An Assessment of Ultrasonic 

Techniques for Far-Side 

Examinations of Austenitic 

Stainless Steel Piping Welds

Office of Nuclear Regulatory Research

NUREG/CR-7113

PNNL-19353
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An Assessment of Ultrasonic 

Techniques for Far-Side 

Examinations of Austenitic 

Stainless Steel Piping Welds

Manuscript Completed: October 2011

Date Published: November 2011

Prepared by

M. T. Anderson, A. A. Diaz, A. D. Cinson, S. L. Crawford,

S. E. Cumblidge, S. R. Doctor, K. M. Denslow, and S. Ahmed

Pacific Northwest National Laboratory

P.O. Box 999

Richland, WA 99352

W. E. Norris, NRC Program Manager

NRC Job Code N6398

Office of Nuclear Regulatory Research

NUREG/CR-7113

PNNL-19353

iii

Abstract

Austenitic stainless steels are used in operating nuclear power plants primary loop piping 

because of their resistance to corrosion, or in locations where high strength and creep 

resistance are required because of elevated temperatures. Austenitic stainless steel piping 

welds are susceptible to stress corrosion cracking so it is essential to perform effective and 

reliable inspections. However, their anisotropic and large-grained structure strongly affects the 

propagation of ultrasound making it difficult to inspect. Another impediment to performing 

satisfactory inspections is that a significant number of weld locations exist where component 

geometry or permanent obstructions only provide suitable access to one side of the weld. 

These ―far-side inspections‖ of piping welds currently are performed on a ―best-effort‖ basis and 

with the exception of single-side dissimilar metal welds, do not conform to ASME Code 

Section XI, Appendix VIII, performance demonstration requirements. The industry, through the 

Electric Power Research Institute (EPRI), is actively involved in improving far-side inspections 

through the use of phased-array probes.

As a result, confirmatory research is being conducted for the U.S. Nuclear Regulatory 

Commission at the Pacific Northwest National Laboratory to assess the capabilities of ultrasonic 

testing (UT) for the examination of austenitic stainless steel piping welds from the far side.

Specifically, studies were conducted to assess the ability of advanced UT techniques to detect 

and determine the size of flaws from the far-side of wrought austenitic piping welds. Far-side 

inspections of nuclear system piping welds are currently performed on a ―best-effort‖ basis and 

do not conform to ASME Code Section XI Appendix VIII performance demonstration 

requirements. To ensure safety in the case of less than 100% examination coverage of a 

particular weld, the staff requires that the licensee demonstrate in its relief request that other 

welds of the same examination category have had 100% coverage with no flaws, and that the 

subject weld has no active degradation mechanism based on operating experience. The staff 

may also require the licensee to improve its examination coverage by improving its examination 

technique (e.g., using certain transducers). The staff approves these less-than-100% 

examination coverage relief requests based on rigorous evaluation to ensure that structural 

integrity of the piping is maintained in accordance with the ASME Code, Section XI. For the 

laboratory work in this study, four circumferential welds in 610-mm-diameter, 36-mm-thick 

ASTM A-358, Grade 304 vintage austenitic stainless steel pipe were examined. The welds 

were fabricated with varied welding parameters; both horizontal and vertical pipe orientations 

were used, with air and water backing, to simulate field welding conditions. A series of saws, 

electro-discharge machined notches, and implanted fatigue cracks were placed into the heataffected zones of the welds. The saws and notches ranged in depth from 7.5% to 28.4% 

through-wall. The implanted cracks ranged in depth from 5% through-wall to 64% through-wall. 

The welds were examined with phased-array technology operating at 1.5 MHz, and with lowfrequency/Synthetic Aperture Focusing Technique (SAFT) methods in the 250–400 kHz 

regime. These results were compared to encoded conventional (monolithic transducer element) 

ultrasonic techniques as a baseline at 2.0 MHz, a frequency typical for examinations of piping 

welds in the field. The examinations showed that while phased-array and low-frequency/ SAFT 

were each able to detect and accurately length-size, but not depth-size, the notches and flaws 

through the welds, phased-array ultrasonic inspection provided the best results, detecting nearly 
iv

all of the flaws from the far side. The ultrasonic results were insensitive to the different welding 

techniques used in each weld.

Based on the laboratory work conducted in this study, phased-array ultrasonic inspection 

provided the best results, detecting nearly all of the flaws from the far side. These results were 

presented at the Fifth International Conference on NDE in Relation to Structural Integrity for 

Nuclear and Pressurised Components in 2006. To better understand acoustic propagation 

through austenitic welds to determine whether the phased-array inspections currently being 

conducted are effective and reliable, sound-field mapping was conducted. At the invitation of 

EPRI, PNNL examined field-removed specimens containing service-induced intergranular stress 

corrosion cracks at the EPRI Nondestructive Evaluation Center, in Charlotte, North Carolina, to 

compare results. Collective results from the activities described above are presented here.
v

Foreword

Article IWB-2500, ―Examination and Pressure Test Requirements,‖ of Section XI, ―Rules for 

Inservice Inspection of Nuclear Power Plant Components,‖ of the American Society of 

Mechanical Engineers (ASME) Boiler and Pressure Vessel Code (BPV Code), provides that 

certain nuclear power plant components be inspected in accordance with Table IWB-2500-1, 

―Examination Categories.‖ In most cases, the table states that essentially 100% of any given 

weld be inspected. The ASME BPV Code is incorporated by reference into Title 10, Part 50, of 

the Code of Federal Regulations, ―Domestic Licensing of Production and Utilization 

Facilities.‖ Thus, licensees are required by law to adhere to the provisions of the ASME BPV 

Code.

Conventional transducers operate at fixed frequencies. The location at which a transducer 

focuses sound is determined by the transducer area and shape, the ultrasound frequency, and 

the sound velocity of the propagation medium. Conventional transducers require access to both 

sides of a weld to properly insonify the weld volume and thus detect any defects that may be 

present. There are a significant percentage of weld locations where component geometry or 

permanent obstructions only provide suitable access to one side of the weld. Multiple 

inspection techniques are sometimes required to adequately inspect certain welds which results 

in additional occupational dose. Various investigations have been performed to determine the 

potential for inspecting welds from one side using phased-array probes, and the work described 

herein helped the NRC evaluate the acceptability of these techniques.

Phased-array probes consist of many small ultrasonic elements, each of which can be pulsed 

individually. By varying the pulsing of the elements one-by-one in sequence, a beam at a 

known angle can be initiated. This beam can be steered electronically thus sweeping through a 

weld like a searchlight. Software allows the data from multiple sweeps to be collected into an 

image showing internal sound reflectors such as counterbores or weld defects. Many studies 

have demonstrated the superiority of phased-array probes over conventional probes.

Nuclear power plant austenitic stainless steel piping welds are susceptible to stress corrosion 

cracking. The U.S. Nuclear Regulatory Commission conducted this confirmatory research at 

the Pacific Northwest National Laboratory (PNNL) to assess the effectiveness and reliability of 

ultrasonic testing for examining austenitic stainless steel piping welds. The assessment by 

PNNL shows that the use of phased arrays provides the best overall results with the highest 

number of far-side flaws being detected. The capability to length-size is well within ASME 

Code-established acceptance criteria. Flaw-tip signals were not observed from the far-side, so 

no estimate of flaw depth was possible. Far-side inspections are performed on a best-effort 

basis. While the ASME Code requires essentially 100 percent inspection of welds, the structure 

of austenitic stainless steel welds in combination with limitations due to geometry results in 

some cases of inspection coverage of less than 50 percent. Accordingly, while the use of 

phased arrays has been shown to be a significant step forward relative to the capability to 

inspect austenitic stainless steel piping welds, further enhancements by the industry will be 

required to improve weld coverage.

vii

Contents

Abstract...................................................................................................................................... iii

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

Executive Summary .................................................................................................................. xv

Acknowledgments.................................................................................................................... xix

Acronyms and Abbreviations.................................................................................................... xxi

1 Introduction.......................................................................................................................1-1

2 Background and Objectives ..............................................................................................2-1

3 Scope of Ultrasonic NDE Assessments ............................................................................3-1

4 The Far-Side, Austenitic Weld Inspection Challenge ........................................................4-1

5 Specimens Evaluated in this Study...................................................................................5-1

5.1 Phase 1: Austenitic Piping Specimen No. 02-24-15...............................................5-1

5.2 Phase 2: Austenitic Piping Specimen No. 3C-022..................................................5-8

5.3 Phase 3: IGSCC - EPRI Practice and Secure Set Specimens .............................5-10

5.3.1 EPRI Performance Demonstration Practice Specimens ............................5-11

5.3.2 EPRI Performance Demonstration Secure Specimens..............................5-12

6 Ultrasonic Probes and Inspection Protocols......................................................................6-1

6.1 Conventional–UT....................................................................................................6-1

6.2 Low-Frequency/SAFT–UT ......................................................................................6-4

6.3 Phased Array–UT .................................................................................................6-12

7 Phase 1 Evaluation...........................................................................................................7-1

7.1 Conventional-UT: Data Analysis and Results.........................................................7-1

7.2 Low-Frequency/SAFT: Data Analysis and Results...............................................7-10

7.2.1 Tutorial on SAFT Data Analysis Process...................................................7-12

7.2.2 Piping Segment 1......................................................................................7-16

7.2.3 Piping Segment 2......................................................................................7-20

7.2.4 Piping Segment 3......................................................................................7-23

7.3 Phased Array UT: Data Analysis and Results ......................................................7-26

8 Phase 2 Evaluation...........................................................................................................8-1

8.1 Conventional UT: Data Analysis and Results.........................................................8-1

8.1.1 Weld One....................................................................................................8-1

8.1.2 Weld Two....................................................................................................8-2

8.1.3 Weld Three .................................................................................................8-5

8.2 Low-Frequency/SAFT: Data Analysis and Results.................................................8-6

8.2.1 Weld One....................................................................................................8-7
viii

8.2.2 Weld Two..................................................................................................8-10

8.2.3 Weld Three ...............................................................................................8-12

8.3 Phased Array UT: Data Analysis and Results ......................................................8-14

8.3.1 Weld One..................................................................................................8-15

8.3.2 Weld Two..................................................................................................8-16

8.3.3 Weld Three ...............................................................................................8-20

9 Discussion........................................................................................................................9-1

9.1 Conventional UT – Phase 1....................................................................................9-1

9.2 Conventional UT – Phase 2....................................................................................9-1

9.3 Low-Frequency/SAFT-UT – Phase 1 ......................................................................9-2

9.4 Low-Frequency/SAFT-UT – Phase 2 ......................................................................9-7

9.5 Phased-Array UT – Phase 1 ...................................................................................9-9

9.6 Phased Array-UT – Phase 2 .................................................................................9-10

10 Longitudinal-Wave Sound-Field Assessments and Simulations......................................10-1

10.1 Results .................................................................................................................10-6

10.2 Theoretical Analysis of Longitudinal-Wave Propagation through Austenitic 

Weld ...................................................................................................................10-12

10.3 Discussion and Conclusions ...............................................................................10-14

11 Phase 3 Evaluation.........................................................................................................11-1

12 Summary and Conclusions.............................................................................................12-1

13 References .....................................................................................................................13-1

Appendix A – Piping Specimen Fabrication Photographs........................................................ A-1

Appendix B – Enhanced Far-Side Analysis Using a Synthetic Aperture Focusing 

Technique ................................................................................................................... B-1

Appendix C – Conventional UT Data and Analysis Images .....................................................C-1

Appendix D – Low-Frequency/SAFT Data and Analysis Images .............................................D-1

Appendix E – Phased Array Data and Analysis Images .......................................................... E-1

Appendix F – Conventional UT Data and Images of Piping Specimen No. 3C-022 ................. F-1

Appendix G – IGSCC Images and Detection Calls from Data on Practice Set 

Specimens ..................................................................................................................G-1
ix

Figures

4.1 Polished and Chemically Etched Side View of an Austenitic Weld 

Microstructure in Wrought Stainless Steel ....................................................................4-2

5.1 Sectioning and Flaw Locations for Pipe Specimen 02-24-15 ........................................5-2

5.2 Cross-Sectional Schematic (circumferential orientation) Showing Typical 

Weld Geometry and Implanted Crack in HAZ ...............................................................5-3

5.3 Cross-Sectional Schematic (axially oriented) Showing Typical Crack Location.............5-3

5.4 Positions and Angles for the Sawcuts in Pipe Specimen 02-24-15 ...............................5-5

5.5 EXAMIX Molds Used to Measure the Length and Depth of the Sawcuts in 

Pipe Specimen 02-24-15 ..............................................................................................5-6

5.6 Experimental Set-up Showing Piping Specimens and Scanner ....................................5-7

5.7 Close-up of Scanner on Pipe Specimen .......................................................................5-7

5.8 Experimental Set-up Showing Piping Specimen and Scanner......................................5-9

5.9 Field-like Welding on Horizontally Oriented Pipe Segments .......................................5-10

5.10 Examples of Specimens as Viewed from the Outer Surface.......................................5-11

5.11 Examples of Specimens as Viewed from the Inner Surface........................................5-12

6.1 Side and Top Views of the 1.5-MHz, Conventional-UT, Shear-Wave Probe 

and Gimbal/Couplant-Delivery Configuration for Translation of the Ultrasonic 

Transducer Across the Specimens ...............................................................................6-3

6.2 The Low-Frequency/SAFT-UT Ultrasonic Data Acquisition System..............................6-7

6.3 Front-end Transducer Configuration, Including Gimbal and Coupling Lines .................6-7

6.4 Transducer-Wedge Configuration for Dual-Element, Pitch-Catch Scanning .................6-9

6.5 Phased Array Data Acquisition System ......................................................................6-13

6.6 The 2.0 MHz TRL Phased-Array Probe ......................................................................6-14

6.7 The 2.0 MHz TRS Phased-Array Probe......................................................................6-15

6.8 1.5-MHz TRL Probe, Configured with Gimbal and Couplant Lines..............................6-16

6.9 TRL 1.5-MHz Beam Model at 45° with the Side View on the Left and Top 

View on the Right .......................................................................................................6-18

6.10 TRL 2.0-MHz Beam Model at 45° with the Side View on the Left and Top 

View on the Right .......................................................................................................6-18

6.11 TRS 2.0-MHz Beam Model at 45° with the Side View on the Left and Top 

View on the Right .......................................................................................................6-18

6.12 Beam Skewing Model for the 1.5-MHz TRL, 10 by 3 Element Probe ..........................6-19

6.13 Mechanical Scanner Mounted on a Ring Section in the Background with the 

Scanner Arm and Probe Extending to the Foreground and Side View of the 

1.5-MHz TRL Probe Next to a Weld ...........................................................................6-20
x

7.1 C-scan Views for Pipe Segment 1 from the Near and Far Side with the 

1.5-MHz, 60° and 70° Conventional Shear-Wave Probes.............................................7-2

7.2 C-scan Views for Pipe Segment 2 from the Near and Far Side with the 

1.5-MHz, 60° and 70° Conventional Shear-Wave Probes.............................................7-3

7.3 C-scan Views for Pipe Segment 3 from the Near and Far Side with the 

1.5-MHz, 60° and 70° Conventional Shear-Wave Probes.............................................7-4

7.4 Sawcut Lengths Measured Using the 6 dB Drop and Loss of Signal 

Techniques with the 1.5-MHz 60° and 70° Probes from the Near and Far Side 

of the Weld Versus the Actual Sawcut Lengths ............................................................7-7

7.5 Thermal Fatigue Crack Lengths Measured Using the 6 dB Drop and Loss of 

Signal Techniques with the 1.5-MHz 60° and 70° Probes from the Near and 

Far Side of the Weld Versus the Actual Crack Lengths ................................................7-8

7.6 Sawcut and Crack Depths Measured Using the 6 dB Drop and Crack Tip 

Techniques with the 1.5-MHz 60° and 70° Probes from the Near and Far Side 

of the Weld Versus the Actual Sawcut and Crack Depths.............................................7-9

7.7 Signal-to-Noise Ratio Plotted Against the Crack Depth for the 60° and 70° 

Far-Side Scans...........................................................................................................7-10

7.8 C-scan Analysis Process............................................................................................7-13

7.9 B-scan End-View Analysis..........................................................................................7-14

7.10 Excel Spreadsheet and Associated Magnified B-scan End-View for S 

Analysis and Detailed Calculations for Length Sizing and Recording of 

Positional Data ...........................................................................................................7-15

7.11 Typical Analysis Screen with a Low Angle (35°) Selected, Showing the Weld 

Counterbore with the TRL Array .................................................................................7-27

7.12 Typical Analysis Screen with a Mid-Range Angle (53°) Selected, Showing the 

Weld Root with the TRL Array ....................................................................................7-27

7.13 Typical Analysis Screen with a High Angle (65°) Selected, Showing the Flaws 

with the TRL Array......................................................................................................7-28

7.14 B-scan Views for Pipe Segment 1 from the Near and Far Side with the 

Longitudinal- and Shear-Wave Arrays ........................................................................7-29

7.15 B-scan Views for Pipe Segment 2 from the Near and Far Side with the 

Longitudinal- and Shear-Wave Arrays ........................................................................7-30

7.16 B-scan Views for Pipe Segment 3 from the Near and Far Side with the 

Longitudinal- and Shear-Wave Arrays ........................................................................7-31

7.17 Sawcut Lengths Measured Using the 6 dB Drop and Loss of Signal

Techniques with the TRL and TRS Probes from the Near and Far Side of the 

Weld Plotted Versus the Actual Sawcut Lengths ........................................................7-33

7.18 Thermal Fatigue Crack Lengths Measured Using the 6 dB Drop and Loss of 

Signal Techniques with the TRL and TRS Probes from the Near and Far Side 

of the Weld Plotted Versus the Actual Flaw Lengths...................................................7-35
xi

8.1 C-scan Views for a Representative Area of Weld Number One Acquired with 

1.5-MHz 60° and 70° Shear Waves from the Far Side..................................................8-2

