows similar results when a LOOT training/prediction procedure is implemented.

The standard deviation of the distribution when the LOO)T procedure Is followed is = 50%o

greater than when all of the data are used to fit the model. The LOOT training scheme may

provide the most conservative estimate of the accuracy of the predictors.

X.-~~~~~~~~~~~~~~~~~~

A study was also carried out to indirectly assess the prediction accuracy of the

multivarlate regression model by comparing the results presented above with those obtained by

a linear regression model. The data set used in this work was developed by renormalizing the

EC inspection results that were used to develop the voltage-based APC model. The

renormalization was performed by referencing all measurements to a single ASME tube

53
20~

is

i 11:

1 0

Io15 1 -'

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

I

. .:...x . : : x H- 7 

. . . . IN .

.,,. ,. ... .:.... .... .. .. . .. .. . . .. .. . ...... . ....... ..... -- l Ei. ..

''''' . .. . . ...... .. ... ... .. ......... ....................... .... ... ......

Figure 3.5.

Residual vs. regression-predicted

failure pressures when all model-boiler

(x) and pulled-tube (*) specimens are

included for training. -10 P

-15i

-20

-25 . . . . . . . .

10 20 30 40 50 1

Regressio Presaum (MPa)

50 70 80

lrr. 100 .r . . r . . r

901

70 I 80

1 30

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

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

..........

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

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

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

I-;

ee.. ............... ....... ..... o... 0

0

0p....

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

so

9 70

1 30

20

la

0 C . .. .... ... .... . P oI....

.0o

-20 .~ -15 - -1 ., 0O 5 10 - *

~~~~~~~~~~~~........ .. ;,,A........-.................. ..... ................

~~~~~~~~~~~~~~~~~~~~~~~. '.'. °..'-..'.'

~~~~~~~~~~~~~~~~. .°° '.... . '... . . 201

10

-2 :i : : 0 -a 0 1 -0 2

Resikwj Pressure (MP&)

(a)

Resldui Pressure (UPa)

(b)

15 20

Figure 3.6. Cumulative distribution of residual failure pressure. Normal distribution function was

assumed, with training and prediction based on (a) all indications and (b) LOOT.

standard. A linear least-squares fit to the logarithm (log 0) of the measured voltage-vs.-fallure

pressure was implemented in accordance with the APC-recommended model. Although a

direct analogy between the APC linear regression and multivariate models cannot readily be

described, comparison of the predicted residual pressures suggests a tighter global prediction

for the multivariate model.

The failure pressures predicted by the APC linear regression model vs. measured failure

pressures are shown in Fig. 3.7. Also shown is the 95% confidence interval for the first order

least-squares fit of the measured-vs.-predicted failure pressures. A linear correlation

coefficient R of 0.88 was calculated for this fit. Figure 3.8 is a plot of the residual failure

pressure vs. regression-predicted pressures. Once again, the distribution of the residual

pressures does not display any distinct pattern. Assuming that the residual failure pressures

are normally distributed, the cumulative distribution of the residual pressure is shown in

Fig. 3.9.

54
s o - .1 .. .... . .. . .. . . .. .. ... . . .. .. . . . . . .

7 0 -... -. ..-.... .. .. . ........ . ... . . .. . . . . . .. . . . . .. ,.. .

o .. : . ........... Figure 3.7.

Regression-predicted vs. measured

so -- ': . ' :.' x-failure pressures for all model-boler (x)

. , . .and pulled-tube (*) Indications, using

40 - - the APC model. Also shown are the

. . x ,. . . upper and lower bounds of the 95% x3 _ ...i.._X -. - , ........... . ' ........... .i.-...--.i-....... .'---...'...x confidence interval based on a linear

least-squares fit to the data.

10 . ... . . .. . .. ,.. .. .... .. .. . .. ...... . . . .

10 20 S0 40 50 60 70 e0

Measured Presse (MPa)

25. ' I . '

20

15

10 - :

0 ... .... ....... ..' ..... ....... ... . ....... ......

..... ......... Figure 3.8.

X x 8 Fx ( # ' x # M 'Residual vs. regression-predicted

failure pressures for all model-boiler (x)

.......... Sx 5 ' x ' .and pulled-tube (*) specimens, using

-10 .............. the APC model.

*15 ,

-1 . . . ... ... . .. . . .. . . . . .. . .. . . . .. . .. . . . ..... . . . . . . .

-20

10 20 30 40 5 o 60 70 80 Regsson Presur (p )

100 gCo .. . . . .. .. . .. .. . . . . . . . .... . . . . . .... . .. .. .. .. . . .

80 . .......

. . ................ Figure 3.9. o

w : ............... .............. Cumulative distribution of APC model

residual failure pressures, based on an .. 40i. V ^ ..... , assumed normal distribution function.

3 0 o. .. . . . .. . . . .. . . .I. . . . .. .... . . . .. . . . ......... . . . ... .. . . .

2 0 .. .. .. .. . . . .. . . . .. . . . .. . . . .... ... I. . . . .I . . . . .........-- -----'---

. wr . ; . B

-2 -20 -15 -10 4 0 5 10 15 20 25

ResWW Pressure (mPg)

55
3.3 Computer-aided Analysis of Eddy Current Inspection Data

Manual analysis of multiple-frequency EC data is a tedious and challenging process. At

present, no qualified technique, manual or automated, exists that provides reliable estimation

of flaw size over a wide range of SG tubing damage. Conventional data analysis methods

become rather subjective when dealing with complex forms of degradation such as stress

corrosion cracking. Signal distortion by interference from internal/external artifacts in the

vicinity of a flaw further complicates discrimination of flaw signals from noise. In this context,

noise refers to any random or coherent signal that is not induced by a flaw.

In comparison to high-speed BC inspections, high-resolution multicoil rotating and array

probes generate enormous amounts of data over comparable scanning lengths. Rotating probe

ISI of SG tubing is thus generally restricted to areas that are historically predisposed to known

damage mechanisms and sections of particular interest that are flagged by the initial BC

examinations. More extensive application of such probes for improving NDE reliability rests, In

part, on automating various stages of the data screening process. Computer-aided data

analysis is a viable means to overcome many of the challenges associated with reliable

processing of data acquired with high-resolution probes.

A key objective of the work has been to characterize flaws in the ANL steam generator

mock-up in order to minimize the need for destructive examinations and to further assist in

parallel studies on prediction of tube structural integrity from NDE estimates of flaw proffle.

Application of the methods developed has been limited so far to analyses of data from rotating

pancake coils; however, many of the fundamental processes described here are applicable to

other probe geometries and coil configurations.

3.3.1 Data Processing Structure

Computer-aided data analysis here refers to software to carry out various stages of

processing raw EC inspection data. Figure 3.10 depicts the basic structure of the algorithm

developed to allow for off-line analysis of data acquired with a commercial EC testing

instrument. Blocks associated with the processes mentioned in this report are numbered. The

output of the acquisition block shown in Fig. 3.10 is the digitized output of a

multiple-frequency inspection system. The EC testing instrumentation and software used at

ANL are a Mlz-30TM (Zetec, Inc.) remote data acquisition unit that is controlled under the

HP-UX-based Eddynet 1

m (Zetec, Inc.) environment The raw inspection data are converted to a

standard file format. All data analysis and signal processing algorithms were written using the

PC-based software MATIAB, which is a high-level scripting language that provides an efficient

environment for data manipulation and computation, together with convenient graphical user

interfaces (GUIs) and graphical displays of the results.

The block diagram shown in Fig. 3.10 can be divided into three basic components: data

conversion, calibration, and analysis. Each block is described in the following sections. In the

conversion stage (block 1) digitized recordings of inspection data that sequentially represent the

in-phase and quadrature signal components are converted to a readable format for off-line

manipulations. In the subsequent stage (block 2), multiple-frequency raw EC data are

calibrated for all the recorded frequency channels. Finally, in the data analysis stage,

(blocks 3-6). calibrated data are processed to produce depth estimates for the flaw indications

in a selected test section of a tube.

56
Calibration

Amplitude scaling [

Phase alignment

: ndustry format

to

Standard format

L

Acqulsition

Multiple frequency

: EC Inspection 1

Pre-processing

Data alignment

Pre-conditionig.

Artffact suppression

Resolution enhancement

la r [14 Visualization

Image..: ,a e9

Strip chart

jssajous

Isometric

Sizing Profile L

rt

Detectlonldentificatlon

SIN calculation

Signal detection X

Rule-based Identification:

I

_A

- .. w .rb ~ . ..

Post-conditioning :

LS multiparameter fitting

* Data segmentation -Ij

Figure 3.10. Schematic diagram showing basic structure of computer-aided data analysis algorithm.

Blocks associated with operations described in this report are numbered.

.3.32 Conversion and Calibration Routines

A series of algorithms has been implemented under the MATLAB environment to carry

out the conversion, normnalation, and reformatting of the raw EC data. An Interactive MATLAB

program. ECMWoL calls the data retrieval and calibration routines through a single GUI control

window. This software currently retrieves and processes inspection data acquired with either

the Zetec MIZ-18 or MIZ-30 EC inspection instruments. With appropriate modifications, the

software could also be used to convert inspection results from other standard commercial

instruments.

The main control window contains virtual instrument controls such as pull-down menus,

editable text, and push buttons that activate various functions for the manipulation of the NDE

results. Data files In Eddynet format are initially read by these codes. The data conversion

routine extracts the essential header information (currently assumed to be from MIZ-18 or

MOZ-30 instruments), such as the number of channels and their associated frequencies, from

the original Eddynet-formatted Mles. The decoded header information contains the frequency

and channel configuration that is subsequently used to sort out raw EC readings. The

reformatted data matrix, along with the header and coil configuration information, is then

57
stored in a user-defined data file. Normalization values of all the available channels are

calculated by using the Inspection results from a calibration standard tube. These values

consist of amplitude scaling factors, phase angle rotations, and null values that are

automatically calculated based on user-defined approximate locations of known indications on

the tube. Although the necessary information for standard normalization procedures is defined

as a set of default values in the program, the code also allows calculation of these values in an

arbitrary fashion. Finally, the raw EC readings are calibrated by another code that is also

activated from the main menu and applies the previously calculated normalization factors to

each new raw data file. To avoid incorrect application of normalization factors, the codes

initially check the header information for consistency between the original calibration standard

tube and new data files that are subsequently loaded.

3.3.2.1 Data Conversion

Eddynet provides inspection data in MIZ-18 and MIZ-30 formats. The header sizes for

MIZ-18 and MIZ-30 are currently preset to contain a fixed number of bytes. Following the

header is a data stream that sequentially contains the horizontal and vertical signal

components for all channels that are stored in 2-byte signed integer format. The necessary

information can be extracted from the header to carry out the data conversion. This

information reveals whether the data are in the MIZ-18 or MIZ-30 formats and the numbers of

channels in the data file. Once these two pieces of information are determined, the conversion

routine can retrieve and convert EC data files to. a format suitable for any future

manipulations.

3.3.2.2 Automated Data Calibration

Calibration of raw EC data plays an important role in any data analysis procedure.

Conventional phase-angle calibration for normalizing EC data is carried out manually and

routinely involves visual alignment of the Impedance-plane signal trajectory (i.e.. Lissajous

trace) with respect to some reference indication. This implies that the extent of background

fluctuations and S/N ratio, as well as analyst judgment, could affect the calibration process

and consequently lead to varying estimates of a desired parameter. Because estimates of flaw

depth from phase angle information of multlfrequency inspection data depend heavily on initial

calibrations, computer-aided data calibration routines play an essential role in uniform and

accurate normalization of raw EC data.

Signals from machined flaws on a calibration standard tube are used to normalize the

amplitude and phase of raw EC data. Amplitude scaling is performed with respect to peak

signal voltage from a known indication on the calibration standard. Alignment of the phase

angle at each frequency is typically made in reference to either a shallow ID or a throughwal

flaw.

Figure 3.11 displays the user interfaces along with intermediate and final stages of

calibration of raw EC readings collected with the 2.92-mm (0.11 5-in.)-diameter primary

pancake coil of a three-coil rotating probe on a 18-notch standard. The code sequentially asks

the user to interactively provide information on the approximate positions of lift-off, a sound

portion of the tube, and an indication for amplitude scaling; it calculates the phase-angle

rotation, null value, and scaling factors for uniform normalization of all the data files that

belong to the same calibration group.

58
Figure 3.1 1.

(C)

Displays of GUIs for (a) main ECTool menu with data calibration submenus, (b) raw

data and trigger channel, and (c) calibrated data in strip chart and Ussajous display

format. Eddy current data were acquired on a 1 8-notch standard with a 2.92-mm

(O.115-n.)-diameter primary pancake coil of a three-coil rotating probe.

59
3.3.3 Signal Processing, Data Analysis, and Display Routines

In reference to Fig. 3.10, the data analysis section (shown by blocks 3-6) has three basic

stages: pre-processing, flaw detection and identification, and post-processing of data. These

blocks consist of various scripts that successively perform the calculation of S/N ratios for all

channels, apply filters for pre- and post-processing of data, and ultimately combine

multiple-frequency information from all processed channels to provide an estimate of flaw

depth for the entire tubing test section. Both frequency and spatial domain filters are

incorporated for signal conditioning, baseline reduction, and resolution enhancement. Initially,

the S/N ratio is calculated from a user-defined approximate location along the trace baseline

and minimum detectable amplitude from a calibration standard tube. Subsequently, this

information is used to implement a series of filters that perform suppression of artifacts,

baseline fluctuations, and signal enhancement or restoration. The filter characteristics take

into account both the coil configuration and the sampling frequency of the inspection data.

The identification of flaws and their origin, i.e., OD or ID, is done through a series of rules that

are applied to the multiple-frequency EC data. Both amplitude and phase relationships among

the processed channels are used at this stage of the process. Finally, the phase information at

multiple frequencies is combined to calculate the flaw depth for the tube test section in

reference to known indications on a calibration standard tube. Multivariate analysis using the

available information from multiple channels is used for estimating the NDE depth profile of

the tube.

A series of algorithms has been implemented as MATLAB scripts to provide the profile of

flaw depth in an SG tube test section from NDE inspection results. These codes are executed

through a user interface tool to automatically process the EC data. Figure 3.12 shows the

main window of the GUI tool ETProf, which incorporates various algorithms for the processing

of raw EC data. Figure 3.12 also shows several forms of graphics currently in place for the

visualization of data at various stages of analysis. Subsequent to calculation of the relative

depth profile, estimated values are converted to percent of tube wall thickness. Reconstruction

of helically scanned data into C-scan format allows for the observation of sizing results from

any azimuth and elevation view angle and for any axial or circumferential cross section of the

tube. Scaling of data in axial and circumferential directions permits direct deduction of the

flaw extent along the tube.

3.3.3.1 Signal Conditioning

Pre- and post-conditioning of data are some of the most fundamental operations in any

signal processing application. The aim of such data manipulations Is to minimize the

contribution from known sources of signal distortion. Coherent noise that is present at

multiple channels occurs commonly in EC testing of SG tubing. Low-frequency baseline

fluctuations and high-frequency noise are also always present, to various degrees, in any EC

inspection data. Typical sources of low-frequency signals, particularly for absolute impedance

coils, are probe wobble, probe wear, temperature fluctuations, and gradual dimensional and

geometry changes of the tube. Tube ID conditions and electronic noise, on the other hand, are

typical sources of high-frequency noise. For both these sources of signal degradation,

frequency domain filters can be designed to effectively suppress unwanted signals. Two types

of filtering have been incorporated into the data analysis routine to reduce the level of such

noise. They consist of standard Fourier domain filters and the more elaborate finite-impulse

60
Figure 3.12. Series of MATLAB-based GUI tools (top four rows) for computer-aided analysis of EC

Inspection results acquired with standard commercial Instruments. Also shown (bottom

two rows) are several graphical display formats such as Image, terrain, and

cross-sectional NDE profiles of tube test section at various stages of analysis.

response (FIR) filters. Filter characteristics are determined from the data sampling rate and

the frequency content of noise. The FIR filters used to serve the same purpose are briefly

described below.

Frequency domain bandpass FIR filters are nonrecursive filters that offer certain

advantages and limitations over recursive infinite-duration impulse response fuR) filters. The

primary advantages are their linear phase, stability, finite-duration transients, and efficiency of

implementation. FIR filters are always stable, but require longer sequences to meet the same

requirements as their HR counterparts. This latter property, a direct consequence of their

simpler transfer function, often translates into more computation time for nonrecursive filters.

As with all frequency-domain filters, when conditions are not Ideal, the increase in S/N ratio

gained by the filtering process is always accompanied by some degree of perturbation of the

amplitude and phase of the original signal.

The time-domain difference equation for a linear filter with input sequence x(k) and

output sequence -y(i could be written In a general form as

61
M N

y(n) = E bpc)x(n - k) + Y alk)y(n - k), (3.4)

kJ=O k-=I

in which the coefficients a and b define the filter response. The transfer function of the HR

filter in z domain, which is the z-transform of the time-domain difference Eq. 3.4, can be

described by

M

Y(Z) F, b(k)Vk

H(z) = X(Z) kON (3=5)

aD + , )zk

where ao a 1 for the case of a FIR filter, and the summation in the denominator of Eq. 3.5

vanishes. The difference between the two filtering tIpes can be seen by examining the filter

frequency response, which is a rational function in z for an ER filter and represents a discrete

Fourier transform for an FIR filter. Regardless of filter type, unless the data sequence in the

Fourier domain is composed of sinusoidal components of the desired signal and noise that are

well separated, the amplitude and phase of the filtered signal will always experience

perturbation with respect to those of the original signal. This is of particular concern in the

case of EC signals in which the frequency content of the entire data sequence is often closely

spaced and thus requires very sharp (high-order) filters to separate real signals from

background fluctuations. Filters for EC signal conditioning often require a sharper transient

and, consequently, a larger number of coefficients, because much of the useful information

from the Fourier domain spectrum is at the low-frequency end. Studies so far have suggested

that frequency domain filters are generally more suitable for improving detection probability.

Phase angle perturbations that are a result of filter application can often lead to more complex

impedance-plane trajectories, which may complicate conventional data analysis procedures.

3.3.3.2 Signal Enhancement and Restoration

Two basic approaches for improving data quality are the signal enhancement and signal

restoration methods. Signal enhancement refines the quality of data without requiring

knowledge of the degradation phenomenon. Signal restoration, on the other hand, is founded

on the assumption that the degradation source is well understood so that an inverse process

could be implemented to recover the true response. The principal objective in both cases is to

treat the data so that the result is more suited for a specific application than the original data.

Although the two approaches are clearly differentiated in most standard data restoration

applications, distinction between the two becomes less apparent in practical applications

associated with the EC testing of SG tubing. This is because of the difficulty in accurately

modeling the underlying degradation phenomenon, which arises from the combined influence

of many factors, and because of the blurring effect of the coil impulse response. Both data

enhancement and approximate restoration schemes have been incorporated into the

multiparameter data analysis algorithm. Comparable results can be achieved by the two

techniques for the processing of EC rotating probe data. Signal restoration, referred to here as

approximate deconvolution or inverse filtering, Is implemented in the frequency domain, while

62
signal enhancement is implemented directly in the original spatial domain. Although

deconvolution-based methods are generally expected to produce more accurate results, they

are also computationally more intensive.

Deconvolution techniques are used in a wide range of signal processing applications,

primarily to recover signals distorted by the sensing environment. In EC inspection

applications, pseudoinverse filters could be effective for enhancement of spatial resolution that

is degraded by the finite spread of the coil-induced field. Better separation of flaw indications

from extraneous signals could, in turn, improve the estimation of flaw depth determined from

the phase-angle information of multifrequency data. Frequency-dependent signal restoration

could also help reduce differences in probe response at different frequencies. This could be

particularly beneficial for multifrequency mixing techniques in which the differences in the coil

impulse response at the base and auxiliary frequencies could significantly contribute to the

level of mix residuals.

Data enhancement and restoration techniques in application to EC signals are limited

because the flow of current in a conducting medium is governed by a diffusion-type equation.

Distortion is not simply a modulation of signal amplitude. Instead, the wave form could

undergo a complete alteration of structure that is not linearly dependent on Its original form.

Another factor that could significantly influence the degree to which a signal can be recovered

is the lack of separation between the spectral content of flaw indications and extraneous

signals. Internal/external artifacts and design discontinuities (conducting and magnetic

deposits, tube dimensional variations, external support structures, etc.) can produce signal

trajectories with spectral components that are close to those from flaw Indications.

Furthermore, practical sampling rates that are typically used for acquisition of ISI data with

rotating probes do not yield the continuous smooth signals that are essential for optimal

restoration by inverse-filtering schemes. Finally, it is important to note that physically

realizable optimal inverse filters are often unstable. For this reason, approximate

deconvolution algorithms may provide the best alternative for real-time restoration of EC flaw

signals for practical ISI applications.

Signal Enhancement- When the exact form of the degradation process can only be

approximated, signal enhancement techniques are often the first choice for the treatment of

data. Because such techniques can be implemented directly in the original domain, they are

computationally efficient. Data enhancement is achieved by applying pre-stored kernels to the

EC C-scan data. Both smoothing and peak restoration are attained In this manner. The

kernels, In effect, perform least-squares polynomial fitting of the data. Because of the

availability of Information around any center pixel, a symmetric convolution process is applied.

Test case results on signal restoration by 2-D kernels are shown in the following discussion.

Signal Restoration- Degradation of the probe's spatial resolution in two dimensions could,

in general, be described by a discrete convolution model. With prior knowledge of the distortion

model, the original signal may then be recovered through deconvolution, either directly by

polynomial division in the spatial domain, or by transformation into the frequency domain.

Test Case Results on Signal Restoration for Machined Specimens- Preliminary evaluation

of the 2-D deconvolution scheme that is integrated into the multiparameter data analysis

algorithm was carried out by using a subset of laser-cut notch specimens. The geometry of all

63
the available laser-cut specimens and associated flaw dimensions is discussed in detail in

Ref. 16.

Figures 3.13 and 3.14 display the results of inverse filtering on two samples from the

subset of laser-cut tubes with one or both types of 2-D Lorentzian kernels. In all cases, the

restored data show improved spatial resolution and reduction of the baseline level. It is

important to note that the primary objective of the proposed deconvolution scheme as a

preprocessing tool for multiparameter analysis of EC inspection results is to maximize the

separation between crack-like indications and extraneous signals with minimum loss of

information from the original data. For all test cases shown here, the results also indicate less

ambiguity in the measurement of the phase angle of restored indications in comparison to

those from the original readings. A more detailed description of analyses of EC inspection

results using bobbin and +Point" probes, as well as multiparameter data analysis, was

provided in our earlier report [161.

Test Case Results on Slgnal Restoration for Laboratory-Grown SCC Specimens- Eddy

current readings on a collection of 22.2-mm (0.875-in.)-dlameter Alloy 600 tubes with

laboratory-produced cracking were analyzed with the multiparameter data analysis scheme

described in this report. Flaws in this small set of six samples consisted primarily of OD axial

cracking and circumferential cracking in the free-span and roll-transition regions, respectively,

plus a single specimen with axial ID cracking at a dented TSP region. Multiple-frequency NDE

data used in this study were acquired with a standard three-coil rotating probe containing a

2.92-mm (0.115-in.) pancake, a mid-range +Point. and 2.03-mm (0.080-in.) high-frequency

pancake coil. The primary pancake coil readings were utilized for analyses. The calibration

standard contained 18 EDM notches of axial and circumferential orientation originating from

the OD and ID of the tube and ranging in depth from 20% to 100% 'W. All notches were

6.35 mm (0.25 in.) in length. Data analysis results are presented as an isometric display of the

relative depth for the flawed region of the tube. Estimated flaw sizes are shown both with and

without signal restoration by Inverse filtering.

Figure 3.15 shows the results of analyses for an axial ODSCC with estimated maximum

depth of -80% 'IW. Comparison of the restored sizing data in Fig. 3.15(b) with that in

Fig. 3.15(a) suggests the crack to be composed of a series of axial segments with the one

segment being nearly separated from the rest. Similar results are shown in Fig. 3.16 for

another specimen with an axial ODSCC flaw. Terrain plots of the relative depth, however,

suggest the orientation of the crack to be approximately at 450 with respect to the tube axis.

Maximum flaw depth in this case was estimated to be - 80% 'W.

Processed data from other examples (161 clearly indicate complete suppression of the roll

transition, which is the dominant signal feature in the amplitude trace. The terrain plots of the

restored data in one case suggested two circumferential cracks, both composed of multiple

segments. The results of analysis for an ID axial crack in a dented region at the TSP

intersection were also provided. In this case, a terrain plot of the defected segment was

constructed from the estimated ID flaw size. Some improvement of spatial resolution followed

signal restoration. Once again, suppression of the TSP and dent signals was clearly evident in

both cases (161.

64
i

40

1n

ckc.4r. 0°°0 - 5 -1 4M 0 0.5 I 1.5

(a)

140.

120, 1 . .. . . . . . . . . . . . .. . .. .. . .. . . . . . . . . . . . .

40

60 50~~~~~~~~~~~~~~0w.

arc..*. O adr. '-1 4.5 0 0.5 1

(b)

Figure 3.13. (a) Original and (b) restored signal for Type-3 laser-cut specimen #5528-3-3 with

nominal OD flaw depth of 40% TW. NDE maximum depth estimates were 42% by

bobbin coil (phase-angle) and 61% by multiparameter analysis of RPC (2.92-mm

[0.11 5-in.] pancake) data.

65
140

120

100

tso,

I eo

40

20 2

140

20

120]

1001

so,

! c]

2.5.. . . .. . .. .

Z~ ~~~~~~~~~~~~~~~~ '', i' . ..' ~~~~~~~~~~ .'' . ....... .,.. 1 .5 -- ,

0.5 .. . ......... .... ,, ... ....... ........ ........

O , -.!............ ........... .... .....

-1 i . ; ; . ; ; .2 -1.5 -1 40.5 0 0.5 1 1.5 2

Hon.

]'7

30

1-0 '- Ax.-dr.

-40 s a

(a)

40

CbC-0. An.-&.

(b)

2 5 .. . . . ... . .. .. . . . ... . . . .........

1402

120:

100_ 15 . . . . . ....

40 0.

20

J00

So~~~~~~~~~~~~~~~~~s

40~ ~ 4

20 2 00 20 Sc10 

2 -1.5 -1 -0. O n5 1 1.5 2 Axdh. H-nqX

(C)

Figure 3.14. (a) Original and (b)-(c) restored signals with 2-D Lorentzian kernels for Type-4 laser-cut

specimen #5469-2-1 with nominal OD flaw depth of 80% TW. NDE maximum depth

estimates were 45% by bobbin coil (phase-angle) and 74% by multiparameter analysis of

RPC (2.92-mm [0.1 15-in.] pancake) data.

66
* I. I~~~~~~~~~~~~~~~~~~~~~~~~~~~~~11C1I

----------- I-------- "II .-

0.8 - : I

0

10 -~~~~~~~~~ ~ Af 50

0*T'-0

(a)

1 I-~~~,__~~r~~ ' I... a S 

0.8I + .. .------- -- --------- ""

(b)

Figure 3.15. Data analysis results for laboratory-grown specimen with longitudinal ODSCC showing

(a) terrain plot of relative OD depth for cracked zone and (b) restored data by inverse

fittering. Maximum crack depth Is estimated to be =80% TW. Isometric display of results

shows finer details in restored data.

67
U I."M* -. . II

----------- ~ ~ -

0.6 ----------- ~ ~ ---- a

-~~~~~ S~~~

0.82

0. ... ... .%

30~~~~~~~~:' 4

0 *~~~~~~~~~~~~6

U -o

(a)

(b

0i6,

-a:~~~~~~~~~~~~~~~~~~~~~~~~~~%

* {~~~~~~~~b)

Figure 3.16. Data analysis results for laboratory-grown specimen with skewed ODSCC showing

(a) terrain plot of relative OD depth for cracked zone and (b) restored data by inverse

filtering. Maximum crack depth is estimated to be =80% TW. Isometric display of results

shows finer details in restored data.

68
3.3.4 Multiparameter Data Analysis

Multivariate data analysis algorithms serve as the foundation for various operations in

the processing of EC data. Three algorithms were investigated in this program and have been

adapted for use in various areas of research associated with the EC testing of SG tubing. They

consist of multiple linear regression (MLRI. principal component regression (PCR, and partial

least squares (PLS) techniques. Conventional least-squares-based regression has been used

primarily in standard artifact suppression schemes, commonly referred to as mixing

algorithms. The more versatile PLS algorithm has been used in a wider range of applications,

including more sophisticated suppression schemes and predictive models.

3.3.4.1 Least Squares Regression and Factor-Based Techniques

Multivariate analysis techniques are often utilized to explore unknown trends in large

quantities of data. The two basic stages in such analyses are generally referred to as

calibration and prediction. In the calibration stage, a data matrix is constructed from the

probe response at various time or frequency slots for a given set of calibration samples. The

calibration phase produces a model that relates the probe response to known values of the

parameters to be estimated. In the prediction stage, the matrix of coefficients and weights that

describes the system response (input-output relationships) is used to estimate the parameters

of interest in new measurement test sets. Having the response matrix (XI and the parameter

matrix [Y], MLR develops a linear combination of variables in lx that minimizes the errors in

reproducing 1. This is accomplished by constructing a relationship between 1X and [Y] such

that

FY] = MXANl + [El, (3.6)

where lfl is the matrix of regression coefficients, and [El is the matrix of errors associated with

the MLR model. The matrix l[i is found by linear regression to minimize the difference between

the true and estimated elements of the parameter matrix in a least-squares sense as

N M N M

jEj= IE)=X E E Xynrn-Ymnri (Ymn_- ,U2 =y Z E2 2 = Xsn (3.7)

n=lm=l n=lm=l

MLR attempts to use all of the available information in the independent variable matrix

for model construction. Both relevant (signal) and irrelevant (noise) information are weighted

equally by this method when input-output relations are constructed. To overcome this

limitation, the two factor-based techniques (PCR and PLS) have also been examined. The

factors are linear combinations of the original variables. Factor-based models attempt to use a

minimal set of factors that best describe the true relationship between the independent and

dependent variable matrix. Because factor-based models do not attempt to directly invert the

covarlance matrix, potential colinearities in data can be accommodated without causing

singularities.

PLS was chosen to construct a multivariate regression model for the EC data. For the

PLS algorithm, a general model would consist of a regression between the scores of the

69
relationships for the dependent variable I[Y and independent variable [XI. The outer

relationships can be written in matrix form as

N

[XI = (TI[P1 + [El = I tnPh + [El (3.8)

n=1

N

[Y] = JUI[Q1 + IF] = E, Unqn + [FIN (3.9)

n=1

with the inner relation defined by

u = bntn, (3.10)

where 1X1, [UL, (FI, and I[Q are the eigenvectors and eigenvalue matrices, and IBI Is the vector

that contains the inner relationships. The principal components are simultaneously calculated

for both blocks. The intention is to minimize HFJ (i.e., NorrrmiF)) and, hence, to construct a

useful relationship between IAI and IYl.

3.3.4.2 Multifrequency Mixing Techniques

Processed channel information from multiparameter mixing is commonly utilized in

analyzing multifrequency EC Inspection data. So-called mixing techniques, originally

implemented electronically on analog signals by using rotation, scaling, and subtraction, were

devised to enhance detection of EC signals of interest by suppressing the effect of unwanted

signals such as background fluctuations and artifacts. Mix frequencies are chosen so as to

provide independent information on the signal to be eliminated. Current mix algorithms

generally use a linear least squares method in which the output (mix channel) signal is a

weighted linear combination of horizontal and vertical components of input signal components

at different frequencies. Because independent mix coefficients can be calculated for each

signal to be suppressed (often based on simulated artifacts on a calibration standard), in

theory an unwanted signal could be eliminated by successive implementation of different sets

of mix coefficients. However, except under Ideal conditions, the mix signal will always contain

residual signals from unwanted artifacts, and the mixing process could affect, to different

degrees, both the amplitude and phase of the signal from defects of interest.

Direct (independent) and indirect (dependent) mixing algorithms have been examined for

suppressing multiple unwanted indications from composite signals. In an indirect mix.

regression coefficients are determined by using a data segment from a simulated artifact such

as a TSP ring on a reference standard tube, which is expected to closely resemble those present

in the actual SGs. The aim of the regression model is to best reproduce the primary/base

frequency signal by combining signal components from auxiliary frequency channel(s). This is

the conventional approach used for the analysis of EC ISI results. Alternatively, independent

mix procedures have the potential to suppress unwanted signals by using multiple-frequency

readings on the same tube. This approach, primarily suggested for suppression of dominant

signal features, is of particular interest when tube standards with simulated artifacts that

resemble field-induced signals are not readily available.

70
Various two- and three-frequency regression fits have been Investigated. Preliminary

results indicated that two-frequency mixes are generally more consistent, so the approach was

to sequentially apply two-frequency mixes for suppression of two unwanted indications at the

same axial location. To improve the mix outcome, several modifications were made to our

standard mix algorithm. These consisted of energy-scaling each trace, resampling in the

frequency domain, and phase-angle referencing. Frequency resampling was done to increase

the number of available independent variables. Phase-angle tracking was done to help reduce

phase ambiguity from multiple application of linear regression coefficients. Based on cases

tested so far, the amplitude renormalization consistently exhibited improved quality of mix

outputs. This finding is believed to be a result of the weighting introduced by the

energy-scaling process.

3.3.4.3 2-D Suppression of Unwanted Signals

Multifrequency mix procedures that involve combining the information from a

primary/base frequency and one or two auxiliary frequency channels have been fairly effective

for the processing of bobbin coil readings. However, mix algorithms developed for 1-d bobbin

coil signals are not as effective for the processing of higher-resolution rotating-probe inspection

results. Unlike the situation for the bobbin coil, mixes for rotating-probe C-scan data are quite

subtle and routinely -fall to produce consistent outcomes. Also, when mix coefficients are

applied to the entire signal trace, significant distortion of both the amplitude and phase-angle

information of the original readings could result. To improve rotating-probe mixes, specially

designed simulated TSP rings (e.g., a 2700 ring as opposed to an axisymmetric 3600 ring used

for bobbin mixes) are often utilized. Nevertheless, the mix results could vary substantially,

depending on the selected portion of signal for the calculation of mix coefficients. For this

reason, mix channel information is not currently used for reporting rotating-probe inspection

results.

Potential alternatives to 1-D conventional mix procedures for RPC data have been

evaluated. Two approaches based on 2-D factor analysis techniques were considered. Both

methods utilize spectral decomposition, also referred to as singular value decomposition (SVD)

or outer product expansion. However, they differ significantly in their analysis of factors. The

first method is based on a standard decomposition at a single frequency and subsequent

selection of appropriate coefficients. The second method uses a least-squares (LS) regression

approach with two frequencies.

Spectral decomposition methods serve as the fundamental numerical technique in a wide

range of applications from solution of a linear system of equations to problems In signal and

image processing such as restoration, compression, denoising, pattern recognition, and feature

extraction. Such transformation algorithms basically attempt to decompose an often

ill-conditioned data matrix to an outer product expansion of left- and right-hand unitary

matrices of orthonormal eigenvectors with a matrix of pseudo-eigenvalues. With the

underlying assumption that the basis functions that span the signal and noise subspaces are

orthogonal, this mapping should then provide improved separation of relevant signal features

from inconsequential trends. Following the decomposition, selection of appropriate factors

allows for the preservation of relevant signal features and, in effect, elimination of background

noise.

71
In practice, selection of appropriate factors often poses the greatest challenge when

spectral decomposition alone is used. This is particularly true for the EC inspection results

with rotating probes because flaw and artifacts at one frequency are often described by

common factors. As a result, it is not always possible to select a set of factors that reasonably

separate consequential indications from artifacts and background fluctuations. Therefore,

factors cannot be applied blindly, and some form of discretion is necessary in deciding which

factors to keep. For this reason, the method of multiple-frequency LS-based factor analysis

provides a more robust and systematic approach to suppression of unwanted signals in

rotating-probe data.

Single-frequency spectral decomposition and a two-frequency decomposition with

LS-based regression were initially tested on simulated composite signals from RPC readings on

a 22.2-mm (7/8-in.)-OD tube standard with machined notches and a 2700 support plate slip

ring. Various combinations of OD and ID notches within the TSP region were produced by

vectorial addition of data segments at various frequencies with notches with the TSP data

segment. The nonaxisymmetric geometry of the TSP ring used here poses a greater challenge

to any signal suppression scheme than do axisymmetric drilled support plates. However, it is

expected that such nonuniform and overwhelming EC indications could be encountered in the

fleld from external structures such as broached support plates.

Suppression of TSP indications at a single frequency by using a standard SVD algorithm

did not produce acceptable results. In several test cases it was impossible to achieve a

reasonable degree of suppression of TSP indications without substantial degradation of the flaw

signal. The method of two-frequency LS-based factor analysis was able to suppress TSP

indications with acceptable degradation of flaw signals. Figures 3.17 and 3.18 show results for

two test cases. In both cases, single and multiple flaws that were arbitrarily placed within the

TSP region were recovered with a high level of suppression of TSP indications. Selection of

primary and auxiliary frequencies for 2-D suppression is analogous to that used by standard

1-D mix procedures. As with conventional 1-D mix procedures, multiple unwanted indications

could be suppressed by successive application of this method.

3.3.4.4 Multivarlate Analysis for Estimation of NDE Profile

Multivariate analysis is also used in the final stage of data analysis for estimation of the

flaw depth. Phase-angle information from multiple channels is used to construct a model that

correlates the NDE results to flaw size and origin. The flaw size and origin are assumed to be

related to the phase-angle information at the different frequencies. Eddy current

measurements on a calibration standard tube are used to determine the regression coefficients

in the relation.

The reliability of a predictive correlation depends heavily on the data used to construct

that model. Both the range and composition of data play crucial roles. With regard to the

analysis of EC data, this means the adequacy of the results is strongly dependent on the range

of flaw size, type, and origin in the calibration standard tube. Sizing accuracy is also strongly

dependent on the availability of multiple frequency data. A minimum of two frequencies is

required for a two-parameter model, but our work has shown that three frequencies generally

produce more consistent results. Although incorporation of a larger number of channels is

theoretically feasible and desirable, acquiring rotating probe data at a greater number of

frequencies is generally prohibitive for practical ISI applications.

72
(a) (b)

Figure 3.17. Demonstration of 2-D signal suppression on RPC traces of (a) simulated raw data at

400 kHz (top) and 200 kHz (bottom) frequencies composed of a 40% TW axial OD notch at

an arbitrarily selected location within a 2700 tube support ring and (b) processed data using

400 kHz primary and 200 kHz auxiliary frequency (top) and 300 kHz primary and 100 kHz

auxiliary frequency (bottom). In both cases, the processed data Indicate substantial

Improvement of SIN ratio over original readings.

mu Il

11101110MM.0, -  --,- -.- - .", , -- - - =

dat poir

1

2 S __ _

-0 500 100 1500 2000 25W d ha polet

3000

dataPoir

data polrt data pout Ii _______________________________________________________________

Figure 3.1 8.

(a) (b)

Demonstration of 2-D signal suppression on RPC traces of (a) simulated raw data at

400 kHz (top) and 200 kHz (bottom) frequencies composed of a 40% TW axial OD and

40% TW circumferential ID notch at arbitrarily selected locations within a 270° tube support

ring and (b) processed data using 400 kHz primary and 200 kHz auxiliary frequency (top)

and 300 kHz primary and 100 kHz auxiliary frequency (bottom). In both cases, the

processed data Indicate substantial Improvement of S/N ratio over the original data.

73
3.3.5 Rule-based Flaw Identification

In conventional EC data analysis of SG tubing, human analysts are trained to

differentiate flaw-induced signals from artifacts by analysis of the behavior of signals, typically

extracted from their Impedance plane trajectory and signal amplitude, through application of a

series of rules. For example, clockwise (CW) and counter-clockwise (CCW) rotation of the

impedance plane signal trajectory is commonly used to determine the flaw origin. When the

phase angles at all frequencies from a w1W flaw are set to a fixed value, indications of OD origin

exhibit CCW rotations, and those of ID origin have smaller CW rotations. In addition to such

well-defined rules for interpretation of signal behavior developed from a fundamental

understanding of the effect of flaws on EC signals, empirical rules based on past experience

and historical lessons learned in dealing with similar SG units are also used.

Analogous to the manual analysis of EC inspection data, computer-aided data analysis

algorithms can be developed based on analysis of characteristic behavior of EC signals as a

function of frequency. Because all available information can be examined rather than a

selected subset in the case of a manual analysis, such algorithms allow more effective

identification of subtle forms of degradation. Computer-based algorithms that imitate some

form of the human decision-maklng process are generally labeled as expert systems.

The automated identification of flaws and their origin through application of a series of

rules is shown in block 4 of Fig. 3.10. The rules are coded as a series of conditional statements

(i.e., IF-THEN) that are sequentially applied to the data. Calculation of S/N from the earlier

stages of the data analysis is used to set a threshold for sorting of signals that are to be

examined. Combined amplitude and phase-based rules are used to better discriminate

between potential flaws and noise. The rules are set to be generally conservative for

identification of flawed regions of the tube. As a tradeoff between conservatism and sensitivity,

they can be adjusted to some degree based on the S/N value for a particular test section. The

current set of rules is valid for a specific coil configuration and range of frequency.

3.3.6 Data Quality

Because eddy current testing is a comparative NDE method, success is highly dependent

on the consistency of data acquisition parameters and proper calibration results. Although the

quality of data affects both detection and sizing, this issue is of particular concern when

quantitative estimates of flaw size are needed. Minimum requirements for acceptable data

acquisition parameters are documented in various EPRI examination technique specification

sheets and NRC generic letters for ISI of SG tubing. These guidelines generally define the

instrumentation setup (coil excitation frequencies, gain setting, cable length, sampling rate,

probe speed, etc.) and calibration procedures to achieve a certain probability of detection (POD)

for a given probe (e.g., bobbin, rotating, and array probes). These techniques have been

primarily aimed at ensuring attainment of an acceptable POD. The requirements on data

quality, for example, in connection with the use of rotating probes for estimation of flaw size

are less well defined. Industry is attempting to define a suitable set of parameters that could

be used to judge data quality. One such parameter is the S/N for a particular section of SG

tubing. Because initial detection of potential flaw signals during manual analysis of EC data is

based on visual identification, defining a particular value or range of values for the S/N may be

appropriate for manual analysis for the purpose of detection. However, data quality that may

be appropriate for detection may not be adequate for reliable estimation of the flaw size.

74
For instance, degradation of S/N arising from baseline- fluctuations and serniperiodic

signals from tube deformations that could significantly reduce visual flaw identification can be

generally dealt with by numerical processing of data. On the other hand, degradation of S/N in

the form of acquisition-related corruption of data and insufficient sampling rate that may not

substantially affect detection may significantly influence the sizing results. Recovery of signal

peaks in a digitized data stream is governed by the Nyquist sampling rate. The relationship

between sample interval Ax and maximum observable frequencyfm=, according to the Nyquist

theorem is given by

fmax 2^x (3.11)

For rotating probes, sufficient sampling rates should be acquired in both axial and

circumferential directions in accordance with the rotational and push/pull speed of the probe.

Although lack of sampling can be remedied to some degree through interpolation of the data,

significant undersampling will always result in loss of information.

3.4 Multiparameter Data Analysis Results

The effectiveness of the multiparameter algorithms for the computer-aided analysis of EC

NDE data has been evaluated in studies on EDM notches, laboratory-grown flaws, and

in-service flaws. Additional examples are found in Ref. 1161. A more detailed description of

sizing results is provided in Section 3.4.2.

3A.1 Graphical Display of Multiparameter Analysis of NDE Data

Figures 3.19 and 3.20 compare conventional data displays, e.g., Eddynetl isometric

plots, with the graphical displays generated by multiparameter analysis to demonstrate

enhancement in spatial resolution provided by the analysis. The multiple-frequency EC data

were acquired with a standard three-coil rotating probe that incorporates a 2.92-mm

(0.11 5-in.) pancake, a mid-range +PointTM, and 2.03-mm (0.080-in.) high-frequency pancake

coil. The isometric plots are amplitude responses from the +Point coil at a single channel. The

processed results are estimated depth profiles obtained by analysis of multifrequency data fom

the mid-range pancake coil. Comparison of estimated profiles with destructive results also

shows improved sizing accuracy by the multiparameter analysis 1161. For the test cases shown

here, conventional analysis of original data generally resulted in underestimates of the

maximum depth.

Figure 3.21 displays the multiparameter analysis results for laboratory-grown

circumferential ODSCC. The analysis results indicate the presence of multiple circumferential

cracks with maximum depths exceeding 60% 1W. Figure 3.2 1(a) shows the flaw image at

different stages of the process. Figure 3.21(b) shows a terrain plot of the estimated flaw depth

from an aritrarily chosen azimuth and elevation angle. Figure 3.21(c) and (d) show axial and

circumferential projections of the flaws.

Figure 3.22 shows similar results for a specimen with a single axial ODSCC. Figure 3.23

shows another example with two deep axial cracks. The dominant indications are two parallel

axial cracks with maximum depths estimated to be w90% 'W.

75
J- 0 0 0 0 0 N F ... .. 1I-~~ -~-~- - --I --I .~..I. ..

(a)

Mnl*.vr~r-j-"fUi I '.- *1..U. Sb

(b)

Figure 3.19.

,00 .-z. - (a) Conventional Isometric plot

------ (Eddynett m) of .Pointm probe 100 --,,, '',,.......... . amplitude response, (b) Image

80 - ----------- tdisplay, and (c) multiparameter sizing

terrain plot for laboratory-grown

g 60 ,,,..--t l circumferential ODSCC. Image

40 ,, + | l t display axes in this case are pixel

* ,, . 19L|,>s. dlocations In axial (horizontal) and

circumferential (vertical) directions.

01 ~ ~ ~ ~ ~ ~ *** Results suggest the presence of

-360O1 parallel cracking with

300 ~ ~ ~ ~ ~ ~ * * maximum depth - 1 00% TW.

Circ. scale (do. 00 Ax. scale (mm)

(c)

3.4.2 Assessment of Sizing Accuracy In Presence of Artifacts

In the study to assess the effect of various artifacts on flaw signals, data on laboratorydegraded specimens were collected with and without the removable collars that simulate

artifacts such as the TSP, sludge, and magnetite deposits that may be present to various

degrees during field inspections. Eddy current data for each test case of a simulated OD

artifact were collected separately, each time scanning the entire batch of tubes. Acquisition

frequencies represent typical primary and auxiliary channels that are used for the inspection of

thin-wall SG tubing. Studies so far suggest that the degree of signal distortion depends

primarily on the artifact geometry and composition, as well as signal amplitude at a given

excitation frequency. Sizing results for the case of plain tube and TSP simulation are in good

agreement. For the data set used in this study, the estimated maximum depth of cracks varied

by <10% of tube-wall thickness for all the test cases considered. As expected, the effect of OD

simulated collars is less for larger amplitude signals and more significant for the smaller

amplitude portions of the flaw.

76
-------~~~~~~~~~~~~~~~~~~~I

(b)

Figure 3.20.

(a) Conventional isometric plot

100 -.-- (Eddynets) of +Pointm probe , ---. * @ ". amplitude response, (b) Image

80 ,- - ,,.. J-- *..*.display, and (c) multiparameter sizing , -- ,-.-.----,--,- terrain plot of same specimen with

if 60 ----- ,----~ | , , "*v*, "i laboratory-grown axial ODSCC. The

Image display axes In this case are

pixel locations In axial (horizontal)

20 --I----- - ------. e, and circumferential (vertical)

. _ A ' } ^^ directions. NDE results suggest the

_ presence of axial cracking with max.

depth > 90% TW. Flaw exhibited no

bobbin and uncharacteristically low

rotating probe amplitude response.

Carc. scale (deg.) 0 0 20

Ax. scale (mm)

(C)

A batch of 31 tubes containing various crack types and sizes has also been utilized to

further assess the capability of computer-aided data analysis algorithms. Flawed tubes in this

set have been augmented with blank specimens having no visible damage, to better test the

techniques. Eddy current readings on these tubes were acquired with and without simulated

artifacts such as the TSP, sludge. magnetite, and copper deposits. The effect of tube

dimensional variations on detection and sizing of indications is also simulated by producing

flaws within the transition zone of mechanically rolled tube sections. Eddy current Inspection

results from samples show a wide range of complexity in flaw geometries, varying from single

axial/circumferential to multiple and mixed-mode cracks. A more detailed description of NDE

results on this set is provided In Ref. 1161.

77
101-0

MM"

1, -

80

-..... I.-...-,...---.l--.-.---,--I(g)----,-.--.--.------.-. I

5"'!z -' AAilk"Mm"79

---- - b-L- M '

100

90

80

s 70o

1 400

30

20

10

0

. __, _,--- _._... _ _ . _- ... _-. .

------ ---v-------:--------;------- - ------ ;--------:---

L............. ----------- - --------.- _ .....

~.4.. . .........------ .....

.... _;,------ _ .^....... ------t---

10 20 30 40 50 60

Ax scab (mm)

10 00----------------------- - .-------

90 -------- ------- ------- ------- ------ . ---

so ---------------;------,-------t---------- -------l-------i70 ------ 4- - ---------------...-- : .:._-__4._-__-_-_-_- 7 

so A---------- : ...... __'. -'_ ---- -----

40 ------4-- ---i--- ^..+.. - .......... so ------- ,------- .l. ......-t

20 ----. 4- . .-.------ -II . . _ ................4

80

0 50 100 14 D 200 250 300 350 arc. scal(de.)

(C) (d)

Figure 3.21. Representative display of data analysis results for specimen with laboratory-grown ODSCC

degradation showing (a) image display, (b) terrain plot, (c) axial projection, and

(d) circumferential projection. Results indicate multiple cracks at several axial locations

around circumference, with maximum depth >60% TW.

3.4.4 Analysis of Field-Degraded Tubes

Eddy current data from a set of six pulled tube sections that were removed from a retired

McGuire SG were analyzed. Pre- and post-pull inspections identified all flaws as ODSCC at

dented TSP intersections. Data were collected with both a bobbin and 3-coil rotating probe

containing a mid-range pancake, +Pointl, and a high-frequency pancake coil. The tubes were

first decontaminated and then were scanned in a glove box with and without a simulated TSP

ring. In addition to an ASME standard, an in-line EDM 18-notch standard was used. This

standard contains a series of 9.52-mm (0.375-in.)-long notches of axial and circumferential

orientation, varying in depth from 20% to 100% 'W, originating from both the tube ID and OD.

Both bobbin and rotating probe data were acquired with procedures similar to those used

during the fleld inspections. Additional data at intermediate frequencies were also obtained for

78
LI t -

i_112t~

U.- 10-0_ 

a .t_ 

v W M J

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

100, .......... I- L----------

L L

so - ---------

:----------

M ......... ......

40 . .......... I---

20 . ..........

200

Orc. scale (deg) a AX Vale irro)

(a) (b)

u _ ~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~Sc

100

80 . .....

S A \ -~~~~~~~~~~~igure 3.22.

80 . ..... *----------,-nk-----4-- ---- Representative display of data analysis

2 40.......t *|~~~~~ results for specim en with

20 .,.......... ; ... , - -- -- laboratory-grown axial ODSCC , , . ,1 5 .showing (a) pre- (top) and

A scat (mm) post-processed (bottom) images of

normalized data, (b) terrain plot and

100 (c) axial and circumferential

80------.-............ -------------- --------- -- projections. Results indicate a

.E- ...... .>15-mm (0.6-in.)-long axial crack with

maximumdepthof=90%oTW.

40 .. . ..... . . -- - -I - - - I

2 0 ------T ---- T- ---- -- - --- ---- ------------- 4 20--'.-a50 100 150 2 -250 300 350

arc. scale (deg)

(C)

further assessments. The multiparameter analysis routines were modified to analyze EC data

from 19.0-mm (0.75-In.)-dameter tubing with a nominal wall thickness of 1.09 mm (0.043 In.).

The flaw depth and extent were also estimated by conventional analysis of the data using the

Eddynetsm software. Initial results from the conventional and multiparameter analyses of the

NDE data are discussed here.

Flaw depth estimates by bobbin coil were obtained using the 550/130 kHz mix process

channel. Estimates of flaw maximum depth and extent by rotating probe were made using the

300 kHz channel. he probe gain had to be reduced during acquisition to avoid saturation of

79
Al I

"it YNO-14 W-ftmo - -- iffir"Vric . 9 '--- V _ W

4 en _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ __ _ _ _ _ _ _ _ _ _ _ __ _ _ _ _ _ _ _ _ _ _ _ _ _ _

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

90 ------------i.. ..... I _+ ........... ......

100- so ------ ----- 4- ------

........... 70 ........... T------------------- so- L

60 ----------- ....

00- ...... 50 ........... ... - ------------I-------

40 f 40 ---------- - ----------- t------

30 ---------- . ........... ------ 20- 1#

20 ........... ----------- ------

0- 10 T .. . .......

0 Ax scale (mm) 0 10 20 30 40 Orc. scale (dogj Ax scale (mm)

(a) (b)

100 ------------------------ ............... ...................

go ------- ------- --i ....... ---- ------- ------- 80 ------- ...... .. I I I Figure 3.23. r.... ...... I------- T.-

Representative display of data analysis 70 -------- .......... ...... I---- ----------------- 1. results for specimen with laboratory-grown r"O ------------------ .. ' ........ 1-11, -------I----------

axial ODSCC showing (a) terrain, and so ........ ............... (b) and (c) cross-sectional views of 40 -------- L------ i- 4_1_1 ------ _1 ------- i_ estimated flaw profiles.. Results show two 30 ...... 4....... ---- ------ 4------- 4- indications with maximum depths >80% TW. 20 -------- ...... ------- ------ -------

1 0 ....... . ........ ------------ ----------------

0 so 100 150 200 250 300 350

Orc. scale (dog.)

(c)

the pancake coil signal from 100% 7W EDM notches. The noise levels in all the pulled tube

test sections were significantly higher (i.e., lower SIN ratio) than in laboratory-produced

specimens. The noise was primarily H)-originated. Comparison of the SIN ratios from all coils

showed that the +PointTm signal was the least affected by the presence of the noise.

Except for one case, all the flaws detected by the surface-riding coils of the rotating probe

were also detected by the bobbin coil. In the one exception (a single test section with no

detectable degradation by bobbin cofl), the rotating probe indicated a >50% TW flaw where a

large signal from a tube deformation appears to have masked the flaw signal for the bobbin

coil. In the cases where a flaw was detected by the bobbin coil, no apparent correlation was

found between flaw depth and the BC signal magnitude (peak-to-peak voltage), which varied

from 0.75 V to >4 V. Maximum flaw depths estimated by +PointTu varied from < 10%'IW to

<1 900A 1W. with axUd crack length:!f ftom -3.5 mm (0. 13 in.) to -10 mm (0.40 In.). The shortest

axial flaw had the largest circumferential extent based on the rotating-probe C-scan display. in

tubes without the simulated TSP collar, there was reasonable agreement between bobbin and

so
rotating probe estimates of maximum crack depth based on phase-angle information for the

majority of detectable flaws.

Flaw depth estimates were also obtained by multiparameter analysis of mid-range

pancake coil data. The minimum necessary modifications of the algorithms, which have been

used to date for 22.2-mm (0.875-in.)-diameter tubing, were made to allow sizing of flaws in

19.05-mm (0.75-in.)-diameter tubing. The rotating probe data were resampled off-line to

produce approximately equal digitization rates in axial and circumferential directions.

Figures 3.24-3.29 show the data analysis results for all the tube sections examined so far.

Terrain plots of flaw size are shown for a small zone around the degraded region of each tube.

Except for one case, the maximum depth estimates are generally in agreement (i.e., to within

+/-20%) with those from the +PointTh coil. It is more difficult to estimate the flaw sizes in

1.09-mm (0.043-in.)-thick wall tubing than in thicker wall 1.27-mm (0.050-in.) tubing. This

effect is due to fundamental limitations that are independent of the coil design and the analysis

method. As a direct consequence of the skin depth effect, the same upper and lower frequency

limits for thinner wall tubing result in smaller ID-OD phase separation at each frequency as

well as a smaller variation of phase angle between the primary and auxiliary frequency

channels. This smaller dynamic range, in effect, reduces the sizing accuracy and increases the

ambiguity in determination of flaw origin.

81
1 ~ ~ ~ ~ ~ ~ ~ ~ ~ ~~~III

X

2

i IC

i a 0 4

I -S a a

I

I a I t

I

i a V.

i 0 M M a

i Vs 0

Figure 3.24.

Data analysis results for McGuire

pulled tube section with axial ODSCC

degradation at dented TSP

intersection. Maximum flaw depth Is

estimated to be <80% TW. Graphics

shows (a) amplitude image display of

the standard and flawed tube at

different stages of the process, and

(b) flaw Image, and (c) terrain plot of

sizing data.

(a) -- -- ow = =I , M. . !.* -, il' I,, 4. IV qi -A -kcma. -

I~~~~~~

& 1 M a U M I A ..*"

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

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

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

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

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

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

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

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

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

Om sc~do de) a 10

Am. scab O)

(b) (c)

82
If: Figure 3.25.

Data analysis results for McGuire

pulled tube section with axial ODSCC

degradation at dented TSP

- ;_intersection. Maximum flaw depth is

estimated to be <40% TW. Graphics

shows (a) amplitude Image display of j i l l the standard and flawed tube at

different stages of the process,

(b) flaw image, and (c) terrain plot of

sizing data.

(a)

_ . . .. .

AX0 ' ''','. . ' '- ' '1~~~~~~~~. .'. '.""'.. ......

MCk. SCO(dog.) Ax.c) Qw)

(b) (C)

21' 1 -

I

I U

I g

A-~~~~~~~~~~~~~~~~~~

83
a

qI

Figure 3.26.

Data analysis results for McGuire

pulled tube section with axial ODSCC

degradation at dented TSP

intersection. Maximum flaw depth is

estimated to be > 90% TW. Graphics

shows (a) amplitude image display of

the standard and flawed tube at

different stages of the process,

(b) flaw image, and (c) terrain plot of

sizing data.

(a) -- m-Mm

mik

I .. .I 1. 1. - . -- ...' ...- -1

t

I

I0

9W I

tm

tm

6 Z

.0.., '-''~~~~~~~~~......".'.".. . ... .. ... 

1W .*.....

20 4 F

Circ. scal (dog.) ° 0 5 wa (rn

(b) (C)

84
Figure 327.

Data analysis results for McGuire

pulled tube section with axial ODSCC

{ I - ffi degradation at dented TSP

Intersection. Maximum flaw depth is

estimated to be <70% TW. Graphics

shows (a) amplitude Image display of

:__ the standard and flawed tube at

| different stages of the process, s *P!~~~~~~~~~~~~~~~(b) flaw image, and (c) terrain plot of

I____________________________________________ .sizing data.

(a)

.= --= =m....

' s z '"""'""'' ~ ~ ~ ~~~~~~~~~~~~~~~~~~~~~~~~~~~....... - . ..... '"'I"" ,'"

i., 4 _ { U. , -", ,,,,., a~~~~~~~~~~~~~~~~~~~~2 .... .....

1~~~~~~~~~I 0

o .2

Crc. scale (dag) 0 0a Om)

(b) (c)

85
-! M!

i

I

a -t ,., I 0 i a I 0

1 AL

i

I

"I

I 11 Ar I a i

i la I

I

Figure 3.28.

Data analysis results for McGuire

pulled tube section with axial ODSCC

degradation at dented TSP

intersection. Maximum flaw depth is

estimated to be <20% TW. Graphics

shows (a) amplitude image display of

the standard and flawed tube at

different stages of the process,

(b) flaw image, and (c) terrain plot of

sizing data.

(a) -- - ------ ;-.-tobii ;r:| f -- is , S.'- i, -.--- j -i , ,-£ o't1~.tri> Ir; -e i'ij j'j

ixR211.1011Y.M1

- - I . --- -

A a ~ ~ a

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

..........

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

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

.... . ......

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

40 . ............. . .......

Cim sta (ft) o 0 An. scale tow)

(b) (c)

86
Figure 3.29.

Data analysis results for McGuire

pulled tube section with axial ODSCC

degradation at dented TSP

43 SO intersection. Maximum flaw depth Is

i a . ^ " w w estimated to be > 50% TW. Graphics

shows (a) amplitude image display of

the standard and flawed tube at .1 .Ij;* .I different stages of the process,

(b) flaw image, and (c) terrain plot of

l _______________________________________________ sizing data.

(a) _

' In-~ ~ ~ ~~~~~.. I*

_I_._ _ _ __-s . .......

40 . ....... -. '... '.

(b) (c)

87
4. Research on Degradation Modes and IntegrItyQ(D. R. Dlercks, K. E. Kaszal S. Malumdar, J. Y. Park, and W. J.Shack

4.1 Production of Laboratory-Degraded Tubes and Characterization of Flaws

Because degraded tubing from operating or retired SGs is difficult and expensive to

obtain and the availability of such tubing is limited, it Is necessary to produce prototypical

degradation in tubes as a part of this program. These tubes are to be used to evaluate NDE

equipment and techniques In the SG mock-up and to conduct pressure and leak-rate testing.

Techniques for Alloy 600 tubes have been developed to produce SCC degradation of various

sizes, depths, and orientations that are prototypical of field-degraded SG tubes. Flaws in tube

specimens were destructively examined and compared with the results of EC NDE.

4.1.1 Materials Properties of Alloy 600

As the test material for the program, Alloy 600 tubing with diameters of 19.1 mm

(0.750 in.) and 22.2 mm (0.875 In.) has been obtained from four sources. Valinco supplied

19.1-mm (0.750-in.) tubing (Heat NX8524) In the as-fabricated ("cold-worked") condition, and

22.2-mm (0.875-in.) (Heat NX8520 and NX8527) tubing In the mill-annealed condition. Heat

NX8520 was supplied with two mill-annealing processes: 3 min at 980'C (1796 0F) for a low

temperature annealing or 2 min at 10600C (1940011 for a high temperature annealing. Material

with the low-temperature annealing is designated as NX8520L, and material with the hightemperature annealing Is designated as NX8520H. Heat NX8527 was mill-annealed for 3 mmn

at 9401C. PNNL provided -110 m (360 ft) of 22.2-mm (0.875-in.) tubing in 3.7-m (12-ft)

lengths. PNNL was unable to Identify the heat number, but the tubing was labeled

'EX-82-1/2675' on all but one piece, which was labeled *HT #9829/L-102." PNNL also

provided cold drawn and annealed 22.2-mm (0.875-in.) tubing (Teledyne Columbla-Summeril

Heat NX7968). EPRI provided 22.2-mm (0.875-in.) tubing (Heat 96834 M-81) that was

mill-annealed at 9271C(17000F). This material (Heat 96834) was mainly used for the crack

evolution studies described In Section 4.2.

Heat analyses were provided by the vendors, and a check chemical analysis was

performed on a sample from the PNNL tubing. These analyses are shown in Table 4.1. In all

cases, the tubing conforms to specification ASTM B 163 for Alloy 600. Room-temperature

tensile properties, hardness, and grain size data were provided by the vendors. However,

additional tests were also conducted by ANL and by an outside vendor. The results,

summarized in Table 4.2, show that, although the ultimate tensile strengths are comparable,

the ductility of the 22.2-mm (0.875-in.) tubes Is significantly greater than that of the 19.1-mm

(0.750-in.) tubes. The room-temperature harnesses and grain size of the Alloy 600 tubing are

also included in Table 4.2. The average grain size is ASTM 8.0 (average linear intercept, 20 pm)

for the 19.1-mm (0.750-in.) tubing (Heat NX8524) and ASTM 8.6 (average linear intercept

16 um) for the 22.2-mm (0.875-in.) tubing and PNNL ID EX-82-1/2675.

88
Table 4.1. Elemental analysis (wt. %) of Alloy 600 tubing material. Diameters of 19.1-mm

[0.750-in.] (heat NX8524) and 22.2-mm [0.875-in.] (heats NX8520, NX8527,

NX7968,96834, and PNNL ID EX-82-1/2675).

Specifications

Element ASIM B163 NX8520 NXB5 2 4 a NX8 5 2 4 b NX8527 NX7968 96834 EX-82-1c

Ni 72.0 min. 75.63-75.77 74.66 74.95 74.85-75.20 75.44 74.81 75.05

Cr 14.0-17.0 15.28-15.40 15.21 14.84 15.33-15.51 15.58 15.85 15.21

Fe 6.0-10.0 7.96-8.03 9.16 9.11 8.49-8.68 8.38 8.16 8.30

Mn 1.Omax. 0.19 0.20 0.20 0.21 0.24 0.27 0.18

Cu 0.5 max. 0.02 <0.01 0.01 <0.01 0.18 0.01 0.17

C 0.15 ma- 0.022 0.022 0.023 0.023-0.025 0.040 0.037 0.035

S 0.015 max. <0.001 <0.001 <0.001 <0.001 0.001 0.001 0.001

Si 0.5max. 0.18-0.21 0.20 0.17 0.13-0.16 0.14 0.30 0.23

Al d 0.21 0.24 0.22 0.23-0.25 d d 0.15

TI d 0.26-0.34 0.29 0.35 0.30-0.36 d d 0.18

Co d 0.02 0.01 0.02 0.01-0.02 d d 0.03

P d 0.004 0.003 0.005 0.004-0.005 d d 0.009

B d 0.002-0.004 0.004 0.002 0.001-0.003 d d 0.002

N d <0.01 <0.01 <0.01 <0.01 d d 0.004

Top of ingot.

bBottom of ingot.

CCheck analysis for tubing from PNNL with a label, EX-82-1 /2675.

dNot specified, not measured, or not reported.

Table 4.2. Room-temperature tensile properties, hardness and grain size of Alloy 600 tubing.

Diameters of 19.1-mm [0.750-in.) (heat NX8524) and 22.2-mm [0.875-in.] (heats

NX8520, NX8527, NX7968,96834, and PNNL ID EX-82-1/2675)

Spec. D"am oy Ave. °v au Ave. au et Ave. et Grain Size Hardness

Heat No. No. (mm) (MPa) PA (Pa) RMPa) (% M) ASIM

NX8520 La 22.2 d d d 696 d d d VH 162

Hb 22.2 d d d 641 d d d VH 146

NX8524 6-1 19.1 296 d 679 d 25.5 d d d

6-2 19.1 314 d 682 d 25.5 d d d

6-3 19.1 315 308 682 681 24.5 25.2 8.0 VH 173±8

NX8527 2-1 22.2 267 629 40.0

2-2 22.2 275 648 41.0

2-3 22.2 270 271 625 634 41.5 40.8

SGMC 22.2 317±14 696±14 VH 179

SGMCC 22.2 179±14 600*14 VH 148

NX7968 22.2 296 639 51.6 - -

96834 22.2 359 734 38.0 9.0 RB 85.0

EX-82-1 15-1 22.2 356 689 44.0

15-2 22.2 356 681 44.5

15-3 22.2 358 357 678 683 41.0 43.1 8.6 VH 184±9

aLow temperature mill-annealed at 980°C (17960 1F) for 3 min.

bHmgh temperature mill-annealed at 10600C (1940°F1 for 2 mmn.

CMilannealing plus 1 100°C 10 mln, and 600°C 48-h heat treatment.

dNot specified, not measured, or not reported.

89
4.1.2 ProductIon of Laboratory Cracked Tubes

Stress corrosion cracks have been produced In 22.2-mm (7/8-in.)-diameter specimens

(Heats NX7968, NX8527, NX8520L, and NX8520H). Cracking is accomplished using a 1 M

aqueous solution of sodium tetrathionate (Na 2O6S4 -2H2O) at room temperature and

atmospheric pressure. The material was first mill-annealed for 3 min at 9400C for Heat

NX8527, 3 min at 9800C (17960 Fl for Heat NX8520L, and 2 min at 10600C (19400F) for Heat

NX8520H. Heat NX8524 was in the as-fabricated (cold-drawn) condition. Before exposure to

the tetrathionate solution, the specimens were heat treated at 6000C (=1 1000F) for 48 h to

produce a microstructure that is susceptible to cracking 1181. Various techniques, including

localized environmental exposure, low applied load, and controlled electrochemical potential,

have been applied in an effort to produce segmented SCCs. During the course of the program,

about nine-hundred 22.2-mm (7 /&-in.)-diameter Alloy 600 tubes were processed to produce

axial ODSC and IDSCC, circumferential ODSCC andl IDSCC, and SCCs in skewed orientations.

Some of the specimens were produced with multiple cracks or combinations of multiple cracks

of different orientations. Most of the specimens were straight tubes, but some were

roll-expanded, dented, or ovalized. Some of the heat-treated tubes have been installed as

uncracked blanks in the NDE SG tube mock-up. The current cracking facility is able to handle

ten individual tube specimens to produce axial SCCs and three specimens for circumferential

SCCs.

To accelerate cracking, some specimens from Heat NX8527 were additionally

solution-heat-treated at 1010-1100'C (1850-2012 0 F) for 10-20 min before the sensitization

heat treatment at 6000C for 48 h. Cracking time ranged from 10-1000 h. The specimens that

received the high-temperature treatment cracked in shorter times than those without it.

However, this solution treatment causes grain growth and a reduction in hardness and flow

stress. Tensile tests were performed and showed the flow stress of the material to be reduced

by =20% due to the heat treatments.

The sensitization heat treatment at 6000C (11 120F) for 48 h did not weaken the material.

Tensile tests were performed for the 22.2-mm (7 / 8-in.) Alloy 600 tube material from NX8520H

and NX8520L.

The solution and sensitization heat treatments may also change the magnetic properties

of the alloy and influence EC measurements. To examine this, two 22.2-mm (

7 /gin.) Alloy 600

tubes (Heat NX8527) with EDM notches have been treated with the solution and sensitization

heat treatments, and EC results were compared before and after the heat treatments. It was

found that sensitized sections (i.e., those containing flaws) could be distinguished from blank

(non-sensitized) sections, unless magnetically biased probes were used. When magnetically

biased probes were used, the signals from sensitized and non-sensitized sections were

indistinguishable. I

To examine the influence of oxide films on Alloy 600 tube specimens on EC response, 10

cracked tubes were oxidized in high-temperature water environments. Of these, five tubes

(SGL1 14, 159, 170, 182, and 187) were oxidized in deionized high-purity water with 8 ppm

dissolved oxygen at 2900C (554 0F) and 8.8 MPa (1280 psi) for 1680 h. Five additional tubes

(SGL127, 149, 169, 189, and 197) were oxidized in deionized high-purity water with 3 ppm

dissolved oxygen at 2900C (5540F) for 120 h, then in a simulated primary water environment

(1200 ppm B and 2 ppm U) at 3200C (6080F) and 12.4 MPa (1800 psi) for 1400 h. Eddy

90
current NDE was performed before and after the oxidation treatment. Results from the EC

examinations of these specimens are described In Ref. [191.

In the early stages of the program, SCCs were produced In Alloy 600 tube specimens in a

caustic solution using a high-pressure autoclave system. Fifteen 19.1-mm (0.750-in.) tubes

(Heat NX8524) and fourteen 22.2-mm (0.875-in.) tubes (Heat NX8527) were cracked by this

technique. The concept, design, and operational procedures of the autoclave system are

described in Refs. 1131 and 1201. Tube specimens were exposed at PWR SG operating

temperatures (or slightly higher to accelerate cracking to aggressive solution containing =60%

NaOH plus CuO and Cu. which are known to induce stress corrosion cracking in Alloy 600.

Components In the facility that operate in contact with the aggressive water chemistry were

fabricated of Hastelloy C-276 or stainless steel (SS). The results of the degradation

experiments for 29 tubes in the autoclave system are summarized In Ref. [19]. Use of this high

pressure technique was discontinued since the development of the low temperature technique

that is simpler and more convenient, and the majority of the tube specimens for the program

were produced by the low-temperature sodium tetrathionate technique. The BC voltages were

found to increase significantly for the oxidized tubes, though the shape of the Lissajous figures

remained unchanged. In contrast, the results for the +Point coil were inconclusive.

4.1.3 Characterization of Laboratory-Cracked Tube Specimens: Specimen Set for Control and

Verification

Twenty-seven laboratory-cracked Alloy 600 tube specimens were characterized by

destructive means, and the results were compared with the EC/NDE measurements. All

twenty-seven specimens are the nominal 22.2-mm (7 /sin.)-diameter tubing. The cracks were

destructively opened by fracture and examined macroscopically and with scanning electron

microscopy. The specimens were heat tinted prior to fracturing to permit discrimination of the

SCC and fracture opening surface. The heat tinting is accomplished for 1/2-1 h at 6000C

(11120F) in an air atmosphere. This tinting makes the SCC appear in darker contrast to the

fracture area. Details of the method used for the destructive examination and comparisons

with the EC/NDE crack profile are given in Ref. 1161. Figures 4.1 through 4.6 are the results of

destructive examination and comparison with the EC NDE evaluation for the 27 laboratoryproduced cracks. The EC/NDE evaluations are overall in good agreement with the destructive

examination for cracks deeper than about 50% TW, while poorer agreements is observed for

shallow cracks.

4.1.4 Characterization of Laboratory-Cracked Tube Specimens: Mock-up Specimens

Seven ANL mockup tube specimens were characterized by destructive means, and the

results were compared with the EC/NDE measurements. All seven specimens were the

nominal 22.2-mm (7 /-n.)-diameter tubing. Four specimens were produced by ANL using the

technique described in the previous section., and three specimens were provided by

PNNL/ West. The cracks were destructively opened by fracture and examined macroscopically

and with scanning electron microscopy. The EC/NDE crack profile was obtained using the

multi-parameter analysis technique. Figures 4.7-4.13 show the results of the destructive

examination and comparison with the EC NDE profiles. The EC/NDE measurements are

overall in good agreement with the destructive examination. However, EC/NDE noticeably

underestimated the crack size for specimen PNNL/W 4-02 (Fig. 4.11), and a minor crack in

specimen PNNL/W 4-05 was not resolved (Fig. 4.13).

91
SOL10I OD CIRCI*

1040

60

(a)

Figure 4.1. Fractography profile and NDE for circumferential IDSCC 1

20 30 40 50 so 70

Crack Length (mm)

(b)

in specimen SGL101.

SGL101 ID aRCd2

100

80

0.

I

so

40

20

0 =~ IL WI MW 0 10 20 30 40 50 6o 70

Length (mm)

(a) (b)

Figure 4.2. Fractography profile and NDE for circumferential IDSCC 2 in specimen SGL1 01.

92
S0L271 OD CIRC

F

120.00 -

100.00

I60.00

680.00

40.00

20.00

0.00 _

20 25 30 35 40 45 50 55 Go

Crack Length (mm)

(a) (b)

Figure 4.3. Fractography profile and NDE for circumferential ODSCC in specimen SGL271.

SGL410 OD AXIAL

.I

120.00

100.00

60.00

60.00

40.00

20.00

0.00 Mirmaw-11BUT _2 0 2 4 6 8

Crack Length (mm)

(a) (b)

Figure 4.4. Fractography profile and NDE for axial ODSCC in specimen SGL410.

10 12

93
SGL412 OD AXIAL

120.00

100.00

I 80.00

6 0.00

9 40.00

20.00

0.00

-2 0 2 4 6

Crack Length (mm)

(a) (b)

Figure 4.5. Fractography profile and NDE for axial ODSCC in specimen SGL412.

a 10

SGL413 OD AXIAL

120.00

100.00

, 80.00

' 60.00

40.00

20.00

0.00

-2 0 2 4 6 a

Crack Length (mm)

(a) (b)

Figure 4.6. Fractography profile and NDE for axial ODSCC in specimen SGL413.

10 12

94
SGL434

120 . .. .... .

Fracdograh

4V 100

so

6.0 a

20

-10 0 10 20

Length (mm)

(a) (b)

Figure 4.7. Fractography profile and NDE for axial ODSCC in specimen SGL434.

80

SGL4A7

I.

120

100

80

60

40

20

0

0 5 10 15 20

Length (mm)

(a) (b)

Figure 4.8. Fractography profile and NDE for axial ODSCC in specimen SGL487.

25 s0

95
SGL488 NOE MP

120

s.o

6d0

20

20

-5 o 5 10

Length (mm)

(a) (b)

Figure 4.9. Fractography profile and NDE for axial ODSCC in specimen SGL488.

15 20

SGL490

0.I

120

100

80

60

40

20

0

_________________________________________________ -5 5 10 15

Length (mm)

(a) (b)

Figure 4.10. Fractography profile and NDE for axial ODSCC in specimen SGL490.

20 25

96
PNNLJW 402

120

100

5 80

e.,60

i-Be ..

2V

20

-6 0 5 10 15 20

Length (mm)

(b)

25 30 35

(a)

Figure 4.11. Fractography profile and NDE for circumferential ODSCC in specimen PNNLW 4-02.

PNNLJW 4.03

120

100

F B0

- 60

la

a 40

20

0.

-5 0 5 10 15 20

Length (mm)

(a) (b)

Figure 4.12. Fractography profile and NDE for circumferential ODSCC in specimen PNNL/W 4-03.

97
PNNLW 4-05

812

40

20

0

-5 o 5 10 15 20 25 30

Length (mm)

'(a) (b)

Figure 4.13. Fractography profile and NDE for circumferential ODSCC in specimen PNNLW 4-05.

4.1.5 Characterization of Fleld-Serviced Tube Specimens: McGuire SG Tubes

Six tube specimens from the retired McGuire SG were characterized by destructive

examination, and the results were compared with the EC/NDE results. The nominal diameter

of the tube specimens is 19.1 mm (0.750 in.). The cracks were destructively opened by fracture

and examined macroscopically and with scanning electron microscopy. Figures 4.14-4.17

show the results of destructive examination and comparison with EC NDE profiles for tube

specimen 4-43-2, 5-51-2, 7-24-3, and 14-55-3. The EC/NDE crack profile was obtained

using the multi-parameter analysis technique, although this technique is still under

development for 19. 1-mm (0.750-In.) diameter tubing. The EC/NDE results are overall in good

agreement with the destructive examination, except for specimen 5-51-2 (Fig. 4.15).

4.1.6 CharacterizatIon of Laser-Notched Tube Specimens

Destructive analyses were performed for 22.2-mm (0.875-in.) Alloy 600 tube specimens

with laser-cut notches of various geometries. The notches were destructively opened during

pressure and leak tests, and by subsequent fracture of ligaments.

Digitized profiles of the laser-cut notches in tube specimens were obtained from

fractography and compared with eddy current NDE measurements. The results for specimen

5528-1-1, 5528-1-2, and 5528-2-1 are shown in Figures 4.18-4.21. Specimen 5528-1-1,

5528-1-2, and 5528-2-1 contain a single-segment axial flaw in each specimen. The average

depth at the mid-section of the laser-cut notch is 70, 75, and 68% TW for specimen 5528-1-1,

5528-1-2, and 5528-2-1, respectively. Specimen 5528-3-2 has a 6-segmented axial cut. The

average depth at the mid-section of the laser-cut notch is 65% TW. The EC/NDE

measurements slightly underestimated the depths of the laser-cut notches, although they are

98
McGuIre 4-43-2

120 . I . I .

rai ograpj 0. Mmm Offset Ugament

100

60

40 / 0 1

I _ Length (mm)

(a) (b)

Figure 4.14. Fractography profile and NDE for axial ODSCC in McGuire tube specimen 4-43-2.

generally in good agreement with the fractographic results. During pressure and leak tests, the

thickness of the tube is reduced due to plastic deformation. However, In the present analyses,

an assumption was made that reduction of the thickness at the laser-cut notch face in the

tube is small and can be neglected. This may explain some of the discrepancy between the

NDE and fractographic results.

The laser-cuts are significantly deeper at the end of notch segments. For all the

specimens, the ends of the laser-cut segments are cut much deeper than the average depth at

the mid-section, i.e.. 100% TW. This is probably caused from longer dwell times at the end of

the segments during the traverse of the laser beam. NDE does not resolve these pin-hole

defects.

4.2 Crack Evolution Studies

An SCC testing facility was constructed for experimental investigation of Pb-induced

cracking in Alloy 600 and Alloy 690 steam generator materials. The system consists of two 8-L

Hastelloy C276 autoclaves, a Hastelloy C276 small-diameter pipe autoclave, loading

mechanism, and electrochemical measurement instruments. The 8-L Hastelloy autoclaves are

rated for a maximum pressure of 13.1 MPa (1900 psi) at 3500C (6621F) and are ASME Code

stamped. These F-L autoclaves will be used for c-ring type static SCC and high-temperature

electrochemistry tests (Fig. 4.22). The small diameter pipe autoclave will be used for the

constant extension rate tests (CERT) (iFg. 4.23). Testing material will be the tubing that has

been used in the other tasks in this program. However, the Electric Power Research Institute

(EPRI) provided some additional materials: 5 m each of 22.2-mm (7 /g-in.)-diameter Alloy 600

tubes of Heat no. 96834 M81 (annealed at 9270C) and 96834 (annealed at 10240C), and 5 m of

22.2-mm (7/S-in.)-dameter Alloy 6901T of Heat no. D2A2201 (annealed at 5790C and aged at

7000C 112920F1 for 15 h). A potentio/galvanostat system was installed that operates

99
F

fis

120

100

80

so

40

20

0

IWvIi-~I.,i k-L -2 0 2 4 a I

LeDngth (mm)

(a) (b)

Figure 4.15. Fractography profile and NDE for axial ODSCC in McGuire tube specimen 5-51-2.

I 10

MUcuire 7-244

I.

120

100

so

60

40

20

0

.-1 0 1 2 3 4 5

Length (mm)

(a) (b)

Figure 4.16. Fractography profile and NDE for axial ODSCC In McGuire tube specimen 7-24-3.

6 7

100
McGulm 14454

120

100

I e

I=.60

To

C i

40

20

0

-2 0 2 4 a

Length (mm)

8 10

(a) (b)

Figure 4.17. Fractography profile and NDE for axial ODSCC in McGuire tube specimen 14-55-3.

120

100

s0

so

40

20

0

. . . . . . .I . .I =EC NDE

. . . . I . . . I . .

5528-1 -1

II

II

10aI

II

I6

aI

II

I11

I9

II

i . . .. . . .. . . . . . . . . . . . . . .

. .

9 .

0 2 4 6 8 10 12

Length (mm)

(a) (b)

Figure4.18. Fractography of axial laser-cut 5528-1-1 and profiles determined by ECINDE and

fractography.

101
a

I

i -

Figure 4.19.

120

100

a.

80

60

40 H

20

0 2 1LA 1 :UJF} ZIIII S 1- X X Length (mm)

(a) (b)

Fractography of axial laser-cut 5528-1-2 and profiles determined by

fractography.

8 10

ECINDE and