8.2 C-scan Views for a Representative Area of Weld Number Two Acquired with 

1.5-MHz, 60° and 70° Shear Waves.............................................................................8-3

8.3 C-scan Views of Crack 4 in Weld Number Two Acquired with 1.5-MHz, 

60° and 70° Shear Waves ............................................................................................8-4

8.4 C-scan Views of Crack 5 in Weld Number Two Acquired with 1.5-MHz, 

60° and 70° Shear Waves ............................................................................................8-4

8.5 C-scan Views of Crack 6 in Weld Number Two Acquired with 1.5-MHz, 

60° and 70° Shear Waves ............................................................................................8-5

8.6 C-scan Views for a Representative Area of Weld Number Three Acquired 

with 1.5-MHz, 60° and 70° Shear Waves from the Far Side..........................................8-6

8.7 C-scan Views for a Representative Area of Weld Number One Acquired with 

400-kHz, 45° S-waves, 45° L-waves, and 60° L-waves from the Far Side ....................8-8

8.8 C-scan Views for a Representative Area of Weld Number One Acquired with 

250-kHz, 45° L-waves and 70° L-waves from the Far Side...........................................8-9

8.9 Same Section of Weld Number One Scanned Using 45° Longitudinal Waves 

at Both 400 kHz and 250 kHz from the Far Side.........................................................8-10

8.10 C-scan Views for a Representative Area of Weld Number Two Acquired with 

400-kHz, 45° S-waves, 45° L-waves, 60° L-waves, and 70° L-waves. The 

10% notches are detected but not the 15% implanted crack.......................................8-11

8.11 C-scan Views for a Representative Area of Weld Number Two Acquired with 

250-kHz, 45° S-waves, 45° L-waves, and 70° L-waves ..............................................8-12

8.12 C-scan Views for a Representative Area of Weld Number Three Acquired 

with 400-kHz, 45° S-waves, 45° L-waves, 60° L-waves, and 70° L-waves..................8-13

8.13 C-scan Views for a Representative Area of Weld Number Three Acquired 

with 250-kHz, 45° L-waves and 70° L-waves..............................................................8-14

8.14 B-scan Views for the Ultrasonic Response for the 10% Notch in a Portion of 

Weld One from the Far Side with the Longitudinal- and Shear-Wave Arrays ..............8-15

8.15 B-scan Views for the Ultrasonic Response from the EDM Notch in Weld Two 

from the Near and Far Side with the Longitudinal-Wave Array ...................................8-17

8.16 A-scan Views for the Ultrasonic Response from the EDM Notch in Weld Two 

from the Near and Far Side with the Longitudinal-Wave Array ...................................8-18

8.17 B-scan Views for the Ultrasonic Response from the EDM Notch in Weld Two 

from the Near and Far Side with the Shear-Wave Array .............................................8-18

8.18 A-scan Views for the Ultrasonic Response from the EDM Notch in Weld Two 

from the Near and Far Side with the Shear-Wave Array .............................................8-19

8.19 B-scan Views for the Ultrasonic Response from the Three Implanted Cracks 

in Weld Two from the Far Side with the Longitudinal-Wave Array...............................8-19

8.20 B-scan Views for the Ultrasonic Response from the Three Implanted Cracks 

in Weld Two from the Far Side with the Shear-Wave Array ........................................8-20
xii

8.21 B-scan Views for the Ultrasonic Response for the 10% Notch in Weld Three 

from the Far Side with the Longitudinal- and Shear-Wave Arrays...............................8-21

10.1 WSS Coupon Used for Sound Field Mapping.............................................................10-2

10.2 Side View Schematic Showing Orientations of PA Probe, Specimen, and 

Pinducer.....................................................................................................................10-3

10.3 Left: Phased Array Probe Mounted and Secured on Austenitic WSS Coupon; 

Right: Coupon Positioning Fixture for Sound Field Mapping ......................................10-4

10.4 Phased-Array Data Acquisition System ......................................................................10-4

10.5 Pinducer Positioned Just Beneath a Coupon End During Beam Mapping ..................10-5

10.6 L-Wave Sound Beam Simulation at 50°, 25-mm Depth Focus; Side View on 

the Left and End View on the Right ............................................................................10-6

10.7 Data Images at 50°, 25-mm Depth Focus, 19 dB, 50 mm Back from WSS 

Side on the Left and from the Weld Side on the Right ................................................10-7

10.8 L-Wave Sound Beam Simulation at 60°, 25-mm Depth Focus; Side View on 

the Left and End View on the Right ............................................................................10-7

10.9 Data Images at 60°, 25-mm Depth Focus, 27 dB, 64 mm Back from WSS 

Side on the Left and from the Weld Side on the Right ................................................10-8

10.10 L-Wave Sound Beam Simulation at 70°, 15-mm Depth Focus; Side View on 

the Left and End View on the Right ............................................................................10-8

10.11 Data Images at 70°, 15-mm Depth Focus, 19 dB, 57 mm Back from WSS 

Side on the Left and from the Weld Side on the Right ................................................10-9

10.12 Sound-Field Sizing Comparison at a −6 dB Level Showing WSS Results with 

Black Ovals and WSS + Weld Results with Magenta Ovals......................................10-11

10.13 Schematic Diagram of Wrought Stainless Steel – Austenitic Stainless Steel 

Composite Media .....................................................................................................10-12

11.1 Example of a ―Yes‖ Call..............................................................................................11-2

11.2 Example of a ―Marginal‖ Call ......................................................................................11-2

11.3 Example of a ―No‖ Call ...............................................................................................11-3

11.4 Detection Calls on the Secure and Practice Data Sets as a Function of PA 

Probe and Near- and Far-Side Inspection ..................................................................11-4

11.5 Signal-to-Noise Ratios for ―Yes‖ Detected Cracks ......................................................11-5

11.6 Cross-Sectional Schematic Showing the Weld Geometry and Location of an 

Implanted Crack in the Weld Heat Affected Zone .......................................................11-6

11.7 Far Side Data from a Specimen with the Cross Section Shown in Figure 11.6 ...........11-7

11.8 Data Images Showing a Flaw Signal Separated from Geometrical Indications ...........11-7

12.1 Far-Side Detection Results for All Techniques Applied on All Thermal Fatigue 

Cracks and Machined Sawcuts and Notches as a Function of Through-Wall 

Flaw Size....................................................................................................................12-2

12.2 Images of Flaws Comparing Results from Conventional-UT, Low-Frequency/ 

SAFT and Phased Array-UT Methods ........................................................................12-3
xiii

Tables

5.1 Implanted Thermal Fatigue Cracks in Austenitic Piping Specimen 02-24-15 ................5-4

5.2 Sawcuts in Austenitic Piping Specimen 02-24-15 .........................................................5-5

5.3 Thermal Fatigue Cracks in Austenitic Piping Specimen 3C-022 .................................5-10

6.1 Final Combination of Scan-Angle-Frequency-Modality for Phase 1 Evaluation 

Using the LF/SAFT Inspection Technique ..................................................................6-11

7.1 Sawcut Lengths in Pipe Specimen 02-24-15 as Measured by the 60° 

Conventional UT Technique .........................................................................................7-5

7.2 Sawcut Lengths in Pipe Specimen 02-24-15 as Measured by the 70° 

Conventional UT Technique .........................................................................................7-5

7.3 Implanted Thermal Fatigue Crack Lengths in Pipe Specimen 02-24-15 as 

Measured by the 60° Conventional UT Technique........................................................7-6

7.4 Implanted Thermal Fatigue Crack Lengths in Pipe Specimen 02-24-15 as 

Measured by the 70° Conventional UT Technique........................................................7-6

7.5 Actual Versus Measured Values for Detection Results of Sawcut F from a 

Tutorial-Based Analysis of the 400-kHz, 45° Shear-Wave Inspection of Piping 

Segment #2, from the Near Side ................................................................................7-16

7.6 250-kHz, Far-Side Detection Results for Piping Segment 1........................................7-18

7.7 400-kHz, Far-Side Detection Results for Piping Segment 1........................................7-18

7.8 250-kHz, Far-Side Detection Results for Piping Segment 2........................................7-21

7.9 400-kHz, Far-Side Detection Results for Piping Segment 2........................................7-21

7.10 250-kHz, Far-Side Detection Results for Piping Segment 3........................................7-24

7.11 400-kHz, Far-Side Detection Results for Piping Segment 3........................................7-24

7.12 Sawcut Lengths in Pipe Specimen 02-24-15 as Measured by the TRL Array .............7-32

7.13 Sawcut Lengths in Pipe Specimen 02-24-15 as Measured by the TRS Array.............7-32

7.14 Thermal Fatigue Crack Lengths in Pipe Specimen 02-24-15 as Measured by 

the TRL Array.............................................................................................................7-34

7.15 Thermal Fatigue Crack Lengths in Pipe Specimen 02-24-15 as Measured by 

the TRS Array.............................................................................................................7-34

8.1 Thermal Fatigue Crack Lengths in Weld Two of Pipe Specimen 3C-022 as 

Measured by the 60° and 70° Conventional-UT Probes................................................8-5

8.2 Thermal Fatigue Crack Lengths in Weld Two of Pipe Specimen 3C-022 as 

Measured by the TRL and TRS Arrays .......................................................................8-20

9.1 Summary of Detection Results for the 250-kHz Low-Frequency/SAFT 

Inspection Technique ...................................................................................................9-3

9.2 Summary of Detection Results for the 400-kHz Low-Frequency/SAFT 

Inspection Technique ...................................................................................................9-3
xiv

9.3 Summary of Length-Sizing Results for the 250-kHz Low-Frequency/SAFT 

Inspection Technique Using the −6 dB Drop Method....................................................9-5

9.4 Summary of Length-Sizing Results for the 250-kHz Low-Frequency/SAFT 

Inspection Technique Using the Loss-of-Signal Method ...............................................9-5

9.5 Summary of Length-Sizing Results for the 400-kHz Low-Frequency/SAFT 

Inspection Technique Using the −6 dB Drop Method....................................................9-6

9.6 Summary of Length-Sizing Results for the 400-kHz Low-Frequency/SAFT 

Inspection Technique Using the Loss-of-Signal Method ...............................................9-7

9.7 Gain (in dB) Settings Used for Scanning to Reach Near-Saturation for 10% 

Notch............................................................................................................................9-8

10.1 Sound Field Mapping Parameter Matrix for Data Acquired from Both Ends of 

the Coupon.................................................................................................................10-2

10.2 Simulated Beam Size .................................................................................................10-9

10.3 Wrought Stainless Steel Coupon End, Longitudinal-Wave Data .................................10-9

10.4 Coarse-Grain (Weld Side) Coupon End, Longitudinal-Wave Data ............................10-10

10.5 Comparison of Peak Signal Response from the Weld to the Base Material ..............10-11

10.6 Material Properties of Base and Weld Metal.............................................................10-13

10.7 Summary of Computational Results .........................................................................10-14

11.1 IGSCC Detection Summary........................................................................................11-4

12.1 Far-Side Length-Sizing Comparison for All Methods Applied......................................12-2
xv

Executive Summary

Research is being conducted for the U.S. Nuclear Regulatory Commission (NRC) at the Pacific 

Northwest National Laboratory (PNNL) to assess the effectiveness and reliability of advanced 

nondestructive examination (NDE) methods for the inspection of light-water reactor components. The scope of this research encompasses many primary system pressure boundary 

materials including cast austenitic stainless steels (CASS); dissimilar metal welds; piping with 

corrosion-resistant cladding; weld overlays, inlays and onlays; and examinations of wrought 

austenitic piping welds. One objective of this work is to evaluate various ultrasonic testing (UT)

methods to assess their ability to detect, localize, and size cracks through coarse-grained 

materials.

This report summarizes an assessment of the capabilities and limitations of various ultrasonic 

techniques for far-side examinations of austenitic stainless steel welds. Periodic inservice 

inspection of nuclear power plants is intended to detect service-related degradation that may 

adversely impact structural integrity, or impair a component’s ability to function as designed. 

Section XI of the American Society of Mechanical Engineers (ASME) Boiler and Pressure 

Vessel Code, entitled Rules for Inservice Inspection of Nuclear Power Plant Components, 

requires that volumetric examinations be performed for many safety-related components, 

including those that comprise the primary coolant pressure boundary. The volumetric 

examination most often applied is ultrasonic testing. Because of the anisotropic microstructural

features found in austenitic coarse-grained materials, UT of stainless steel components 

continues to be a challenge to the industry. This is especially true for detection and 

characterization of service-induced flaws, including stress corrosion cracking (SCC), that may 

initiate on the far side of welds where access for examination may be limited to the near side of 

the weld.

PNNL conducted ultrasonic studies in a three-phased approach, to assess far-side inspection in 

specimens containing implanted thermal fatigue cracks and machined notches. From the first 

two phases of work, results were favorable in that nearly all of the flaws were detected. 

Reported here are the results of this initial work, including conventional, phased-array, and lowfrequency/Synthetic Aperture Focusing Technique (SAFT) ultrasonic techniques. The data 

show that with the use of precise acquisition techniques and rigorous analytical methods, farside detection and length-sizing for surface-breaking circumferential flaws is quite possible in 

austenitic piping welds. However, no flaw-tip diffracted signals could be observed from the far

side of the weld; therefore, time-of-flight through-wall depth sizing may not be feasible.

Efforts in Phases 1 and 2 show that, while variation in peak amplitudes occur randomly along 

the length of shop and field welds, no regional effects in acoustic responses from flaws on the 

far side of welds were observed. In other words, no relationship between circumferential weld 

microstructures (due to varied welding positions) and acoustic transmission quality of the weld 

could be ascertained. These results are shown to be valid for implanted thermal fatigue cracks 

and machined reflectors, with both exhibiting moderately planer shapes. No other serviceinduced flaws, for example, inter-granular or primary water stress corrosion cracking, were 

available during Phase 1 and Phase 2 laboratory trials. 
xvi

Of the UT technologies tested through Phase 2, the ultrasonic phased-array method provided 

the best overall results with the highest number of far-side flaws being detected, and lengthsizing capability well within ASME Code-established acceptance criteria. The conventional UT 

method also showed good detection and sizing effectiveness, but this is primarily due to the 

automated data acquisition and extensive analyses performed; it is not expected that manual 

UT examinations with real-time interpretation would perform at this level. The low-frequency/

SAFT method explored the aspects of using very long wavelengths over a range of frequencies 

from 250–400 kHz (for insensitivity to coarse-grained weld structures) and their impact on flaw 

detection performance and sizing accuracy. Reduced performance was noted for this 

application. The results indicate that one must determine the highest frequency that can 

effectively penetrate the coarse-grained microstructure with minimal distortion and use this 

frequency for the inspection, coupled with signal processing such as SAFT.

To better understand acoustic propagation through austenitic welds using phased-array ultrasonic methods, sound-field mapping was conducted, and results were used to guide improved 

designs for more effective phased-array probes to address this inspection problem. The 

longitudinal (compressional) wave modality showed little degradation as the sound field propagated through the weld material from the far side. The sound-field maps provided in this report

clearly show that the longitudinal wave mode provides suitable penetration and the sound field 

maintains ample coherence and signal strength as it propagates through the austenitic weld. 

Only at the highest angles of insonification (70°) coupled with a shallow focal depth (15 mm) do 

the longitudinal waves start to exhibit slight degradation to the field shape, coherence, and 

structure. Beam partitioning and scattering of the compressional-wave energy is not evident as 

the L-wave sound fields appear to be relatively insensitive to the effects of the austenitic weld 

material at the frequency evaluated. The results from this evaluation show that use of 

longitudinal waves, coupled with the appropriate frequency (wavelength) relative to the size of 

the grain boundaries and weld structure, can provide the foundation for an effective inspection.

The initial activities were followed by a Phase 3 effort to employ ultrasonic phased-array 

inspection on specimens containing service-induced inter-granular stress corrosion cracking 

(IGSCC) at the Electric Power Research Institute (EPRI) NDE Center, in Charlotte, 

North Carolina. For the EPRI exercise, phased-array data were acquired on a series of 

specimens containing service-induced IGSCCs. True-state location information for all flaws 

was provided during the data analysis phase. The crack lengths were variable and PNNL only 

examined the specimens for circumferentially oriented cracks (cracks that are parallel to the 

weld). Flaw depth information was not provided and, once again, flaw-tip signals were not 

detected, so no estimate of flaw depth was possible.

The investigation was focused on assessing current phased-array ultrasonic techniques for 

austenitic welds. Detection results should be considered as a best-effort, given equipment and 

inspection conditions; that is, the arrays were not specifically optimized for the weld conditions 

and the weld crowns were not removed, which would have provided greater access for each of 

the phased-array modalities and probes applied. However, as a result of the sound-field 

mapping activities that occurred at the end of Phase 2, a 1.5-MHz transmit-receive longitudinal 

probe was designed for application to far-side examinations, in an attempt to obtain a balance 

between effective penetration through the austenitic weld and suitable resolution for flaw 

detection.
xvii

Because the true-state location data were known for the flaws under examination, the results 

described here should not be considered representative of a blind test, only an assessment of 

state-of-art phased-array capabilities to detect actual service-induced IGSCC. Data were 

acquired with two refracted longitudinal wave phased-array probes, 1.5 MHz and 2 MHz, and 

one shear-wave probe at 2 MHz from both the near and far side of the weld center relative to 

the flaw. The far-side detection rate for longitudinal mode inspections was nearly 70% while the 

detection rate for shear mode inspections was 52%. Signal-to-noise values for the detected 

flaws were good. Shear wave sound fields are significantly partitioned and attenuated (primarily 

by scattering) as they propagate through the austenitic weld material, while the longitudinal 

wave sound fields generally maintain coherence with reduced scattering. Both shear and 

longitudinal data tended to be cluttered with geometrical and material reflections making signal 

discrimination the primary challenge in flaw detection. Flaw detection would be improved if the 

weld crowns were removed because it would provide better access to the far-side region of 

interest and enable the use of probes with larger footprints; thus allowing large acoustic 

apertures for better sound field generation. The use of lower frequencies (longer wavelengths) 

may provide better penetration and reduce scattering effects, while more effective approaches 

for peaking signals may improve signal discrimination. The improvements cited above coupled 

with more accurate profiling of the complex part geometry in the critical area of examination, 

could collectively enhance the detection results from what is reported here.