* ~~~~~~120

100

SO

JDIum IIIH IllIllh- I0

(a)

Fractography of axial laser-cut 5528-2-1

fractography.

2 4 a a 10 12 14 18

Length (mm)

(b)

Figure 4.20. and profiles determined by EC/NDE and

102
120

100

40 B

_~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~

20

0

(a)

Figure 4.21. Fractography of axial laser-cut 5528-3-2 and profiles

fractography.

6 10 I5

Length (mm)

(b)

determined by ECINDE and

(a) (b)

Figure 4.22. Eight-L autoclaves used for c-ring type static SCC and high-temperature electrochemistry

tests.

103
(a) (b)

Figure 4.23. Small-diameter Hastelloy pipe autoclave system used for CERT.

electrically isolated signals from the ground. Preliminary functional tests have been performed

for the systems. Silver-silver chloride reference electrodes were fabricated. Accuracy and

stability of the reference electrode were evaluated by measuring the potential with respect to a

standard calomel electrode (SCE) at room temperature. A reference electrode that was

fabricated with a 32 mA/mm2 oxidation current outperformed the others. The measured

potential remained stable within 4.4% (2 mV) for about 850 h at room temperature (Fig. 4.24).

As part of functional tests, electrochemical polarization tests were performed for Alloy 600 tube

specimens (Heat No. 96834 M81) in a near neutral (pH = 6.0) aqueous solution of 0.4 M

sodium sulfate at 250, 1000C, and 2501C (77, 212, and 4820F). Figure 4.25 shows the results

of the polarization tests. The polarization curve, in general, shifts to higher potential and

current ranges as the temperature of the solution increases. The corrosion potential is about

-0.6, -0.4, and -0.1 V normal hydrogen electrode (NHE) at 25°, 1000.C and 2500C, respectively.

The potential range for anodic passivation is significantly reduced at higher temperatures.

Anodic polarization tests were performed for specimens from a reference Alloy 600 (Heat 96834

M81) in aqueous solutions of 0.4 M Na2SO4 at pH 3 and pH 6.8 at room temperature

(Fig. 4.26). The corrosion potential is -2.57 V standard hydrogen electrode (SHE) in acidic

solution (pH 3.0) and -0.60 V SHE in near neutral solution (pH 6.8). respectively. The primary

activation current Is 10- 3 A/cm2 In pH 3.0 and 3 x 10-4 A/cm2 in pH 6.8 solution.

Electrochemical impedance (ECI), which provides Information for characterization of the

corrosion and passive layer at the specimen surface [21, 221, was measured for the Alloy 600

(Heat 96834 M81) specimens in a near-neutral (pH = 6.8) 0.4 M Na2SO4 at room temperature.

The measurements were performed in a frequency range of 103-105 Hz. Impedance and

current versus electrochemical potential are plotted for 0.02 and 100 Hz (Figs. 4.27 and

104
80

70

CD

.0S

60

60

40

30

20

10

1.5I

MU

=, 0.5

a.

-0.

0 200 400 600 800 1000

Time (h)

Figure 4.24.

Accuracy and stability of Ag/AgCl

reference electrodes that were

manufactured with different oxidation

current.

Figure 4.25.

Electrochemical polarization curve for

Alloy 600 specimen (Heat No. M81) in

0.4 M sodium sulfate aqueous solution

at 25, 100, and 2500C.

........ ........................ -6-- 25C

.- 6 18:81 Cyclic Polarization In O.4M

2 4 *

250-C

100-cI

-..- ... I...... -t

-8 -7 -6 -5 -4 -3 -2 -1

Log Current Density (APcm 2 )

4.28). High and the maximum impedance are observed in the passive range potentials at both

0.01 and 100 Hz. Additional ECI measurements will be performed at the other environmental

conditions (potential, pH, Pb concentration, and temperature), along with electrochemical

polarization tests.

105
1.6

U)

(a .5

'a

-0.5

- 9--.

(. --

: . ... P

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

* ... ,.

1 . 1 .~~~~~%. -. I ..

Figure 4.26.

Anodic polarization of Alloy 600 (Heat

96834 M81 annealed at 926.70C) In

0.4 M Na2SO4 at pH 3.0 and pH 6.8.

-1 I

-5.I

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

5 -5 -4.5 -4 -3.5 .3

Log Current Density (A/cm2)

Impedance Modulus (Ohm)

-2.5

0 5000 110 1.5 104 2 104

1.5

U,

ft0co

I

0.5

0

-0.5

Figure 4.27.

Electrochemical Impedance and

current at 0.02 Hz for Alloy 600 (Heat

98834 M81 annealed at 928.7°C) In pH

6.8 and 0.4 M Na2SO4 aqueous

solution at room temperature.

-1

Log current (Alcm2)

106
Log Current (Al cm2) -8 -7 -6 -5 -4 -3

1.5. ... ......... ........

I--u Impedance Modulus | W1 Log Current l

iW Figure 4.28.

I) 0.5. . . __..r.__.,,.._..,,_Electrochemical Impedance and

> 0.sf ^--current at 100 Hz for Alloy 600 (Heat

% ._,__.W 96834 M81 annealed at 926.7CC) in pH

CD 0, 6.8 and 0.4 M Na2SO4 aqueous

solution at room temperature.

0 -0.5 ' **1

2 3 4 5

Impedance Modulus(Ohm)

107
4.3 Tube Integrity Tests under Design Basis Conditions

Two facilities capable of testing steam generator tubes under design-basis conditions

were constructed as part of the program. One of these is a high-temperature Pressure and

Leak-Rate Test Facility (temperatures up to 3430C 1650'F!. pressures of up to 21 MPa

(3000 psi], and pressurized-water flow rates up to 1520 L/min (400 gpml) for conducting tube

failure and leak-rate tests under prototypical SG operating conditions. The other Is a roomtemperature High-Pressure Test Facility (pressures up to 52 MPa 17.500 psi] and flow rates up

to 48.4 L/min [12.8 gpmJ) that operates at high pressures to determine the structural margins

associated with flawed tubes. The High-Pressure Facility was also expanded to include a

containment module designed for testing radioactively contaminated field-pulled tubes.

4.3.1 High-Temperature Pressure and Leak-Rate Test Facility

The ANL High-Temperature Pressure and Leak-Rate Test Facility is used to obtain data

on the failure pressures, failure modes, and leak rates of flawed tubing at temperatures up to

3430C (650011, pressures of up to 21 MPa (3000 psi), and pressurized-water flow rates up to

1520 L/min (400 gpm). This facility was designed to be able to maintain pressures up to those

expected in design-basis accidents for flow rates approaching 1000 L/min (250 gpm). More

details on the design and operation of the facilities are given in Ref. 23.

The facility has a large heated blowdown vessel with water inventory of =760 L (200 gal)

that is able to provide high flow rates; piping and valves sized to minimize pressure drop in the

supply line to the flawed tube; nitrogen cover-gas pressurization in the blowdown vessel with

computer feedback control to maintain pressure even at high flow rates; a downstream backpressure regulator valve to control tube pressure differential, thereby minimizing nonprototypical two-phase flow from entering the tube and permitting rapid pressure control; and

a tube test module containment vessel. This test module can hold tubes of =0.15-1.0 m

(6-40 in.) in length and can also be fitted with support stings to simulate support plate

constraints on flawed tubes. A photograph of the facility Is shown in Fig. 4.29.

Diagrams of the facility and test module are shown in Figs. 4.30 and 4.31, respectively.

The pressurized and heated blowdown vessel holds -760 L (200 gal) of water and thereby

provides adequate water inventory for ramping the tube pressure differential P1-P2 and testing

tube specimens that contain preexisting throughwall flaws. Tube pressure differential ramp-up

Is produced through a back-pressure regulator valve (V6) located downstream from the test

module containment vessel. The pressure differential can also be ramped up by increasing the

gas overpressure in the blowdown vessel. Both modes of operation can be carried out under

manual or computer control. In the computer-controlled mode, the pressure can be

programmed to simulate a prototypical transient pressure history. During a test, the

blowdown vessel is maintained at a constant pressure P3 up to 21 MPa (3000 psi) by use of

regulated nitrogen cover gas pressurization from a 30-MPa (4400-psi) tank system of 0.9659

ML (34,110 standard ft3) capacity. The differential pressure P1-P2 across the specimen wall is

ramped up in a series of quasi-steady plateau increases through feedback control of valve V6,

which controls pressure P2. The pressure P2 can also be increased linearly with time at

differing rates. The flow rate Q, the pressure differential PI-P2, and the temperature are

recorded as a function of time.

108
~~ZFigure 4.29

ANL Pressure and Leak-Rate Test Facility.

A turbine flow meter is located immediately upstream of the test module to monitor the

leak rate. Load cells in the support legs of the blowdown vessel are used to monitor the mass

of the water in the blowdown vessel as a function of time, thereby providing a backup to and

cross check of the flow meter. Any flow through a flaw is routed outdoors through a large pipe

to minimize back pressure and is discharged into a large muffler for noise suppression.

Tests In this facility do not use bladders. The flow capacity Is adequate to test axial flaws

as large as 25-38 mm (1.0-1.5 in.) in overall length and to drive them open by increasing

pressure while the flaw is opening up. The differential pressure across the specimen is

controlled by changing the back pressure on the tube. The pressure In the blowdown vessel

during a test is always maintained sufficiently above the saturation pressure associated with

the desired water temperature to prevent water flashing to steam in the water supply piping

upstream of the test specimen. For flaws that do not leak initially or leak at a very low rate,

the pressure difference across the specimen is raised in 0.34-1.38 MPa (50-200 psi) increments

with a 5-10 s hold at each pressure plateau. If the leak rate increases, the pressurization rate

is correspondingly Increased. In general, the pressurization rate will not exceed 0.69 or

1.38 MPa/s (100 or 200 psi/s). If major ligament tearing occurs for a large flaw with a sudden

jump in flow rate to 380 L/min (100 gpm) or more, the test pressure is then set to a constant

value. The actual pressure will initially vary, but a steady-state pressure and corresponding

flow rate can be obtained through feedback control of the computerized gas pressurization

valves. The test Is shut down automatically when the vessel water inventory drops to a

minimum level.

109
Vented pressure

Gas-pressurized

blwdownvessel:

757 L Water

20.7MPa at 343oC.

Water storage &

conditiong tank: low Bldg. 212 H-wing

pressure & high bay south wag

temperature.

Vented pressure

Flawed SG tube release .

0

7helmal

N2 Cover 1as Load cells (3)

Figure. 4.30. High-Temperature Pressure and Leak-Rate Test Facility.
,1o1.6x50.$ 1

mmrconcentdc 203.2 x101.6m

Sch. 160 concentuic Sch.

/ reducer 160 reducer

blind raised face flange with 25.4-

mm female P-NPT thread for

Swagelock male tube connector, fa

bored 25.4-mm hole thru

wed steam generator tube 203.2-mm 2500-lb

blind flange

Figure 4.31. Tube test module for Pressure and Leak-Rate Test Facility.