Several insights were gained from the subject study that could be used by industry as it moves 

forward to improve far-side inspections, such as further modeling of the inspection process 

(weld microstructures, acoustics, etc.) to facilitate the design of improved search units, 

additional laboratory measurements on IGSCC and axial-oriented flaws, and metallographic 

studies to help interpret and correlate weld microstructures with the findings of the modeling and 

measurement activities. These activities would strengthen the technical basis for performancedemonstration requirements.

xix

Acknowledgments

The work reported here spans activities conducted over several years and was sponsored by 

the U.S. Nuclear Regulatory Commission (NRC). This work was conducted under two NRC Job 

Code Numbers, initially starting under Y6604 and transitioning to N6398. Deborah Jackson was 

the NRC project manager for the Y6604 work and Wallace Norris is the current NRC project

manager for work conducted under N6398. The Pacific Northwest National Laboratory (PNNL) 

would like to thank Ms. Jackson and Mr. Norris for their involvement and guidance, technical 

direction, and attention to detail throughout the course of this effort. PNNL also thanks

Mr. Jeff Hixon (NRC intern that rotated to PNNL) for his participation and technical support 

throughout the initial phase of the study. The work reported here was conducted at PNNL in 

Richland, Washington, and PNNL staff appreciated the opportunity to work side-by-side with 

NRC personnel. We also thank Mr. Jim Davis (NRC) for his guidance and support of this work.

In addition, the authors express their sincere gratitude to Guy Maes and Patrick Tremblay at 

ZETEC, Inc. for their willingness and ongoing engineering support with phased-array system 

functional and operational issues, as well as their guidance and input regarding phased-array 

probe design and manufacture.

PNNL would like to thank several personnel at the Electric Power Research Institute (EPRI)

NDE Center for enabling the work reported here. We appreciate Greg Selby’s original invitation 

to collect and review IGSCC data on the practice specimens, and for making his staff available 

to PNNL during this activity. During the data collection and analyses, Carl Latiolais, Doug Kull, 

John Langevin, and Lief Esp provided their support to assist PNNL with set up and true-state 

review. Mark Dennis facilitated the final contract to allow PNNL to review secure data. We also 

appreciate the comments provided by EPRI staff on the previous draft sections of this report.

At PNNL, the authors wish to express their gratitude to Mr. Stan Owsley for designing the 

laboratory platform and peripheral mechanics for effective configuration of the piping segments 

for laboratory examination. The authors would like to thank Marino Morra for providing initial 

support to this effort in coordinating the cutting and machining of pipe segments for specimen 

fabrication. Appreciation is given Mr. Jake Tucker, Mr. Eric S. Andersen, and Mr. George 

Schuster for their technical support, advice, and assistance on this project. The authors are 

also grateful to Lori Bisping, Angie Dickson, and Earlene Prickett for their efforts in supporting 

program procurements, contractual issues, financial tracking, and various other administrative 

duties pertinent to this effort. Finally, the PNNL technical team would like to extend their thanks 

to Kay Hass for her ongoing support, attention to detail, and technical editing expertise in 

preparing and finalizing this document. 

PNNL is operated by Battelle for the U.S. Department of Energy under Contract DE-AC05-

76RL01830.

xxi

Acronyms and Abbreviations

A/D analog-to-digital

ASME American Society of Mechanical Engineers

ASTM American Society for Testing Materials

BWR boiling water reactor

CASS cast austenitic stainless steel

CD compact disc

CSS cast stainless steel

dB decibels

DMW dissimilar metal weld

DVD digital video disc

EDM electro-discharge machined

EPRI Electric Power Research Institute

FSH full screen height

HAZ heat affected zone

ID inner diameter

IGSCC intergranular stress corrosion crack

ISI inservice inspection

LF/SAFT low-frequency/Synthetic Aperture Focusing Technique

LOS loss of signal

LWR light water reactor

NDE nondestructive examination

NPP nuclear power plant

NRC U.S. Nuclear Regulatory Commission

NRR Nuclear Reactor Regulation (Office of)

OD outer diameter

PA phased array

PA-UT phased array ultrasonic testing

PC personal computer

PD performance demonstration

PDI performance demonstration initiative

PNNL Pacific Northwest National Laboratory

PWSCC pressurized water stress corrosion crack

RF radio frequency

RMSE root mean square error

SAA signal amplitude anomaly

SAFT synthetic aperture focusing technique
xxii

SNR signal-to-noise ratio

SS stainless steel

TFC thermal fatigue crack

TRL transmit-receive longitudinal

TRS transmit-receive shear

UT ultrasonic testing

WSS wrought stainless steel

WCL weld centerline
1-1

1 Introduction

The overall objective of the work at PNNL is to determine the effectiveness and reliability of 

ultrasonic inspection techniques on light water reactor (LWR) components containing cast 

stainless steel material and other coarse-grained components with dissimilar metal welds, 

piping with corrosion-resistant cladding, and examinations of austenitic piping welds. This 

report summarizes the technical evaluations performed to address a limited portion of the 

objective, specifically aimed at assessing the capabilities and limitations of ultrasonic techniques 

for examining flaws from the far side of austenitic stainless steel (SS) welds. 

Throughout many industrial sectors, efforts are underway to evaluate the potential for new 

nondestructive examination (NDE) methods to reliably detect structural degradation in aging 

components while controlling costs and limiting inspection times. In particular, the nuclear 

industry, through the Electric Power Research Institute (EPRI), is actively involved in 

determining the feasibility of applying new NDE methods for challenging inspections at 

operating nuclear power plants. The purpose of this report is to provide the results of laboratory 

investigations at the Pacific Northwest National Laboratory (PNNL) to determine capabilities of 

ultrasonic phased-array (PA), low-frequency (LF)/Synthetic Aperture Focusing Technique 

(SAFT), and conventional technologies as applied to the far-side inspection of wrought 

austenitic piping welds. 

Austenitic stainless steels are generally used for applications where resistance to corrosion, or 

high strength and creep resistance, are required at elevated temperatures, such as in the 

primary piping of a light water reactor (LWR). Ultrasonic techniques have routinely been used in 

various industrial applications for over 50 years and yet, cast or welded austenitic components 

remain difficult to reliably and effectively examine. Over the past 20 or so years, significant 

progress has been made towards improving ultrasonic inspection techniques. Typically, 

conventional ultrasonic Testing (UT) methods are used for inservice inspection (ISI), but these

methods are often times hampered by the anisotropic, large-grained texture of the austenitic 

weld metal.

ISI requirements dictate that piping welds in the pressure boundaries of LWRs are subject to 

volumetric examination based on the requirements of Section XI of the American Society of 

Mechanical Engineers (ASME) Boiler and Pressure Vessel Code. The volumetric examination 

may be either radiographic or ultrasonic. For inservice examinations at operating reactors, 

background radiation and access limitations may preclude the use of radiography. Hence, 

volumetric examination of austenitic welds in selected LWR components are generally 

performed using ultrasonic examination. The purpose of ultrasonic testing/inservice inspection 

(UT/ISI) of piping and pressure vessels is to reliably detect and characterize material 

degradation that may develop during operation. The detection of potential degradation typically 

results from analyzing ultrasonic echo waveforms scattered from inherent material artifacts. 

Austenitic weld material exhibits a macro-crystalline structure that is established during the 

solidification process. Typically, the austenite phase forms long columnar grains based on the 

weld bead size and direction of maximum heat loss during solidification and cooling processes 

(Moysan et al. 2003). Due to the physical properties of these grains, many inspection problems 
1-2

can arise, and ultrasonic flaw responses are often obscured by noise and scattered energy from 

the material structure.

Austenitic welds exist in many LWR piping systems. A large number of welds are the terminal 

ends of piping systems and consist of pipe-to-component welds. Access to each side of the 

weld may be limited, so a need exists to develop effective and reliable far-side inspection 

techniques for these components. Far-side austenitic weld inspection procedures continue to 

perform unsatisfactorily due to the coarse microstructure that characterizes these materials. 

The large size and orientation of the anisotropic grains, relative to the acoustic pulse 

wavelength, strongly affects the propagation of ultrasound by causing severe attenuation, beam 

skewing, changes in velocity, and scattering of ultrasonic energy (Macdonald et al. 2000). 

Refraction and reflection of the sound beam occur at the grain boundaries, from root conditions, 

counterbore, and at the weld fusion lines, resulting in defects being incorrectly characterized, 

specific volumes of material not being examined, or both. When coherent reflection and 

scattering of the sound beam occur at grain boundaries or from weld geometry, ultrasonic 

indications occur that are difficult to distinguish from signals originating from flaws. Conducting 

far-side inspections of austenitic welds, where signal-to-noise ratios are relatively low, can be 

confusing, unpredictable, and unreliable during conventional ultrasonic examinations.

PNNL conducted ultrasonic studies in a 3-phased approach. In Phases 1 and 2, technical 

assessments were conducted for far-side inspection in specimens containing implanted thermal 

fatigue cracks and machined notches in welds fabricated under variable conditions. In Phase 3, 

the focus of the technical evaluation was aimed at examining field-removed piping welds 

containing service-induced inter-granular stress corrosion cracking (IGSCC). Initially, PNNL 

chose to evaluate two advanced ultrasonic inspection approaches, and compare the 

performance of these techniques against conventional UT methods. The phased-array 

approach was selected because industry is evaluating this technology for several applications 

and anecdotal information indicates that there appear to be ―windows‖ where effective 

inspections can be made through coarse-grained microstructures. It is not known if these 

―windows‖ are related to inspection frequency, inspection angle, or some other microstructural 

feature. The use of phased-array technology provides the capability to rapidly assess the effect 

of multiple inspection angles. The LF/SAFT approach was also chosen because of previous 

success on cast stainless steel (CSS) and the fact that 400-kHz wavelengths are much larger 

than the austenitic weld microstructure, which should minimize the impact of the microstructure 

on the sound field. At the same time, PNNL also documented the performance of a 

conventional, or monolithic element, pulse-echo inspection method as employed to pass the 

austenitic weld performance demonstration administered at the EPRI NDE Center and 

developed for ISI, in order to baseline the standard inspection for comparison with these 

advanced techniques.

The focus of this NUREG report targets far-side examination assessments conducted at PNNL 

over several years, in support of, and sponsored by, the United States Nuclear Regulatory 

Commission (NRC). Much information was gathered from lessons learned in detection and 

characterization of flaws during Phases 1 and 2 (Anderson et al. 2003; Anderson et al. 2006; 

Anderson et al. 2008). From these assessments, efforts were conducted to better understand 

sound-field propagation through austenitic welds. These activities focused on modeling and 

simulation efforts coupled with sound-field mapping exercises to determine austenitic weld 
1-3

effects (coherence/distortion, shape, etc.) on sound fields as a function of frequency. These 

laboratory studies were followed by a Phase 3 evaluation aimed at employing optimized phased 

array-UT methods for examination of IGSCC in field-removed specimens. This NUREG 

provides a summary of a multi-phased laboratory effort, and in Section 2, the historical 

background of these efforts is provided, with specific technical objectives of these assessments 

being defined. Section 3 describes the scope of these evaluations, and Section 4 provides a 

summary of the inspection challenges encountered in examining wrought stainless steel (WSS) 

components from the far-side, and through austenitic welds. Section 5 provides all of the 

technical details associated with the specimens evaluated in this study, while Section 6 

describes the ultrasonic probes used to employ conventional, LF/SAFT and phased-array UT 

techniques. Sections 7 and 8 summarize the data and results of Phase 1 and Phase 2 

evaluations, respectively, for all three ultrasonic inspection techniques. Section 9 provides a 

technical discussion of the results obtained from the work in Phases 1 and 2. Section 10 

describes the longitudinal-wave sound field assessments and simulations work. Section 11 

provides a summary of the data and results from the Phase 3 evaluation at the EPRI NDE 

Center, while Section 12 provides the summary and conclusions from all work performed at 

PNNL on far-side examination effectiveness. Section 13 provides a listing of references cited in 

this report. 

2-1

2 Background and Objectives

The United States Nuclear Regulatory Commission has sponsored PNNL to conduct a project 

titled ―Reliability of Nondestructive Examination for Nuclear Power Plant Inservice Inspection.‖ 

The objectives of the project are to (1) evaluate the accuracy and reliability of nondestructive 

examination methods used for the inservice inspection of nuclear power plant (NPP) systems 

and components; (2) provide information for the Office of Nuclear Reactor Regulation (NRR) to 

assess the adequacy of proposed industry changes to ISI programs; (3) provide recommendations to the staff of the NRC to improve the effectiveness and adequacy of ISI programs; 

(4) evaluate the effectiveness of ISI techniques in detecting cracks in reactor coolant system 

components (e.g., dissimilar metal welds [DMWs], Alloy 52 weld overlays); and (5) provide 

technical assistance on NDE and related issues to the NRC program offices on an as-needed 

basis.

The results obtained from this project will assist the NRC by supporting their technical reviews 

and will serve as confirmatory results for determining the effectiveness and adequacy of NDE 

and ISI methods and programs presently used in operating nuclear power plants. Recent 

experience with cracking in reactor vessel penetrations and piping welds has increased reliance 

on the ability of NDE methods to detect and characterize service-induced flaws over a range of 

materials, geometries, and configurations. The ability to detect and characterize flaws varies 

with the selected NDE method. With the industry trend toward reducing inspection time, 

radiation exposure, and the number of examinations, the effectiveness and reliability of NDE 

methods becomes more important.

The research described in this report has been organized into three phases. The goal of 

Phase 1 was to test the capabilities of advanced ultrasonic technology on a limited set of 

welded specimens, and evaluate the ability of these UT techniques to characterize machined 

reflectors and fabricated cracks, while establishing a baseline for comparison against the results 

of conventional UT methods. The pipe samples were examined from several directions, both 

with the weld crown in place and with the weld crown machined off. In Phase 2, the information 

gained in Phase 1 was used to evaluate more challenging conditions and assess any affects 

presented by variations in welding position or procedure. Finally, in Phase 3, lessons learned 

from the first two phases were incorporated to provide an improved ultrasonic inspection 

protocol for examination of service-induced IGSCC using the most promising technique(s)

evaluated in earlier phases. 

Phase 1 and 2 evaluation efforts were conducted at PNNL and involved the application of 

conventional transducers, and use of LF/SAFT and ultrasonic phased-array methods on 

specimens containing implanted thermal fatigue cracks and machined reflectors. From these 

studies, it was concluded that PA inspection provided the best results, detecting nearly all of the 

flaws from the far side of the welds. These results were presented at the Fifth International 

Conference on NDE in Relation to Structural Integrity for Nuclear and Pressurised Components

in 2006. This led to an invitation to examine field-removed specimens containing serviceinduced IGSCC at the EPRI NDE Center, in Charlotte, North Carolina.

3-1

3 Scope of Ultrasonic NDE Assessments

The scope of the work reported here was defined by NRC guidance and is focused toward 

evaluating UT inspection capabilities in wrought stainless steel (WSS) components with 

austenitic welds, as examined from the far-side. PNNL was tasked with acquisition and 

procurement of materials, fabrication of flawed specimens, design and procurement of 

specialized ultrasonic probes, and conducting data acquisition and analyses on these 

specimens. The scope of these activities was maintained over the course of evaluating the 

flawed specimens, with a focus on replicating materials used in the primary coolant headers in 

boiling water reactor (BWR) recirculation systems. In the first two phases of the work, vintage 

material was used in the fabrication of specimens; in particular, Schedule 80, American Society 

for Testing Materials (ASTM) 358 Grade 304 WSS, with nominal 610-mm (24.0-in.) outer 

diameter, and a wall thickness of approximately 36 mm (1.4 in.). Also, Phase 1 and 2 

specimens were welded together using single-Vee girth welds and filler materials typical to 

nuclear applications such as ASME SA-240, Type 304 and E308-L. In Phase 3 of the 

assessment, the specimens examined at the EPRI NDE Center were field-removed piping 

segments removed from several BWR primary recirculation systems containing service-induced 

IGSCC.

The technical evaluation reported here included results and analyses of data acquired using 

three primary ultrasonic techniques:

 Conventional Ultrasonic Inspection 

 Low-frequency/SAFT-UT

 Phased-Array Ultrasonic Inspection

These examination approaches were used to evaluate inspection performance on a set of 

implanted/fabricated thermal fatigue cracks (TFCs), machined reflectors, and service-induced 

IGSCC, to evaluate detection, localization, and sizing performance as a function of frequency 

and other critical inspection parameters. The work consisted of design and acquisition of 

sound-field simulations and mapping assessments for various LF/SAFT and phased-array

probes, probe configurations, and wedges; development of specific PA focal laws; and use of 

the latest phased-array UT (PA-UT) data acquisition system (Tomoscan III) for implementing 

examinations. The experimental parameters of the work reported here include: 

 Varied inspection frequencies

– 1.5 MHz to 2.25 MHz (Conventional UT)

– 250 kHz to 400 kHz (LF/SAFT UT)

– 1.5 MHz to 2.0 MHz (Phased-Array UT)

 Varied inspection angles

– 60° and 70° (Conventional UT)

– 30° to 70° in 5° increments (LF/SAFT UT)

– 30° to 70° in 1° increments (Phased-Array UT)
3-2

 Varied inspection modalities

– Longitudinal and shear-wave modes (LF/SAFT and PA-UT)

– Shear-wave mode only for Conventional-UT

 Varied orientations to the weld

– Examinations were documented from both near- and far-sides of the welds

During data analysis of the piping specimens evaluated here, ultrasonic flaw detection and 

sizing performance were quantified against the true state, and assessments of the UT 

capabilities for flaw localization/positioning versus true state were conducted. In addition, an 

assessment of sound-field continuity and signal strength (baseline material noise analysis) 

through the austenitic weld and WSS parent material was performed. Metrics used for the 

analyses included signal-to-noise ratio (SNR); root-mean-square error (RMSE); and true-state 

values for flaw length, depth, and axial and circumferential locations relative to the fullpenetration piping welds.