Another type of test is conducted to determine whether flaws will exhibit time-dependent

opening under constant temperature and pressures associated with normal and main steam

line break (MSLB) conditions. In this test, a 8.3 MPa (1200 psi) pressure differential is held

from one to several hours (depending on leak rate and water Inventory), and the flaw is

monitored for signs of leak initiation or increased flow rate. After this hold period, the tube

pressure differential is increased in 0.69 or 1.38 MPa (100 or 200 psi) increments up to MSLB

pressure. At each step, the pressure is held constant for a period of a few seconds to several

minutes, depending upon the flow rate. After the MSLB pressure is reached, the pressure is

held for an extended time (dictated by remaining water inventory), and the leak rate monitored.

4.3.2 Room-Temperature High-Pressure Test Facility

The Room-Temperature High-Pressure Test Facility can produce pressures up to

51.7 MPa (7500 psi) and is used to investigate the design margin associated with flawed tubes.

Because the facility has a 48.4 L/min (12.8 gpm) flow capability at 51.7 MPa (7500 psi), tubes

with small throughwall flaws can be tested without the use of a bladder. For somewhat larger

cracks for which bladders and foils are required to maintain pressure, the small leaks that may

occur during tube pressurization are less likely to result in an aborted test.

This facility can also be used for long-duration tests of crack stability and for jetimpingement erosion tests on adjacent tubes, since water can be continuously supplied from

an outside source during a test. The apparatus is also capable of measuring very low flow

rates (considerably less than the 0.04 L/min 10.01 gpml lower limit for the Pressure and LeakRate Test Facility) through the use of flaw leak collection and timed measurement of water

mass. Additionally, real-time ultra-high-speed video recording (2000-12,000 frames per

second) of crack and jet dynamic behavior can be done during testing. A special containment

module for testing radioactively contaminated field-pulled tubes has also been built.

A schematic diagram of the facility is shown in Fig. 4.32, and a photograph of the facility

is shown in FIg. 4.33. Pressure is provided by a constant-rpm, single-acting triplex, constantdisplacement pump driven by a 60-hp electric motor that provides a constant 48.4 L/min

(12.8 gpm) flow of water at up to 51.7 MPa (7500 psi). Pump-generated pressure pulsation Is

<0.34 MPa (<50 psi). The pump has two overpressurization protection safety relief valves, one

located on the pump, and the other on the 139-L (40-gal) water accumulator suction tank. The

111
Pressure

Transducers (2)

Makeup Relief DivScrewr

Water Control Unloader Flow

I- _High- Valve Valve Meter

43 ~Pump li

i.I ;. 4>-_ _ _ _

e C LH Specimen E

Drain Protective Enclosure

Figure 4.32. Schematic diagram of High-Pressure Test Facility pressurizer and associated components.

lever operated hydro-diverter control valve originally supplied with the pressurizer was modified

to work in series with a screw unloader valve that allows more positive, finer control of

pressurization. The pressurizer water pump system can operate with a continuous supply of

water from the building water system.

The tube pressurization rate can be varied from quasi-steady-state (<0.7 MPa/s or

<100 psi/s) to 7-21 MPa/s (1000-3000 psl/s). which encompasses the rates given in EPRI

guidelines for burst testing to 70 MPa/s (10.000 psi/s). Pressurization over the full range of

the system. 0-52 MPa (0-7500 psi), is possible at all of these pressurization rates.

The module for testing of clean tubes (Fig. 4.32) consists of a 0.46-m (18-in.)-diameter x

1.02-m (40-in.)-long stainless steel pipe with transparent LexanTl plates at the ends of the

module to confine the water jet spray produced by a lealdng tube.

Figure 4.34 shows a schematic of the containment module for testing field-pulled tubes.

This module is also supplied with pressurized water and controlled from the High-Pressure

Facility. It Is essentially a 380 L (100 gal) stainless steel holding tank that captures all the

water that flows through the tube and flaw under test. Thus, any particulate contamination

released from the tube will be captured. Upon completion of a test, the water in the tank is

filtered to trap any particulate before the water is released. A breather/overflow vent prevents

module pressure buildup during testing and Is fitted with a filter to prevent any release of

particulate contamination. A level sensor alarm in the containment module Is used to Indicate

that the maximum water fill level has been reached, and the test should be terminated.

Whenever possible, tests are performed without bladders or reinforcing foils and under

quasi-steady pressurization rates to avoid complications in interpreting data and to maximize

information on flaw behavior. If flaws do not fail unstably in the initial test without a bladder,

112
Overall view of major components of High-Pressure Test Facility. Visible are the water

pump pressurizer, test module, tube support, video system, and 3000-L (800-gal) tank for

conducting tests on field-pulled tubes. Also shown is Lexanm personnel-protection barrier.

a bladder and possibly a backup foil are inserted In the tube to permit continuation of the test.

The most commonly used bladder is hard Tygon tubing with a 3.2-mm (1/9-in.) wall thickness.

The tubing is slightly smaller than the inside diameter of the specimen and is slipped inside

the flawed tube and sealed in each end with a Teflon plug (Figs. 4.35 and 4.36). For cracks

longer than 12.7 mm (1/2 in.), Stage 11 tests usually also require a 0.13-mm (O.005-in.)-thick

brass foil to be inserted between the inner wall of the SO tube and the outer surface of the

bladder. The foil extends 6.3 mm (0.25-in.) beyond the extremities of the flaw. The foil and

113
t

r 100 gal. Holdg tank T

(not pressurlz vessel) retentin tank

Lockable valatz ater

to drain filter

Figure 4.34. Containment module for testing radioactively-contaminated field-pulled tubes.

Figure 4.35.

Side view of tube with bladder and

bored plug seal being Installed.

114
U: Figure 4.36.

A i mEnd view of tube with bladder and plug

installed.

bladder are lubricated with a multipurpose spray lubricant to reduce friction between the foil

and the SG tubing. To avoid extrusion of the bladder/foil through the flaw, the tube is

pressurized at a rate of 3.5-7 MPa/s (500-1000 psi/s).

4.3.3 Test Results for Machined Flaws

All machined flaws were fabricated from 22.2-mm (7/8-in.)-diameter Alloy 600 SG tubing.

Electro-discharge machining and laser cutting were used to fabricate single notches (both axial

and circumferential), circular holes, and multiple flaw geometries of various types. Laser

cutting generates tighter flaws than can be achieved with EDM.

The flawed tubes were all 0.56-m (22-in.) long, with the flaw centered at 0.15 m (6 in.)

from the end of the tube fitted with a welded plug. The unplugged end of the tube was held in

place in the test module vessel of the Pressure and Leak-Rate Test Facility or the High-Pressure

Test Facility by a Swagelok fitting with a hole that permits flow to enter the flawed tube.

4.3.3.1 EDM Circular Holes

A series of tests was conducted on 22.2-mm (7/8-in.)-diameter SG tubes In the Pressure

and Leak-Rate Test Facility using room- and high-temperature water. The tubes had EDM

sharp-edged, circular throughwall holes The tests covered hole sizes with diameters ranging

from 0.40-6.35-mm (1/64-1/4-in.) and were initially used to check out the facility in bringing it

online.

Leakage from the SG tubes occured over a wide range of ratios of crack length (L) to crack

opening widths (h), as shown in Fig. 4.37. Most of the database used in formulating and

validating the current multiphase models for crack leak rate comes from tests simulating flowthrough SCCs in thick-wall piping. In these cases, L/h is typically >100. As shown in

FIg. 4.37, under MSLB conditions for cracks in steam generator tubes longer than about 6 mm

(0.25 in.), L/h is <100, i.e., these cracks are more open than SCCs in piping. The work of

Amos and Schrock 1241 and others shows that L/h strongly Influences whether choking can

occur and also the importance of friction and roughness effects. In addition, flashing of a

metastable liquid depends on a relaxation time, which implies that the actual wall thickness as

well as L/h may be important in determining when flashing occurs within the crack. For flows

through a relatively large orifice such as a 0.79-mm (1/32-in.)-diameter hole at 17.2 MPa

(2500 psi), the mean velocity of the jet is -120 m/s (-400 ft/s). Thus, the flow through the

115
10000

1000

100 Figure 4.37. Variation of the ratio of flow path

length to the crack opening width

10 with crack length for normal

operating and main steam line break

~~ I~~Lh Normal conditions.

1 - ~~Uh MSLB\

0.1 I p , * * .

0.1 0.2 0.3 0.4 0.5 0.6 1

Crack Length (in.)

tube wall requires -1 x 10-5 s. The time to travel across the 6.35 mm (0.25 in.) gap to the

adjacent tube Is -5.2 x 10-5 s. Because these times are short, in some cases, flashing may not

occur within the crack or possibly not even completely in the gap between tubes.

The flow rates for the 0.794-, 3.175-, and 6.35-mm (1/32-, 1/8- and 1/4-in.)-diameter

orifices at the nominal differential pressure of 17.2 MPa (2500 psi) were 2.76, 56.0, and

216 L/mln (0.73, 14.8, and 57.0 gpm). respectively. The average jet velocity through the

3.175-mm (1/s-in.)-diameter orifice at 17.2 MPa (2500 psi) was 112 m/s (366 ft/s). Such jet

velocities can produce significant lateral thrust, and the specimen with a 6.35-mm (1/4-in.)-

diameter orifice was bent -100 from its axis as a result of this thrust. No bending was observed

for the specimens with smaller orifices.

The standard con-elation for Incompressible flow through a sharp-edged circular orifice is,

Q=C .A.(p) (4.1)

where Q is the flow rate, C is the orifice discharge coefficient, A is the cross-sectional flow area.

Ap is the pressure difference across the orifice, and p Is the density of the fluid. Roomtemperature flow rates were predicted quite well using a value of 0.6 for C. The orifice model

predicts a temperature effect on flow rate through the dependence of fluid density on

temperature. The density of water at 21 and 2820C (70 and 540lF) is 1.00 and 0.745 g/cm3

(62.3 and 46.51 lb/ft3 ), respectively. In two separate tests, the same tube tested at a pressure

differential of 8.72 MPa (1265 psi) and a temperature of 2820

C (54001i exhibited a flow rate of

3.03 L/min (11.45 gpm). With C = 0.6, Eq. (4.1) predicts a flow rate of 3.04 L/min

(11.51 gpm), In very good agreement with the experimental value. The orifice correlation

predicts a flow rate of 37.7 L/min (9.95 gpm) for this same 3.175-mm (1/8-in.)-diameter hole

under the same pressure differential for room-temperature water, in good agreement with the

experimental value of 37.1 L/mln (9.8 gpm).

116
For the 0.79,- 3.18-, and 6.35-mm (1/32-, 1/8-, and 1/4-in.)-diameter orifices tested at a

pressure difference of 17.2 MPa (2500 psi) and room temperature, the orifice model predicts

flow rates of 3.33, 53.0, and 212 L/min (0.88, 14.0, and 56.0 gpm), respectively, which are

again in good agreement with the experimental values.

Additional tests were performed with orifice diameters of 3.18, 1.59, 0.79, and 0.40 mm

(1/8, 1/16, 1/32, and 1/64 in.) at room temperature (21PC or 700F) and elevated temperature

(2820C or 5400F) for pressure differences of 0-17.2 MPa (0-2500 psi). The ratio of the flow path

length (L) (wall thickness of the tube =1.27 mm 10.05-in.1) to hole diameter (h) for these hole

sizes is 0.4, 0.8, 1.6, and 3.2, respectively. Even the smallest of the orifices is still very open

compared to most cracks.

For the elevated-temperature tests, the internal tube pressurization was held at 17.2 MPa

(2500 psi), and the back pressure on the outside of the tube was varied from zero to a few MPa

(several hundred psi) to determine if the flow was choked. For all four orifice sizes, changes in

the back pressure produced changes in hole flow rate, indicating no choking. The orifice flow

model predicts the flow rates to within 8% of the experimental values in all the tests. In the

high-temperature tests, the water does not appear to flash to steam until after it exits the

orifices.

4.3.3.2 Axial EDM OD Notches

A series of tests on tubes with single 38.1-, 25.4-, 12.7-, and 8.9-mm (1.5-, 1.0-, 0.5-,

and 0.35-in.)-long axial EDM OD notches of various depths was performed at room and

elevated temperatures In the Pressure and Leak-Rate Test Facility. All flaws were nominally

0.20 mm (0.0075 in.) wide.

Three tubes C3EA8OX1.5, T4EA8OXI.5, and T7EA8OX1.5) with geometrically identical

EDM flaws were tested at room temperature. These axial OD flaws were 38.1-mm (1.5 in.)-

long, 0.20-mm (0.0075-in.)-wide, and 80% TW. Two of these tubes had the same as-received

heat treatment, and the remaining tube was sensitized (solution heat treated at 11000C for

10 min and aged at 6000C for 48 h) before testing.

The behavior of the three tubes was quite similar, and test details are therefore presented

only for tube T4EA8OX1.5. For this tube, the pressure difference across the tube wall was

increased In a series of pressure steps starting at 0.69 MPa (100 psi) in Increments of 1.38 MPa

(200 psi) up to 10.3 MPa (1500 psi), then in increments of 0.69 MPa (100 psi) up to 13.8 MPa

(2000 psi), and finally in increments of 0.34 MPa (50 psi). At -17.22 MWa (2497 psi), the flow

through the flaw abruptly increased from zero to a sustained leak rate of -1470 L/min

(388 gpm) in less than 1 s. The ANL ligament failure correlation4 predicted a failure pressure

of 17.24 MPa (2500 psi) for this flaw. The facility automatically shut down because of low

water inventory in the blowdown vessel. The facility was originally designed to sustain a leak

rate of at least 760 L/min (200 gpm), and this design limit was exceeded in this test. Upon

removal from the test module, the tube was found to have been bent by the jet thrust, and the

crack had opened up to a large hole.

Table 4.3 summarizes the failure pressures, flow rates, and posttest rupture geometries

for tests T3EA8OX1.5, T4EA8OX1.5, and 77EA80X1.5. The rupture geometries were obtained

by an analyses of digital photographs of the failures that project the actual three-dimensional

117
Table 4.3. Results from pressure and leak-rate tests on tubes with 38.1-mm (1.5-in.)-long 80% TW axial

EDM flaws. Tube T3EA80X1.5 was subjected to a sensitization heat treatment before

testing, and the remaining two tubes were tested in as-received mill-annealed condition.

Failure Flow Opening Overall Maximum

Pressure, Rate, L/min Area, Length, Width,

Tube ID MPa Ipsil [gpml CM2 lin. 21 cm/ [in.] cm/ [in.]

T3EA8OX1.5 13.8 1605 8.31 3.30 1.83

120001 14241 [0.5071 11.301 10.721

T4EA80X1.5 17.2 1469 7.41 3.35 1.60

124971 13881 10.4521 11.321 10.631

T7EA80X1.5 16.5 1453 7.24 3.38 1.57

[23981 13841 0.4421 11-331 [0.621

geometry onto a plane. The failure pressures, leak rates, and rupture geometry data for the

two tubes with the same heat treatment (T4EA80X1.5 and T7EA80X1.5) are in close

agreement. The sensitized tube (T3EA8OX1.5) exhibited a reduced failure pressure, an

increased flow area, and a larger opening. This heat treatment, which was used In the initial

stages of this program to increase susceptibility to SCC, clearly weakens tubes. The failure

behavior is consistent with the lower flow stress observed for this heat treatment 1251.

Forty-two tests were performed on shorter EDM OD notches. These notches were all

0.19 mm (0.0075 In.) wide and had four axial lengths, namely 6.35, 12.7, 19.1, and 25.4 mm

(0.25, 0.5, 0.75, and 1.0 in.). Flaws depths explored were 60, 80, 90, and 100% TW.

Room-temperature tests were run on specimens with 12.7-mm (0.5-in.) and 19.1-mm

(0.75-in.)-long, 80 and 100% TW flaws without a bladder, with a bladder, and with a bladder

plus backup foil to help understand the effects of bladders and foils on failure pressures.

These tests are summarized In Table 4.4. For Specimen OM120, which was tested without a

bladder, the throughwall ligament failed at a pressure of 26.9 MPa (3900 psl), which is in good

agreement with the predicted value of 25.5 MPa (3700 psi). No discernible unstable tearing

occurred at the notch ends. A 2.4-mm (3/32-in.) or 3.2-mm (1/8-in.)-thick hard Tygon bladder

was used for the tests with a bladder.

Tests with bladders were performed for specimens with 80% 'W flaws of 12.7 mm

(0.5 in.) (OM121; Fig. 4.38) and 19.1 mm (0.75-in.) (OM123) lengths, respectively. The tests

were conducted at a nominal pressurization rate of 13.8 MPa/s (2000 psl/s). Clearly, both

flaws experienced significant tearing at both ends of the original notches. The

experimental/predicted unstable burst pressures were 29.0/30.3 MPa (4200/4400 psi) and

21.9/22.8 MPa (3180/3300 psi), respectively.

A 12.7-mm (0.5-in.)-long 60% 'IW flaw, unlike an 80% '1W flaw of the same length, will

exhibit unstable tearing or bursting simultaneous with the failure of the throughwall ligament.

Such flaws can be tested with and without a bladder to assess the effect of bladders on burst

data. Tube OM I13 was tested without a bladder. The flaw burst unstably without a bladder at

118
Table 4.4. Summary of room-temperature pressurization tests conducted using High-Pressure Test

Facility on tubes with axial EDM notches to determine effect of Internal Tygon tube bladder

and 0.13-mm (0.005-in.)-thick brass backup foil on failure pressure. All tests with Internal

bladders were conducted at a pressurization rate of 13.8 MPWas (2000 psys); the remaining

tests were conducted at a quasi-steady-state pressurization rate.

Flaw Geometry Unstable Burst Pressure

Spec. Length, Depth, Internal Bladder and Experimental, Predicted,

ID mm (in.) %1W Backup Foil MPa (psi) MPa (psi)

OM120 12.7 (0.5) 80 None 26.9a (3900) 25.5a (3700)

OM121 12.7 (0.5) 80 2.4-nun (3/32-in.)-thick 29.0 (4200) 30.3 (4400)

bladder, no foil

OM123 19.1 (0.75) 80 2.4-mm (3/32-in.)-thick 21.9 (3180) 22.8 (3300)

bladder, no foil

OM1 13 12.7 (0.5) 60 None 40.7 (5900) 38.6 (5600)

OM112 12.7 (0.5) 60 2.4-mm (3/32-in.)-thick 41.4 (6000) 38.6 (5600)

bladder, no foil

OM102 12.7 (0.5) 100 2.4-mm (3/32-in.)-thtck 29.6 (4300) 30.3 (4400)

bladder, no foil

OMlOl 12.7 (0.5) 100 3.2-mm (l/8-in.)-thick 29.6 (4296) 30.3 (4400)

bladder, no foil

OM133 12.7 (0.5) 100 3.2-mm (1/8-in.)-thick 30.0 (4350) 30.3 (4400)

bladder plus foil extending 6.35 mm (0.25 in.)

beyond notch

OM134 12.7 (0.5) 100 3.2-mm (1/8-in.)-thlck 30.7 (4450) 30.3 (4400)

bladder plus foil

extending circumferentially 900 to either

side of notch

aPressure for failure of throughwall ligament; tube not tested to unstable burst.

Figure 4.38.

Posttest appearance of specimen

OM121 (12.7-mm [0.5-in.J-long 80%

TW axial EDM notch), tested with

bladder at a pressurization rate of

13.8 MPa/s (2000 psIs). Flaw burst

occurred at 29.0 MPa (4200 psi), as

compared to the predicted failure

pressure of 30.3 MPa (4400 psi).

119
40.7 MPa (5900 psi), as compared to the predicted burst pressure of 38.6 MPa (5600 psi).

Tube OM1 12 had the same flaw geometry, but was tested with a bladder. The tube failed at

41.4 MPa (6000 psi), which Is close to the predicted value of 38.6 MPa (5600 psi) and to the

failure pressure observed In the test without a bladder. Use of the bladder caused significantly

more unstable tearing of the notch ends than in Tube OM113.

Tests were also performed on tubes with 12.7-mm (0.5-in.)-long 100% TW EDM axial

notches. Such flaws cannot be unstably burst in the High-Pressure Test Facility without the

use of a bladder. This flaw geometry was tested with two different bladders. Tube OM102 was

tested with a 2.4-mm (3/32-in.)-thick bladder, and tube OMI01 was tested with a 3.2-mm (1/8-

in.)-thick bladder. The bladder was extruded through the flaw opening. The tube burst at

29.6 MPa (4300 psi), close to the predicted value of 30.3 MPa (4400 psi), but only a very slight

amount of unstable tearing occurred.

Tube OM101 (Fig. 4.39) was tested with the thicker bladder. The thicker bladder did not

extrude through the flaw, in contrast to tube OM102, and significantly more flaw bulging and

unstable tearing took place at the notch ends. Tube OM101 burst at 29.6 MPa (4296 psi).

which is good agreement with the predicted value of 30.3 MPa (4400 psi), despite the more

extensive tearing with the thicker bladder. Thus, the proper use of bladders produces a unique

value for the unstable burst pressure for a given flaw geometry, irrespective of the extent of

tearing. However, the resulting leak areas can be drastically different.

Backup foils are sometimes used to prevent premature bladder extrusion through

throughwall flaws and resulting excessive water leakage that would prevent pressurization to a

sufficient level to burst a tube. To evaluate the influence of foil backups, two tubes with

identical flaw geometries (axial notch 12.7-mm 10.5-in.- long and 1000% TW were tested at a

pressurization rate of 13.8 MPa/s (2000 psi/s) with 3.2-mm (1/8-in.)-thick bladders and two

designs of 0.13-mm (0.005-in.)-thick brass foil backups. Specimen OM133 (FIg. 4.40) had a

foil that extended 6.35 mm (0.25 in.) beyond the defining dimensions of the axial notch, in

conformance with EPRI guidelines. Specimen OM134 had a larger foil that extended

circumferentially 90° to either side of the notch and 12.7 mm (0.5 in.) beyond the ends of the

notch. The use of a larger foil can be advantageous when dealing with SCC flaws that cover a

large tube area and have additional secondary cracking that makes a precise determination of

flaw extent by pretest NDE or dye penetrant uncertain. In both cases, the internal tube surface

was lubricated with a multipurpose spray lubricant to reduce friction between the foil and the

tube.

Test OM133 was performed with a small foil, and OM134 with a large foil. Both flaws

exhibit more unstable tearing than test OM101, which used only a bladder (see Fig. 4.39).

However, the burst pressures of 30.0 MPa (4350 psi) for Specimen OM133 and 30.7 MPa

(4450 psi) for Specimen OM134 are close to the observed failure pressure of 29.6 MPa

(4296 psi) for Specimen OMI01 tested without the foil. Even though these tests show that the

foil coverage area has not influenced the failure pressure, this behavior could change if larger

foils were used or tubes that fail at considerably lower pressures were tested, in which case the

greater extent of foil might produce local strengthening and erroneous flaw-opening pressure

data. Likewise, thicker foils or foils made of stronger material such as stainless steel could

also produce unreliable data.

120
Figure 4.39 Top (left) and side (right) views of tube OMI01 with a 12.7-mm (0.5-in.)-long 100%W TW

EDM axial OD notch, tested with a 3.2-mm (1/8-in.)-thlck bladder at a pressurization rate

of 13.8 MPa/s (2000 psVs).

Figure 4.40.

Specimen OM133 with 12.7-mm (0.5-in.)-iong

100% TW OD axial EDM notch tested with

3.2-mm (1/8-in.)-thick hard bladder and small

backup foil at a pressurization rate of 13.8 MPals

(2000 psil/s). Failure pressure was 30.0 MPa

(4350 psi).

Room-temperature failure tests were also conducted on nine tubes with the same shortlength (6.35-mm I0.25-in.) long) EDM OD axial notches having depths of 60, 80 and 90% 1W.

The tests were conducted with and without bladders and foil backups at various pressurization

rates. The results from the nine tests are summarized In Table 4.5. The 60% TW flaws

(Specimens OM 1I8 and OMi 19) were tested at fast pressurization rates without and with a

bladder, respectively, and the flaws opened within 2.1 MPa (300 psi) of each other. These flaws

were predicted to burst unstably at 45.4 MPa (6600 psi) before tearing, but the actual failure

pressures were found to be somewhat higher. In addition, Specimen OMI18 exhibited no

unstable tearing, but OM 1 9 (tested with a bladder) exhibited significant tearing many times

longer than the original flaw length. This effect of the bladder on the flaw opening area is in

agreement with the observations for Specimens OM1 13 and OM1 12 described above.

Specimens OM 116 and OM 117, both of which had 80% TW flaws, were tested at low and

high pressurization rates without and with bladders, respectively. Specimen OM1 16 exhibited

no unstable burst and opened at 37.7 MPa (5475 psi), whereas OM1 17 tested with a bladder

exhibited unstable burst at 43.1 MPa (6252 psi). This flaw geometry was predicted to tear at

37.2 MPa (5400 psi) and to go unstable at 45.5 MPa (6600 psi). Specimen OM1 16, which could

not be pressurized sufficiently rapidly to induce unstable burst because of the absence of a

bladder, tore near the predicted value. Specimen OM I17, tested with a bladder went unstable

at a pressure -2.4 MPa (350 psi) lower than the predicted value for unstable burst. Again, the

flaws exhibit greatly different flaw opening areas.

121
Table 4.5 Test results for flawed tubes containing 6.35-mm (0.25-in.)-iong EDM OD axial notches.

Initial Flaw Tearing Unstable Burst Predicted

Pressuri- Pressure, Leak Pressure, Leak Ligament

Speci- Flaw zation MPa Rate, MPa Rate, Tear/Burst

men Depth, Rate, (psi) L/min (psI) L/min Press.,

ID (% TWI MPa/s (gpm) (gpm) MPa (psi)

(pSl/s)

OM1 18 60 13.6 49.6 46.9 d d 48.3 (7000)/

(1970) (7200) (12.8) 45.5 (6600)

OM119a 60 20.7 4 7 .6 a >4 9 a 4 7 .6 a e >49 a.e 48.3 (7000)/

(3000) (6 90 0)a (13)a (6 90 0 )a.e (>13)a.e 45.5 (6600)

OM116 80 Quasi- 37.7 12.5 d d 37.2 (5400)/

steady (5475) (3.31) 45.5 (6600)

OM117a 80 21.4 43.1a >4 9 a 4 3 .

1 axe >4 9 a~e 37.2 (5400)/

(3110) (6 25 2 )a (13)a (62 5 2 )ae (>13)a.e 45.5 (6600)

OM138 90 Quasi- 30.0 4.2 4 2 .7 d 4 5 .

0 d 27.6 (4000)/

steady (4350) (1.1) (6 2 00 )d (ll.9 )d 45.5 (6600)

OM139 90 5.2 28.3 2.8 43.8d 4 5 .4 d 27.6 (4000)/ (750) (4100) (0.75) (63 5 6 )d (1 2 .0 )d 45.5 (6600)

OM140 90 22.1 33.8 4 4 .8 d 4 5 .4 d 27.6 (4000)/

(3200) (4900) (64 9 2 )d (12 .0)d 45.5 (6600)

OM141b 90 4.9 30 .3 b 4 .5 b 4 5 .0b-e 4 8 .4 b.e 27.6 (4000)/

(705) (44 0 0)b (1 .2)b (6 52 0 )b~e (12. 8)b.e 45.5 (6600)

OM142c 90 6.3 2 7.6 c 4 .9C 41.5ce 48.4C.e 27.6

(916) (40 08)C (1.3)C (6026)C.e (12.8)C.e (4000)/

45.5 (6600)

arested with 2.38-mm (3/32-in.)-thick bladder.

brested with 3.18-mm (1/8-in.)-thlck bladder.

0 rested with 3.18-mm (1 /8-ln.)-thlck bladder with foil backlng.

dNo unstable burst.

eUnstable burst.

Three of the five 90% TW flaws (Specimens OM138, OM139, and OM140) were tested

without bladders at progressively increasing pressurization rates of quasi-steady-state, 5.2,

and 22.1 MPa/s (750 and 3200 psi/s), respectively. None of these flaws exhibited unstable

burst. This flaw geometry was predicted to first tear at 27.6 MPa (4000 psi) and then to go

unstable at 45.5 MPa (6600 psi). The specimens actually tore open at pressures of 30.0, 28.3,

and 33.8 MPa (4350, 4100, and 4900 psi), respectively. Thus, the results for the tests

conducted at the two lower pressurization rates were in good agreement with the predicted

values, but the specimen tested at the highest pressurization rate tore open at 6.2 MPa

(900 psi) above the predicted value. This dependence of flaw opening pressure on

pressurization rate has also been observed in testing of 19.1-mm (0.75-in.)-long axial notches,

as described at the end of this section. For all three specimens, the leak rates upon flaw

opening were below the facility maximum of nominally 45 L/min (12.8 gpm), and the flaws

were therefore pressurized to higher levels of 42.7, 43.8, and 44.8 MPa (6200, 6356, and

6492 psi), respectively. At this point, the flow rates reached the facility maximum, and further

pressurization to unstable burst (-45.5 MPa or 6600 psi) could not be achieved.

122
This same flaw geometry was tested with a bladder (Specimen OM141) and a bladder/foil

combination (Specimen OM 142). Both of these specimens went unstable and exhibited greatly

increased flaw opening areas. Both of these flaws first tore open near the predicted valued of

27.6 MPa (4000 psi) and were then pressurized to higher levels. Specimen OM141, tested with

only a bladder, exhibited unstable burst at 45.0 MPa (6520 psi), which is very near the

predicted value of 45.4 MPa 46600 psi). However, Specimen OM142, which used a foil backup.

burst at 41.5 MPa (6026 psi). The reason for this behavior is not clear.

To supplement the preceding on the influence of pressurization rate on a short flaw, three

identical 80% TW, 19.1-mm (0.75-in.)-long axial notches (OM107, OM122, and OM109) were

tested in three different pressurization scenarios. A low quasi-steady-state pressurization rate

resulted in ligament tearing at 21.4 MPa (3100 psi), pressurization at 13.8 MPa/s (2000 psi/s)

produced ligament tear at 24.2 MPa (3510 psi), and pressurization at 48.3 MPa/s (7000 psi/s)

produced ligament tear at 26.4 MPa (3830 psi). The predicted ligament tearing pressure for

this flaw geometry is 20.7 MPa (3000 psi), which is in good agreement with the quasi-steady

pressurization rate test OM 107. All three initially identical flaws exhibited ligament tearing but

no discernible unstable tearing at the notch ends. In addition to the increase in ligament tear

pressure with pressurization rate, the flaws exhibit a slight increase in maximum notch

opening width with increasing pressurization rate. This behavior is similar to that for the

6.35-mm (0.25-in.)-long axial notches presented above.

A nine-tube set of 22.2-mm (7/8-in.)-diameter Alloy 600 tubes was prepared with pairs of

EDM 6.35 or 12.7-mm (0.25 or 0.5-in.)-long axial notches and was tested in the High-Pressure

Test Facility. The flaws were either 70% (for tube OM153) or 80% (remaining eight tubes) 'W

separated by various full-wall-thickmess ligaments of lengths 0.25, 1.25, and 2.5 mm (0.010,

0.050, and 0.100 in.), all with an initial width of 0.19 mm (0.0075 in.). The flaws were either in

axial alignment or shifted circumferentlally parallel to each other to give the required ligament

size.