Finally, the scope of this effort also entailed modeling, simulation, and laboratory measurements 

for sound-field mapping, to better understand acoustic propagation through the austenitic welds, 

and to quantify spatial sound-field characteristics through the weld. These activities focused on 

such variables as incident angle, position from the weld centerline, and frequency, using the 

longitudinal (compressional) wave modality.
4-1

4 The Far-Side, Austenitic Weld Inspection Challenge

Austenitic steels are generally used for applications where resistance to corrosion, or high 

strength and creep resistance, are required at elevated temperatures, such as in the primary 

piping of an LWR. ASME Code Section XI requires volumetric examinations of these welds in 

the pressure boundary of LWRs. These examinations are either radiographic or ultrasonic. 

Often, background radiation and access limitations prevent the use of radiography, thus 

ultrasonic inspection must be employed. Conventional ultrasonic approaches are typically used 

for ISI; however, many austenitic welds are pipe-to-component welds where outer-diameter 

(OD) access to either side of the weld may be limited. In many cases, these access limitations 

often preclude near-side access, thereby requiring a far-side inspection.

Ultrasonic techniques have been routinely used in industry for over 50 years and, yet, cast or 

welded austenitic components can be difficult to reliably and effectively examine using 

ultrasound. Over the past 20+ years, significant progress has been made to improve inspection 

techniques. Assuring the structural integrity of reactor components requires nondestructive 

examinations during fabrication, and appropriate inservice examinations, to detect the presence 

of potentially significant fabrication defects or service-related degradation. Successful 

volumetric examination for selected components is predicated upon the use of satisfactory 

austenitic welding procedures; however, some of the most common arc-welding processes

often exacerbate the inspection problem. Typically, conventional ultrasonic methods are often 

hampered by the anisotropic, textured structure of the austenitic weld. Figure 4.1 illustrates a

typical austenitic weld microstructure.

Far-side austenitic weld inspection procedures continue to perform unsatisfactorily due to the 

coarse microstructure that characterizes these materials (Macdonald et al. 2000). These weld 

inspections continue to yield poor results due to the large size of the dendrites and the 

orientation of the anisotropic grains (relative to the acoustic pulse wavelength). Austenitic weld 

material exhibits a macro-crystalline structure that establishes itself during the solidification 

process. Typically, the austenitic phase forms long columnar grains that grow along the 

directions of maximum heat loss during the cooling process (Moysan et al. 2003). Due to the 

physical properties of these grains, many inspection problems can arise, and ultrasonic flaw 

responses are often obscured by noise and scattered energy from the material structure. At the 

frequencies required to resolve reflections (i.e., detect) from fatigue cracking and IGSCC on the 

far side of a weld, these conditions lead to severe attenuation (primarily scattering), sound beam 

re-direction and/or partitioning, changes in acoustic velocity as a function of spatial position on 

the component, and refraction/reflection of sound energy at grain boundaries, root conditions, 

counterbore, and weld fusion lines. The complexity inherent to this challenging inspection 

scenario translates into a lower signal-to-noise ratio (SNR), difficulties in signal (echo) 

discrimination, and the potential for incomplete insonification of the component. 
4-2

Figure 4.1 Polished and Chemically Etched Side View of an Austenitic Weld 

Microstructure in Wrought Stainless Steel

Because austenitic welds are used in many nuclear reactor piping systems and access may be 

limited, there exists a need to develop effective and reliable far-side inspection techniques for 

varied components. 
5-1

5 Specimens Evaluated in this Study

5.1 Phase 1: Austenitic Piping Specimen No. 02-24-15

The specimens used in Phase 1 were made with ASTM A-358, Grade 304 vintage austenitic SS

piping segments having nominal 610-mm (24.0-in.) outside diameter and approximately 36-mm 

(1.4-in.) wall thickness. This type and size of piping is typical of that used during original 

fabrication of primary coolant headers in BWR recirculation systems. The specimen was made 

by joining two cylinders of this material with a single-Vee girth weld, using procedures and filler 

material (ASME SA-240, Type 304) typical of those used in nuclear applications. The pipe was 

shop-welded with air backing in the horizontal position. The pipe was rotated during the welding 

process to allow the welder to always remain in the most advantageous position above the 

weld, in the 1G ―down-hand‖ orientation.

During the welding process several thermal fatigue cracks were implanted into the heat-affected 

zone (HAZ) of the girth weld. The TFCs were generated by thermal cycling of a tension 

specimen to produce fatigue cracks of a specific size, removing the crack coupon from the 

tension bar, then in-situ fusing these coupons during the girth-welding process. This technique 

allows the crack face to be characterized via precise mechanical measurements and 

photographed prior to implantation, and has been shown to produce cracks in the as-welded 

condition within ± 1.0 mm (0.04 in.) of the specified length and height in the piping material. 

Also, this technique enables the vendor (FlawTech) to produce various geometrical and welding 

conditions that simulate those found during field welding; for example, counterbore, innerdiameter miss-match, weld root, fusion anomalies, and weld crown configurations. Vendorsupplied photographs of the piping material, welding and flaw implantation processes, and 

implanted flaw configurations are included for review in Appendix A. The implanted TFCs have 

crack opening dimensions ranging from 25–250 μm (1–10 mils). Figure 5.1 shows where the 

flaws (TFCs and additional sawcuts) were placed and how the pipe was sectioned. Figures 5.2

and 5.3 depict idealized cross sections of the specimens, and give reference dimensions for 

implanted crack characteristics. The TFC dimensions and locations for those used in this study 

are shown in Table 5.1.

It has been experimentally determined (Schuster et al. 2010) that cracks implanted into a 

sample in this way have an ultrasonic response that is typically 5–6 decibels (dB) lower than a 

notch of the same depth using 45° shear waves. These cracks most closely resemble lowcycle, high-stress mechanical fatigue cracks. The cracks are straight, with no branching, have 

mid-sized to large crack opening dimensions, and have a high degree of surface roughness.

This is in contrast to high-cycle, low-load mechanical fatigue cracks that are much tighter and 

have less surface roughness. Service-induced thermal fatigue cracks are usually tighter than 

the cracks used in this study, possibly nonlinear in shape, occasionally branched, and thus 

more difficult to detect using ultrasound than the implanted cracks. IGSCC, which is a likely 

failure mechanism for austenitic SS, are highly variable in nature, but are often highly branched 

and tight, resulting in an increased level of difficulty for detection and more challenging with 

regard to acquiring detectable tip responses. Depending on the crack morphology, an IGSCC 
5-2

can be more difficult (Lemaitre 1994; Gruber et al. 1995) or comparable (Heasler and Doctor 

1996) to detect using ultrasound than a thermal fatigue crack of the same size.

Figure 5.1 Sectioning and Flaw Locations for Pipe Specimen 02-24-15
5-3

Figure 5.2 Cross-Sectional Schematic (circumferential orientation) Showing Typical 

Weld Geometry and Implanted Crack in HAZ

Figure 5.3 Cross-Sectional Schematic (axially oriented) Showing Typical Crack Location
5-4

Table 5.1 Implanted Thermal Fatigue Cracks in Austenitic Piping Specimen 02-24-15

Crack Designation A B C D E

Crack Orientation Circ. Circ. Circ. Axial Circ.

Crack Length - L 

[± 1.0 mm (0.040 in.)]

10.7 mm

(0.42 in.)

30.5 mm 

(1.2 in.)

43.6 mm 

(1.715 in.)

13.3 mm 

(0.525 in.)

33.8 mm 

(1.33 in.)

Through-wall Depth - H

[± 1.0 mm (0.040 in.)]

5.0 mm 

(0.195 in.)

14.9 mm 

(0.585 in.)

21.5 mm 

(0.845 in.)

6.6 mm

(0.26 in.)

16.5 mm 

(0.65 in.)

% Wall Thickness 15 43 64 19 48

Aspect Ratio 2.3 2.1 2.1 2.1 2.1

Circumferential Location - D (from 0°) 30° 65° 165° 270° 330°

The piping specimen was labeled 02-24-15, a 0° reference datum point was inscribed, and each 

crack arbitrarily assigned an alphabetical designation. Welding and crack implantation were

performed with the piping material in the as-formed (cylindrical) shape, then the specimen was 

axially sectioned into three, 120° segments for ease of handling during the experimental work. 

The weld crown was machined off to allow improved access for ultrasonic transducer placement 

during scans from the outside surface of the pipe. 

In addition to the implanted TFCs, eight sawcuts were introduced into the heat-affected zone of 

the girth weld to provide standard ultrasonic reflectors. The sawcuts were made using a 4-in.-

diameter radial saw blade and range from 2.5 mm (0.1 in.) deep to 10.2 mm (0.4 in.) deep. 

Four of the sawcuts are at right angles to the inner surface of the pipe and four follow the 35° 

angle of the weld HAZ. The exact lengths and depths of the sawcuts were determined by 

applying EXAMIX (a commercially available vinyl polysiloxane molding material), allowing the 

material to harden, removing the molds, and then measuring them. This process was repeated 

three times for each sawcut to better quantify experimental error. The sawcut dimensions and 

locations are given in Figure 5.4 and Table 5.2. A photograph of the EXAMIX positives is 

shown in Figure 5.5.
5-5

Figure 5.4 Positions and Angles for the Sawcuts in Pipe Specimen 02-24-15

Table 5.2 Sawcuts in Austenitic Piping Specimen 02-24-15

Sawcut Designation A B C D

Sawcut Orientation Straight Cut Straight Cut 35° Angle Straight Cut

Length – L [± 0.4 mm (0.016 in.)] 32.8 mm

(1.29 in.)

65.2 mm

(2.57 in.)

36.2 mm

(1.43 in.)

54.1 mm

(2.13 in.)

Through-wall Depth – H [± 0.4 mm (0.016 in.)] 2.7 mm

(0.11 in.)

10.2 mm

(0.40 in.)

2.5 mm

(0.10 in.)

6.8 mm

(0.27 in.)

% Wall Thickness 7.5 28.4 7.1 18.8

Aspect Ratio 12 6 15 8

Circumferential Location – D (from 0°) 22.5° 45° 85° 150°

Sawcut Designation E F G H

Sawcut Orientation Straight Cut 35° Angle 35° Angle 35° Angle

Length – L [± 0.4 mm (0.016 in.)] 43.7 mm

(1.72 in.)

59.7 mm

(2.35 in.)

57.3 mm

(2.25 in.)

68.4 mm

(2.69 in.)

Through-wall Depth – H [± 0.4 mm (0.016 in.)] 4.3 mm

(0.17 in.)

7.0 mm

(0.28 in.)

6.3 mm

(0.25 in.)

9.3 mm

(0.37 in.)

% Wall Thickness 12 19 18 26

Aspect Ratio 10.2 8.5 9.1 7.4

Circumferential Location – D (from 0°) 185° 210° 285° 310°
5-6

Figure 5.5 EXAMIX Molds Used to Measure the Length and Depth of the Sawcuts in Pipe 

Specimen 02-24-15

The three axial pipe segments were placed end-to-end on a wheeled cart to facilitate access to 

both LF/SAFT and PA-UT data acquisition stations. Mechanical scanning was performed by 

modifying the scanner to accommodate conventional transducers, LF/SAFT transducers or PA

probes, with water being used as a coupling medium (see Figures 5.6 and 5.7 showing a typical 

experimental configuration).
5-7

Figure 5.6 Experimental Set-up Showing Piping Specimens and Scanner

Figure 5.7 Close-up of Scanner on Pipe Specimen
5-8

5.2 Phase 2: Austenitic Piping Specimen No. 3C-022

The specimen used in Phase 2 was made using the same vintage material used in the Phase 1 

work—Schedule 80, ASTM 358 Grade 304 WSS having a nominal 610-mm (24.0-in.) outer 

diameter and a wall thickness of approximately 36 mm (1.4 in.). The specimen was made by 

joining four cylinders of this material with three single-Vee girth welds, using procedures and 

filler material (E308 L) typical of those used in nuclear applications.

The piping specimen was labeled 3C-022, and a 0° reference datum point was inscribed. The 

weld crown was machined off to allow improved access for ultrasonic transducer placement 

during scans from the outside surface of the pipe. The pipe specimen was placed vertically in a 

water bath to recycle couplant for conventional UT, LF/SAFT, and PA-UT data acquisition 

systems. Again, mechanical scanning was performed by modifying the scanner to 

accommodate conventional-UT transducers, LF/SAFT transducers, or PA probes, with water 

being used as a coupling medium (see Figure 5.8 showing typical experimental configuration). 

Reference marks were indented 75 mm (3 in.) to the right of the 0° reference point to designate 

the starting point of scans. 

Girth welds number one and number three were made with the pipe held horizontally in a fixed 

position. The cylinders were welded in a field-like fashion, with the welder starting at the top of 

the weld and working his way down to the bottom on each side of the pipe. This process is 

shown in Figure 5.9. Girth weld number one was welded with air backing, and girth weld 

number three was welded with water backing. After the welding process was completed, a 

10% through-wall, 3.6-mm (0.14-in.) deep notch was machined 360° around one side of the 

weld on the inner diameter in the heat affected zone. This notch provided a full-circumferential,

uniform ultrasonic reflector.

Girth weld number two was made with the pipe held vertically with air backing. During the 

welding process for weld number two, three cracks were implanted into the HAZ of the girth 

weld. These cracks were implanted by the same vendor and in the same way as the cracks in 

specimen 02-24-15. After the welding process, three electro-discharge machined (EDM)

notches were implanted into one side of the weld on the inner diameter in the HAZ. The three 

EDM notches are all 10% though-wall, 3.6-mm (0.14-in,) deep, and begin and end 25 mm 

(1.0 in.) from each implanted crack. The EDM notches are designated flaws 1, 2, and 3 and the 

cracks are designated cracks 4, 5, and 6. The flaw dimensions and locations for cracks 4, 5, 

and 6 are shown in Table 5.3.
5-9

Figure 5.8 Experimental Set-up Showing Piping Specimen and Scanner
5-10

Figure 5.9 Field-like Welding on Horizontally Oriented Pipe Segments (photo courtesy 

of FlawTech Inc.)

Table 5.3 Thermal Fatigue Cracks in Austenitic Piping Specimen 3C-022

Crack Designation 4 5 6

Crack Orientation Circ. Circ. Circ.

Crack Length – L [± 1.0 mm (0.040 in.)] 3.6 mm (0.14 in.) 7.1 mm (0.28 in.) 10.7 mm (0.42 in.)

Through-wall Depth – H [± 1.0 mm 

(0.040 in.)]

1.8 mm (0.07 in.) 3.6 mm (0.14 in.) 5.3 mm (0.21 in.)

% Wall Thickness 5 10 15

Aspect Ratio 2.0 2.0 2.0

Circumferential Location – D (from 0°) 45° 165° 285°

5.3 Phase 3: IGSCC - EPRI Practice and Secure Set Specimens

In Phase 3 of this study, a number of specimens from the EPRI performance demonstration set 

were made available to PNNL for ultrasonic examination. These specimens were field-removed 

piping segments taken from several U.S. boiling water reactor primary re-circulation systems 

and contain service-induced IGSCC. Some of the specimens were part of a practice set and 

some were part of a secure set used for blind performance demonstration tests. PA data were 

acquired on both sets using identical techniques during multiple visits to the EPRI NDE Center 

in Charlotte, North Carolina. The practice data were taken back to PNNL for development of an 

analysis approach and the secure data set was held at EPRI for subsequent data analyses.
5-11

5.3.1 EPRI Performance Demonstration Practice Specimens

A number of welded specimens containing service-induced intergranular stress corrosion cracks 

were available for examination at the EPRI NDE Center. These specimens varied in 

configuration with different weld crown geometries, counterbore, weld root conditions, etc. 

Figures 5.10 and 5.11 show some of the variation in specimens as viewed from the outside and 

inside of the pipe segments. The red arrows painted on two of the specimens show the flow 

direction so that upstream and downstream sides of the welds are identifiable. Specimens had 

a nominal 71.0-cm (28.0-in.) diameter and 2.0–4.1 cm (0.8–1.6 in.) wall thickness. A mapping 

of the flaws was provided, showing flaw location in the circumferential direction and axial 

position as upstream or downstream from a weld center line (WCL). However, a WCL was not 

marked on the specimen so axial position for acquired data was referenced relative to the 

center of the weld crown. True-state location and length-sizing information was given, but no 

true-state depth information was available. The examinations for this activity focused on 

detection of inner-diameter (ID)-connected cracks, oriented circumferentially (parallel to the 

weld).

Figure 5.10 Examples of Specimens as Viewed from the Outer Surface
5-12

Figure 5.11 Examples of Specimens as Viewed from the Inner Surface

5.3.2 EPRI Performance Demonstration Secure Specimens

The secure specimens were similar to the practice specimens but access was limited to the 

outer surface only. These specimens are used by EPRI in administering performance 

demonstration evaluations so the specimen number and flaw information are purposefully not 

revealed to maintain the integrity of the test set. The flaw true-state was provided to PNNL 

personnel during proctored data analysis activities at the EPRI NDE Center and the secure truestate information and data acquired were not allowed to leave the testing room.
6-1

6 Ultrasonic Probes and Inspection Protocols

In Phases 1 and 2 of the work reported here, the far-side examination assessment included 

evaluations of all three ultrasonic inspection techniques—conventional-UT, LF/SAFT and 

PA-UT. In Phases 1 and 2, the PA-UT technique focused on the use of 2.0-MHz probes, but 

employed both longitudinal- and shear-wave modalities. In Phase 3 of the evaluation, only the 

PA-UT approach was assessed, and the focus of the data acquisition included a 1.5-MHz 

transmit-receive longitudinal (TRL) probe in addition to the 2.0-MHz probes used in the earlier 

phases. 

6.1 Conventional–UT

As a benchmark for comparison, it is useful to understand the common industry practice for the 

far-side inspection of austenitic welds. The ASME Section XI Code, Appendix VIII, and 

10CFR50.55a require successful demonstration of the far-side inspection of austenitic welds.

The test administered by EPRI may be biased in that flaws are selected based on the fact that 

conventional techniques can successfully detect them. This test is still very challenging 

because, even when the flaws have been selected in this manner, the far-side inspection is very 

difficult due to the geometry and metallurgical signals that must be correctly classified. EPRI 

NDE Center staff indicated that they recommend (and this is what many inspectors employ) the 

following approach for a candidate to pass the performance demonstration test:

 The first probe to use is a dual shear-wave transducer having two 10 mm by 18 mm (0.4 in. 

by 0.7 in.) piezo-composite elements, operating at 1.5 MHz and focused for 80% of the wall 

thickness with an inspection angle of 60° to 70°. 

 After indications are found, evaluate with a 60° dual longitudinal probe operating in the 1- to 

2-MHz range with a 2-MHz most commonly used.

The images of the raw data from the 60° longitudinal LF/SAFT inspections shown in Appendix B 

are similar to detection results that are expected for a conventional, dual-element inspection. It 

should be noted that these results are not exactly the same as the results that would be 

obtained when using the EPRI NDE Center recommendations; however, they are similar and 

illustrate the challenge the inspector must deal with regarding the detection of far-side austenitic 

weld flaws. The high-angle, shear-wave modality is effective because most of the piping welds 

are of a single "V" design and, thus, only a small amount of weldment must be traversed by the 

sound field near the inner diameter (ID) of the component, there are no mode-converted signals, 

and one may be able to obtain a corner trap from the flaw. However, one of the major problems 

with the inspection is the access limitation produced by the weld crown. Because, for most 

welds, the probe cannot be scanned on top of the weld, it is very difficult to get the full echo 

dynamic response for far-side flaws. Acoustic signals are reflected from existing flaws but only 

a limited portion of the echo dynamic curve is obtained and at the same time there can be many 

signals of geometric and metallurgical origin. The crown also provides limits on the ability to 

skew the transducer to assess the angular dependence of flaw signals from those produced by 

the geometry and metallurgical conditions. Refracted longitudinal probes are used for the 

assessment of indications. However, with refracted longitudinal waves, there is a problem of 
6-2

mode conversion at the inner surface producing many more signals that may overlap in time 

with signals coming from the indication or from the metallurgical and geometrical conditions, 

which can create problems in classification of the signals. Also, tip-diffracted signals may not be 

reliably detected, so time-of-flight depth sizing is improbable.

Automated, single-element, conventional shear-wave ultrasonic techniques were employed on 

the wrought-austenitic SS specimens 02-24-15 and 3C-022, for comparison with the PA-UT and 

LF/SAFT techniques. The conventional ultrasonic system at PNNL consisted of the SAFT data 

acquisition software, a PNNL pipe scanner, a Ritec SP-801A square-wave pulser, and a Ritec 

BR-640A receiver. The four Krautkramer probes that were used in this study were chosen 

based on discussions with EPRI staff to identify probes that were most commonly used by 

inspectors during manual austenitic stainless steel performance demonstrations. The probes 

include two 1.5-MHz transducers, one with a diameter of 9.5 mm (0.375 in.) and one with a 

diameter of 12.7 mm (0.5 in.), and two 2.25-MHz transducers with the aforementioned 

diameters. The transducers were mounted on wedges to provide 60° and 70° shear waves in 

the material. The scans were completed from both sides of the weld to examine the effects of 

the weld microstructure on the UT response from implanted cracks and notches. The laboratory

scans differ from standard manual ultrasonic examinations of pipes in that the scans were taken 

using an automated scanner and the data were stored digitally. This method was chosen to 

provide images for a more direct comparison with the PA-UT and LF/SAFT data.

In Phase 1 of the study, the specimen was cut into three segments to allow easier handling and 

scanning. The first section of specimen 02-24-15 was scanned with all four probes at both 60° 

and 70° incident shear waves, and the weld was examined from both the near- and far-side, for 

a total of 16 scans. The raster scan path length in the x-direction (axial direction, perpendicular 

to the weld) was 152 mm (6.0 in.) allowing the probe to traverse over the machined weld crown 

while digitizing the radio frequency (RF) data at a sampling rate of 20 MHz. The scanner would 

then index (step) 1.3 mm (0.05 in.) in the y-direction (circumferential direction, parallel to the 

weld) and then translate the ultrasonic probe a distance of 152 mm (6.0 in.) parallel and 

adjacent to the preceding stroke. A total of 508 mm (20.0 in.) of travel in the y-direction was 

completed for each scan on these pipe welds, and required about 40 minutes per search unit to 

scan (total scan time for all probes was approximately 10 hours, 40 minutes).

Preliminary screening of these initial data sets was performed immediately after the scanning 

process to determine which transducer(s) provided the best signals and highest SNR 

performance. After examining the 16 scans of pipe section one, it was determined that the 

1.5-MHz probes performed marginally better than the 2.25-MHz probes, and the larger 12.7-mm 

(0.5-in.) diameter transducers performed much better than the 9.5-mm (0.375-in.) transducers. 

Figure 6.1 provides views of the conventional-UT probe configuration with associated gimbaling 

and couplant connections, with the 1.5-MHz, 12.7-mm (0.5-in.) diameter probe.
6-3

Figure 6.1 Side and Top Views of the 1.5-MHz, Conventional-UT, Shear-Wave Probe and 

Gimbal/Couplant-Delivery Configuration for Translation of the Ultrasonic 

Transducer Across the Specimens

After this analysis was conducted, the total number of scans was reduced to four scans per 

weld. The 1.5-MHz, 12.7-mm (0.5-in.) diameter 60° and 70° shear wave scans were completed 

on the near and far sides for pipe sections two and three, and for all of Phase 2. Therefore, only 

the 1.5-MHz results will be discussed in subsequent sections of this report.

In Phase 2 of the study, conventional-UT methods were applied using much longer raster scan 

patterns for volumetric coverage of the three circumferential welds provided by specimen 

3C-022. Each of the three welds was scanned using 60° and 70°, 1.5-MHz shear waves. The 

data were stored electronically and analyzed using the SAFT 2002 software package. For weld 

number one (as described in Section 5), the 10% notch was scanned using 1.5-MHz, 60° and

70° shear waves from the far side of the weld. The probes were raster-scanned over the pipe 

specimen, with 152-mm (6.0-in.) scan-strokes in the x-direction (axial direction, perpendicular to 

the weld) and indexed (stepped) in 1.3-mm (0.05-in.) increments for a total of 1966 mm 

(77.4 in.) around the outside circumference of the pipe specimen (parallel to the weld), which 

produced a total of 1548 lines (strokes) per data set. Each scan was started at a distance of 

76 mm (3.0 in.) after the centerline of the longitudinal seam weld and indexed 360° around the 

pipe. The scan duration was approximately 4 hours per scan, for a total of 16 hours of scanning 

for weld number one.

Weld number two was scanned from the near side with the 1.5-MHz, 60° shear-wave probe and 

from the far side with the 60° and 70° shear-wave probes. The probes were raster-scanned 

over the pipe specimen, with 152-mm (6.0-in.) scan-strokes in the x-direction (axial direction, 

perpendicular to the weld) and indexed (stepped) 2118-mm (83.4-in.) around the outside 

circumference of the pipe specimen (in the y-direction, parallel to the weld). The probes were 

indexed in 1.3-mm (0.05-in.) increments, for a total of 1668 lines per data set. Again, each scan 
6-4

was initiated at a distance of 76 mm (3.0 in.) before the centerline of the longitudinal seam weld, 

indexed 360° around the pipe, and then continued indexing an additional 152 mm (6.0 in.) to 

76 mm (3.0 in.) beyond the longitudinal weld centerline. The 152 mm (6.0 in.) of overlap in each 

scan provides two built-in fiduciaries in the scans as the longitudinal seam appears as a ―dead 

zone‖ in the scans. These scans took approximately 4 hours each to complete, for a total of 

16 hours of scanning.

Finally, weld number three was raster-scanned using identical x-direction scan-strokes as in the 

previous welds, but were indexed in the y-direction (circumferential direction, parallel to the 

weld) 2118 mm (83.4 in.) around the outside circumference of the pipe specimen. The probes 

were indexed in 1.3-mm (0.05-in.) increments, for a total of 1668 scan lines per data set. Once 

again, these scans were initiated at the same position as the other two welds, starting at a 

distance of 76 mm (3.0 in.) before the centerline of the longitudinal seam weld, indexed 360° 

around the pipe, and then continued indexing an additional 152 mm (6.0 in.) to 76 mm (3.0 in.) 

beyond the longitudinal weld centerline. The scans took approximately 4 hours per scan, for a 

total of 16 hours of scanning.

6.2 Low-Frequency/SAFT–UT

The low-frequency/SAFT inspection technique uses a zone-focused, high-bandwidth, lowfrequency (400-kHz) inspection protocol coupled with the SAFT technique (Doctor et al. 1996). 

The examination protocol is based on the premise that there exist sufficient differences between 

the characteristics of coherently scattered ultrasonic energy from grain boundaries and 

geometrical reflectors from that of the scattered ultrasonic energy from surface-breaking fatigue 

and stress corrosion cracks in wrought austenitic piping welds. PNNL’s empirical approach 

relies on the notion that acoustic impedance variations at the grain boundaries (and in the HAZ 

along the weld fusion lines) can be minimized by using lower frequencies (longer wavelengths), 

and the degree of coherent energy scattered from these boundaries should be inconsistent as a 

function of frequency, insonification angle, scan direction, and the amplitude of returning 

signals. The LF/SAFT approach is directed toward detecting the corner-trap response from the 

surface-breaking crack as a function of time and spatial position. If the frequency is low 

enough, the examination is less sensitive to geometry and the effects of the microstructure and 

the probability of detection increases for surface-breaking cracks. The tradeoff is resolution. 

However, with the addition of SAFT signal processing, the examination can be performed at 

lower frequencies while maintaining sufficient resolution and sensitivity to effectively detect 

cracks. Therefore, by using multiple examination frequencies and incident angles, and 

inspecting from both sides of a weld, the LF/SAFT technique invokes a composite approach for 

detection, localization, and sizing of cracks in these types of materials.

The examination process is further enhanced by the addition of low-frequency, piezo-composite, 

high-bandwidth search units that enable the inspector to use a wide range of examination 

frequencies centered around 250 kHz and 400 kHz. The zone-focal characteristics of the 

Krautkramer-AGFA dual-element search units provide optimal insonification of the inner surface 

(ID) over an incremental range of incident angles for pipe wall thicknesses of approximately 

3.56 cm (1.4 in.).
6-5

PNNL staff conducted a preliminary ultrasonic evaluation of the LF/SAFT inspection method on 

a set of three sectioned segments (specimen 02-24-15) of a wrought 304 SS pipe with an 

austenitic weldment and a limited array of cracks and notches. From this work, a more 

extensive study was proposed, and the LF/SAFT approach was evaluated in more detail as a 

result of recommendations derived from earlier work. Extensive laboratory work using this novel 

approach has been performed at PNNL and other sites in the field prior to the study described 

here. This inspection technique has evolved to focus on inspection problems associated with 

thick-section, reactor-coolant piping that exhibit coarse-grained microstructures or multi-layered 

boundary conditions such as centrifugally and statically cast pipe, elbows, and dissimilar metal 

welds. Due to the difficulties encountered in far-side examinations through austenitic welds, this 

technique was identified as having potential for improved inspection performance due to the 

utility of longer wavelengths, thereby reducing the attenuative effects of the weld material. This 

study is an extension of previous evaluations and focuses on applying the inspection 

methodology to far-side examinations in wrought piping with austenitic welds, where a suite of 

fabricated flaws and cracks have been introduced for performance evaluation and comparison 

to other inspection methods.

"Synthetic aperture focusing" refers to a process in which the focal properties of a largeaperture focused transducer are synthetically generated from data collected over a large area 

using a small transducer with a divergent sound field. The processing required to focus this 

collection of data has been called beam-forming, coherent summation, or synthetic aperture 

processing. The resultant image is a full-volume focused characterization of the inspected area

(Doctor et al. 1996).

SAFT-UT technology is able to provide significant enhancements to the inspection of austenitic 

welds and other anisotropic, coarse-grained materials. The resolution of an imaging system is 

limited by the effective aperture area; that is, the area over which data can be generated, 

collected, and processed. SAFT is an imaging method that was developed to overcome some 

of the limitations imposed by large physical apertures, and has been successfully applied in the 

field of ultrasonic testing. Relying on the physics of ultrasonic wave propagation, SAFT is a very 

robust technique (Doctor et al. 1996).

Using a pitch-catch configuration for typical data collection throughout these activities, the 

transducer was positioned on the surface of the specimens, and RF ultrasonic data were 

collected. As the transducer was scanned over the surface of the specimens, the A-scan record 

(RF waveform) was amplified, filtered, and digitized for each position of the transducer. Each 

introduced flaw produced a collection of echoes in the A-scan records. The unprocessed or RF 

data sets were then post-processed using the SAFT algorithm, invoking a variety of full beamprocessing angles (between 6° and 24°) to evaluate and optimize the temporal/spatial 

averaging enhancement. If the reflector is an elementary single-point reflector, then the 

collection of echoes will form a hyperbolic surface within the data-set volume. The shape of the 

hyperboloid is determined by the depth of the introduced flaws in the specimens and the velocity 

of sound in the specimens. This relationship between echo location in the series of A-scans 

and the actual location of the introduced flaws within the specimens makes it possible to 

reconstruct a high-resolution, high signal-to-noise ratio image from the acquired raw data.
6-6

If the scanning and surface geometries are well known, it is possible to accurately predict the 

shape of the locus of echoes for each point within the test object. The process of coherent 

summation for each image point involves shifting a locus of A-scans, within a regional aperture, 

by predicted time delays and summing the shifted A-scans. This process may also be viewed 

as performing a temporal/spatial matched filter operation for each point within the volume to be 

imaged. Each element is then averaged by the number of points that were summed to produce 

the final processed value. If the particular location correlates with the elementary point 

response hyperboloid, then the values summed will be in phase and produce a high-amplitude 

result. If the location does not correlate with the predicted response, then destructive 

interference will take place and the spatial average will result in a low-amplitude value; thus, 

reducing the noise level to a very small value. SAFT processing is also effective at removing 

constant-time signals that are not near the front surface, and sub-volume selection during data 

analysis readily removes any residual near-surface signals (Doctor et al. 1996).

The LF/SAFT-UT inspection system is an automated, computerized UT imaging system 

designed and developed by PNNL, and evolved from a system originally employed by NRC 

Region 1 inspectors as a mobile SAFT-UT inspection system. The LF/SAFT-UT system is a 

field-tested system, consisting of several subsystems and components. The Panametrics 

pulser used for this study is a high-voltage model 5058 pulser-receiver unit that produces highvoltage spike pulses for broadband excitation of the transducer. This unit provides damping 

control, step attenuation and gain control, high-pass and low-pass filtering options, and a 

variable pulse voltage control of up to 900 volts. This unit provides the capability to exploit the 

high bandwidth of the piezo-composite transducers and allow utility of both 250-kHz and 

400-kHz examination frequencies. This unit also includes an embedded dual-stage receiver 

that consists of a low-frequency, low-noise pre-amplifier in series with a high performance, 

60-dB low-noise amplifier for conditioning and amplification of the received UT signal 

responses. A stand-alone 866-MHz personal computer (PC) contained in this subsystem fulfills 

the role of data acquisition control, where the operator invokes the LF/SAFT scanning software 

menus and adjusts the scanning parameters prior to initiating a scan. The analog-to-digital 

(A/D) converter resides on this PC, which also controls the motor drivers and pipe scanner. 

This computer is directly connected to the data analysis PC. Unprocessed ultrasonic data 

streams from the data acquisition PC to the data analysis PC where it resides for SAFT postprocessing, analysis, and archival. The data processing and storage subsystem performs SAFT 

processing of the raw (unprocessed) UT data using a Dell Precision 650 high-performance PC 

with dual high-speed processors (2.4 GHz each). This PC platform contains both read/write 

compact disc (CD) and digital video disc (DVD) drives for storage of digital data. This PC also 

contains two flat-panel 19-in. monitors for displaying the SAFT-UT processed data (A-scans, 

B-scans, C-scans, and D-scans). The photographs in Figure 6.2 illustrate the LF/SAFT 

inspection system in its configuration as used in this study.
6-7

Figure 6.2 The Low-Frequency/SAFT-UT Ultrasonic Data Acquisition System

The low-frequency ultrasonic data acquisition system used in this study allowed ultrasonic data 

to be efficiently acquired under rapid, low-noise conditions, for a number of sectioned pipe 

segments and weld conditions using a field-ready automated pipe scanner. The LF/SAFT 

automated pipe scanner shown in Figure 6.2 was used for accurate and smooth scanning of the 

search unit through a specified number of points in the X-Y plane, while maintaining low-noise 

conditions and a constant coupling of the transducer to the surface. The pipe scanner was 

configured with tailored gimbals for transducer attachment and translation over the surface of 

the specimens. A network of specially fabricated drip lines attached to a couplant reservoir with 

a peristaltic pump was used to provide a constant stream of water on the specimen surface for 

improved coupling and transducer motion. Figure 6.3 shows photographs of the pipe scanner’s 

front-end configuration.

Figure 6.3 Front-end Transducer Configuration, Including Gimbal and Coupling Lines. 

Colored water was used as couplant to facilitate assessing the presence of

air bubbles at the transducer wedge-to-component interface.
6-8

The LF/SAFT system’s 866-MHz control PC contains the A/D converter used to sample 

(digitize) the data. The motor controller electronics initialize a trigger output to the A/D and 

pulser-receiver unit which, in turn, was used to sync the remaining instrumentation. The 

received signal responses (echoes) were initially amplified by a low-noise, low-frequency, 

ultrasonic preamplifier, and then bandpass filtered using the Panametrics 5058 to allow suitable 

amplification while reducing extraneous low-frequency noise components under 100 kHz and 

higher-frequency noise components over 500 kHz. 