Table 4.6 summarizes the results on the pairs of flaws. Data on the pressure at initiation

of throughwall leakage and unstable burst, if obtained, are presented. Figures 4.41-50 show

posttest photographs for some of the flaws. The flaws were fist tested without bladders using

quasi-steady pressurization consisting of raising the pressure in 0.69 MPa (100 psi) increments

with a 5 s hold at each pressure plateau. The testing was stopped when the flaw first opened,

as indicated by the abrupt onset of flow. None of the flaws experienced unstable tearing at the

end of the quasi-steady pressurization testing. Most of the flaws exhibited simultaneous

tearing of the part-through ligaments at the base of the notch and of the ligament separating

the segments in the first phase of testing. In specimens OM152, OM159, and OM162, the

part-through ligaments failed, but the separating ligament remained intact (see Figs. 4.44 and

4.47).

All the flaws consisting of two 12.7-mm (0.5-in.)-long notches leaked the maximum

48.8 L/min (12.9 gpm) capacity of the facility upon opening. Hence, pressurization could not

be sustained after the flaws opened. In contrast, the 6.35-mm (0.25-in.)-long flaws when

opened in Phase 1 testing were capable of sustained pressurization because the flow rates were

less than the facility maximum. The sustained flow rate and pressure for these flaws are

presented in Table 4.7.

123
Table 4.6 Pressures at first flaw opening and unstable burst for a nine-tube set of 22.2-mm (7/8-In.)-

diameter Alloy 600 tubes containing pairs of 6.35- or 12.7-mm (0.25- or 0.5-in.)-long axial

EDM notches of 70 and 80% TW separated by various lengths of full-wall-thickness

ligaments.

Flaw Geometry Pressure at

Notch Ligament Flaw Pressure at Unstable

Specimen Lengths, Length, Depth, % First Leak. Burst,

No. mm (in.) mm (in.) 1W MPa (psi) MPa (psi)

OM149 12.7 (0.5) 2.5 (0.10) 80 27 .Oa (3920)

OM150 12.7 (0.5) 0.25 (0.01) 80 19.Oa (2760)

OM151 12.7 (0.5) 1.25 (0.05) 80 23.Oa (3340) -

OM152 12.7 (0.5) 2.5 (0.10) 80 18 .9 aab.c( 2 7 4 4 ) 23.5 (3405)

OM153 6.35 (0.25) 0.25 (0.01) 70 3 4 .8 a (5048) -

OM159 6.35 (0.25) 2.5 (0.10) 80 3 4 .

8 a~bc (5050) 38.3 (5550)

OM160 6.35 (0.25) 1.25 (0.05) 80 3 2 .3 a (4675) -

OM161 6.35 (0.25) 0.25 (0.01) 80 2 8 .3 a (4100) -

OM162 6.35 (0.25) 1.25 (0.05) 80 31.4ab.C (4560) 32.5 (4710)

aNo unstable tearing at this pressure.

bOnl bottom of notches tore open; no ligament tearing.

CFlaw was retested with bladder and 13.8 MPa/s pressurization rate to cause unstable tearing.

Figure 4.41.

Specimen OM149 with two aligned axial notches

[12.7-mm (0.5 In.) long, 2.54-mm (0.100-in.)

ligament, 80% TWJ after Phase 1 testing.

Figure 4.42.

Specimen OM150 with two shifted axial notches

112.7-mm (0.5-In.) long, 0.25-mm (0.010-In.)

ligament, 80% TW] after Phase 1 testing.

124
Figure 4.43.

Specimen OM151 with two shifted axial notches

[12.7-mm (0.5 in.) long; 1.27-mm (0.050-in.)

ligament, 80%h TWJ after Phase 1 testing.

Figure 4.44.

Specimen OM152 with two shifted axial notches

[12.7-mm (0.5 in.) long; 2.54-mm (0.100-in.)

ligament, 80% TW] after Phase 1 testing.

Figure 4.45.

Specimen OM160 with two aligned axial notches

[6.35-mm (0.25-in.) long, 1.27-mm (0.050-in.)

ligament, 80%h TWI after Phase 1 testing.

Figure 4.46.

Specimen OM1 61 with two shifted axial notches

[6.35-mm (0.25-in.) long, 0.25-mm (0.010-in.)

ligament, 80% TWI after Phase 1 testing.

Figure 4.47.

Specimen OM162 with two shifted axial notches

6.35-mm (0.25-in.) long, 1.27-mm (0.050-in.)

ligament, 80%h T1q after Phase 1 testing.

Figure 4.48.

Specimen OM152 with two shifted axial notches

12.7-mm (0.50-in.) long, 2.54-mm (0.100-in.)

ligament, 80% TW] after Phase 2 burst testing.

125
Figure 4.49. Figure 4.50.

Specimen OM159 with two shifted axial notches Specimen OM162 with two shifted axial notches

[6.35-mm (0.25-in.) long, 2.54-mm (0.100-in.) [6.35-mm (0.25-In.) long, 1.27-mm (0.050-in.)

ligament, 80% TW] after Phase 2 burst testing. ligament, 80% TW] after Phase 2 burst testing.

Table 4.7. Sustained pressure and leak rate during Phase 1 testing after flaw opening of two paired

6.35-mm (0.25-in.)-long axial notches.

Flaw Geometry

Notch Ligament Flaw Sustained Sustained

Specimen Lengths, Length, Depth, % Pressure, Leak Rate,

No. mm (in.) mm (in.) 1W MPa (ps) L/min (gpm)

OM153 6.35 (0.25) 0.25 (0.01) 70 5.4 (780) 47.69 (12.60)

OM159 6.35 (0.25) 2.5 (0.10) 80 33.2 (4820) 26.08 (6.89)

OM160 6.35 (0.25) 1.25 (0.05) 80 3.9 (560) 48.45 (12.80)

OM161 6.35 (0.25) 0.25 (0.01) 80 13.8 (2000) 46.29 (12.23)

OM162 6.35 (0.25) 1.25 (0.05) 80 29.0 (4200) 32.89 (8.69)

In a second phase of testing, Specimens OM152, OM159, and OM162 were retested with

a bladder and a 13.8 MPa/s (2000 psi/s) pressurization rate to obtain information on unstable

burst pressure. Figures 4.48-4.50 show the burst flaw openings.

The data in Table 4.6 furnish some insight on how pairs of part-TW axial notches

separated by a ligament interact. The data for Specimens OM149 (27.0 MPa) and OM152

(23.5 MPa) show that, for 12.7-mm-long flaws with 2.5-mm ligaments, the axially aligned flaws

are stronger than the circurnferentially shifted notches with ligaments of the same size.

Similarly, for paired 6.35-mm-long notches with 0.25-mm ligaments, i.e., Specimens OM153

(34.8 MPa) and OM161 (28.3 MPa), the axially aligned flaws are stronger. The same behavior is

also seen for paired 6.35-mm-long notches with 1.25-mm ligaments, i.e., Specimens OM160

(32.3 MPa) and OM162 (31.4 MPa).

Table 4.8 shows data for six tests from Table 4.6 selected to Illustrate the effect of

ligament width on the failure pressures. As can be seen, the opening pressure for each flaw

length increases to a value that is increasingly less sensitive to ligament size. This finding

126
Table 4.8. Opening pressures of six flaws as a function of separating ligament size for paired 6.35- and

12.7-mm (0.25- and 0.50-in.)-long interacting circumferentlally shifted axial notches, all 80%

TW.

Individual Ligament Opening

Specimen Notch Length, mm Size, Pressure,

No. (in.) mm (in.) MPa (psi)

OM161 6.35 (0.25) 0.25 (0.01) 28.3 (4100)

OM162 6.35 (0.25) 1.27 (0.05) 31.4 (4560)

OM159 6.35 (0.25) 2.54 (0.10) 34.8 (5050)

OM150 12.7 (0.50) 0.25 (0.01) 19.0 (2760)

OM151 12.7 (0.50) 1.27 (0.05) 23.0 (3340)

OM152 12.7 (0.50) 2.5 (0.10) 23.5 (3405)

implies that the flaws are interacting less. Furthermore, the limiting pressure for each flaw

length is approaching the predicted opening pressures for single flaws of 6.35- and 12.7-mm

length, which are 34.5 and 20.7 MPa (5000 and 3000 psi), respectively.

4.3.3.3 Laser-Cut Notches

Additional data on flaw interaction were obtained from tests on twenty-four 22.2-mm

(7/8-in.)-diameter Alloy 600 tubes with a variety of laser-cut OD flaws consisting of simple

single axial and circumferential notches, as well as complex OD patterns of axial and

circumferential notches separated by various ligament sizes. Ten generic flaws were tested

with the geometries shown in Figs. 4.51 and 4.52. The laser-cutting technique permits

considerably more complex flaws to be fabricated than are possible using EDM. The laser-cut

flaws are also considerably tighter than the minimum 0.19-mm (0.0075-in.)-wide notches that

can be achieved by EDM.

The 24 laser-cut flaws were tested in two stages. The first stage consisted of pressurizing

at a quasi-steady-state rate without the use of internal bladders. The pressurization protocol

consisted of starting at an internal pressure of 0.34 MPa (50 psi) and increasing the pressure

in 0.69 MPa (100 psi) increments after a 10 s dwell at each pressure plateau. This process

continued until ligament tearing occurred, and the resultant increases in flow made it

impossible to continue the tests. The specimens that exhibited only stable ligament tearing

and opening in the first-stage tests were retested with a bladder and foil backups, If necessary.

The results from the 24 Stage 1 tests are summarized in Table 4.9. Pre- and post-test

photographs of some representative flaws of the ten types are shown in Figs. 4.53-55.

Table 4.9 also gives pretest EC estimates of the maximum flaw depth. All of the laser-cut flaws

exhibited pinhole laser burn-through at the ends of each notch. These pinholes were too small

to be detected by EC techniques, but they leaked water at 0.34 MPa (50 psi). At a pressure of

6.9 MPa [1000 psi), the leak rate through these pinholes was typically in the range of 0.04-

0.4 L/min (0.01-0.1 gpm or 14.4-144 gal per day).

127
=

Notch width

as narrow as

possible

Type 2 specimen

Notch Length -I I Notch widffi = =~~~~~~~a narrow as

T possible

8046 'Troughwafl notch

Ugament Width /

Type 2 specimen

Ligament Width

I *1

Notch Length

Type 3 specimen

id Notch Length q

C== i = Ligament

.===== C=== - width I = Ligament width

I Notch length __

esial

Type 4 specimen Type 5 specimen

Figure 4.51. Configurations of notches In various types 1-5 of laser-cut tube specimens.

128
I

I

U

Axial

Type 6 specimen

180°0

U8

14

1200

600

00

-60°

-1200

-180°

UIgament

width AM

1800

1200

60°

00

-600

-1200

Axial -180°

' Ugament Axial width

Type 7 specimen Type 8 specimen

1800

1800

8 900

-900

-1800

I

900

00

-900

Axdal

-180°

Axial

Type 9 specimen Type 10 specimen

Figure 4.52. Configurations of notches in various types 6-10 of laser-cut tube specimens.

129
Table 4.9. Test results from Stage 1 quasi-steady-state testing of flawed tubes containing OD laser-cut

notches (without bladders). All notches had nominal depths of 80% TW except for

Specimens 5528-3-3 and 5469-2-2, which had nominal depths of 40%O TW. All tubes leaked

water through laser bum pinholes at the ends of the laser-cut notches at 0.34 MPa (50 psi)

except Specimen 5528-2-2.

Max. Sust. Press. Flow Rate

Nominal Pressure after after

Number NDE Flaw Flaw Ligament at Ligament Ligament L4 Tearing.

Specimen Flaw of Depth Length. Width. mm Tearing. Tearing, MPa L/mln

Number Type Notches (% TW) mm (in.) (In.) MPa (psi) (PSI) (pm)

5528-1-1

5528-1-2

5528-1-3

5528-114

5528-2-1

5528-2-2

5516-4-3

5516-4-2

5528-3-1

5528-3-2

5528-3-3a

5528-3-4

5469-2-1

5469-2-2a

5469-2-3

54692-4

5469-3-1

5531-3-1

5469-3-3

5469-3-4

5469-4-1

5469-4-2

5469-4-3

1 1 75 6.35

(0.25)

1 1 74 6.35

(0.25)

1 1 76 8.89)

(0.35

1 1 73 8.89)

(0.35

1 1 69 12.7

(0.5)

1 1 61 12.7

(0.5)

2 2 72 12.7

(0.5)

2 2 70 12.7

(0.5)

3 6 70 12.7

(0.5)

3 6 73 12.7

(0.5)

3 6 61 12.7

(0.5)

4 6 76 12.7

(0.5)

4 6 74 12.7

(0.5)

4 6 60 12.7

(0.5)

5 6 79 12.7

(0.5)

5 6 78 12.7

(0.5)

6 6 80 12.7

(0.5)

6 6 79 12.7

(0.5)

7 2 75 360°

NA

NA

NA

NA

NA

NA

0.25

(0.01)

0.13

(0.005)

0.13

(0.005)

0.25

(0.01)

0.25

(0.01)

0.13

(0.005)

0.25

(0.01)

0.25

(0.01)

0.25

(0.01)

0.51

(0.02)

0.25

(0.01)

0.51

(0.02)

0.13

(0.005)

0.25

(0.01)

0.13

(0.005)

0.25

(0.01)

NA

NA

36.6

(5310)

37.1

(5380)

31.7

(4600)

32.4

(4700)

34.5

(5000)

38.6b

(5 60 0 )b

33.8

(4900)

33.8

(4900)

29.0

(4200)

30.0

(4350)

35.2

(5100)

35.9

(5200)

32.8

(4750)

37.6

(5450)

33.5

(4860)

39.3b

(500)b

28.3

(4110)

34.3

(4980)

51.4

(7450)

51.7

(7500)

39.2

(5680)

38.3

(5550)

30.8

(4475)

31.3

(4540)

36.5

(5300)

37.0

(5360)

29.8

(4320)

31.2

(4530)

3.5

(500)

1.4

(200)

7.2

(1040)

6.5

(940)

10.7

(1550)

8.5

(1240)

3.7

(540)

2.7

(390)

6.0

(870)

3.8

(550)

2.1

(800)

1.0

(150)

13.3

(1930)

5.2

(760)

0.4

(60)

did not fail

0.3

(50)

0.3

(50)

0.7

(100)

0.7

(100)

18.1

(4.78)

13.8

(3.651

32.6

(8.6)

30.5

(8.05)

48.1

(12.7)

48.4

(12.81

48.1

(12.7)

48.1

(12.7)

48.1

(12.7)

48.1

(12.7)

48.4

(12.8)

48.4

(12.8)

48.4

(12.8)

48.4

(12.8)

48.4

(12.8)

48.4

(12.8)

46.9

(12.4)

48.4

(12.8)

49.2

(12.99)

did not fall

48.8

(12.9)

48.8

(12.9)

48.8

(12.9)

48.8

(12.9)

7

8

8

9

2 74 3600

6 70 360°

6 70 360°

2 69 180° x 0.5

5469-44 10 2 72 180° x 0.5

aLaser-cut notches had nominal depths of 40% 7W.

bSpecimen ehibited unstable tearing without internal bladder.

130
Figure 4.53. Appearance of Specimen 5516-4-3 with Type 2 flaw before (left) and after (right) Stage I

testing.

Figure 4.54. Appearance of Specimen 5528-3-3 with Type 3 flaw before (left) and after (right) Stage 1

testing.

- Ab~~~~~-

Figure 4.55. Appearance of Specimen 5469-2-2 with Type 4 flaw before (left) and after (right) Stage 1

testing.

131
The six Type 1 flaws (Fig. 4.51) were single axial notches of three lengths (6.35, 8.9, and

12.7 mm 10.25, 0.35, 0.5 In.), and all were nominally 80% TW. The two Type 2 flaws

(Fig. 4.51) consisted of two aligned axial notches, each 6.35 mm (0.25 In.) In length and 80%

TW, separated by a complete wall thickness ligament of lengths 0.25 or 0.13 mm (0.010 or

0.005 in.). During testing, progressive sequential tearing was not evident; both types of

ligaments tore simultaneously at 33.8 MPa (4900 psi).

The three Type 3 flaws (Fig. 4.51) consisted of six aligned equal-length axial notches

separated by complete-wall-thickness ligaments. The total length of the composite flaw

(notches plus ligaments) was 12.7 mm (0.5 in.). For two of the tubes, the flaws were 80% TW,

and the interspersed ligaments were 0.13 or 0.25 mm (0.005 or 0.010 in.) in length. For the

third tube, the six notches were only 40% 'W, and the interspersed ligaments were 0.25-mm

(0.010-in.) long. Again, in the tests on these specimens, progressive sequential tearing of the

ligaments was not evident. For this type of flaw and the dimensions tested, the length of the

ligaments separating the notches had little influence on failure pressure, but the depth of the

notches exerted a noticeable influence.

The three Type 4 flaws (Fig. 4.51) consisted of six nonaligned, side-step, equal-length,

axial notches in two axial rows of three aligned flaws, each separated by complete-wallthickness ligaments. The overall length of the composite flaw was 12.7 mm (0.5 in.). In two of

the three tubes, the individual notches were 80% TW, and the interspersed ligaments were

0.13 or 0.25 mm (0.005 or 0.010 In.) in length. In the third tube, the Individual notches were

only 40% 1W, and the interspersed ligaments were 0.25-mm (0.010-in.) long. Figure 4.55 show

the pre- and post-test flaw appearances of Specimen 5469-2-2, which had the 40% 'W flaw

depth. The post-test photograph indicates no unstable tearing at the notch ends and shows

that the ligaments between the notches and those at the bottom of each notch have torn.

Again, all of the ligaments tore simultaneously at the failure pressure of 37.6 MPa (5450 psi),

and the depth of the notches exerted a stronger influence on the failure pressure than did the

length of the interspersed ligaments.

The two Type 5 flaws (Fig. 4.51) were composed of six individual notches of equal length

in two parallel offset rows of three notches each. The laser-cut notches in both specimens were

80% 1W. and the ligament lengths were either 0.25 or 0.51 mm (0.010 or 0.020 in.).

Figure 4.56 show the pre- and posttest appearances of Specimen 5469-2-4. which had the

0.51-mm (0.020-in.) ligaments. The post-test photograph shows that ligament tearing occurred

along only one of the unaligned axial rows, and in contrast to the previous tests, unstable

tearing also occurred at the ends of flaw. This unstable tearing occurred without the use of an

internal bladder. Progressive sequential tearing was not evident during testing; all ligaments

tore simultaneously at 39.3 MPa (5700 psi). The other flaw of this type (Specimen 5469-2-3)

with the thinner (0.25-mm 10.010-in.J ligament between notches did not tear unstably, and It

opened at the considerably lower pressure of 33.5 MPa (4860 psi).

The two Type 6 flaws (Fig. 4.51) consisted of six axial parallel notches of equal 12.7-mm

(0.50-in.) length and 80% TW depth with ligament widths of 0.25 or 0.51 mm (0.010 or

0.020 in.). For Specimen 5531-3-1, which had the 0.51-mm (0.020-in.)-long ligaments,

ligament tearing occurred in only one of the parallel notches, and no unstable tearing occurred

at the ends of notch. The flaw tore open at 34.3 MPa (4980 psi). The other flaw of this type

(Specimen 5469-3-1) with the thinner (0.25-mm 10.010-in.J)-long ligament between the notches

behaved similarly but opened at the considerably lower pressure of 28.3 MPa (4110 psi).

132
Figure 4.56. Appearance of Specimen 5469-2-4 with Type 5 flaw before (left) and after (right) Stage 1

testing.

The two Type 7 flaws (Fig. 4.51) consisted of two parallel 3600 continuous circumferential

notches of 80% TW depth separated by ligament widths of 0.13 or 0.25 mm (0.005 or

0.010 in.). Specimen 5469-3-3 had the thinner 0.13-mm (0.005-in.) ligament, and ligament

tearing occurred around the entire tube circumference, resulting In complete separation of the

tube into two pieces. Failure occurred at 51.4 MPa (7450 psi), which Is near the maximum

53.1 MPa (7700 psi) operating pressure of the facility. The other flaw of this type

(Specimen 5469-3-4) had the thicker (0.25-mm 10.010-in-l) ligament and could not be opened

at the maximum facility pressure of 53.1 MPa (7700 psi.

The two Type 8 flaws (Fig. 4.51) consisted of six equal-length circumferential notches of

80% TW depth in two circumferential parallel rows having three notches each, separated by

ligaments widths of 0.13 or 0.25 mm (0.005 or 0.010 in.). In Specimen 5469-4-1, which had

the thicker (0.25-mm 10.010-in.1) ligament, ligament tearing occurred around the entire tube

circumference, resulting in complete separation of the tube into two pieces. Failure occurred

at 39.2 MPa (5680 psi). The other flaw of this type (Specimen 5469-4-2) with the thinner

(0.13-mm 10.005-in.J) ligament opened at the slightly lower pressure of 38.3 MPa (5550 psi).

The single Type 9 flaw (Fig. 4.51) consisted of two intersecting notches, one extending

1800 around the circumference and the second perpendicular notch extending 12.7 mm

(0.5 in.) in the axial direction. The two notches were 80% TW and formed a tee. Liganent

tearing occurred primarily along the 12.7-mm (0.5-in.)-long axial notch and only slightly along

the circumferential notch. The flaw tore open at 30.9 MPa (4475 psi).

The single Type 10 flaw (Fig. 4.51) consisted of two intersecting notches similar to those

in the Type 9 flaw. However, the notches in the Type 10 flaw were configured to form an el

rather than a tee. For this geometry, ligament tearing again occurred primarily along the

12.7-mm (0.5-in.)-long axial notch and only slightly along the circumferential notch. The flaw

tore open at 31.3 MPa (4540 psi). very near the failure pressure for the Type 9 flaw.

Most of the tubes from the 24-tube set were subjected to Stage 2 testing after the above

Stage 1 tests were completed. Initially, the Stage 2 tests were conducted with only a bladder

(3.2-mm 11/8-in.1-thick Tygon tubing) and no foil backup. Seven tubes were tested in this

fashion, namely Specimens 5528-1-1, 5528-1-2, 5528-1-3, 5528-1-4. 5518-3-1, 5516-4-2, and

133
5516-4-3. Al flaws after this were tested with a similar bladder and a 0.13-mm (0.005-in.)-

thick brass foil backup. The results of the Stage 2 testing are summarized In Table 4.10.

4.3.3.4 Pressurization of Flaws to Obtain Specified Openings

The Paul Sherrer Institute (PSI) of Switzerland requested that ANL produce axial fishmouth flaw openings of specified sizes In Type 304 stainless steel (SS) tubes for subsequent

testing at the Institute. Although only three fish-mouth specimens were needed, PSI provided

10 tubes to permit experimentation in developing the appropriate production technique. Each

of the tubes had an OD of 18.8 mm (0.741 In.) and an ID of 16.6 mm (0.655 in.) and contained

100% 'SW EDM notches with an initial 0.20-mm (0.0075-in.) width. Four of these ten tubes

had a notch length of 20 mm (0.8 in.), three had a length of 33 mm (1.3 in.), and three had a

length of 48 mm (1.9 in.). ANL was to expand these flaw openings to maximum widths of 10,

5.6, and 4.4 mm (0.4. 0.22, and 0.17 in.). respectively, by Internal pressurization. Producing

'flaws to order" provided a unique test of the capability to predict crack opening and failure

behaviors.

The pressurization tests were carried out using internal bladders to prevent excessive

leakage through the flaw and reinforcing circumferential brass foils, typically extending 2700

around the circumference. to prevent extrusion of the bladder through the flaw opening during

pressurization. Figure 4.57 shows that the calculated crack opening displacements (CODs)

were in reasonable agreement with test data for the 20.3- and 33.0-mm (0.8- and 1.3-in.)-long

notches. The COD for the 48.3-mm (1.9-in.)-long notch was significantly overestimated.

However, since the COD vs. pressure curves are fairly steep for all three notches, the final

pressures recommended were quite close to the actual pressures applied during the tests.

Overall, the values for the target notch width were met within reasonable limits trlable 4.11).

4.3.4 Laboratory-Grown SCC

In contrast to rectangular machined notches of constant depth, laboratory-grown SCCs

are irregular in shape and have variable depths along their lengths. They often consist of a

family of crack segments in different planes rather than consisting of a single planar crack.

Nineteen leak and burst tests have been performed on laboratory-grown SCC flaws. Flaw

geometries include ID and OD axial and ID and OD circumferential flaws. The crack

geometries for the flaws were estimated prior to testing by multiparameter and conventional EC

analysis. Where possible, post-test fractography was also performed to confrm the NDE

estimates, but in cases where the burst pressure was high, the resultant large deformations of

the tube made the post-test analysis difficult. An additional 31 tube set, which was used for

detailed studies on the accuracy of NDE crack sizing and characterization, was also tested.

The purpose of the tests on these tubes was not to characterize the failure and leakage

behavior of the tubes, but to moderately pressurize the tubes to slightly open the flaws. This

step simplified destructive examination of the specimens and allowed accurate profiles of the

cracks undistorted by the large deformations that often occur when tubes are tested to failure.

The initial tests on laboratory-grown SCC flaws were performed on five 22.2-mm

(7/8-in.)-diameter Alloy 600 tubes with laboratory-produced axial ODSCC of nominal length

12.7 mm (0.5 in.). Four of the tubes, SGL-177, SGL-195, SGL-104, and SGL-219, were

cracked at ANL in an aqueous solution of sodium tetrathionate. The fifth tube (W 2-10)

134
Table 4.10. Test results from Stage 2 testing of flawed tubes containing OD laser-cut notches. All

notches had characteristics that existed after Stage 1 testing (from Table 4.9).

Pressurization Maximum Pressure,

Test Flaw Rate Burst M[Pa (psi)

Number Type M[Pa/s (psi/s) (Yes/No)

5528-1-1 1 12.6 (1823) Y 41.2 (5980)

5528-1-2 1 13.9 (2023) Y 40.3 (5850)

5528-1-3 1 12.9 (1865) Y 33.4 (4850)

5528-1-4 1 14.4 (2084) Y 32.1 (4660)

5528-2-1 1 No Stage2 -

5528-2-2 1 No Stage 2 -

5528-3-2a 3 13.3 (1928) Y 30.4 (4409)

5 52 8-3_ 3ab 3 10.0 (1450) Y 28.6 (4146)

5528-3-4a 4 12.2 (1776) Y 30.0 (4346)

5469-2_1a 4 13.4 (1950) Y 31.9 (4630)

5469-2-2abb 4 10.3 (1497) Y 33.4 (4850)

5469-2-3a 5 13.3 (1925) Y 32.2 (4669)

5469-2-4a 5 13.3 (1930) Y 26.8 (3893)

5469-3-ja 6 13.5 (1953) Y 32.2 (4675)

5531 3-ja 6 13.2 (1909) Y 29.0 (4200)

5469-3-3 7 Stagel opened -

5469-3-4c 7 14.2 (2066) N 53.1 (7700)

5469-4-1 8 Stage 1 opened - -

5469-4-2 8 Stagel opened - -

5469-4-3 9 Stagel opened - -

5469-4-4 10 Stagel opened - -

'Specimen tested in Stage 2 with internal bladder plus foil. All other Stage 2 tests were conducted with

an internal bladder only, except for Specimen 5489-3-4. which was tested in Stage 2 with neither a

bladder nor a foil.

bSpecimen exhibited unstable tearing in Stage 1 testing without internal bladder.

cSpeclmen did not open in Stage 1 testing, and therefore was tested in Stage 2 with neither a bladder nor a foil.

It did not fail at the maximum system pressure in Stage 2.

p (MPa)

0 4 8 12 16 20

0.s 5. . . I . ,

in.1.3 

0.4 - 19h.

8 0.2 _ 0

0

0.1

0

40.5 0 0.5 1 1.5

p (ksi)

12

10

Figure 4.57.

B Calculated variations of crack opening

i displacements with pressure. Symbols

6 - denote measured COD values after the

0

8 tests.

4

2

0

2 2.5 3

135
Table 4.11. Flaw opening widths and internal pressures for PSI tubes.

Notch Crack Opening Length, Displacement, Internal Pressurea,

Test mm mm (in.) MPa (ksi)

No. (In.) Target Actual Calculated Actual Remarks

1 20 - - 16.8b 17.6 Intentionally pressurized

(0.8) (2.43) (2.55) to unstable burst

2 20 10.0 8.3 16.9 17.6 Slight tearing

(0.8) (0.4) (0.326) (2.45) (2.55)

3 20 10.0 5.3 16.9 16.0

(0.8) (0.4) (0.210) (2.45) (2.32)

4 20 10.0 10.4 16.9 15.0 Slight tearing

(0.8) (0.4) (0.411) (2.45) 2.18

5 33 5.6 - 7.9 11 .6c Inadvertently tested to

(1.3) (0.22) (1.15) (1.68) large unstable tearing

6 33 5.6 7.0 7.9 8.1

(1.3) (0.22) (0.277) (1.15) (1.17)

7 33 5.6 5.1 7.9 6.9

(1.3) (0.22) (0.201) (1.15) (1.00)

8 48 4.4 5.5 3.1 4.5

(1.9) (0.17) (0.218) (0.45) (0.65)

9 48 4.4 4.0 3.1 4.0

(1.9) (0.17) (0.158) (0.45) (0.58)

10 48 4.4 - 3.1 - Not tested

(1.9) (0.17) (0.45)

Calculated Internal pressure to produce target COD and actual internal pressure applied to produce

observed COD.

bCalculated internal pressure required to produce unstable burst.

cActual internal pressure applied to produce inadvertent unstable burst.

contained an axial ODSCC flaw produced by Westinghouse using a doped steam procedure.

Tubes SGL-177 and SGL-195 were tested at room temperature, and tubes SGL-104 and SGL219 were tested at 2820C (540'F). The Westinghouse tube W 2-10 was tested at both

temperatures.

Dye-penetrant test results on the flaws are shown in FIgs. 4.58-4.61. Eddy current

examinations indicated that all four ANL-produced SCC flaws had variable crack depths along

their lengths, with maximum depths of 75-95%. The Westinghouse tube exhibited a similar

crack depth profile. All the flaws have small TW regions, which were identified by bubble

testing. They were not seen by EC techniques because of their tightness and very small axial

extents, and the dye-penetrant inspections revealed flaw branching for all of the flaws. The EC

BC voltages indicated that the five flaws fell into two categories. The EC BC voltage for the

Westinghouse tube was 4.5 V. very close to the 4.0 V exhibited by the ANL tube in Test SGL177. The other three ANL tubes had higher EC voltages, in the range of 6-10 V.

The test results on these five tubes are summarized in Table 4.12. Although all the

specimens showed gas leakage under very low pressures, measurable leakage of water was not

observed until the pressures reached 13.3 MPa (1928 psi) or greater. This is reported in

136
secdary lk mn l

sacaldary leak

I

I

mawn Mak

m"conday ak

4ji8 1216

j1b-24

Figure 4.58.

Pre-test Image of Specimen SGL-177

with flaw highlighted by dye penetrant

and digital image processing. Two

regions of TW penetration are

indicated by bubble testing.

Figure 4.59.

Pre-test image of Specimen SGL-195

with flaw highlighted by dye penetrant

and digital image processing. Two

regions of TW penetration are

indicated by bubble testing.

Figire 4.60.

Pre-test image of Specimen SGL-21 9

with flaw highlighted by dye penetrant

and digital Image processing. A single

region of throughwall penetration was

Indicated by bubble testing.

j - , . t.

- ;..

137
Figure 4.61.

Pre-test dye-penetrant digital image of

Westinghouse ODSCC tube produced

using doped steam (Test T505CATW2-

10). 010 1wo**414

: it

;->..£

* ,W,,-

Table 4.12 Summary of results

ODSCC cracks.

from tests on four Argonne tubes and Westinghouse tube with axial

Flow Rate at

Overall Pretest Initial Leak Constant Initial/Final

Flaw Test Air Leak Pressure/Flnal Tearing Opening

Spec. Length, Temp, at 0.28 Test Pressure, pressure, Pressure.

Number mm 0C MPa MPa MPa L/mln.

SGL- 12.2 R.T. yes 16.9/19.3 16.9 0.04-0.26/

177 1.67

SGL- 13.7 R.T. yes 14.7/15.5 14.7 <0.04/32.6

195

SGL- 10.4 282 yes 16.2/17.2 no 21.2/23.5

104 tearing

SGL- 14.0 282 yes 13.3/16.2 16.2 3.7/

219 14.0-39.0

W 2-10 12.4 R.T. yes 17.2/18.6 17.2 0.04-

0.068/0.12

W 2-10 12.4 282 yes 18.6 18.6 0.30-0.72

Table 4.12 as the Initial leak pressure." Another phenomenon that was observed in most of

the tests was an Increase in the leakage while the flaw was held at a constant pressure. Tils

occurred even in room temperature tests and shows that, with irregular crack geometries.

time-dependent failure of some portion of the crack ligament is possible. The pressures at

which time dependent increases in the leakage rates were observed are reported in Table 4.12

as the 'constant tearing pressure." Post-test photographs of Specimens SGL-195 and -177 are

shown in FIgs. 4.62 and 4.63.

Additional tests were performed on 14 tubes that contained a wider range of flaw

geometries, including OD axial and circumferential cracks and combinations of these types, as

well as OD flaws in the roll transition.

Table 4.13 summarizes the results from the 14 tests, all on tubes with OD flaws.

Specimens SGL-226 and SGL-363 contained complex flaws with combinations of axial and

138
gaining

P1_PX'rTW-O_. P -1-1.11 I  - -- I

Figure 4.62.

Photograph of failed region of

Specimen SGL- 95 after testing.

28 i 4 

Figure 4.63.

Photograph of failed region of

Specimen SGL-1 77 after testing.

circumferential cracks, including cracking in the roll transition for Specimen SGL-363.

Specimens SGL-413, -425, -480, -493, -494, -719, -728, -730, -731, and -734 contained axial

flaws, and Specimens SGL-099 -and 158 contained circumferential cracks, based on dyepenetrant determinations. The tabulated maximum depths and the EC voltages represent

pretest characterizations using standard EC signal analysis procedures and software.

The Information on pressure at first leak comes from visual observation of the flawed tube

through observation windows for the first signs of water leakage. For Specimens SGL-158,

-413. -425, -480, 493, -728. -731, -226, and -363, first leak occurred In the form of one drop

of water issuing from the flaw over a period of several minutes. The frequency of drop

formation increased with increasing pressure, and the leaks often passed through being a

small droplet spray, followed by a continuous leak stream. The tests were terminated when the

flaw opened enough that the flow reached the maximum 48.4 L/min (12.8 gpm) capacity of the

pressurizer pump. The tabulated maximum pressure for the tests represents the pressure at

which the flow capacity of the facility was reached.

For Specimens SGL-099. -494, -719, and -734, the pressure at first leak and maximum

pressure achieved were the same or differed only slightly. These tests were characterized by

139
Table 4.13. Results from pressure tests on 14 SCC flaws

Max. Pressure at First Max. Pressure

Test Flaw Flaw BC Leak Attained,

No. Type Depth. Voltage MPa (psi) MPa (psi)

099 OD/Circ 20 3.25 51.0 (7400) 51.0 (7400)

158 OD/Circ 60 4.32 24.8 (360 0 )a.b 31.4 (4560)

413 OD/Axial 90 1.57 12.4 (18 0 0)b 24.6 (3570)

425 OD/Axial 0 2.14 16.6 (24 00 )b 30.3 (4350)

480 OD/Axial 0 3.08 6.2 (9 0 0)b 15.2 (2200)

493 OD/Axial 0 3.84 0.55 (80 )b 17.6 (2550)

494 OD/Axial 0 0.54 31.7 (4600) 33.7 (4880)

719 OD/Axial 55 0.83 29.0 (4200) 29.7 (4300)

728 OD/Axial 90 1.97 14.5 (2 10 0)b 20.7 (3000)

730 OD/Axial 100 1.10 37.9 (5500) 41.7 (6050)

731 OD/Axial 90 3.12 0.28 (4 0)a~b 21.4 (3100)

734 OD/Axial 55 1.71 34.5 (5000) 36.5 (5300)

226 OD/Axial/Circ 90 - 12.1 (1 75 0)b 26.2 (3800)

363 OD/Axlal/Circ/ 90 - 18.6 (2 70 0)b 24.7 (3580)

Roll Trans.

aPre-test air bubble leak test at 0.28 MPa (40 psi) indicated throughwaU pinhole.

bFtrst leak occurred as one drop of water issuing from the flaw over a priod of several minutes.

essentially no leak prior to sudden rapid flaw opening, yielding a leak rate >48.4 L/min

(12.8 gpm). All of the flaws tested in this 14-tube series exhibited significant tearing, with

varying degrees of secondary cracking evident. Some tubes also exhibited evidence of unstable

tearing without the use of bladders.

Specimen SGL-226 contained an ODSCC flaw that exhibited both axial and

circumferential cracking, based on pretest dye-penetrant and NDE examinations. A posttest

photograph of the flaw opening (Fig. 4.64) shows a -6.35-rmm (0.25-in.)-long axial crack

intersected by a -25-mm (1-in.)-long circumferential crack. This flaw first began to leak at

12.1 MPa (1750 psi), with a rate of approximately one drop every few minutes. By the time the

pressure had reached 20.7 MPa (3000 psi), the leak rate had increased to 4.39 L/min

(1.16 gpm). The test was terminated at 26.2 MPa (3800 psi) when the flaw suddenly opened

and the flow rate exceeded the facility capacity.

Specimen SGL-363 contained an ODSCC flaw at a roll transition that, based on pretest

dye-penetrant and NDE examinations, was made up of a complex series of

axial/circumferential cracks. FIgure 4.65 shows the posttest photograph of the flaw opening, a

large hole with both circumferential and axial ligament tearing and considerable secondary

cracking. This flaw began to leak at 18.6 MPa (2700 psi), with a rate of approximately one drop

every few minutes. The flaw abruptly opened at 24.7 MPa (3580 psi), and the flow rate again

exceeded the facility capacity.

140
iSui in nniniiiiunutni II IIIIIHI 1111 11111 11111101 I Figure 4.64.

Post-test appearance of complex

ODSCC flaw in Specimen SGL-226

exhibiting both axial and

circumferential cracking.

Figure 4.65.

Post-test appearance of complex

ODSCC flaw In Specimen SGL-363

exhibiting a large opening resulting

from both axial and circumferential

cracking.

4.3.5 SCC from Operating Reactors

Pressure and leak-rate tests were performed on six McGuire tubes that were

decontaminated and Inspected using EC NDE. These tubes are designated 7-24-3. 5-51-2.

4-43-2. 14-55-3. 14-55-5, and 39-57-2. The NDE results identified short ODSCC flaws in the

TSP region. The tubes were tested In the High-Pressure Test Facility using room-temperature

water without bladders under a quasi-steady pressurization rate consisting of successive

0.7 MPa (100 psi pressure increases with Interspersed 5-s dwell periods up to the facility

maximum pressure of 51.7 MPa (7500 psi).

The results for these tests are presented In Table 4.14. Four of the six flaws did not burst

nor was any leak detected, and testing was terminated near the facility maximum pressure.

Note that there was no access to the tubes to observe if very small leakage (e.g., single drops at

the tube OD surface) was occurring. Flaws in two of the tubes, Tubes 4-43-2 and 39-57-2.

opened suddenly and yielded significant flow rates at the pressures Indicated In Table 4.14.

These failures occurred at pressures lower than estimated by structural analysis based on the

NDE estimates of flaw size.

141
Table 4.14. Results from pressure and leak-rate tests on tubes from McGuire SG. All flaws were

in the TSP region.

Leak Rate at Max. Predicted

Pressure at Pressure, Opening

Tube First Leak. L/min G MPa Pressure, Visual Flaw

ID MPa (psi) (gpm a psi) MPa (psi Characteristics

7-24-3 (no leak) 0.0 0 51.6 62.0 Bulge; hairline circ.

(0.0 0 7.490) (9,000) crack 6.35-mm

(0.25-in.) long;

secondary cracks.

5-51-2 (no leak) 0.00 51.0 66.2 Bulge; hairline axial

(0.0 0 7,400) (9,600) crack 6.35-mm

(0.25-in.) long;

secondary cracks.

4-43-2 49.0 (7,100) 32.2 0 49.0 54.5 Bulge. 2 axial cracks

(8.5 0 7,100) (7,900) with ligament

between; total length

10 mm (3/8 in.).

14-55-3 (no leak) 0.0 0 51.2 64.8 No bulge; hairline

(0.0 0 7,425) (9,400) axial crack 5-mm

3/16-in.) long, secondary cracks.

14-55-5 (no leak) 0.0 0 50.9 72.4 No bulge; axial crack

(0.0 0 7,380) (10,500) 6.35-mm (0.25-in.)

long.

39-57-2 36.2 (5,250) 28.0 0 36.2 61.4 Bulge; hairline axial

(7.4 0 5,250) (8,900) crack 10-mm (3/8-

in.) long; secondary

cracks.

A preliminary posttest visual examination of Tube 7-24-3 showed a puckered, or raised,

area on the outside of the tube in the TSP region. The puckered area was oriented

circumferentially and had associated small secondary cracks, also with circumferential

orlentatations. The overall extent of the puckered region was nominally 6.35 mm (0.25 in.).

Tube 5-51-2 exhibited a fine hairline axial crack, -6.35-mm (0.25-in.) long, in the TSP region,

with less puckering and some small secondary cracking. All of the other tube flaws were axial.

The flaw In Tube 4-43-2, which opened suddenly at a pressure of 49.0 MPa (7,100 psi) and

exhibited a leak rate of 32.2 L/min (8.5 gpiJ, had two axial flaws separated by a ligament, with

an overall length of -10 mm (3/8 in.). All of the flaws except that in Tube 39-57-2 did not leak

or open up until the internal pressure exceeded the 3AP safety margin. Also, each flawed

region consisted of a main flaw plus secondary flaws that also apparently contributed to the

overall behavior. This finding is similar to that observed for the short laboratory-grown SCC

flaws. These flaws have been destructively examined, and the results have been reported in

Section 4.1.3.

142
4.4 Jet Erosion Under Design Basis Conditions

Two basic erosion damage mechanisms, droplet Impact and cavitation, are believed to be

associated with fluid jets. In the case of jets comprised of water droplets, the impact of the

droplet gives rise to fluctuating stresses in the target, which produce fatigue damage. In the

case of cavitation, bubbles form in the liquid stream and then collapse on impact, again

producing fluctuating stresses and fatigue in the target.

Tests have been conduced to determine the susceptibility of steam generator tubes to

erosive damage from impacting jets of superheated water or steam leaking from adjacent tubes.

Most of the tests were performed using jets leaking from circular holes, though a few tests also

used leaking tubes with laboratory-produced SCC flaws. In all cases, the distance between the

leaking tube and the target tube was 6.35 mm (0.25 in.), and the tubes were oriented to

produce an Impact angle of 900 on the target tube. All tests were conducted In.the hightemperature Pressure and Leak-Rate Test Facility. The test duration was generally based on

the hole leak rate, which is a function of the hole area and the pressure. Test times are

constrained by the 760 L (200 gal) blowdown vessel capacity, though longer test times can be

achieved by shutting down the facility, refilling the blowdown vessel, and restarting.

The target specimens were fabricated from 22.2-mm (7/8-in.)-diameter Alloy 600 tubing

with a nominal wall thickness of 1.27 mm (0.050 in.). A cold-discharge machining method was

used to avoid any mechanical or thermal working of the material. The target specimens were

6.35-mm (0.25-in.) wide and 38.1-mm (1.5-in.) long, with four legs that attach to-the leaking

flawed tube to hold the target at a distance of 6.35 mm (0.25 in.) from the jet exit. Each leg is

fastened to the flawed tube by four spot-welded fastening straps, and the escaping jet impacts

the concave surface of the target specimen (FIg. 4.661.

Erosion tests were performed using jets from 1.59-, 0.79-, and 0.40-mm (I/16-. /32- and

/&4-in.)-diameter EDM circular holes. The jets were produced using a tube internal pressure

of 17.2 MPa (2500 psi). The external pressure for most of the tests was zero, approximating

MSLB conditions, but a few tests were performed with other external pressures. The test

duration was typically -2 h. A range of water temperatures was used to assess the influence of

degree of subcooling on erosion behavior.

The jet impact zone of a test with a 0.79-mm (1/32-in.)-diameter hole, a temperature of

2820 C (5400F), an Internal pressure of 16.7 MPa (2428 psi), and a back pressure of 3.7 MPa

(530 psig) is shown in Fig. 4.67. Erosion of the hole in the leaking tube was also verified.

High-magnification SEM examinations of the jet impact zone verified the presence of erosion

suggested in this optical photograph. The depth of material removed can be estimated using a

scanning phase-shift three-dimensional surface profiler microscope. This Instrument produces

an Image in which color or shading represents height or depth relative to a reference location.

Figure 4.68 shows the phase-shift Image of the jet impact zone In Fig. 4.67. The depths of

various features for each scan referenced relative to the uneroded surface of the specimen are

rendered in different shades in the photograph. The deepest erosion penetration is found to be

260 jum, or -21% of the tube wall thickness.

The results for the erosion tests on circular orifices are summarized In Tables 4.15-4.17.

The tests In Table 4.15 involved a significant amount of subcooling. An additional 16 erosion

143
Figure 4.66.

Setup for jet-impact erosion test

showing EDM hole in Alloy 600 tube

and Alloy 600 impact specimen

fastened by spot-welded straps to the

tube. There is a 6.35-mm (0.25-in.)

gap between the jet exit and the impact

specimen.

Figure 4.67.

Close-up view of 0.79-mm (1/32-in.)

hole In jet-impact specimen after

265 min of testing at 2820C (5400F)

and 17.2 MPa (2500 psig) with 3.7 MPa

(530 psig) back pressure.

tests were conducted with 0.79 and 0.40-mm (1/32- and I/6-in.)-diameter hole sizes

trables 4.16 and 4.17, respectively), to more thoroughly explore the parameters influencing jetimpingement erosion under prototypical conditions. The tests featured reduced levels of

subcooling using water temperatures of 135-3081C (275-587F).. Most of the tests were of 2 h

duration, but two of the 0.79-mm (1/32-in.)-diameter-hole tests lasted 4 and 6 h, respectively.

Figure 4.69 shows a plot of jet-induced erosion depth as a percent of specimen wall

thickness as a function of temperature for both hole sizes. For the sake of completeness, a

data point from testing a 1.59-mm (1/16-in.)-diameter hole with 2820C (5500 F) water, which

yielded 3.2% coupon penetration depth, is included. This plot indicates that impact erosion

depth varies with water temperature, and the erosion rate is substantially reduced with

decreasing and increasing subcooling on either side of the peak damage temperature range.

Examination of the jet impact zones shows that, as expected, for large amounts of subcooling,

the impact zone is about the same diameter as the jet. For very small amounts of subcooling,

the size of the impact zone is substantially larger than the initial jet diameter, indicating that

flashing has occurred. The zone of peak erosion corresponds to a situation in which the

144
Figure 4.68.

Depth of the erosion zone shown in

Fig. 4.67, as determined using

interferometric techniques.

Table 4.15. Erosion data from jet-impact erosion tests utilizing jets from various sizes of EDM

holes impacting Alloy 600 erosion coupons positioned 6.35-mm (0.25-in.) from the

hole exit at 2820C and room temperature.

Tube

Jet Hole Water Internal/External Test Erosion Depth,

diameter Temperature, Pressure, Duration, % Wall

nun (in.) 0C (OF) MPa (psig) min Thickness

1.59 (1/16) 282 (540) 16.7/0.04 34.1 (slight surface

(2428/6) burnishing)

0.79 (1/32) 21 (70) 17.2/0 240 (slight surface

(2500/0) burnishing)

0.79 (1/32) 282 (540) 16.7/3.7 265 21

(2428/530)

0.40 (1/64) 282 (540) 17.3/0 198 9

(2502/0)

impact zone is the same diameter as the jet. This suggests that the primary mechanism for

erosion Is cavitation due to vapor bubble formation at the onset of two-phase conditions.

Erosion rates are low in the purely single phase regime (large subcooling) and in the two-phase

region where flashing occurs either in the tube wall or in the region between the tube and the

target (small subcoolng).

145
Table 4.16. Erosion data from jet-impact erosion tests utilizing jets from 0.40-mm (1&4-in.)-diameter

EDM holes Impacting Alloy 600 erosion coupons positioned 6.35-mm (0.25-in.) from the

hole exit for various water temperatures. For all tests, the tube internal pressure was

16.8 MPa (2432 psig) with zero back pressure.

Coupon Water Temp., Test Duration, Erosion Depth,

No. 0

C (CF]) min % WaU Thickness

21 289 (553) 123 1.8

24 239 (463) 123 11.9

26 209 (408) 123 18.6

28 182 (360) 123 18.6

29 135 (275) 123 5.9

Table 4.17. Erosion data from jet-impact erosion tests utilizing Jets from 0.79-mm (16rin.)-diameter

circular EDM holes impacting Alloy 600 erosion coupons positioned 6.35-mm (0.25-in.) from

the hole exit for various water temperatures. For all tests, the tube Internal pressure was

16.A MPa (2432 nain) with zern back nrassura.

Coupon Water Temp., Test Duration, Erosion Depth,

No. 0C (OF7 min % Wall Thickness

8 278 (532) 123 28.0

9 307 (586) 123 4.6

10 281 (537) 369 28.1

11 308 (587) 246 7.9

12 292 (558) 124 14.0

13 271 (519) 123 27.8

14 287 (548) 127 13.0

15 259 (499) 123 14.0

17 288 (550) 123 15.2

18 233 (451) 124 17.0

19 180 (356) 123 15.6

The temperature at which the peak erosion rate occurs varies with hole size. The peak

erosion rate for the 0.40-mm (l/wi-In.)-diameter hole occurs in the range of 182-2100 C (360-

4100F1, while that for the 0.79-mm (l/32-in.)-diameter hole occurs In the range of 271-2880C

(520-550 0F). This behavior Is consistent with the primary damage mechanism being cavitation

and reflects the degree of flashing of water to steam in the hole, with the smaller-diameter hole

producing more flashing at a given temperature and consequently lower erosion. In addition,

the temperature at which flashing becomes significant Is lower for the smaller-diameter hole.

For the orifice flows considered here, the erosion rates peak for degrees of subcooling

typical of the cold leg, where, however, cracking is much less likely to occur. Erosion rates are

much lower for conditions more typical of the hot leg. The trends observed with decreasing

orifice size also suggest that the margin Is even larger for actual stress corrosion cracks, since

they typically would be even tighter than the 0.40-mm (l/64-in.)-diameter hole.

146
25

,,, 25 / A Figure 4.69.

Jet erosion penetration depth as a

2 - Jo - ~ so/ function of water temperature for 0.40-

1 * 0 (triangles), 0.79-(circles), and 1.59-mm

15 - / z o (cross) (164-, 1 &-, and 1 As-In.)-

10 X diameter us \ EDM holes. For all tests, the

c 10 - / X \ _ tube internal pressure was 16.8 MPa 0 -0

0 1/32 onfice 0 (2432 psig) with zero back pressure.

5 _ i-8-- - 1164 onfic \ Most tests were 123-mn duration.

x 1/16 orifice

250 300 350 400 450 500 550 600

Temperature (F)

Tests were planned to validate these results on more prototypical flaws, but to date only

one test has been performed with a laboratory-produced ODSCC axial flaw (SGL-818). The

flaws and the associated jets at 17.2 MPa (2500 psi) were first evaluated in the roomtemperature High-Pressure Test Facility, where visual observation of the flaw jet behavior and

very accurate determination of flow rate are possible.

Specimen SGL-818 was Identified by NDE to be axial OD and nominally 10-mm (0.4-In.)

long, with greater than 95% penetration. At 17.2 MPa (2500 psi) and room temperature, the

leaking jet was fan-shaped and somewhat rotated out of axial alignment, with a base length at

the tube of 3.2 mm (1/s in.); at 6.35 mm (1/4 in.) above the tube surface (where the coupon

would be located In the erosion test), its length had spread to 6.3 mm (1/4 in.). The jet was

nonuniform over its breadth. The measured leak rate for the flaw was 1.1 kg/min (2.5 lb/min)

or 1.1 L/mln (0.3 gpm).

An erosion test at 2880C (5500F) and an internal pressure of 17.2 MPa (2500 psi) was

then performed with this flaw. After 123 min of testing, no significant erosion of the target

specimen was detected, with only very slight burnishing observed. Because this flaw was

much tighter than the EDM holes, it would likely have higher frictional losses and, hence, a

greater tendency for flashing under the same pressure and temperature conditions. Based on

the orifice results, flashing results in a decreased erosion rate.

The SCC flaw in Specimen SGL 818 was opened in the High-Pressure Test Facility by

subjecting it to a pressure of 21.7 MPa (3150 psi) to achieve a higher flow rate. It was then

used in a jet-impact erosion test under the same conditions as previously. The erosion coupon

showed more burnishing than in the first test, but the erosion depth was still <1%. It appears

that the flaw remained tight enough that flashing is still occurring either in the tube wall or in

the gap between tube and coupon or both.

147
4.5 Tube Failure Models under Design-Basls and Severe Accident Conditions

4.5.1 Characterization of Alloy 600 Tubing Material

4.5.1.1 Room-Temperature Mechanical Properties

Room-temperature tensile tests were conducted by two outside vendors on two heats of

22.2-mm (0.875-in.)-diameter Alloy 600 tubing, one purchased from Valinco (heat no.

NX 8527) and a second heat obtained from PNNL. The results from these tests have been

reported in detail in Ref. 13. The PNNL tube (ID EX-82-1/2675) is both stronger and more

ductile than the Valinco tube (Heat No. NX 8527).

Alloy 600 tubes with diameters of 19.1-mm (0.750-in.)-diameter (Valinco heat no.

NX 8524), 22.2-mm (0.875-In.)-diameter (Valinco heat no. NX 8527) and 22.2-mm (0.875-in.)-

diameter (PNNL ID EX-82-1/2675) were sent to an outside vendor for pressure testing. Three

unflawed tubes for each diameter were pressure tested with room-temperature water to failure.

Results from pressure tests conducted at ANL are compared with those conducted at PNNL 1261

in Table 4.18, which also includes the average tensile properties for the various heats. The

failure pressure for the 19.1-mm (0.750-in.)-diameter Valinco tube is comparable to that of the

19.1-mm (0.750-in.) diameter PNNL tube, although the PNNL tube has 16% higher yield

strength. Both the failure pressure and tensile properties of the 22.2-mm (0.875-in.)-diameter

PNNL heat EX-82-1/2675 are significantly higher than those of the other 22.2-mm (0.875-in.)-

diameter PNNL heats. Tensile properties of the 22.2-mm (0.875-in.)-diameter Valinco tube are

slightly lower than the PNNL heats, even though its failure pressure is about 5% higher than

those reported by PNNL.

Elevated-Temperature Mechanical Properties

Elevated-temperature tensile and creep rupture data were collected from the literature.

The flow stress (average of yield and ultimate tensile strengths in MPa as a function of

temperature in 'C) for the INEEL rod samples is given by 1271

[564.4 -0.4546T + 1.5055x10-3T2 -1.9907xIo 4 T3 for 200C • T s 727C. (4.2)

l4308.9-11.381T+1.03)CIO 2T2 -3.1734xlO6T3 for 7270C <T 11000C

A plot of the flow stress curve for various product forms is given in Fig. 4.70. Note that at high

temperatures (>6000C), the flow stresses of various product forms tend to converge.

The Larson-Miller equation for creep rupture time is [131

log1oa = 4.31±0.10 - 1.43 x 1 0 4 Pim for a S39.3 MPa (4.3a)

logoa = 5.030.13 - 1.81 x 1 0 4 Pim for a >39.3 MPa (4.3 b)

where the time to rupture tR is then given by

148
Table 4.18. Room-temperature failure pressures and axial tensile properties of unflawed Alloy 600 tubing

tested by ANL [13] and PNNL [26].

Heat

No.

Valinco

(NX 8524)

Heat C

PNNL

(EX-82-1/2675)

Heat B

Heat E

Heat F

Valinco

(NX 8527)

Tested

by

ANL

PNNL

ANL

PNNL

PNNL

PNNL

ANL

Dian.,

mm (in.)

19.1

(0.750)

19.1

(0.750)

22.2

(0.875)

22.2

(0.875)

22.2

(0.875)

22.2

(0.875)

22.2

(0.875)

Wall

Thickness,

mm (in.)

1.09

(0.043)

1.09

(0.043)

1.27

(0.050)

1.27

(0.050)

1.27

(0.050)

1.27

(0.050)

1.27

(0.050)

Avg.

Fail. Press.,

MPa (ksi)

68.5 (9.94)

67.1 (9.73)

72.4 (10.50)

65.1 (9.44)

65.8 (9.54)

64.1 (9.30)

69.4 (10.07)

-

Avg. Yield

Strength. MNPa

(ksi)

308 (44.7)

361 (52.4)

357 (51.8)

328 (47.6)

300 (43.5)

274 (39.7)

271 (39.3)

-

Avg. Utimate

Tensile Strength,

MPa (ksll

681 (98.8)

674 (97.8)

683 (99.1)

641 (93.0)

681 (98.8)

665 (96.4)

634 (92.0)

600

500

le

a.

400

300

200

100

0

D. PNNL (19 mm) 80 a PNNL (22 m) 7 . ANL(19 lun) 70

0 o L (22 mm) 60

50

.. ............... _ --------- 4 0 . \\ ~~40

.30

20

10

0

600 800 1000 1200

T ("C)

Figure 4.70.

1 Flow stress curves for various

i product forms of Alloy 600. _

0 200 400

"m - 15

tR =1 T (4.4)

with tR In h and T In K. The Larson-Miller plot is shown In Fig. 4.71.

4.52 Low-Temperature Failure Models

The failure pressure of unflawed tubes can be described In terms of a simple plastic

collapse model:

Pb = ha

Rm (4.5a)

149
W | | >tK ~~~~~~~~~~Figure 4.71. iF y - -1.S1OE-4x + 5.0298Fgr 4.1

1 t.50_ . _ 1__ Bilinear fit to existing creep data for

5 1.50 -- Fe - 0.9M Alloy 600. The three low outliers were

1.00 not included In the statistical analysis.

0.50_y _ _i 2E-4 2 x 4.3108

12000 14000 16000 18000 20000 22000 24000 26000 28000

L*- Pmramd w (C.15.6 Temperature In K)

where

Ca = Flow stress = k(Su + SY), (4.5b)

Pb = failure pressure of unflawed tube,

Rm = mean radius of tube,

h = wall thickness of tube,

k = flow stress factor,

Su= ultimate tensile strength, and

y= yield strength.

Substituting the tensile and failure pressure data from Table 4.18 Into Eqs. 45a and .b yields

the following the values of k for the three heats:

Vallnco (19.1-mm diameter, heat no. NX 8524): k = 0.57,

Vallnco (22.2-mm diameter, heat no. NX 8527): k = 0.63, (4.6)

PNNL (22.2-mm diameter, ID EX-82-1/2675): k = 0.57.

These are reasonably consistent with the k values for Alloy 600 reported in the literature,

which vary between 0.5 and 0.6. The value of k for the 22.2-mm (0.875-in.)-diameter Valinco

heat appears to be slightly on the high side.

4.5.2.1 Throughwall Axial Cracks

- The most widely used model for predicting the pressure to cause unstable ductile (plastic

collapse) failure of a tube with a throughwall axial crack is that originally proposed by

Hahn et al. 1281, later modified by Erdogan 1291:

150
oh = p Pcr MR, m

(4.7a)

where

m = 0.614 + 0.481A + 0.386exp(-1.25X). (4.7b)

= [12( - v2)]4 = ;r. , (4.7c)

and Pb is the failure pressure of an unflawed tube (Eq. 4.5a) and 2c is the axial crack length.

4.5.2.2 Part-throughwall Axial Cracks

For part-throughwall axial cracks, the pressure required to fail the remaining throughthickness crack tip radial ligament can be calculated by an empirical equation (referred to as

the BCL equation) developed by Riefner et al. 1301:

Ps= ___ (4.8a) mpRm mp1

where

mp = I, 1~~~~~~~~~~~~~~~~~~

MP a

-l-h

and a is the crack depth. It should be emphasized that Eq. 4.8a gives only the pressure

required to fail the remaining radial ligament. The stability of the resulting throughwall crack

can be analyzed using Eq. 4.7a. If Pcr > Psc, the throughwall crack is stable. Although the

crack will leak, it will not increase in length without a further increase in pressure. If Pcr <

Psc, the resulting crack will be unstable and will rapidly increase in length without any

additional increase in pressure.

Comparison with experiments shows that Eqs. 4.8a-b tend to underestimate the ligament

rupture pressures of short and deep cracks 1251. Under the auspices of an NRC-sponsored

steam generator integrity program, PNNL 1261 conducted a series of pressure tests on Alloy 600

steam generator tubes containing part-through axial slots. The PNNL tube test data were

re-analyzed (including re-measurements of the flaws by posttest fractography) by ANL, and the

following equation was proposed for predicting ligament rupture pressure (referred to as the

ANL equation) 1251:

laa

mp = a, (4.8c)

-h a

151
where

a =I+ ( ) -1 and (4.8d)

is a constant - 1.

Except for short and deep cracks, Eq. 4.8c predicts similar failure pressures as Eq. 4.8b. The

ANL equation has been validated for short and deep slots with a series of tests on Alloy 600

tubes with EDM slots (Ig. 4.3) and laser-cut slots (Fig. 4.4).

4.5.2.3 Throughwall Circumferential Cracks

Failure loads of tubes with a single circumferential crack depend critically on the bending

constraint imposed externally on the tubes. The two extreme cases are the free-bending case

and the fully constrained case. In most cases, the tube support plates (TSPs) offer substantial

restraint against bending, and the failure loads are generally closer to the fully constrained

case.

Ranganath and Mehta 1311 calculated failure loads for throughwall circumferential cracks

by plastic limit load (collapse) analyses, which were based on earlier work by Kanninen et al.