The preamplified and conditioned signal responses were then amplified by the second stage 

wideband Panametrics 5058 amplifier and sent to the control PC’s A/D converter for digitizing. 

All signals (excitation pulses and received signal responses) were monitored using a LeCroy 

9450 digital oscilloscope, providing the capability to view the trigger and sync pulses, the 

excitation pulse before and after amplification, and the received signal response just prior to 

digitization, simultaneously. The LeCroy oscilloscope also provided linear averaging, and the 

capability to analyze the frequency characteristics of the A-scan data. A linear averaging 

scheme was implemented with the RF ultrasonic A-scan data being averaged eight times per 

repetition, to minimize the effects of motor noise and random electronic noise, and to increase 

the unprocessed SNR. The data acquisition system was capable of acquiring low-frequency, 

low-noise, ultrasonic data within a region of 0 dB to 100 dB total system gain.

After storage of the RF ultrasonic data sets to CD media (with the option for DVD archival), the 

data was signal processed using SAFT. The raw data sets were post-processed using a variety 

of beam processing angles, ranging from 6° to 24°. The SAFT software provides the user the 

capability to view the entire ultrasonic data set (three-dimensional array of points) in twodimensional slices, by viewing color-enhanced composite images that depict slices of the threedimensional array in the X-Y plane (C-scan view), the Y-Z plane (B-scan end-view or D-scan 

view), and X-Z plane (B-scan side view). The X-axis defines the axis of motion for translation of 

the transducer, perpendicular to the circumferential weld axis. The Y-axis defines the 

translation of the transducer in the circumferential direction, parallel to the centerline of the weld. 

The Z-axis defines the depth into the specimen from the contacting top surface of the specimen, 

and is correlated to the time-of-flight data.

In 2002, preliminary assessments with the LF/SAFT inspection system on wrought piping with 

austenitic welds resulted in a number of recommendations. These lessons learned were 

accommodated in subsequent efforts. Of primary significance was a re-design of the 

transducers used for the inspection of these components. The originally designed lowfrequency, variable-angle transducer used for thicker-section cast piping components did not 

provide a suitable sound field for the thinner-section 304 WSS piping. PNNL staff worked with 

Krautkramer-AGFA staff in developing the necessary specifications for fabricating highbandwidth, piezo-composite, low-frequency transducers and a set of associated wedges for 

application to this component geometry. A set of 10 dual-angled wedges providing inspection 

angles of 30°, 45°, 50°, 55°, 60°, 65°, 70° (L-wave in 304 SS) and 45°, 60°, 70° (shear-wave in 

304 SS) were acquired for the initial data acquisition activities.

The transducers were fabricated to operate as dual (pitch-catch), isolated (to reduce cross-talk), 

high-bandwidth (at the −6 dB, half-power point), low-frequency, piezo-composite transducers. 

The nominal and peak operating frequencies were to be approximately 350 kHz and the −6-dB 
6-9

bandwidth for both transducers were greater than 90%. This provided a suitable operational 

frequency range for employing both 250 kHz and 400 kHz as the two primary examination 

frequencies for this study. To reduce the acoustic ―footprint‖ on the specimen surface and 

minimize the transducer size, the housings were specified to be no larger than 2.54 cm (1.0 in.) 

× 3.81 cm (1.5 in.) rectangular housings. In addition, the wedge sets were to be configured with 

an appropriate roof angle to provide a zone-focus (cross-over point where receive and transmit 

sound-field zones coincide) in the range of approximately 4.45–7.62 cm (1.75–3.0 in.) along the 

sound field propagation path (i.e., 4.98 cm at 45° in 304 SS). The photographs in Figures 6.3

and 6.4 show the dual-element, pitch-catch transducer configuration employed.

Figure 6.4 Transducer-Wedge Configuration for Dual-Element, Pitch-Catch Scanning

Because the inspection technique takes advantage of two frequencies and two wave-modes, 

the wavelength (and effectively the system’s resolution capability) will vary as a function of 

frequency and wave-mode. The longitudinal-wave acoustic velocity in 304 SS is 5.790 mm/μs 

(0.228 in./μs), while the shear-wave acoustic velocity is 3.180 mm/μs (0.125 in./μs). The 

equation for calculating wavelength is given by

λ = C / f (6.1)

where λ is the wavelength, C is the appropriate acoustic velocity for the specific wave-mode 

employed, and f is the frequency. Typically the calculation is made using velocity units of 

cm/sec or mm/sec and frequency units of Hz or cycles/sec. In this manner, the wavelength 

calculation results in a familiar unit length. For 250-kHz longitudinal (L)-wave inspection, the 

wavelength is 2.32 cm (0.91 in.), and for shear-wave inspection the wavelength is 1.27 cm 

(0.50 in.). For 400-kHz L-wave inspection, the wavelength is 1.45 cm (0.57 in.), and for shearwave inspection the wavelength is 0.79 cm (0.31 in.). Near the weld root, the typical thickness 

of the specimens ranged from 3.53–3.61 cm (1.39–1.42 in.). A 10% through-wall notch would 

have a depth of approximately 0.36 cm (0.14 in.). Thus, the smallest wavelength employed (at 

400 kHz in the shear-wave mode) is still over two times larger than a 10% through-wall crack or 

notch.
6-10

Another key element in understanding the resolution capability of the LF/SAFT inspection 

technique is to know the characteristic zone focal dimensions of the transducers. The 

intersection (or cross-over point) in the component of the transmit and receive transducers is 

where the sound-field energy is quasi-focused and the ―zone‖ where the two transducers 

interact. The dimensions of the zone-focal area are functions of the frequency, the size of the 

transducer element, the spacing and roof-angle (geometric configuration and alignment of the 

side-by-side pair of transducers), the incident angle, and the acoustic velocity of both the wedge 

and the material being examined. A number of calculations were conducted at the −6 dB points 

to better understand the lateral and depth dimensions (Y axis and Z-axis relative to the 

coordinate plane of inspection) of the zone-focused sound field. 

 For 250-kHz 45° L-wave configuration, the sound field overlap in the Y-axis is approximately 

4.1 cm (1.6 in.) and the overlap in the X and Z-axes is approximately 3.0 cm (1.2 in.). 

 For 250-kHz 70° L-wave configuration, the sound field overlap in the Y-axis is approximately 

3.3 cm (1.3 in.) and the overlap in the X and Z-axes is approximately 2.0 cm (0.8 in.). 

 For 400-kHz 45° L-wave configuration, the sound field overlap in the Y-axis is approximately 

3.0 cm (1.2 in.) and the overlap in the X and Z-axes is approximately 2.5 cm (1.0 in.). 

 For 400-kHz 70° L-wave configuration, the sound field overlap in the Y-axis is approximately 

2.8 cm (1.1 in.) and the overlap in the X and Z-axes is approximately 1.5 cm (0.6 in.). 

This illustrates the slight convergence and ―zone-focusing‖ effect inherent in the design of the 

transducer configuration. These dimensions are generally much larger than the depth size of 

the introduced flaws in the piping segments examined for this study, and much larger than the 

length of its cracks or sawcuts, with the exception of the 360° notches.

A general scanning procedure was developed to provide consistent and repeatable data 

acquisition and to provide technical support staff with routine guidelines for conducting data 

acquisition and data processing/analytical tasks. It is recognized that this is not an ASME Code 

qualified procedure, and no calibration on machined reflectors was conducted, but follows the 

general approach of such a procedure. The following process was used:

 Invoke scanning setup menu on control PC.

 Configure proper transducer/wedge into spring-loaded gimbal and adjust for optimal 

translation.

 Calculate required two-way time-of-flight for adjustment of sampled time-window.

 Modify header files and parameters for transducer, material, sampling, and scan pattern.

 Examine A-scan trace on computer monitor.

 Verify proper coupling and flow rate for drip lines.

 Slowly scan transducer over largest indication (flaw or crack) and verify proper gain settings 

on instrumentation, setting the total system gain just below saturation of the highest 

amplitude signal to be found.

 When all parameters have been set, outline scan aperture.
6-11

 Print out header file.

 Initiate scanning.

 Complete documentation of instrument settings.

 When file is completed, verify proper data acquisition by a quick analysis of the raw UT data 

file and store to hard drive and to CD media.

 Initiate SAFT processing on separate PC platform for analysis and data storage.

 Repeat step 1 until data acquisition activity is completed.

Initially, it was thought that all scan-angle-frequency combinations could be acquired, SAFTprocessed using various beam processing angles, and analyzed individually by wave modality 

to compare and contrast the results to the true state of the specimens. However, the large 

volume of scans, the time for data acquisition, and required time for post-processing and 

analysis of the entire set of scan combinations could not be achieved given the schedule and 

resources available for this portion of the study. To address detection performance issues and 

provide the best possible conditions for evaluating the LF/SAFT technique, the full complement 

of scans were conducted on one of the sectioned pieces and an initial analysis was performed 

to determine which combination of modalities, angles, and frequencies would provide the most 

meaningful and useful data. Using the results as guidance from these initial tests on piping 

segment number two, it was determined that the combination of inspection parameters depicted 

in Table 6.1 would be implemented for each pipe segment over the remainder of the study.

Table 6.1 Final Combination of Scan-Angle-Frequency-Modality for Phase 1 Evaluation 

Using the LF/SAFT Inspection Technique

Piping 

Segment ID #

Inspection

Frequency (kHz)

Orientation 

of Scan

Examination 

Incident Angle Wave Mode

All 250 Far-side 45°, 70° Longitudinal

All 250 Near-side 45° Shear

#2 250 Near-side 60°, 70° Shear

#2 250 Far-side 45°, 60°, 70° Shear

All 400 Near-side 45°, 60°, 70° Longitudinal and Shear

All 400 Far-side 45°, 60°, 70° Longitudinal and Shear

In Phase 2, where the focus was on assessing specimen 3C-022, each of the three welds was 

scanned using 400-kHz and 250-kHz probes. The welds were examined using 45° shear and 

longitudinal waves, 60° longitudinal waves, and 70° longitudinal waves at both frequencies. The 

probes were raster-scanned over the pipe specimen, with 152.4-mm (6.0-in.) scan strokes in the 

x-direction (axial direction, perpendicular to the weld) and indexed 2118 mm (83.4 in.) around 

the outside circumference of the pipe specimen in the y-direction (parallel to the weld). The 

400-kHz probes were indexed or stepped in 1.3-mm (0.05-in.) increments and the 250-kHz 

probes were indexed or stepped in 2.5-mm (0.10-in.) increments. Thus the index increments 

represented approximately 1/10th of a wave length and an over-sampling in the lateral plane but 

these small steps produced a more detailed data image for analysis. Again, as described for 
6-12

the conventional-UT scans, each scan was initiated at a distance of 76.2 mm (3.0 in.) before the 

centerline of the longitudinal seam weld, indexed 360° around the pipe, and then continued 

indexing to 76.2 mm (3.0 in.) beyond the longitudinal weld centerline. The 152.4 mm (6.0 in.) of 

overlap in each scan provides two built-in fiduciaries in the scans as the longitudinal seam 

appears as a ―dead zone‖ in the scans. The data were stored electronically and analyzed using 

SAFT post-signal processing software.

The SAFT procedure used on the vertical pipe section (3C-022) was the same as used for 

processing the Phase 1 pipe specimens. The SAFT beam-angle used for the processing was 

12° and the beam entry diameter was 12.7 mm (0.5 in.). No skipping was used so all of the 

data was processed independent of threshold level.

6.3 Phased Array–UT

Recent advances in electronics miniaturization, computer processing capabilities, and 

fabrication methods for ultrasonic transducers have enabled phased-array technology to 

become a viable approach for many field applications. The geometrical design of a PA 

transducer is typically a function of specific implementation variables; that is, geometries such 

as linear, annular, circular, or matrix designs are developed to address a particular ultrasonic 

application need (Poguet et al. 2001). However, the basic premise for all PA transducers 

involves a set of small, individual piezoelectric elements that are independently driven. 

Although these elements may be pulsed individually, or in groups, to simulate conventional 

transducer excitation, the real strength of this technique lies in the capability of the system to 

electronically delay each of these elements during both generation and reception of ultrasonic 

sound fields. The wave-fronts produced by subsets of elements interfere within the inspected 

component to produce a resultant, phase-integrated ultrasonic wave. This is commonly referred 

to as beam forming. The PA system can, therefore, steer and focus the integrated ultrasonic 

beam within the component.

Generally, all PA ultrasonic systems are computer-controlled enabling software to define the 

groups and sequences of elements being electronically delayed. Parameters such as the 

number of individual elements in a virtual element, the delay sequence for firing of the virtual 

elements, element amplitudes, and the delays in reception are programmed into the system 

operating software. The setting of these generation and reception parameters for a particular 

response is called a focal law. Because practically all aspects of the sound beam are being 

controlled electronically, many iterations, or sequences of iterations, can be run in nearly real 

time. This allows a single array to inspect a component with variable inspection angles and 

focusing depths almost simultaneously. For instance, depending on the array design and the 

component thickness, a one-dimensional linear array, with major axis oriented normal to a pipe 

weld, may interrogate close to an entire planar cross section of the weld by sweeping through a 

series of inspection angles without having to mechanically move the transducer toward and 

away from the weld. Theoretically, an entire pipe weld can then be examined with a single 

circumferential scan motion. Most PA systems capture, digitize, and store the ultrasonic data, 

which enhances repeatability and permits off-line analysis and imaging. Specialized computer 

software and hardware (including upgraded RAM and hard disc memory) are needed to support 

analyses of PA data. Also, because of the variations of images that can be produced, a 
6-13

systematic method should be employed when performing these analyses by using a 

standardized procedure.

The PA system at PNNL consists of a Tomoscan III 32-channel instrument, produced off-theshelf by R/D Tech, Inc. (now ZETEC). This is a standard piece of commercial NDE equipment 

used for PA applications in several industrial sectors. This type of system is used by General 

Electric to inspect reactor internals and at the EPRI NDE Center for a variety of applications 

including far-side procedure development and initial true-state characterization of Performance 

Demonstration Initiative (PDI) specimens. The instrument can be programmed (by development 

of focal laws) to control up to 32 channels for transmission and 64 channels for reception of 

ultrasonic signals. It has 12-bit logic and operates through a local Ethernet connection to a 

standard desktop computer. Figure 6.5 shows the Tomoscan III and associated data acquisition 

computer used during the trials.

Figure 6.5 Phased Array Data Acquisition System

Two PA probes were used in the initial work, a transmit-receive longitudinal wave probe and a 

transmit-receive shear-wave probe. The TRL wave array was specially designed for near- and 

far-side applications in austenitic material ranging from about 13 to 38 mm (0.5 to 1.5-in.) in 

thickness. It consists of two pairs of 2D arrays, each array having 14 × 2 elements. One array 

pair is used for transmitting, the other for reception of ultrasonic signals. There exists a slight 

angular pitch, or roof angle, between the planes of the transmit and receive pairs. This provides 

mechanical focusing by setting the nominal cross-over point between transmit and receive rays 

at a certain depth in the material. Also, by using a pair of arrays for each transmit and receive 
6-14

segment, electronic lateral beam skewing is also possible so that the cross-over point may be 

slightly adjusted for a specific application, within the thickness range in this probe design. Both 

the transmit and receive arrays are inclined so that a nominal 50° longitudinal beam will be 

produced, if no focal law delays are introduced.

The array is integrally mounted on a wedge within the probe housing to produce the sound-field 

orientations described above. The overall contact cross section on the surface of the specimen, 

or footprint, of the probe has been made as small as possible to facilitate placement adjacent to 

the weld crown. This is approximately 25-mm (1.0-in.) square. A ―skirt‖ is fixed to the housing 

with tubes for applying a coupling medium (water was used for these trials). The TRL array 

probe has a nominal frequency of 2.0 MHz (1.9 MHz center) and is highly damped exhibiting 

approximately 70% bandwidth. The TRL array has an average wavelength of 3.1 mm (0.12 in.) 

in stainless steel. The aperture for each transmit and receive pair is 20 mm (0.79 in.) × 10 mm 

(0.39 in.). The pitch of elements in each array (distance between element center lines) is 

1.4 mm (0.05 in.). See Figure 6.6.

Figure 6.6 The 2.0 MHz TRL Phased-Array Probe

The transmit-receive shear wave (TRS) array consists of two 24-element 2D arrays in a similar 

arrangement as the TRL array. The TRS array has a roof angle that mechanically focuses the 

shear-wave beam by setting a cross-over point with a depth of 30 mm (1.18 in.). The TRS 

array, as with the TRL array, has a ―skirt‖ to allow for efficient coupling. The footprint of the 

probe is 50 mm (2.0 in.) along the active direction and 30 mm (1.18 in.) in the passive direction. 

The transmit and receive arrays are inclined to provide a nominal 55° shear wave in stainless 

steel if no focal laws are used to steer the beam. The system has a nominal frequency of 

2.0 MHz (2.14 MHz center) and has a −6 dB bandwidth of 65%. The TRS array has an average 

wavelength of 1.4 mm (0.06 in.) in stainless steel. The aperture for each transmit and receive 

pair is 32 mm (1.26 in.) in the active direction and 10 mm (0.39 in.) in the passive. The pitch of 

elements in each array is 1.33 mm (0.05 in.). See Figure 6.7.
6-15

Figure 6.7 The 2.0 MHz TRS Phased-Array Probe

Prior to the trials, focal laws were developed for the TRL and TRS arrays and programmed into 

the TomoView acquisition software. The focal laws were developed to provide ultrasonic 

longitudinal beam angles from 30° to 70°, at 1° increments for the TRL array in the piping 

specimens and from 40° to 70°, at 1° increments for the TRS array. This resulted in the sound 

field being swept through several discrete beam angles in near real time along the entire length 

of the linear scan. In other words, for each axially oriented cross section of material, data were 

acquired from 30° to 70° for the TRL and from 40° to 70° for the TRS, while the linear scans 

progressed circumferentially. 