1321. For an unconstrained (free-to-bend) tube with a throughwall crack of angular length 20

and no applied primary bending stress, the critical failure pressure is

Per = 2-Rh (1 _ _ 2 t (4.9al

where the angular location of the neutral axis is given by

(i- 2 n) (4.9b)

Equation 4.9a Is applicable to the case where the tube is completely free to bend. In the

opposite case of total constraint against bending, a criterion based on maximum shear stress

in the net section as proposed by Cochet et al. [331 can be used to calculate the instability limit

pressure:

Pcr =2(y2 -1)(ly2-Ol (4.10a)

where

(4.10b)

Rg

The following thin-shell, untaxial approximation to Eq. 4. 10a Is often used for predicting failure

of steam generator tubes that are fully constrained against bending:

152
Test Ligament Rupture Pressure (MPa)

g~7

14

is

1

i

1 5Z u 0 a3 au 'IV 430 D 00

................... .... 550.

o 0.25 inJ60Y* 50 o 0.5 iJ60°

* 0.25 hJE0°/6

A 0.5 .1J80% 45

v 0.75 l 4 .0E

8235 nJ90% 0 401

a 30

* 3=1 ~~~~~20~

Y ~~~~~~EDM notche 5

! 3 4 5 6 7 8

Test Ugament Rupture Pressure (ksi)

Figure 4.72.

Observed vs. predicted ligament rupture

pressures for tubes with EDM notches

using k = 0.55 and 0 = 1.

2

Test Failure Pressure (MPa)

30 35 40 45

7

X 6.5

IDI

1 6

L. 5.5

2

if 5

4.5

a,, s . II . . .,I .g r . . I

Lser cut nolches/Fractogra

45 0.

40

35'i

30

Figure 4.73.

Observed vs. predicted ligament rupture

pressures for tubes with =80% deep lasercut notches using k = 0.55 and 0 = 1.

4

4 4.5 5 5.5 6 6.5 7

Test Faflure Pressure Ms

Pcr = R M (1- 0). n (4.10c)

As noted previously, the tube support plates offer a significant but not total restraint

against bending. This condition tends to increase the failure pressure to somewhere between

those predicted by Eqs. 4.9a and 4.10a (or 4.10c). A detailed analysis of the failure of a

partially constrained tube with a throughwal circumferential flaw was presented In Ref. 34. It

was shown that for a tubesheet-to-TSP span of 0.67 m (26 in.), which is typical of many

generators, the critical pressure is much closer to the fully constrained case than the freebending case.

4.5.2A Part Throughwall Circumferential Cracks

The pressure for through-thickness crack tip radial ligament rupture (Psc) is generally

expressed in terms of a stress magnification factor (mp) and is obtained by equating the

magnified axial stress in the ligament to the flow stress,

153
2ah

Psc = (4.11) RmmP

As in the case of throughwall cracks, the ligament rupture pressure for circumferentially

cracked tubes, i.e.. the value of the magnification factor mp, depends strongly on the degree of

restraint against bending. The free bending and the completely constrained case are relatively

easy to analyze (see below). The failure loads are expected to be much closer to the completely

constrained case than the free bending case. The discussion here assumes that the tubes are

either completely constrained or are completely free to bend.

Fully Constrained Case

This would also include the case for an unrestrained tube containing two symmetrical

cracks. The whole section containing the crack (or cracks) Is subjected to uniform axial tensile

stresses with the ligament (or ligaments) being subjected to stress intensification. If the stress

in the ligament (or ligaments) Is expressed as 1/m times the stress in the rest of the section

containing the crack (or cracks), the ligament stress (clig) can be calculated from a simple

equilibrium of axial forces,

llJ = fm( aI m)] (4.12a)

lg= 2h n.1_a 1

where

n1 for a single crack

I2 for two symmetrical cracks

If we define mp as the ratio between the ligament stress and the axial stress in the

unflawed tube, mp is given by

P r r n@Y a A (4.12b)

[m +tAl h -mJ]'

Originally, Kurihara et al. 1361 used an empirically obtained expression for m (with n = 1),

(a)"(n8) (4.13)

Since Eq. 4.13 does not have the correct limiting behavior (i.e., m does not tend to 0 when a/h

tends to 1 for all 0), it was modified to have the same form as in the axial crack case, i.e. [41,

a

M 1-- h (4.14a) a

154
where

N=+ ) (4.14b)

and A and y are fitting parameters. The test results on Type 304 stainless steel reported by

Kurihara et al. 1351 can be predicted fairly well by assuming X = y = 0.2 141.

The original set of test data on four-point bending failure tests on pressurized partthrough circumferentlally cracked Type 304 stainless steel pipes at room temperature used by

Kurihara et al. 1351 are predicted (both failure modes and moments) somewhat better by the

current model than by the Kurihara model 141.

Free-Bending Case

In this case a part of the section containing the crack will, in general, be subjected to

compressive stress. As a result, Eq. 4.12a has to be replaced by

= pR 1 (415a)

2h~~~~

where the angle P defining the location of the neutral axis Is given by

= sin[1f- [ I ( for - e, (4.15b)

and Eq. 4.12b has to be replaced by

mp t 3 ( s X h )] 0 (4.15cJ

with m defined by Eq. 4.14a and n is the number of cracks.

4.52.5 Equivalent Rectangular Crack

So far the flaw has been considered a single crack that is rectangular in shape. In reality,

stress corrosion cracks are multiple, nonplanar flaws that are complex in shape and have

ligaments between the segments. To predict the through-thickness crack tip radial ligament

rupture pressure (i.e., pressure beyond which leakage will occur) for complex-shaped multiple

cracks, we have used the concept of an 'equivalent rectangular crack." The procedure starts

with a crack depth profile that has been determined either by eddy current or by fractography

technique (e.g., Fig. 4.74, dashed line).

To determine the equivalent rectangular crack, a series of candidate equivalent

rectangular cracks is considered with different lengths (e.g., Fig. 4.74, solid line). The

155
equivalent depth for the candidate equivalent rectangular crack is determined by computing

the average depth contained within the crack length being considered and ignoring the depth

profile outside (i.e., equating the hatched areas in Fig. 4.74). Out of all the candidate

equivalent rectangular cracks, the one with the lowest ligament rupture pressure (i.e., highest

mp) is defined as the equivalent rectangular crack. Upon ligament rupture, leakage occurs

through the 100% throughwall portion of the crack whose length is assumed to be equal to the

length of the equivalent rectangular crack. The procedure has been automated in an Excel

spreadsheet.

4.5.2.6 Multiple Notches

Tests were conducted for tubes with multiple notches in various configurations. Details

are given in Ref 36.

Two Part-Throughwall Axial EDM Notches

The through-thickness radial ligament rupture pressures for tubes with two partthroughwall axial notches, type 2 (axial ligament) and type 4 (circumferential ligament) as

depicted in Fig. 4.75, were calculated by the equivalent rectangular approach for various notch

lengths and inter-notch ligament widths. Tests were conducted in the High-Pressure Test

Facility on specimens without bladders and were interrupted after the through-thickness

ligament ruptured, as evidenced by onset of leakage (Stage 1). Comparison of predicted vs.

experimental ligament rupture pressures, given in Figs. 4.76a and 76b, shows that the

equivalent rectangular crack approach can predict the through-thickness ligament rupture

pressures reasonably well. Except for the cases Identified in Figs. 4.76a and 76b, the throughthickness radial ligament and the inter-notch ligament ruptured simultaneously. The internotch ligament width beyond which the two notches do not interact (from the viewpoint of

pressure for through-thickness radial ligament rupture) depends on the notch depth as well as

notch length. A 3-mm (0.l-in.)-wide ligament increases the through-thickness ligament

rupture pressure by =30%, with the axial ligament having a slightly higher strengthening effect

than the circumferential ligament. For smaller ligament widths, both types of ligaments (2 and

4) have comparable strengthening effect.

Two 100%-throughwall axial EDM notches

The pressures for inter-notch ligament rupture for two 100%-TW notches were calculated

by nonlinear, finite deformation FEA. The calculated (by ANSYS and ABAQUS) variation of

average ligament thickness vs. pressure for two 6-mm (0.25-in.)-long. 100% TW notches

separated by a 2.5-mm (0.1-in.)-wide circumferential ligament is shown in Fig. 4.76a.

Figure 4.77b shows the variation of calculated (axial ligament) and experimental ligament

rupture pressure with ligament width for two 6-mm (0.25-in.) and two 13-mm (0.5-in.)-long,

100%-'fIW notches. Also shown in the figure by dashed lines are unstable burst pressures for

25-mm (1-in.) and 13-mm (0.5-in.)-long, 100%-TW notches. It is evident that the unstable

burst pressures of a 13-mm (0.5-in.) and 25-mm (1-in.)-long, 100% TW notches are not

increased by introducing a full thickness ligament with width less than about 1.3 mm

(0.05 In.).

156
100 . . . . . . . . . . . . . .

ECINDE or

J/1h. '- fmrAnMrnhv - 80 / [ ti'prfie

60

t; 40 : , NCandidate'l Figure 4.74.

a, . , equivalent Determination of equivalent rectangular

~~2 0 I ~~~rectangular c a k 20 XaIfk crack.

-20 , a

-10 -5 0 5 10

Relative Axial Position (mm)

Mutiple axial laser-cut notches

A series of tests was conducted on Alloy 600 tubes with 2 to 6 part-throughwall axial

laser-cut notches in various configurations (Fig. 4.75). The ligament widths varied between

0.13 and 0.25 mm (0.005 and 0.010 in.). Predicted vs. observed pressures for throughthickness ligament rupture are shown in Fig. 4.78. Although the rupture pressures are

reasonably predicted by the equivalent rectangular crack method for the case of two notches,

they are overestimated by -30% for the case of six notches.

4.5.3 Leak Rate Models for Single-Phase Flow

The leak rate model based on simple orifice flow through a crack with an opening area A

and a coefficient of discharge Cd = 0.6 is as follows:

Q = 0.6A (4.16a)

where A is crack opening area, Ap is pressure differential, and p is mass density of water. This

equation has been used successfully to predict leak rates in experiments conducted at room

temperature and 2820C on specimens with EDM notches and stress corrosion cracks at ANL

down to a leak rate of =0.44 L/min (=0.1 gal/min). The correlation will break down at lower

leak rates for very tight stress corrosion cracks, particularly at elevated temperatures where

flashing may occur inside the cracks, but the conditions for such a limit have not yet been

established. Expressing Q in L/mln, A in mm2, Ap in MPa, and p in kg/mi,

Q = 50.9AXg (4.16b)

or. using English units, i.e., Q in gpm, A in in2, Ap in psi, and p in lb/ft3 ,

157
Axial or Circ. Ugament Wdth (mm)

-2 0 2 4 8 8 In 12 14

-6.5 -

.0

1.5

a:

4.5-

.41_

| 3.5L

-0.1

. _ _ _ _ . _ . _ , if

S

40I

C.

353

30j

qo

-5

= 4.5

. 4

a

9 3.5

3

12.5

2

Axdal or Clrc. Ugament Wdth (mm)

-2 0 2 4 6 8 10

O Axial ligament|

1 VCIm-flgament l;xO.5'x8O%-

*V 2x0.5"x80% -

* arc. 11g. not ruptured -

Ixi Ox80%

.... ,....I... I..*.1...

0.

305

25 .

201

.4

i

, I.... . . I . . I., l , . ,,, ,,, 25,, ,ic

0 0.1 0.2 0.3 0.4 0.5 0.6

Axdai or Circ. igament Width (in.)

(a)

j I

0.1 0 0.1 0.2 0.3

Axiai or rc. Uigament Width (m.)

(b)

Oi

Figure 4.75. Predicted and observed rupture pressures for through-thickness ligaments as a function of

axial or circumferential ligament widths for two part-throughwall EDM notches 6-mm

(0.25-In.) and 13-mm (0.5-in.) long.

Pressure (MPa)

-1. 7 8910 20 30 40 5060

S_ 0.0g . . ,

4 t 2x0.2

0.05 t

c 0.04 :

I 0.03

C0.02

< 0.01

1.4

1.2

0.8 S

0.6 3

D

0.4 9

Axdal or Cic. Ligament Widfth (mm)

0 0.5 1 1.5 2 2.5 3 3.5 I

I.

is

I

6

5

4

3

2

1

- 40

35

It -30

f 25

20

15

I' 10

5

0

0.15

0.

I.

is

I 0 X

1 Pressure (ksi) 10 0.05 0.1

Axdal or Ch:. Ligament Width (In.)

Figure 4.76.

(a) (b)

Variation of average ligament thickness with pressure for two 6-mm (0.25-in.)-long, 100%

throughwall notches separated by 1.3-mm (0.05-in.)-wide ligament and variation of

ligament rupture pressure with ligament width for two 13-mm (0.5-)-long and two 6-mm

(0.25)-long, 100% throughwall notches.

158
L- - I

Notch length I1

b-1 80% TW notch

Ugament width/

01oz A o~z

Ligament width

-41--

C= = == = = C=

1- -I

Notch length

Type 2 specimen

(a)

= =*= '- 4 Ugament oh lnthwidth

,,,~ Notch length |. S

Type 3 specimen

(b)

=--------- 'Ligament width

r . I~~~~~~~~~~~~~~~~~~~~~~~~~~~

. ---

is Notch lengt~h

Axia

I

Type 4 specimen

(c)

= 1. Ligament width

.- - - --- - I"

Notch length

I'4 -Pi

Axs

Type 6 specimen

Type 5 specimen

(d)

(e)

Figure 4.77. Geometries for multiple axial notch specimens.

159
Test Ugament Rupture Pressure (MPa)

30 35 40 45

I

Z 5

i 4

;

la

: 45 =

I

_ 40 L

:ci.

_ 35 c

I _ 30 7

.2

7 1.

Figure 4.78.

Observed vs. predicted ligament

rupture pressures for multiple lasercut notches.

j

4 5 6

Test Ugament Rupture Pressure (ksi)

Q = 180.2A, AP. (4.16c)

The density of water is 997.5 kg/m 3 (62.27 lb/ft3 ) at room temperature and 735.3 kg/M3

(45.9 lb/ft3 ) at 2820C.

4.5.3.1 Crack Opening Area

Axial cracks

Zahoor's model 1371 for crack opening area A of an axial throughwall crack in a thinwalled tube is

A= =2zcV 0Gr/ E.

where

a = hoop stress,

E is the Youngs modulus,

(4.17a)

V. = l+ 0.64935.4 - 8.9683x 1034e + 1.3 3 8 7 3X 10A 4e, (4.17b)

R and h are the mean radius and thickness of tube, (4. 17c)

160
CC = + a ] (4.17d)

F = 1+ 1.2987X2 - 2.6905x10-2 4 + 5.3549x 1OA 6, (4.17e)

X2 =c 2 /Rh, (4.170

and ac is the yield strength, and c is the crack half-length.

The flaw opening area for an EDM notch was obtained by adding the initial flaw area

(equal to flaw length multiplied by flaw width) to the crack opening area given by Eq. 4.17a.

Crack opening areas calculated by Eq. 4.17a were found to agree well with FEA results [4]. The

predicted leak rates are also in good agreement with experimentally measured leak rates at

room temperature for the as-received tube (Fig. 4.79a) and heat-treated (flow stress reduced by

20%) tube (Fig. 4.79b). Continuously measured leak rates in 100% TW and 90% NW EDM

notches are compared with predicted leak rates in Figs. 4.80a and 80b, respectively, at room

temperature. Note that both the ligament rupture pressure and subsequent leak rates are

predicted well for the 90% part-throughwall notch (Fig. 4.80b).

The results for-a leak rate test at 2820C on an as-received tube with a 12.5-mm (0.5-in.)-

long throughwall EDM notch are given in Fig. 4.81. Again the calculated leak rate Is

reasonably close to the measured value. Thus, the standard smooth-edge-orifice leak rate

equation (Eqs. 4.16a-c) appears to predict the leak rates through EDM notches at room

temperature and 2820C very well. The difference in leak rate between room temperature and

2820C can be accounted for simply by the difference in the density of water.

Circumferential cracks

The opening areas of circumferential cracks can be calculated from the Paris/Tada model

138]:

A p=R2 B. E

(4.18a)

where

B = lC2 + 0.16)e 4 for 0 <;Le < 1 (4.18b)

09.02 +O.lAee2 + 0.30Ae3 + 0.03Xe4 for 11 )e S 5

e= Oe( (4.18c)

161
1400

1200

1000

g 800

ED 600

2 400

1 200

0

-200

Figure 4

. . . I . r T . I I .

As-received tubs

7 25.4 mm EDM notch (200C) 1

o Test x Calulated with

measured COA

Predicted

: , , , . , . , . . . . . . I . . . .. . .

120

100

_~80

860

3 40

I 20

0

-20 _ .

5

1.79.

0 5 10 15 20 -5 0 5 10 15 20

Ap (MPa) Al (MPa)

(a) (b)

Calculated (solid line) and experimentally measured (symbols) leak rates at 20°C (680F) for

as-received and heat-treated 22-mm (0.875-in.)-diameter tubes with (a) 25.4 mm (1 in.) and

(b) 12.7-mm (0.5-in.) throughwall axial EDM notches. Cross symbols (x) In Fig. 4.79a

denote calculated leak rates using posttest measured crack opening areas.

Ap (MPa)

0 2 4 6 8 10 12

tp (MPa)

-2 0 2 4 6 8 10

20

i15

.3J 10

~1 5

70

60

50

40

30

20

10

0

14

12 :

- '10 -

1 6

s1 2

1 0

-2

-0.5

50

40

30 ,

20 1

10 A

O X1

0 0.5 1

AP (ksl

1.5 2 0 0.5

Ap (ks)

1 1.5

(a) (b)

Figure 4.80. Calculated and measured pressure vs. leak rate plots for tests (a) OM169 (100% TW) and

(b) OM171 (90%h TW), both conducted on as-received tubes at 20°C.

162
60

50

- 40

D 30

cI20

6 10

0

. ~ ~ ..... ~ srcie tube ..... . . . .

A

12.711

Pre(

, ., 1 1 . .. I - r| Xs-received tube/i

nm EDM notch (2820C)

licted

| x Test]

I . . .. , I . ., , ,. . , , ,

Figure 4.81.

Calculated (solid line) vs.

experimentally measured (symbol) leak

rates In as-received 22-mm (0.875-in.)-

diameter tubes with a 12.7 mm (0.5 In.)

flaw at 2820C (5400F).

l -10 r

-g

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

5 0 5 11

Ap (MPa)

0 15 20

e= + Fm2p2R2 / (8h2ay2)

F +0.150 .1.5 forA52

m 0.8875 + 0.2625X for 2 •5

X R) 0.5

(4.18d)

(4. 18e)

(4.18X

The Paris/Tada model 1381 is restricted to small-scale yielding (X • 5) and does not

account for the bending effects that become important at pressures that induce large-scale

plasticity In the crack plane. As a result, the COA calculated by this model is independent of

the span of the tube or support conditions.

4.5A Pressure and Leak Rate Tests on Specimens with Stress Corrosion Cracks

Pressure and leak-rate tests have been conducted on specimens with laboratorygenerated SCCs and on tubes pulled from a retired McGuire SG. Crack depth profiles for all

the specimens were estimated by pretest multiparameter NDE and, in some cases, were verified

by posttest fractography. Equivalent rectangular crack lengths and depths were developed

from both the NDE profiles and those determined by fractography and used to predict the

ligament rupture pressure and leak rate. Because the remaining ligaments of part-lW SCCs

rarely rupture along the full length of the crack during the initial ligament rupture event,

specimens generally undergo incremental ligament rupture with increasing pressure before the

cracks become unstable. We have tentatively generalized the equivalent crack approach to

predict incremental ligament rupture after the Initial ligament rupture event 1391.

163
Laboratory-Generated Stress Corrosion Cracks

Two types of heat treatments were used for generating the stress corrosion cracks in the

SG tubes. Early in the program, specimens were subjected to high-temperature annealing

prior to introduction of the stress corrosion crack in the laboratory. This annealing treatment

resulted in a loss of about 20% in flow stress. Later on, SCC specimens were not subjected to

the high-temperature annealing treatment. As a result, the flow stresses of these specimens

were virtually undiminished from those of the as-received tubes.

Crack depth profiles measured by pretest EC/NDE and posttest fractography for a

specimen (SGL-480) with high-temperature annealing are shown in Fig. 4.82a. The predicted

and measured leak rates at room temperature as a function of pressure for this specimen are

shown in Fig. 4.82b. Crack depth profiles and leak rate vs. pressure plots for Specimen SGL493, which also was exposed to high-temperature annealing and tested at room temperature,

are shown in Figs. 4.83a and 4.83b, respectively. The predictions of leak rate based on

fractography tend to be more accurate than those based on EC/NDE. The onset of leakage

often occurs at a lower pressure than predicted.

Crack depth profiles and leak rate vs. pressure plots for Specimen SGL-731, which was

not subjected to high-temperature annealing but was tested at room temperature, are shown in

Figs. 4.84a and 4.84b, respectively. The corresponding plots for Specimen SGL-822 are shown

in Figs. 4.85a and 4.85b, The predictions for SGL-822 are less satisfactory than the other

cases. Although fractography indicated this crack to be very deep with several 100% TW

segments, the crack behaved in a much stronger fashion than predicted.

4.5.5 High-Temperature Failure Models (Severe Accident Conditions)

Two types of failure models were used for predicting ligament rupture pressures of tubes

with part-throughwall flaws under severe accident conditions, i.e., the flow stress model and

the creep rupture model [251.

4.5.5.1 Flow Stress Model

In the flow stress model, it is assumed that, for any arbitrary history of stress a(t) and

temperature T(t) failure occurs at a temperature T and stress a whenever the following failure

equation is satisfied, independent of stress-temperature history:

O = am (4.19)

mp

where aMr Is the flow stress at temperature T. and mp is a stress magnification factor that

accounts for the crack.

164
120.00 Ap (MPa)

0 5 10 15 20

100.00

t 80.00

a 40.00

20.00

0.00

20

15

10

c 5

0

60

40 j

20 9

O A

0 0.5 1 1.5 2 2.5 3 3.5

-6 0 5 10 15 20 hp(ksi)

Crack Length (mm)

(a) (b)

Figure 4.82. (a) Crack depth profiles (measured by pretest EC and posttest fractography), and

(b) measured (symbols) and predicted (based on NDE and fractography depth) pressure

vs. leak rate plots for Specimen SGL-480.

120.00

100.00

.so.00

F

60.00

20.00

0.00-

.5

Figure 4.83.

Ap (MPa)

0 5 10 15 20

20

15

I 10

M S C1

Q

60

40 .

20 j

c

xM ,' p: o * ^e~~1 r \ :o

! . . . had~~~~~~~~~mcogra^h

-0.5 0 0.5 1 1.5 2 2.5 3 3.5

0 5 10 Is 20 AP (W)

Crack Length (mm)

(a) (b)

(a) Crack depth profiles (measured by pre-test EC and post-test fractography), and

(b) measured and predicted (based on NDE and fractography depth) pressure vs. leak rate

plots for specimen SGL-493.

165
120.00 Ap (MPa)

-4 0 4 8 12 16 20 24

100.00

80.00

2 60.00

is40.00

20.00

20

15

i 10

M5

0

-5

60

40j

20

O j

o.oo I : I I , I 1.. i -1 0 1 2 3 4

-5 0 5 10 15 2A Ap(ksi)

Axial Location (mm)

(a) (b)

Figure 4.84. (a) Crack depth profiles (by EC/NDE and fractography) and (b) predicted and measured

pressure vs. leak rate plots of Specimen SGL-731.

120

100

hp (MPa)

0 4 8 12 18 20

10

8

a0

30

80

so0

40_

20

0-

-5

I

a

to

B

4

2

20 1

10 j

0 0 A

-2 r .... I....I . . I .I I ,, ,I

-0.5 0 0.5 1 1.5 2 2.5 3

0 5 10 15 20 Ap(Wi)

Axial Location (mm)

(a) (b)

(a) Crack depth profiles (by ECINDE and fractography) and (b) predicted and measured

pressure vs. leak rate plots for Specimen SGL-822.

Figure 4.85.

166
4.5.5.2 Creep Rupture Model

A rigorous creep rupture analysis of flawed tubes is very complex. Therefore, the

approximate linear time fraction damage model (used in Section III of ASME Code) was

extended to flawed tubing:

Jtf dt (4.20)

fotR(T, mPO)=

Unlike the flow stress model, the predicted failure pressures by the creep rupture model

are dependent on the time-temperature-stress history. The creep rupture model was validated

by failure tests on specimens with part-through axial flaws of various lengths and depths

conducted under constant and ramped Internal pressure and temperature histories [251. In all

cases, predictions of the creep rupture model were closer to experimental results than those of

the flow stress model.

4.5.5.3 Severe-Accident Test Conditions

The severe accident sequences that appear to produce the greatest risk of steam

generator tube failure are those in which the reactor pressure vessel fails to depressurize, but

depressurization does occur on the secondary side. Even in these cases, preliminary

investigations suggest that failures are likely to occur in the hot-leg nozzle or the inlet

surge-line nozzle, leading to depressurization of the reactor system, which prevents failure of

steam generator tubes 140, 411. However, such calculations are subject to large uncertainties,

and the NRC is pursuing additional studies to better understand the progression of such

sequences, the temperature of the steam generator tubes during such accidents, and the

behavior of steam generator tubes (both unrepaired and repaired) at the high temperatures

associated with such accidents. At these high temperatures, plastic deformation is likely to be

much more extensive than at normal reactor operating temperatures, and creep effects may no

longer be negligible.

4.5.5.4 Unrepaired Tubes

The behavior of flawed steam generator tubing during severe accidents has recently been

considered in an INEEL 1401 and an EPRI 1411 report. The severe-accident scenario under

investigation involves a total station blackout with a SG secondary-side atmospheric dump

valve that has stuck open, resulting in loss of feedwater and secondary-side depressurization.

The pressure difference across the SG tube wall remains relatively constant at 16.2 MPa

(2.35 ksQ). Estimates of the maximum tube temperatures presented in these reports are shown

in F~gs. 4.86 and 4.87. In these reports, the failure of unflawed tubing and other components

such as the surge line nozzle is described in terms of creep damage failure. In contrast, both

analyses assume that failure of flawed steam generator tubing in severe accidents can be

described in terms of a flow stress model, Eq. 4.19, by taking the flow stress to be a function of

temperature. With this assumption, the failure pressure of a flawed tube depends only on the

flaw geometry and temperature and is independent of the loading rate or the detailed

time/temperature history.

Although the flow stress model is a straightforward extension of a model that has been

well verified at normal operating temperatures, the use of such a model at high temperatures

167
poses conceptual difficulties. Although Alloy 600 shows little strain hardening at high

temperatures, its flow stress is a function of not only temperature but also strain rate.

Intuitively, failure pressure or temperature would be expected to be truly rate-independent if

the pressure or temperature ramps were sufficiently rapid so that time is insufficient for creep

to influence the deformation or damage of the ligament and the crack tip ligament fails by

plastic instability because the ligament stress exceeds the maximum load-carrying capability of

the material. At the other extreme, If the pressure or temperature ramps were sufficiently slow

(in the limit, a constant-pressure or constant-temperature hold), failure should be controlled by

creep processes, and a failure model based on creep rupture would be more appropriate.

To determine which of the two models is more accurate for severe-accident transients,

tests were conducted at ANL on constant pressurized tubes (16.2 MPa or 2.35 ksl) subjected to

both the INEEL- and EPRI-calculated temperature ramps. The actual ramps for the tests are

indicated by the dashed lines in Figs. 4.86 and 4.87. The EPRI-calculated temperature history

indicates a dropoff in temperature after =223 min, corresponding to a predicted rupture of the

SG hot leg nozzle. Because the ANL tests were concerned only with the validation of a failure

model, this event was ignored, and the tests were ramped to failure at a rate of 20C/mIn

(3.60F/mln) after a 2-h hold time at 6670C (12330 F). As shown in Figs. 4.88 and 4.89, the

creep rupture model is more accurate than the flow stress model for the temperature ramps

considered.

4.5.5.5 Electrosleeved Tubes

Behavior of Electrosleeved tubes during a severe accident is of interest to NRC, because It

has the potential of being used as a repair technique for steam generator tubes. Although the

nanocrystalline Electrosleeve has superior mechanical properties compared to Alloy 600 at

reactor operating temperatures (Fig. 4.90), It loses strength at high temperatures because of

grain growth. This behavior raises concern as to its integrity during severe accidents.

Two approaches were developed for predicting failure of Electrosleeved tubes 1421. The

first is an approximate analytical procedure based on a linear-damage-rule hypothesis

analogous to that often used to analyze creep failures. The second is a more mechanistically

based approach that accounts for the loss of flow stress due to grain growth. Initial studies

showed that the failure temperatures predicted by the two approaches were comparable, but

the mechanistic model was used for most of the analyses. The reference geometry selected for

tests and analyses is shown In Fig. 4.91.

After the initial development of the analytical model, Framatome Technologies, Inc. (FII)

provided failure data from six tests on internally pressurized tubes that were subjected to a

variety of temperature ramps simulating those expected during the most challenging severe

accidents, I.e., "high-dry" sequences that can occur when the core melts, the primary side

remains at high pressure, and the secondary side is depressurized. The initial comparison

showed that the predicted failure temperatures were within 35-70OC (63-1260 F) of those

observed in the experiments. FTI also provided 11 Electrosleeved specimens containing

100% 'W notches in the parent tubes that were subsequently tested at ANL using the severe

accident transient designated as case 6RU in NUREG-1570 (Fig. 4.92) 1431. The results from

these tests were used to refine the input parameters in the analytical model. Based on the

refined model and more accurate data on the test geometries, the predicted failure

168
1200

1000 ANL mulation 1.75 Ctmin

800 7.51C/mln Figure 4.86.

INEEL-calculated severe-accident

SW 3.756C/rnin temperature ramp and simulated ramp

3.7:C/min used In ANL severe-accident tube400 , rupture tests.

200 -,,,D , **...., ..

0 50 100 150 200 250 300 350

line (mdn)

700 - EPRI calculated ramp 14.80°Clnin 2 d

650 .... ANL simulation2.C

850 -FANLidmulatlon 28.24C/mln Figure 4.87,

P EPRI-calculated severe-accident

e 5 10.60°Clmn temperature ramp and simulation

i 500 . / used in ANL severe-accident tube5.37°C/mln rupture tests. In the ANL simulation,

; 450 - the temperature Is held constant at

400 . ,/ 6700C (1240°F) after =217 min rather

than following the subsequent drop

,O -shown In the calculated curve.

160 170 180 190 200 210 220 230

Time (min)

850 .. . .

8 Open Symbol - EPRI Ramp Filled Symbol - INEEL Ramp

0 19 mm Figure 4.88.

c, * Unflawed Creep-rupture-model predicted vs.

' 750 Tubes observed failure temperatures for high- ,, eb/ temperature failure tests conducted

with the INEEL and EPRI temperature

8 700 ramps.

650

650 700 750 800 850

Experimental Failure Temperature (°C)

169
ago

P

O.

i 800 -

E

- 750 I...

X 700

X 650 Z

650

1000 -

800_

0!

E, 600 _

2 _ 0

8 400

0

i 200 .

0 -

Figure 4.89.

Flow-stress-model predicted vs.

observed failure temperatures for hightemperature failure tests conducted

with the INEEL and EPRI temperature

ramps.