For Phase 1 scans of the three segments of pipe specimen 02-24-15, each linear scan was 

approximately 480 mm (19.0 in.) in length to facilitate scanning while avoiding edge effects from 

the specimens. Large data files on the order of 45–60 megabytes resulted for each linear scan. 

This results in 50.8 to 76.2 mm (2.0 to 3.0 in.) of weld not available for scanning near the edges 

of each specimen. The Phase 2 scans on specimen 3C-022 went completely around the pipe 

specimen circumference, with 100 mm (4.0 in.) of overlap, and were approximately 2060 mm 

(81.0 in.) in length. These scans resulted in data files of 220–250 megabytes in size.

For both arrays and both specimens (in Phases 1 and 2), data were acquired by positioning the 

probe with the mechanical scanner adjacent to the weld (illustration shown in Appendix E, on 

page E.1); the sound beam was directed toward the weld, perpendicular to the weld centerline. 

Then linear scan motions were performed along the weld from both the near and far sides, with 

respect to the implanted crack locations. Several linear scans were performed from the far side; 

the first with the probe positioned as close as possible to the toe of the weld crown. Additional 

far-side scans farther from the weld centerline were also performed for several flaws.
6-16

For Phase 2 assessments on specimen 3C-022, each weld was scanned using both TRL and 

TRS phased arrays. The welds were scanned with the front edge of the arrays directly against 

the weld toe and with the front edge of the arrays 6.4 mm (0.25 in.), 12.7 mm (0.5 in.), and 

19 mm (0.75 in.) back from the weld toe. The scans were initiated at a distance of 51 mm 

(2.0 in.) back from the longitudinal seam (L-seam) weld, traveling circumferentially, 360° around 

the pipe, and continuing until the L-seam was scanned again. By running over the L-seam twice 

a beginning and end fiduciary was put into the data. This locator worked well, as the L-seam 

appeared as a ―dead‖ zone in the scan. Welds one and three were scanned using both the TRL 

and TRS arrays and were only scanned from the far-side. Weld two was scanned from both the 

near and far side using both arrays.

For the assessments carried out in Phase 3, an additional 1.5-MHz TRL wave PA probe was 

introduced into the data acquisition protocol. Therefore, in Phase 3, three different PA probes 

were employed—two TRL probes designed to operate at 1.5 MHz and 2.0 MHz, respectively, 

and a 2.0-MHz TRS wave probe. All three probes were specially designed for near- and farside applications in austenitic material ranging from approximately 13 to 38 mm (0.5 to 1.5 in.) 

in thickness. The 1.5-MHz probe is shown in Figure 6.8. The 2.0-MHz probes have integral 

wedges that were contoured to fit the appropriate pipe curvature for these specimens. The 

1.5-MHz probe was designed specifically for the far-side IGSCC inspection. It contains three 

elements in the secondary direction, which allows for improved lateral (side-to-side) beam 

skewing.

Figure 6.8 1.5-MHz TRL Probe, Configured with Gimbal and Couplant Lines

The 2.0-MHz TRL array was designed for near- and far-side applications in thinner pipe 

sections. This probe allowed for data collection in tight geometrical configurations. This smaller 

size generally provided insonification of the far side even with a weld crown present. Skew 

angles of 10 degrees were possible with this array. Skewing was not possible with the 

2.0-MHz TRS probe.
6-17

The 1.5-MHz TRL array consists of two, 3-element by 10-element matrix arrays with a 62% 

bandwidth at −6 dB. This array was designed with a non-integral wedge allowing change out of 

the wedge for inspecting pipes of varying diameters or flat plates. Its footprint is approximately 

50 by 50 mm (2.0 by 2.0 in.). The larger size and increased number of elements in the lateral 

direction provides improved beam forming and skewing but limits its application in tight 

geometrical conditions. This TRL array has a wavelength of 3.8 mm (0.15 in.) in stainless steel, 

at its average center frequency of 1.5 MHz. Skew angles of 10 and 20 degrees were 

possible with the larger number of elements in the secondary axis of this probe.

Focal laws were developed for the TRL and TRS arrays and programmed into the TomoView 

acquisition software. The focal laws were developed to provide ultrasonic longitudinal beam 

angles from 30° to 70° at 1° increments for the 2.0-MHz TRL array in the piping specimens, and 

from 40° to 70° at 1° increments for the 1.5-MHz TRL. Shear-wave focal laws were developed 

for the 2.0-MHz TRS array to provide beam angles from 40° to 70° at 1° increments. This 

resulted in the sound field being swept through many discrete beam angles in near real time at 

each position along the entire length of the linear scan. In other words, for each axially oriented 

cross section of material, data were acquired from 30° to 70° or from 40° to 70°, while the linear 

scans progressed circumferentially.

Sound-field modeling results are summarized in Figures 6.9 through 6.11 for the 1.5-MHz TRL, 

2.0-MHz TRL, and 2.0-MHz TRS probes in stainless steel. The left portion of each figure shows 

an idealized sound field as a function of depth, in a side view, for a 45° inspection angle. The 

designed focal depth is noted in the image by a black crossed circle. A top view of the sound 

field at the focal depth is shown in the right portion of each figure. Note that the top view color 

scheme has been normalized to show the full color range, white to red, in the view. Reference 

lines separated by 2 mm are displayed in Figures 6.9 through 6.11 to show scale. Theoretical 

spot sizes as measured from these simulations at the red level, approximately 3 dB points, are 

3.4 by 2.7 mm (0.13 by 0.11 in.) for the 1.5-MHz TRL probe, 5.0 by 4.3 mm (0.20 by 0.17 in.) for 

the 2.0-MHz TRL probe, and 2.9 by 1.7 mm (0.11 by 0.07 in.) for the 2.0-MHz TRS probe. The 

spot sizes at the yellow, approximately 6 dB points, are 9 by 6.9 mm (0.35 by 0.27 in.) for the 

1.5-MHz TRL probe, 11.2 by 9.0 mm (0.44 by 0.35 in.) for the 2.0-MHz TRL probe, and 7.2 by 

2.6 mm (0.28 by 0.10 in.) for the 2.0-MHz TRS probe.

These beam plot images show that the 1.5-MHz TRL and 2.0-MHz TRS arrays are operating 

near the center of their focal region at a depth of approximately 30 mm (1.18 in.). At this depth 

the 2.0-MHz TRL array is performing beyond its ideal focal plane as shown in the side view of 

the beam plot in Figure 6.10; however, in practice, good signals at the 30-mm depth were 

detected and recorded. The advantage of this 2.0-MHz TRL probe is its small footprint, which 

allows better sound field access to the far side of the weld in the presence of a limiting weld 

crown or other geometrical condition. On at least one of the EPRI pipe specimens, this was the 

only probe that could effectively collect data on both sides of the weld due to geometrical 

constraints.
6-18

Figure 6.9 TRL 1.5-MHz Beam Model at 45° with the Side View on the Left and Top View 

on the Right. The top view is at the focal depth of 30 mm. Vertical and 

horizontal lines with 2-mm separation are shown for reference.

Figure 6.10 TRL 2.0-MHz Beam Model at 45° with the Side View on the Left and Top View 

on the Right. The top view is at the focal depth of 30 mm. Vertical and 

horizontal lines with 2-mm separation are shown for reference.

Figure 6.11 TRS 2.0-MHz Beam Model at 45° with the Side View on the Left and Top View 

on the Right. The top view is at the focal depth of 32 mm. Vertical and 

horizontal lines with 2-mm separation are shown for reference.
6-19

Additional beam-field modeling for the 1.5-MHz TRL array shows the ability of the probe to form 

a coherent beam at skew angles up to 20°. The modeling results are shown in Figure 6.12. 

Skewed data can improve detection of IGSCC because it can be more sensitive to cracks 

branching out of the primary plane of the flaw and also be less sensitive to planar geometrical 

reflectors such as counterbore.

Figure 6.12 Beam Skewing Model for the 1.5-MHz TRL, 10 by 3 Element Probe

For the three probes and all Phase 3 IGSCC specimens, data were acquired with a manual 

scanner. The scanner was mounted on a curved section or ring positioned adjacent to the 

specimen. The 1.5-MHz TRL probe and scanner are shown in Figure 6.13. Data were acquired 

with the sound beam directed perpendicular to and pointing toward the weld, while the manual 

scanner was moved in the circumferential direction. The scanner encoder was calibrated to the 

specimen to give circumferential position information. Several linear scans were acquired from 

both the near side and the far side when specimen configuration allowed. A line scan with the 

probe positioned as close as possible to the weld crown center was acquired first. Additional 

scans with added offsets from this close position were then acquired from both sides when 

possible, typically at offset distances of 0.64 cm (0.25 in.) and/or 1.27 cm. (0.5 in.). The line 

scans proceeded clockwise with respect to the flow and the flow direction was noted by an 

arrow on the specimens. Skewed data were also acquired because the flaws were true 

IGSCCs and typically exhibited branching out of the plane of the flaw parallel or axial to the 

weld. Data obtained at skewed angles can be sensitive to the crack branching, and when 

combined with the normal data, may improve the flaw detection. Skew angles of ±10° and ±20° 

were used with the 1.5-MHz TRL probe. The linear TRS probe was not skewed and the 

2.0-MHz TRL probe was skewed at ±10°. Data were acquired such that each file contained only 

a single skew angle. The default skew was 0°.
6-20

Figure 6.13 Mechanical Scanner Mounted on a Ring Section in the Background with the 

Scanner Arm and Probe Extending to the Foreground (left) and Side View of 

the 1.5-MHz TRL Probe Next to a Weld (right)
7-1

7 Phase 1 Evaluation

The Phase 1 evaluation was focused on assessing the far-side inspection performance of three 

ultrasonic NDE techniques—automated conventional-UT, low-frequency/SAFT-UT, and phasedarray UT. The focus of the Phase 1 evaluation was aimed at far-side detection and characterization of fabricated thermal fatigue cracks and machined reflectors in the wrought austenitic, SS 

pipe specimen 02-24-15.

7.1 Conventional-UT: Data Analysis and Results

Automated conventional shear wave ultrasonic techniques were employed on the wroughtaustenitic stainless steel specimen 02-24-15 for comparison with the advanced phased-array 

and low-frequency/SAFT techniques. These scans differ from standard manual ultrasonic 

examinations of pipes in that the scans were taken using an automated scanner and the data 

were stored digitally. This method was chosen to provide images for a more direct comparison 

with the phased-array and LF/SAFT data. As described in Section 5, pipe specimen 02-24-15

consists of two 610-mm (24-in.) diameter, 36-mm (1.4-in.) thick pieces of ASTM A-358, 

Grade 304 vintage austenitic stainless steel. The specimen was cut into three segments to 

allow easier handling and scanning. Preliminary screening of data sets was performed 

immediately after scanning of the first pipe segment to determine which transducer(s) had the 

best signals and highest SNR performance. The 12.7-mm (0.5-in.), 1.5-MHz probe was 

identified as the transducer of choice for evaluation of conventional UT, reducing the total 

number of required scans significantly. Thus, 1.5-MHz, 60° and 70° shear-wave scans were 

completed on the near and far sides for pipe sections two and three of specimen 02-24-15. All 

of the B-scan end-views and C-scans are contained in Appendix C. Appendix C also contains 

the detailed analysis for each crack and saw-cut. Only results from the 1.5-MHz, 12.7-mm 

(0.5-in.) diameter, 60° and 70° shear-wave scans will be discussed in this section.

Figure 7.1 shows the C-scan images for pipe segment 1 taken using the 1.5-MHz probes and 

angles described above from the near and far side of the weld. Pipe segment 1 contains 

thermal fatigue cracks A (15% through-wall) and B (43% through-wall), and sawcuts A 

(7.5% through-wall straight cut), B (28.4% through-wall straight cut), and C (7.1% through-wall 

35° angled cut). To summarize the results, sawcut A was detectable from the far side with both 

60° and 70° shear waves. From the far side, with the 60° probe, sawcut A is detected as a 6 dB 

increase in the response versus the weld root response. Using the 70° probe from the far side, 

the response from sawcut A is not mixed in with the response from the weld root and is 11 dB 

above the average local noise. The 60° response from thermal fatigue crack A, which is only 

15% through-wall, is difficult to resolve from the weld root response. In contrast, the 70° shear 

waves are able to clearly detect crack A. Sawcut and crack B are very easily detected with the 

60° and 70° shear waves through the weld. Sawcut C gives a weak response and is only 

detectable from the near-side of the weld using 60° shear waves.
7-2

Figure 7.1 C-scan Views for Pipe Segment 1 from the Near and Far Side with the 

1.5-MHz, 60° and 70° Conventional Shear-Wave Probes

By comparing the gains used during scanning and the peak responses from the same reflectors 

when viewed on the near and far sides of the weld, it was determined that the weld metal 

absorbed or redirected 11 dB of 60° ultrasound (4 dB near-side vs. 15 dB far-side gains to 

obtain the same amplitude response from sawcut B) and virtually no absorption for the 70° 

ultrasound (17 dB gain for the same response from sawcut B from both near and far sides).

Figure 7.2 shows the near-side and far-side C-scan images for pipe segment 2 taken using the 

12.7-mm (0.5-in.), 1.5-MHz probe mounted on wedges that provide 60° and 70° shear waves in 

the material from the near and far side. Pipe segment 2 contains crack C (64% through-wall) 

and sawcuts D (19% through-wall straight cut), E (12% through-wall straight cut), and F 

(19% through-wall 35° angled cut). The term ―straight cuts‖ refers to sawcuts that were made 

90° to the inner surface, and the 35° angled cuts follow the HAZ as shown in Section 5, 

Figure 5.4. Sawcuts D, E, and F returned strong signals from the near and far side and were 

easily detected. This is to be expected, as they are all at least 19% through wall. Crack C, 

which is a 64% through-wall implanted thermal fatigue crack, was also easily detected.

The weld in segment 2 exhibited less sound absorption characteristics than segment 1. Once 

again, the 60° scan showed more absorption, with the 60° probe seeing a 4 dB drop through the 

weld and the 70° probe seeing only a 1 dB drop through the weld.
7-3

Figure 7.2 C-scan Views for Pipe Segment 2 from the Near and Far Side with the 

1.5-MHz, 60° and 70° Conventional Shear-Wave Probes

Figure 7.3 shows the near-side and far-side C-scan images for pipe segment 3 using the 

12.7-mm (0.5-in.), 1.5-MHz probe. Pipe segment 3 contains cracks D (19% through-wall axially 

oriented) and E (48% through-wall), and sawcuts G (18% through-wall 35° angled cut) and H 

(26% through-wall 35° angled cut). Crack E and sawcut H are easily detected by both probes. 

Sawcut G appears as a series of closely spaced smaller indications from the far side of the weld 

for both the 60° and 70° scans. Crack D, which is axially oriented, is not detected with any of 

the probes from the near or far side.

The sound absorption characteristics for the weld in segment 3 were similar to segments 1 

and 2. The 60° scan showed the strongest sound loss through the weld, with the 60° probe 

seeing a 7 dB drop through the weld and the 70° probe seeing a 4 dB drop through the weld.

Once it was determined which cracks could be detected, the next step was to length- and depthsize the flaws. Each detected flaw and sawcut was length-sized four times, once from each 

angle (60° and 70°), for the near and far side, and with two techniques, the 6 dB drop and loss 

of signal (LOS) methods. The length-sizing data is contained in Tables 7.1 through 7.4 and is 

represented graphically in Figures 7.4 and 7.5. To summarize, the 6 dB drop technique slightly 

undersizes the sawcuts but is very accurate for the implanted fatigue cracks. The LOS 

technique is very accurate for the sawcuts and slightly oversizes the implanted TFCs. 

The root mean square error (RMSE) values for each column were calculated following these 

steps:

 Subtracting each empirical value from its corresponding true-state value.

 Squaring each of these differences.
7-4

 Calculating the mean of the squared differences.

 Calculating the square root of this mean.

The formula summarizing these steps is 

where x = (empirical value – true-state value). The RMSE values provide a measure of the 

magnitude of deviation from true state.

The LOS length measurements have RMSE values ranging from 4 mm (0.16 in.) to 10.8 mm 

(0.43 in.), which is better than the 19.1-mm (0.75-in.) RMSE value required to satisfy ASME 

Code Section XI, Appendix VIII, length-sizing accuracy requirements. The 6 dB drop technique 

has RMSEs ranging from 2.2 mm (0.09 in.) to 19.4 mm (0.76 in.), and, while close, does not 

always meet ASME Code requirements. Note that the ASME sizing criteria are based on blind 

test results and the PNNL examinations are non-blind, thus represent more of a ―best case‖

scenario. 

Figure 7.3 C-scan Views for Pipe Segment 3 from the Near and Far Side with the 

1.5-MHz, 60° and 70° Conventional Shear-Wave Probes
7-5

Table 7.1 Sawcut Lengths in Pipe Specimen 02-24-15 as Measured by the 60° 

Conventional UT Technique

Near Side Far Side

Sawcut Actual Length 60° −6 dB 60° LOS 60° −6 dB 60° LOS

A 32.8 mm 17.8 mm 40.6 mm 29.2 mm 34.3 mm

(1.29 in.) (0.70 in.) (1.60 in.) (1.15 in.) (1.35 in.)

B 65.1 mm 52.1 mm 69.9 mm 45.7 mm 55.9 mm

(2.57 in.) (2.05 in.) (2.75 in.) (1.80 in.) (2.20 in.)

C 36.3 mm 15.2 mm 19.1 mm 8.9 mm 8.9 mm

(1.43 in.) (0.60 in.) (0.75 in.) (0.35 in.) (0.35 in.)

D 54.1 mm 36.8 mm 52.1 mm 41.9 mm 57.1 mm

(2.13 in.) (1.45 in.) (2.05 in.) (1.65 in.) (2.25 in.)

E 43.7 mm 31.7 mm 41.9 mm 36.8 mm 52.1 mm

(1.72 in.) (1.25 in.) (1.65 in.) (1.45 in.) (2.05 in.)