700 750 800

Experimental Failure Temperature (IC)

850

Figure 4.90.

Flow stresses of Alloy 600 and

Electrosleeve material as a function of

temperature.

_ _ _

0 200 400 600 800

T (

0C)

Alloy 600 tube

Figure 4.91.

Reference geometry for Electrosleeved

steam-generator tube with axial crack.

0.97 mm

(0.038 in.)

Electrosleeve

170
temperatures in the FTI and ANL tests are within 150C (270F) of the observed failure

temperatures (Fig. 4.93). 1421. The test results indicate a leveling-off of failure temperature

with crack length beyond 51-76 mm (2-3 in.), which is consistent with finite element analysis.

Tests were also conducted on Electrosleeved tubes with part-throughwall notches in the parent

tubes. An FEA-based model was used to predict failure temperatures of Electrosleeved tubes

with part-throughwall cracks in the parent tubes. The results showed that, depending on the

crack length, the failure temperatures of 80% deep cracks can be 40-800C (72-1440 F) higher

than those of 100% 'W cracks. 1421.

The model predicts that Electrosleeve repairs on tubes with throughwall axial cracks

s25 mm (I In.) and throughwall 360° circumferential cracks will survive all high-dry"

sequences, including the Case 6RU transient based on the temperature histories for such

sequences presented in NUREG-1570. Tubes with throughwail axial cracks of any length but

with depths s80%Y are also predicted to survive these transients.

4.5.6 High-Temperature Leak Rate and Jet Impingement Erosion

NUREG-1570 presented the results of a study by the NRC to assess the risk implications

of steam generator tube ruptures (SGTRs) induced by severe-accident conditions. One of the

phenomena considered in NUREG-1570 was the potential for the propagation of tube failures

during design basis and severe accidents. In the event of a tube failure, escaping jet of

superheated steam could entrain particles from the core debris and impinge on a neighboring

tube. The particle-laden Jet could then erode through the neighboring tube, leading to another

escaping jet and the possibility of a cascading failure event in the steam generator. Some

analyses based on high-temperature fossil-plant experience were performed in NUREG-1570 to

estimate the time required for a jet from a failed tube to induce failure in an adjacent tube.

These analyses were, at best, scoping estimates. Additional analyses and tests were performed

to obtain more accurate estimates of the potential for failure propagation by jet impingement.

Details may be obtained from Ref. 44.

4.5.6.1 Computational Fluid Dynamics Calculations of High Temperature Jets

Detailed computational fluid dynamics (CFD) calculations were performed at the

University of Maryland to obtain a better understanding of the jet velocities and potential for

particle entrainment that could be associated with leaks In SG tubes under severe accident

conditions.* Calculations were done primarily for a square array of tubes. The distance to the

adjacent tube was taken as 8.23 mm (0.32 in.), which corresponds to an L/h value for the jet

of -16. Comparison with the results shows that this value is less than the predicted L/h ratio

for cracks <6.35 mm (0.25 in.) in length for temperatures up to 7500C. The fluid properties

were taken as those of air, and ideal gas behavior was assumed. The use of air rather than

steam will lead to an underprediction of velocities by about 25%. The baseline calculations

were performed for stagnation pressures of 8 and 16 MPa (1160 and 2320 psi), a stagnation

temperature of 1175 K (9021C), and a slit opening of 0.5 mm (0.02 in.). Because the pressure

in the tube is much higher than on the secondary side of the steam generator, the jet expands

rapidly and becomes supersonic. However, as the flow approaches the adjacent tube, it must

* Steam Generator Jet Impingement Calculations," Memorandum to Nilesh C. Chokshi,

USNRC/RES from John H. Flack USNRC/RES, December 26, 2000.

171
900

800 _

a

. 700 .

0

a 600 L E 1

0

500

400 L~

180

Figure 4.92.

Time-temperature history used for

testing Electrosleeved tubes with

notches in the parent tubes at ANL.

190 200 210 220

Time (min)

230 240 250

850

- 800

750 -

I2 700-

1 650

:5 0

600

Figure 4.93.

Observed vs. predicted failure

temperatures of ANL and FTI tests

using actual notch and Electrosleeve

geometry and, for the FTI tests, actual

test temperature ramp.

650 700 750 800

Test Failure Temperature (OC)

850

decelerate, and a detached shock wave forms in front of the impacted tube. The position and

strength of the detached shock wave depends on the stagnation pressure associated with the

jet flow. The CFD results showed that just in front of the shock the jet has expanded from

:0.25 mm in width to - ±2.5 mm. The maximum fluid velocity along the centerline is

1400 m/s (4600 ft/s). The velocity across the shock is 300 m/s (985 ft/s), and the fluid

velocity decreases rapidly as the jet approaches the impacted tube. For the aerosol particles of

interest, which have masses <10-12 kg, the particle motions are expected to track the fluid

motions very closely, as seen from the CFD results in Fig. 4.94.

172
D0 tP-@ "evsFigure 4.94.

O10 0 iF4 Kg Comparison of velocity of small

SW r particles ._ m o~12e entrained . .... In flow with fluid

M M101 .. velocity. Small particles with low

400 _ _ mz10-"'Kmass follow fluid flow very closely.

Baseline Celailatuon

0 1 2 3 4 5 e 7 8

X (mm)

4.5.6.2 Crack Opening Area

To determine the erosion rate due to jet impingement during severe accidents, the crackopening area as a function of time must be estimated 1441. A simple model was developed to

calculate the crack-opening area as a function of time and temperature during severe

accidents. It is derived by analogy to elastic-plastic solutions for a power law hardening

material 145] and Is applicable to materials that obey a power-law creep rate, i.e.,

E = A~nn,..l t A&\ ~~~~~~~~~~~~~~~~~~

where Uis the creep rate, a is the applied stress, and A and n are constants. The crack

opening displacement solution for the elastic-plastic model gives a crack opening rate of:

8= Ach2 (0, n)(ma)n, (4.22)

where h2 is a function tabulated in Ref. 45, c is the crack half-length, a is hoop stress due to

pressure, and m is the bulging factor (Eq. 4.7b). The model was validated with a number of

high-temperature tests on tubes with two symmetrical circumferential notches under axial

tensile loading. 1441. An example Is shown in Fig. 4.95a. The predicted crack opening area as

a function of axial crack length for a tube subjected to the severe-accident transient Case 6RU

at final temperatures of 700 and 7500C (1292 and 1382OF1 is shown in Fig. 4.95b.

4.5.6.3 Erosion Rate Due to Jet Impingement

During severe accidents, typical conditions involve a maximum temperature of 700-9000C

(1292-16520 F), a maximum pressure gradient across the tube wall of 16.2 MPa (2350 psi). and

an escaping jet consisting of superheated steam plus 10-20% H2. The particulates in the gas

jet arise from the volatilization of low-melting-temperature core materials. The volatiles

agglomerate at the relatively low temperatures found in the reactor coolant piping system.

Results from earlier severe accident analyses showed that the particles in the jet were primarily

173
13 -Axial Load=1 360 kg , 0.8

- Test /7 0.7

25 - ~~~~CR-106 j 0.6

D2 /- - 0.5

5 0.4

)I ~~~~~~~~0.3

15 0.1

o f .. I ,...,,, ,,. I ... . I. .. .. .I... . of

500 550 600 650 700 750 800

T (°C)

I 100

Iz 1 0-

0

Axial Crack Length (mnm)

5 10 15

Final T=750rC

_.1.1-1.1 1 ....Final T=700I C

100 I

10 I

0.1

0.01!

0 0.1 0.2 0.3 0.4 0.5 0.6

Axial Crack Length (in.)

Figure 4.95.

(a) (b)

Predicted and experimentally measured (symbol) notch opening displacement for two

6.4-mm (0.253-in.) circumferential notches loaded axially and subjected to the Case 6RU

(Fig. 4.92) temperature transient and predicted clack-opening area with crack length at final

temperatures 700 and 7500C (1292 and 1382 0F) for a tube subjected to Case 6RU

transient.

Ag (-100 g/m3 ), with significantly lesser amounts of In 2 03 , CsMoO4, SnO2, CsI, and other

species (these results are typical of Westinghouse PWRs). The median particle diameter is

3 1.5 tm, with the bulk of particles having diameters < 3 ism.

A test program was Initiated at the University of Cincinnati to determine erosion rates of

Alloy 600 at high temperatures by particles with sizes and characteristics, including hardness

and density, similar to those expected in a reactor accident. Ni and NIO particles were chosen

initially as surrogate particles for the tests. Because the surrogate particles are somewhat

larger and a higher hardness at temperature than the real particles, they should give realistic

to somewhat conservative estimates of erosion rates. Initial tests showed that erosion was

most severe at an angle of 30°, in agreement with previous literature data on the erosion of

Alloy 600 (and other metals), and this angle was used in all of the subsequent tests.

Tests were conducted at 700 0C (1292 0F) and a 30° impact angle using 3-7 Im Ni erodent

at velocities of 91-549 m/s (300-1800 ft/s). Four test series (A, B. C, and D) were conducted at

different times to determine the reproducibility of the data. The results from test series A and

B indicate that the erosion rate increases markedly with increasing particle velocity, and that

for a given velocity, the rate is independent of total particle loading (fluence). While test series

A and B yield consistent results, the data from the subsequent test series C indicate no erosion

at 91 and 183 m/s (300 and 600 ft/s) and a substantially lower erosion rate at 549 m/s

(1800 ft/s). The reason for these discrepancies is not clear. Overall, however, the erosion rates

are relatively low in all cases. To determine the rate of thinning of the target, it is convenient to

convert the rate of mass loss into a rate of volume loss by dividing by the density of the target

D. R. Diercks, OMinutes of Expert's Meeting on Jet Impingement and Leak Rates from Steam

Generator Tubes During Severe Accidents,' held at Argonne National Laboratory (ANL),

Nov. 19, 1999; minutes published Dec. 6, 1999.

174
material, which in this case is Alloy 600, and to note that the erosion rate (based on the results

from test series A and B) can be expressed in terms of a power-law dependence on velocity:

E = 8.0 x 10-12V 2. 4 3 , (4.23)

where the erosion rate E is in cm3 /g and denotes the rate of loss of material per gram of

impinging particles, and V is the velocity of the impinging particles in m/s.

To determine the thinning rate of tubes, erosion data must be coupled with the results of

the CFD analyses of the jet behavior discussed above:

dh = E * D, (4.24) dt

where dh/dt is the wall thinning rate in cm/s, E is the erosion rate in cm3/s. and 0 is the

mass flux of particles in g/cm2 /s. If we make the conservative assumptions that the impact

velocity is =200 m/s, and the erodent particles are 3-7 Fm diameter Ni particles, then the

power-law estimate of the erosion rate of Alloy 600 given by Eq. 4.23, together with the

estimate of the mass flux (0.52 g/cm2 /s), suggests that the maximum wall thinnng rate will be

less than (8.0 x 10-12 x 2002.43 x 0.52 x 36000) = 0.06 mm/h (2 mil/h). Because of

conservatism in the analyses, the bounding rate of 0.06 mm/h (2 mll/h) should be applied only

for that portion of the history for which the temperature is >7000C. Thus, even after a crack

has opened by creep at high temperatures, failure of an adjacent tube would take >10 h. Thus,

jet impingement is very unlikely to play any significant role in the failures.

175
5 Integrated Model for Steam Generator Integrity Assessments

The Nuclear Regulatory Commission (NRC) is developing a "performance-based"

regulatory framework to assure steam generator tube integrity. Instead of just meeting

prescribed rules on allowable flaw sizes for tubes that have been inspected, evaluations and

assessments are necessary to show that adequate levels of Integrity can be maintained until

the next scheduled outage. This will require the use of a comprehensive validated model that

integrates all important aspects of integrity evaluations, starting from in-service inspection

results and ending with a total leak rate at the end of an operating cycle under various

assumed conditions. Furthermore, the probabilities of single and multiple tube ruptures under

various reactor coolant conditions must be quantified. The nuclear industry has developed

codes and procedures to carry out these assessments. ANL/CANTIA is intended to provide a

tool that could be used for an independent assessment of steam generator integrity.

Under contract from the Atomic Energy Control Board of Canada IAECB, now the

Canadian Nuclear Safety Commission]. Dominion Engineering, Inc. 1461 developed a Monte

Carlo-based code called CANTIA (CANDU Tube Inspection Assessment) to simulate the effects

of tube inspection and maintenance strategies on the safe operation of CANDU design nuclear

steam generators. The integrity and leak rate models in CANTIA were specifically Intended for

the CANDU steam generators in Canada. A decision was made to incorporate into CANT1A the

integrity and leak rate models that have been developed at ANL under the Integrated Steam

Generator Tube Integrity Program (ISGTIP-2) sponsored by the NRC. The original CANTIA was

written in Visual Basic 3.0, and a copy of the program, together with the user's manual, were

obtained by ANL from the AECB. The program was updated to Visual BASIC 6.0 to run in a

Windows NT environment and was then modified to incorporate the integrity and leak rate

models developed at ANL. The basic Monte Carlo simulation portion of the code, together with

much of the Input format, was retained unchanged in the modified CANTIA code, renamed

ANL/CANTIA. For a detailed description of ANL/CANTIA. see Ref. 47.

5.1 Description of Original CANTIA

The CANTIA code uses Monte Carlo techniques to determine probabilities of steam

generator tube failures under accident conditions and primary-to-secondary leak rates under

normal and accident conditions in the future. An initial flaw size distribution is input, and the

program determines future flaw size distributions. The effect of inspections of the steam

generator tubes at future points in time, and subsequent removal of defective tubes from the

population, on the probabilities of failure and leak rates and the future flaw size distributions

is calculated. Thus, the effect of different inspection and maintenance strategies on the

probability of tube failure and primary-to-secondary leak rate can be determined.

CANT1A is designed to handle various defect types (including circumferential and axial

SCCs, frets, and pits), though only one type of degradation can be modeled at a time. All tubes

modeled in a trial are treated as being equally susceptible to the chosen degradation mode, so

that the same distributions of growth rates, flaw Initiation, etc., are used for all the tubes in the

trial. If different distributions are believed to describe the degradation in different parts of the

steam generator, then separate simulations should be run with the appropriate distributions

for the portion of generator being modeled.

176
The CANTIA failure model is of the probabilistic fracture mechanics type. It calculates

the time-dependent flaw size and number distribution so that the probability of failure or the

rate of leakage can be estimated. It uses a Monte Carlo approach; each important parameter is

treated as a random variable with known or predictable median behavior and a known or

predictable distribution of behavior (i.e., variation) around the median. Both the median and

variance of each random variable are specified by the user, or, if desired, a fixed (deterministic)

value can be used. For each Monte Carlo trial, the value of each variable is chosen randomly

in accordance with the selected distribution function, or the fixed value is used, so that the

trials reflect the variability of the actual situation. A large enough number of trials (chosen by

the user) are then run for each analysis to provide a stable statistical distribution of the

results.

For each Monte Carlo trial, CANTIA determines flaw sizes, growth rates, inspection

results, material properties, etc. for each tube, and tracks the progression of the flaws In each

tube throughout the model time period. After all trials are run, the conditions from each trial

at each time of interest are compiled to determine the distributions of flaw sizes, Inspection

results, leak rates, and tube failures at the times of interest.

5.2 New Features Added In ANiCANTIA

5.2.1 InItial Flaw Size Distributions

A major feature added in ANL/CANTIA is the ability to treat the growth of twodimensional cracks, I.e., cracks are allowed to grow (by stress corrosion) in length and depth.

Thus, the input In ANL/CANTIA requires an initial flaw depth distribution as well as an Initial

flaw length distribution. Currently, the depth is taken as the primary dimension, whereas the

initial length distribution is entered in terms of a distribution in the aspect ratio

(length/depth). The distribution for the aspect ratio can be any of the five statistical

distributions: beta, gamma, Gumbel, lognormal, and Weibull. Alternatively, a fixed

deterministic ratio can be chosen.

5.2.2 Flaw Growth Model

The basic flaw considered in the original CANTIA is a one-dimensional flaw, e.g., an

infinitel long shallow part-throughwall crack with a stress intensity factor K given by

K ==BSsze, where the user must specify constants B and S. In the ANL/CANTIA, we

consider an initially elliptical two-dimensional surface flaw of length 2c and depth a.

ANL/CANTIA then computes, by numerical integration (4th order Runge-Kutta), the growth (by

stress corrosion) of the crack both through the wall and in length, up to the point where it

becomes a throughwall crack. CANTIA then follows the growth (by stress corrosion) of the

throughwall crack until unstable rupture.

5.2.3 Growth Rate of Stress Corrosion Crack

In ANL/CANTIA we have provided two models for SCC growth rates: one due to Scott 1481

and the other due to Ford and Andresen 1491. It is assumed that the same equation can predict

crack growth rates In the thickness and the length directions.

177
5.2.4 Failure Models

In addition to the failure models provided in the original CANTIA, ANL/CANTIA includes

options for the failure of axial cracks the ANL ligament rupture model [41 and the

unstable-rupture model due to Erdogan 1291. A circumferential crack failure model, based on

a modification of the work of Kurihara et al. 1351, was also added.

5.2.5 Leak Rate Models

In addition to the leak rate models available in the original CANTIA. ANL/CANTIA

provides a leak rate model based on simple orifice flow through a crack with an opening area A

and a coefficient of discharge Cd = 0.6:

Q=0.6Aj2p (5.1)

where A is the crack opening area, Ap is the pressure differential, and p is the mass density.

This model has been used successfully to predict leak rates In experiments conducted at room

temperature and 2821C on specimens with EDM notches and stress corrosion cracks at ANL

down to a leak rate of =0.44 L/min (=O.1 gal/mnn). The correlation will break down at lower

levels of leak rate for very tight stress corrosion cracks, particularly at elevated temperatures

where flashing may occur inside the cracks; but such a limit has not yet been established. The

crack opening area is calculated from a model developed by Zahoor 1371 and validated by finite

element analyses and tests at ANL.

5.2.6 Residual Stress

In ANL/CANTIA, the peak residual stress value (mean and variance) is entered by the

user. Currently, the residual stress is idealized as a membrane stress that is reduced linearly

with increasing crack depth until it is zero at 100% crack depth. Unless the initial crack

depths are very deep, pressure stresses alone cannot grow the stress corrosion cracks through

the tube wall without the assistance of residual stresses. However, the current residual stress

model in ANL/CAN7IA is highly simplified and should be made more realistic in the future.

5.3 ApplIcations of ANUCANTIA

5.3.1 Effect of POD

The change in the conditional probability of one or more tube ruptures with time is

plotted in Fig. 5.1 for the case when the probability of detection (POD) = 0.9. Variations of

distributions of crack depth and crack length with time for the same POD are plotted in

Figs. 5.2a and 5.2b, respectively. Similar results are plotted in Fig. 5.3 and Figs. 5.4a and 5.4b

for a POD = 0.6 and in Fig. 5.5 and Figs. 5.6a and 5.6b for a POD of 0.01.

178
e

I-K

a:

0

E

0

b-0D

0

C

.0

I0

IEL

Figure 5.1.

Variation of probability of burst of one

or more tubes with time for Initial

residual stress = 30 ksi, POD = 0.9.

and Inspection Interval of 10,000 h.

Ilo 5

Time (h)

0.5 0.5

0.4

0.3

0.2

0.1

0

0.4

0.3

K

0.2

0.1

n 0 0.5 1

Crack Length On.)

1.5 2

Crack Depth

(a) (b)

Figure 5.2 Variation of the distribution of (a) nondimensional crack depth and (b) crack length with time

for initial residual stress of 30 ksi and POD = 0.9.

e 0.015

S_ 0.01

12

0

0

0 0.005

0

.0 0

2

0.. 0

Figure 5.3.

Variation of probability of burst of one

or more tubes with time for initial

residual stress = 30 ksl, POD = 0.6,

and Inspection Interval of 10,000 h.

I.

2 10 4 104 6 104

lime (h)

8 104 1 105

179
1

0.8

0.6

. v s .

x

Bin size

, I IiK

I I

211 POD R

:1 %'. 400%Pi

Ij

- - .

=0.05 [= -No bnaiation

- - - Case 1

- -Case 2

30,000 h

br TSPAD Flaws

ugging Criteria

l Inspection

K

0.4

0.2

- I Ua 0.2 0.4 0.6 0.8 1 W0 0.5 1 1.5 2

Crack Depth Crack Length (in.)

(a) (b)

Variation of the distribution of (a) crack depth and (b) crack length with time for initial

residual stress of 30 ksi and POD = 0.6.

Figure 5.4.

0.

:3

I0

0

0

Z.

0.5 -

0.4

0.3

0.2

0.1

0

0

Figure 5.5.

Variation of probability of burst of one

or more tubes with time for Initial

residual stress = 30 ksi, POD = 0.01,

and inspection interval of 10,000 h.

2104 410k 6104 8104

Time (h)

1 lo,

0.5 _

0.4

0.3

0.2 _

0.1

01

0.5

0.4

0.3

0.2 _

0.1 _

I .

0.2 0.4 0.6 0.8 I1 0.5 1 1.5 2

* Crack Depth Crack Length (hi.)

(a) (b)

Figure 5.8. Variation of the distribution of (a) crack depth and (b) crack length with time for initial

residual stress of 30 ksi and POD = 0.01.

180
As expected. the probability of one or more tube rupture increases with decreasing POD.

Peaks and valleys in Fig. 5.3 represent the effect of plugging on the probability of tube rupture.

The probability of tube rupture is very small for POD = 0.9. For POD = 0.6, the probability of

one or more tube rupture peaks at an intermediate time. In the case of POD = 0.01, the

probability of one or more tube rupture increases monotonically with time. The locations of the

peak in the probability of crack depth and crack length change very little with time or POD.

The spikes in the probability of crack length at zero crack length represent plugged tubes.

Similarly, the spike in the probability at a crack depth of 100% In Fig. 5.6a represent cracks

that become throughwall. As expected, the probability for plugging is essentially zero when the

POD = 0.01.

Next, we explored the difference between assuming a fixed probability of crack detection

(POD = 0.6) and assuming a POD curve determined for TSP/ID flaws in the recent ANL mockup

round robin on the probability of one or more tube ruptures. Although the ANL round-robin

POD was best-fitted with a linear logistic curve (not available as an option in the current

ANL/CANTIA), the POD curve can also be approximated with a lognormal curve, which is one

of the options in ANL/CANTIA. This lognormal curve for POD was used in the calculations and

100% tube inspection was assumed. A steam generator with 10 susceptible tubes, each

containing a single crack, was considered. A lognormal distribution of initial crack depth was

assumed with a fixed initial length to depth aspect ratio of 40. No new cracks were assumed to

initiate during the lifetime of the steam generator. The Scott model with a conservative

threshold stress intensity factor (Kth = 5 ksl-4in) was used to determine crack growth rates.

For the purpose of this exercise, we assumed APnorm = 1.3 ksl, APMSLB = 2.5 ksi, T = 2880C, an

in-service-inspection interval of 10,000 h with a 40% plugging criterion, and a total life of

80.000 h. Also, an Initial membrane residual stress of 30 kst was used, and the residual stress

was assumed to relax linearly with crack depth, reducing to zero when crack depth = 100%.

One-thousand simulations were used for the Monte Carlo analyses. The variations of

distributions of crack depth and crack length with time for the assumed PODs are plotted In

Figs. 5.7a and 5.7b, respectively. Because of plugging, the depth and length at peak

probability decrease slightly with time. Also, plugging causes a spike in the probability

distribution of crack length at zero crack length. Evolution of the probability of one or more

tube ruptures with time is plotted in Fig. 5.8 for both POD curves. The discontinuous drops 'in

the probability of rupture curve are caused by plugging of tubes. The probability of tube

rupture is basically zero (< 0.001) beyond two equivalent full power years (EFPY) of operation if

the full POD curve is used. On the other hand, If a fixed POD = 0.6 is used, the probability of

tube rupture does not reduce to zero until 9 EFPY of operation. Of course, these conclusions

are valid only for cases where no new cracks are allowed to initiate during the lifetime of the

steam generator.

As expected, the probability of one or more tube ruptures is greater when a fixed POD Is

used than when the full POD curve is used. The crack depth and length distributions are quite

similar for both POD curves, although the probability of plugging (spike at zero crack length in

Fig. 5.7b) is greater when the full POD curve is used,

181
0.5 0.6

['13,Wsk, "-''.O'S'lm ......... F 1h

0.4

0.3

0.2

80i000h 0.5 80,

0.4 POD=0.6

DAD Flaws and POD for TSP/I

'~0.3

g Criteria . 40% Plugging

;pecton 0.2 100% Insp

lation I ANointlik

0.1

0

0.8 1 0 0.2 0.4 0.6 0.8 1 1.2

Crack Lergth (In.)

]h

D Flaws

p Crlterta

ecton

tlWon

1.4 1.8

0.1

0

0 0.2 0.4 0.6

Crack Depth

Figure 5.7.

(a) (b)

Variation of the distribution of (a) crack depth (bin size = 0.05) and (b) crack length (bin

size = 0.05 In.) with time using POD curve for TSP ID flaws and a constant POD = 0.6.

Initial residual stress Is 20.7 MPa (30 ksi).

0.015

'ua

0.01 I0

2 0 0.005

0T

I 0

Q (

Figure 5.8.

Variation of probability of rupture of one

or more tubes with time in effective full

power years (EFPY) for a constant

POD= 0.6 and a POD curve for

TSPAID flaws. No new flaw initiation

was assumed.

r.

2 4 6

EFPY

8 10 12

5.3.2 Effect of Crack Initiation

Until now no new crack Initiation with time has been assumed. To explore the effect of

crack Initiation on failure probabilities, we assumed two cumulative Welbull distributions for

crack initiation times, denoted by Case 1 and Case 2 in Fig. 5.9. Case 1 represents a steam

generator that has very low crack initiation resistance, while Case 2 represents an steam

generator with greater initiation resistance. Each new crack initiated has a deterministic initial

depth of 0.2 of the wall thickness and deterministic initial aspect ratio of 40 (i.e., initial

length = 1.0 cm or 0.4 in.). For the purpose of calculation, a steam generator with 50

susceptible tubes was considered, and Initially 10 tubes were assumed to contain a crack with

a lognormal size distribution. The evolution of the distributions for the number of cracks with

time for the two cases is plotted in Figs. 5. 10a and 5. 10b. The evolution of the distributions of

182
1

0.8

=0.6

ED

~30.2 0

Figure 5.9.

Two cases of cumulative Weibull

distributions assumed for crack

initiation times.

0 2104 4104 6104 8104

lime (h)

1 105 1.2 10

0 10 20 30 40 50 60 70 0 10 20 30 40 60

No. of Flaws No. of Flaws

(a) (b)

Figure 5.10. Evolution of the distribution for the number of flaws with time for Weibull Initiation (a) Case 1

and (b) Case 2.

crack depth for various cases is presented In Figs. 5.1 la and 5.1 lb, and similar presentations

for crack length are given In Figs. 5.12a and 5.12b. It Is evident that allowing new cracks to

initiate changes the distributions slightly compared to a case without new crack Initiation.

However, for the POD curve (TSP/ID flaws) and plugging criteria (40%) assumed (with 100%

inspection), the distributions are not significantly different for the two cases in which new

cracks are allowed to Initiate. The probability of one or more tube ruptures or for leak rate >4

L/min (>1 gpm) under normal operation for all cases was calculated to be very small (< 0.001).

183
1

0.8

0.6

0.4

0.2

0

.Bin skze_0.05 No Ihitiaton

-Case I

--- Case 2

-0 S |30,000 h

POD for TSP/ID Flaws

40% Plugging Criteria

100% Inspection

x

1

0.8

0.6

0.4

0.2

0

Bin size=).05 No kiatkmon

_ I Case I

: I; ---- Case 2

iI 100000 0 h

POD for TSPMD Flaws

40% Plugging Criteria

100% Inspection

: ''I

0 0.2 0.4 0.6 0.8

Crack Depth

1 1.2 0 0.2 0.4 0.6 0.8

Crack Depth

1 1.2

(a) (b)

Figure 5.11. Distributions of crack depth calculated without initiation and with Weibull initiation Case 1

and Case 2 at (a) 30,000 and (b) 100,000 h.

0.5

0.4

0.3

A

0.2

Bin size=0.05 n. . ..

.I-----Case I --- Case 2

30,000 h

POD for TSPiD Flaws

I \ 40% Plugging Criteria

100% Inspection

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6

Crack Length (in.)

(a)

0.5

0.4

0.3

0.2

0.1 0.1

0 0

0 0.2 0.4 0.6 0.8 1 1.2

Crack Length (n)

(b)

1.4 1.6

Figure 5.12. Distributions of crack length calculated without initiation and with Weibull initiation Case 1

and Case 2 at (a) 30,000 and (b) 80,000 h.

184
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188
NRC FORM 335 U. & NUCLEAR REGULATORY COMMISSION 1. REPORT NUMBER

PM) 03*eOdbyNRC. AdWO &M, Rev.

3OM10= BIBLIOGRAPHIC DATA SHEET NUREG/CR-6804

____________________________________________ANL-02/28

2. TEAND PMSTME

Second U.S. Nuclear Regulatory Commission International

Steam Generator Tube Integrity Research Program a DATE REPORT PUBMS

Final Project Summary Report MONTH YEAR

September 2003

4, FIN OR GRAN KAMBER

W6487

S.6AUTHOR T. AYPE OF REPORT

S. Bakhtiarl, K. E. Kasza, D. S. Kupperman, S. Majumdar, Technical; Topical

J. Y. Park, W. J. Shack, and D. R. Diercks 7.PERP cOVEED msw.DOW

a&PERXXUNB ORAGkTIO-KE ADP(N4antMsfEorNDb;S#: US Ac*-eVo-*1 &nn0n*wcftdrp

now mdmefefen

Argonne National Laboratory

9700 South Cass Avenue

Argonne, IL 60439

&~SPONSORMORGANVAION-NAME AMD ADORESS (VARCM Wme Io* opd MCe E*it. M CIM= orRegbM U.S. Mx~wRep" CwoVsshr

Division of Engineering

Office of Nuclear Regulatory Research

U. S. Nuclear Regulatory Commission

Washington, DC 20555 -0001

10. SIPLEMENTARY NOlar

J. Muscara, NRC Project Manager

11. ABSTRACT POO waf or Iss)

This report summarizes work performed under the USNRC Steam Generator Integrity Program at

Argonne National Laboratory. The areas addressed by the program include assessment of procedures

and equipment used for In-service (ISi) inspection of steam generator (SG) tubes; recommendations for

criteria and requirements to improve the reliability and accuracy of ISI; validation and Improvement of

correlations and models that are used to evaluate integrity and leakage of degraded SG tubes;

and validation and Improvement of correlations and models that predict the generation and progression

of degradation in SG tubes as a function of aging.

12 KEY lODlSC twt phrasr tlgit ' 1 h ban a spJ 13. AVALASILY BrATUEENT

Steam Generator Unlimited

Tubes 14. SECURMCLAssFICATION

Nondestructive Evaluatibn (Th pea)

Eddy Current Testing Unclassified

Inservice Inspection . .__RO___

Stress Corrosion Cracking Unclassified

Alloy 600, Inconel 600 15.NUAEROFPAMES

1.PRICE

NRCFORa3M=P-
4~~~~~~

I'~~~~~~~~~~~~~~~~~~~~~~~~

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paper

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