F 59.7 mm 39.4 mm 50.8 mm 48.3 mm 59.7 mm

(2.35 in.) (1.55 in.) (2.00 in.) (1.90 in.) (2.35 in.)

G 57.3 mm 52.1 mm 57.1 mm 43.2 mm 54.6 mm

(2.25 in.) (2.05 in.) (2.25 in.) (1.70 in.) (2.15 in.)

H 75.2 mm 39.4 mm 73.7 mm 52.1 mm 73.7 mm

(2.96 in.) (1.55 in.) (2.90 in.) (2.05 in.) (2.90 in.)

RMS Error 19.4 mm 7.7 mm 16.6 mm 10.8 mm

(0.76 in.) (0.30 in.) (0.65 in.) (0.42 in.)

Table 7.2 Sawcut Lengths in Pipe Specimen 02-24-15 as Measured by the 70° 

Conventional UT Technique

Near Side Far Side

Sawcut Actual Length 70° −6 dB 70° LOS 70° −6 dB 70° LOS

A 32.8 mm 40.6 mm 40.6 mm 29.2 mm 30.5 mm

(1.29 in.) (1.60 in.) (1.60 in.) (1.15 in.) (1.20 in.)

B 65.1 mm 48.3 mm 61.0 mm 48.3 mm 57.1 mm

(2.57 in.) (1.90 in.) (2.40 in.) (1.90 in.) (2.25 in.)

C 36.3 mm 47.0 mm 47.0 mm Not Not

(1.43 in.) (1.85 in.) (1.85 in.) detected detected

D 54.1 mm 41.9 mm 49.5 mm 35.6 mm 54.6 mm

(2.13 in.) (1.65 in.) (1.95 in.) (1.40 in.) (2.15 in.)

E 43.7 mm 29.2 mm 39.4 mm 24.1 mm 44.5 mm

(1.72 in.) (1.15 in.) (1.55 in.) (0.95 in.) (1.75 in.)

F 59.7 mm 47.0 mm 58.4 mm 47.0 mm 55.9 mm

(2.35 in.) (1.85 in.) (2.30 in.) (1.85 in.) (2.20 in.)

G 57.3 mm 30.5 mm 50.8 mm 35.6 mm 58.4 mm

(2.25 in.) (1.20 in.) (2.00 in.) (1.40 in.) (2.30 in.)

H 75.2 mm 57.1 mm 67.3 mm 62.2 mm 69.8 mm

(2.96 in.) (2.25 in.) (2.65 in.) (2.45 in.) (2.75 in.)

RMS Error 15.9 mm 6.5 mm 16.1 mm 4.0 mm

(0.63 in.) (0.26 in.) (0.64 in.) (0.16 in.)
7-6

Table 7.3 Implanted Thermal Fatigue Crack Lengths in Pipe Specimen 02-24-15 as 

Measured by the 60° Conventional UT Technique

Near Side Far Side

Crack Actual Length 60° −6 dB 60° LOS 60° −6 dB 60° LOS

A 10.7 mm 12.7 mm 17.78 mm 11.43 mm 12.7 mm

(0.42 in.) (0.42 in.) (0.70 in.) (0.45 in.) (0.50 in.)

B 30.5 mm 31.75 mm 41.91 mm 26.67 mm 30.48 mm

(1.20 in.) (1.25 in.) (1.65 in.) (1.05 in.) (1.20 in.)

C 43.6 mm 39.37 mm 54.61 mm 45.72 mm 57.15 mm

(1.72 in.) (1.55 in.) (2.15 in.) (1.80 in.) (2.25 in.)

E 33.8 mm 30.48 mm 50.8 mm 34.29 mm 49.53 mm

(1.33 in.) (1.20 in.) (2.00 in.) (1.35 in.) (1.95 in.)

RMS Error 2.94 mm 12.15 mm 2.23 mm 10.43 mm

(0.12 in.) (0.48 in.) (0.09 in.) (0.41 in.)

Table 7.4 Implanted Thermal Fatigue Crack Lengths in Pipe Specimen 02-24-15 as 

Measured by the 70° Conventional UT Technique

Near Side Far Side

Crack Actual Length 70° −6 dB 70° LOS 70° −6 dB 70° LOS

A 10.7 mm 13.97 mm 16.51 mm 12.7 mm 15.24 mm

(0.42 in.) (0.55 in.) (0.65 in.) (0.50 in.) (0.60 in.)

B 30.5 mm 27.94 mm 36.83 mm 26.67 mm 30.48 mm

(1.20 in.) (1.10 in.) (1.45 in.) (1.05 in.) (1.20 in.)

C 43.6 mm 41.91 mm 55.88 mm 44.45 mm 52.07 mm

(1.72 in.) (1.65 in.) (2.20 in.) (1.75 in.) (2.05 in.)

E 33.8 mm 33.02 mm 46.99 mm 34.29 mm 41.91 mm

(1.33 in.) (1.30 in.) (1.85 in.) (1.35 in.) (1.65 in.)

RMS Error 2.28 mm 9.98 mm 2.22 mm 6.29 mm

(0.09 in.) (0.39 in.) (0.09 in.) (0.25 in.)
7-7

Figure 7.4 Sawcut Lengths Measured Using the 6 dB Drop and Loss of Signal 

Techniques with the 1.5-MHz 60° and 70° Probes from the Near and Far Side 

of the Weld Versus the Actual Sawcut Lengths
7-8

Figure 7.5 Thermal Fatigue Crack Lengths Measured Using the 6 dB Drop and Loss of 

Signal Techniques with the 1.5-MHz 60° and 70° Probes from the Near and 

Far Side of the Weld Versus the Actual Crack Lengths

Depth-sizing the flaws proved to be much more difficult than length-sizing. While identifying the 

corner trap signals in the data was not difficult, no tip signals could be positively identified in any 

of the scans. When looking at the B-scan end views, no crack tip energy could be discerned. 

The A-scans associated with each flaw were then analyzed for possible crack tip responses. 

Some of the A-scan images in Appendix C contain what look like crack tip signals. Standard 

techniques were applied to each of these crack tip indications to determine the depth of the 

crack based on these possible tip signals. Many of the potential tip signals did not have the 

proper echo-dynamic response that is characteristic of crack tips. Therefore, it was determined 

that no crack tip signals were present in the data. These apparent crack tips are the result of 

high noise from the coarse-grained weld metal and the weld root response.
7-9

It is important to note that it was challenging to know that these apparent crack tips were merely 

noise. Figure 7.6 shows the depths calculated using the 6 dB drop technique and using the 

apparent crack tips. Figure 7.6 includes calculated depths for both sawcuts and cracks. The 

6 dB drop depths tended to size near 5 mm (0.2 in.), independent of the actual crack depth. 

The depths calculated from the apparent tip signals have no relation to the actual crack depth, 

and have an RMSE of 5 mm (0.2 in.).

Figure 7.6 Sawcut and Crack Depths Measured Using the 6 dB Drop and Crack Tip 

Techniques with the 1.5-MHz 60° and 70° Probes from the Near and Far Side 

of the Weld Versus the Actual Sawcut and Crack Depths

Figure 7.7 shows the corner-trap SNR for the four detected cracks (A, B, C, and E) in Phase 1 

and the three cracks (cracks 4, 5, and 6; 5%, 10%, and 15% through-wall, respectively) from 

Phase 2 plotted versus the crack depths. The corner-trap SNR for the cracks on the far side of 

the weld ranged from 2–5, with no strong relationship between the crack depth and the SNR.
7-10

Figure 7.7 Signal-to-Noise Ratio Plotted Against the Crack Depth for the 60° and 70° 

Far-Side Scans

7.2 Low-Frequency/SAFT: Data Analysis and Results

Data analysis was conducted using an established protocol and referencing specific crack 

identification and length-sizing criteria. Within any given planar view (B-scan side view, B-scan 

end view, or C-scan view), the analysis focused on an established criteria for identifying or 

rejecting regions as cracked in the material, where ultrasonic signal amplitude anomalies 

occurred. Signal amplitude anomalies (SAAs) were used as the primary ultrasonic markers for 

determining detection capability. The criteria described here are based upon acquiring SAAs 

above an established amplitude threshold in the ultrasonic data as a function of spatial position 

on the image (axial and circumferential position relative to the true-state). SAAs occurring in the 

volumetric space surrounding the weld root were fully insonified and examined from the far side. 

SAAs occurring outside of this region were not examined as part of this study.

The following criteria were used in the assessment of SAAs detected in the SAFT-processed 

images. SAAs may occur above, on, or below the back-surface line appearing on the SAFTprocessed images. Differences in position of ±0.51 cm (±0.2 in.) above or below this line are 

not deemed significant, and SAA position will be measured from the peak amplitude point. 

When an SAA occurs in this region, its entire shape will be included in the boxed examination 

region of the image. When the back-surface line is positioned accurately, corner trap signal 

returns from cracking in the material will result in SAAs that lie just on the back-surface line, with 

overlap of the SAA above and below the line. The accuracy of the back-surface line is a 
7-11

function of the material velocity and transducer delay, and the actual position of a SAA relative 

to this line is a function of incident angle, acoustic velocity, frequency, wavelength, and zone 

focal dimensions of the transducer. Because material velocity varies with spatial position and 

incident angle in the material, inaccuracies must be accommodated. The specimen 

documentation data were used to determine nominal wall thickness values. SAAs should have 

some degree of characteristic shape (elliptical) to them, with somewhat smooth contours on 

their edges, as opposed to random, blotchy, scattered amplitude blips that appear with little 

symmetry and rough contours. Also, SAAs should have reasonable and proper orientation with 

respect to the examination incident angle of the insonifying beam. A perpendicular orientation is 

normal for opposite surface-breaking defects. Once again, differences in lateral position of 

±0.51 cm (±0.2 in.) or less are not deemed significant. 

The 12° beam processing angle was used for analysis purposes here, as processing angles 

greater than 12° did not significantly improve image enhancements and processing angles lower 

than 12° did not provide the necessary focusing and SNR enhancements to the imaged data. 

The higher the angle, the more time-intensive the SAFT process becomes. So the decision to 

stay with a 12° beam processing angle for all data sets was made early in the study. The SAFT 

data analysis protocol used for identifying SAAs was initiated by generating projected B-scan 

side view images and correlating the weld centerline position and back-surface line position to 

the image based upon component thickness information and scan parameter settings relative to 

transducer spatial position on the component. In this manner, the analyst uses positional 

information to accurately employ the SAFT-UT "boxing" feature to magnify SAAs in the area of 

interest. It should be noted that the SAFT presentation images are defaulted to use dimensions 

in English units only.

Next, the composite B-scan end view was generated to provide an orthogonal perspective that 

yields the circumferential perspective, and again the back-surface line position was used for 

analysis of the image as well. The analyst can then crop and magnify SAAs from both views 

and analyze characteristics and features as a function of view perspective, position relative to 

the weld centerline and back-surface line, orientation to the sound field incident angle, 

characteristic shape, etc. At this point, the various longitudinal- and shear-wave data sets were 

analyzed and features were identified. SAAs and pertinent commonalities between 

perspectives for determining root condition and geometrical reflectors were noted. SAAs 

present in the various views were then further reduced using both 6 dB and 3 dB drop methods 

(clipping the view’s dynamic range by this value) and length sizing was then recorded using the 

6 dB drop and LOS methods. The 3 dB clipping was only used to further reduce background 

and examine redundancy between modes where only the most prominent high-amplitude 

responses were evident. Depth sizing was initially attempted, but calculations based upon 

known depths of sawcuts could not be correlated with any method as tip-diffracted signals were 

not detected. Amplitude drop methods were ineffective because the sound-beam dimensions 

were always larger than the flaws being examined. Depth information was not recorded for this 

inspection technique.

It should be noted that the dynamic range of each scan was configured by translating the 

transducer over the area containing the introduced flaw that provided the largest amplitude 

responses and setting system gain and A/D sensitivity values accordingly. This provided a 

means to use the full dynamic range of the digitizer without saturating the highest amplitude 
7-12

responses. However, each scan varied regarding system gain, resulting in variable ranges for 

total system gain between scans. In general, the results are consistent between all SAFT 

analyses in that the images are always normalized to the signal responses of interest within the 

data set being evaluated. All signal-to-noise calculations were conservatively made using the 

highest amplitude background noise levels from nearby zones of the specimen being evaluated.

The summary analysis documentation for each notch, crack, and sawcut as a function of 

inspection modality, incident angle, and frequency from the far side is provided in Appendix D.

In the first phase of this evaluation, the low-frequency/SAFT inspection technique was employed 

for examination of a sectioned, Grade 304 WSS pipe (specimen 02-24-15) containing an 

austenitic weld. This specimen was sectioned into three pieces, with each 120° segment of the 

full pipe section laid on a platform in horizontal fashion and secured for scanning with the 

automated pipe scanner using a magnetic scanner track. In Section 6, the photograph on the 

right hand side in Figure 6.2 illustrates the Phase 1 configuration for data acquisition.

Here in this section, a tutorial is provided for the reader to follow and understand the process 

used in compiling detection results from the data and interpreting the SAFT-processed images. 

Because this was an iterative process that was undertaken for each individual SAA occurring in 

the SAFT-processed images, the reader is directed to Appendix D for the large volume of data, 

images, and spreadsheets for detailed information on the remaining analyses. A discussion of 

the results of this study will be described after the tutorial provided here.

7.2.1 Tutorial on SAFT Data Analysis Process

This tutorial uses images, positional data (axial, circumferential, and depth values), and the 

detection and length-sizing criteria described earlier in the document as the basis for the 

analysis procedure. The example provided for the tutorial is one of the benchmark data sets 

that were used for verification of scanning parameters, procedures, and other variables. In this 

case, the 400-kHz, 45° shear-wave inspection of piping segment 2, from the near side, was 

used for illustrating the procedures, imaging combinations, and calculations employed for 

conducting the analysis.

The raw, unprocessed ultrasonic C-scan image is analyzed first. This top-view visual is a 

composite view of the peak-amplitude signal responses in the X-Y plane. From this, any 

evidence of transducer de-coupling, electronics failures, or other problems with the data 

acquisition system can immediately be identified, and the data can be re-taken if necessary. 

Also, the operator looks for signs of signal amplitude saturation, and when satisfied with the 

integrity of the ultrasonic data, the data set is then SAFT-processed using a 12° beam 

processing angle. From an analysis of the SAFT-processed C-scan image, SAAs are then 

identified and isolated using a boxing feature to segregate and magnify SAAs of interest. This 

C-scan analysis process is visually illustrated in Figure 7.8. For this tutorial, the focus will be 

directed at the SAA highlighted in the yellow box.
7-13

Figure 7.8 C-scan Analysis Process. From left to right, the unprocessed ultrasonic data 

in C-scan view. Next, the 12° SAFT-processed data in C-scan view. Finally, 

the processed C-scan view depicting an isolated SAA using the “box” 

feature embedded in the SAFT analysis menus.

From these initial views, information relative to axial position and length information can be 

retrieved. Also, application of specific rules defined in the detection and length-sizing criteria 

can begin to be evaluated.

Next, a similar process using the B-scan end-view (sometimes referred to as a D-scan view) is 

then employed. The B-scan end-view is an orthogonal planar representation of the B-scan sideview and represents a composite view of the peak-amplitude signal responses in the Y-Z plane. 

This process allows the analyst the opportunity to view the raw unprocessed data, the 12° 

SAFT-processed data, and the isolated indications of interest (SAAs) from a perspective that 

highlights the depth information while maintaining the circumferential aspect of the indications. 

Quite often these views are examined simultaneously in a side-by-side fashion to better 

understand the enhancement and reduction of various features between the processed and 

unprocessed data sets. Using these images, the detection and length-sizing criteria is again 
7-14

reviewed and applied to the SAAs, and magnification of the isolated indications is performed for 

a more detailed analysis. Figure 7.9 illustrates the B-scan end-view analysis. Again, the SAA 

identified by the yellow box (and corresponding to the same SAA discussed in the C-scan 

analysis and shown in Figure 7.8) is the focus of this discussion.

Figure 7.9 B-scan End-View Analysis. From top to bottom, the unprocessed ultrasonic 

data in B-scan end-view. Next, the 12° SAFT-processed data in B-scan endview. Finally, the processed B-scan end-view depicting an isolated SAA 

using the “box” feature embedded in the SAFT analysis menus.

After review and application of pertinent detection and length-sizing criteria (SAA shape, 

orientation, and position), the magnified B-scan end-view is examined. This view is analyzed 

against the criteria established for SAA shape, spatial position, and other parameters defined in 

the detection and length-sizing criteria, and a determination is made as to the status of this 

SAA. If the SAA is determined to represent a crack (crack, notch, or sawcut in this case), then 

signal-to-noise measurements are computed, length sizing is conducted using both LOS and a 

−6 dB drop method, and spatial position is recorded from interactively placing the cursors at 
7-15

specific locations on the SAA and recording the X, Y, and Z coordinates respectively. The 

resultant image is cut and pasted into an Excel spreadsheet and length sizing and signal-tonoise calculations are recorded in this fashion. The magnified B-scan end-view embedded on 

the Excel spreadsheet with the associated calculations is shown in Figure 7.10.

Figure 7.10 Excel Spreadsheet and Associated Magnified B-scan End-View for S

Analysis and Detailed Calculations for Length Sizing and Recording of 

Positional Data

From the various spreadsheets that contain each SAA and the associated calculations, tables 

are then created where the resultant detection and length-sizing data can be reviewed and 

compared to true-state conditions. For a comprehensive review of each Excel datasheet for all 

SAAs recorded in this study, please refer to Appendix D. As described earlier, the analysis 

process is an iterative activity, where the analyst must conduct this detailed procedure for each 

SAA that satisfies the detection and length-sizing criteria, documenting the results as discussed 

here. The data in the Excel spreadsheet is then translated into tabular format and compared 

with results from other incident angles, examination frequencies, and wave modalities. For this 

particular SAA, the measured and actual values for this sawcut were determined as shown in 

Table 7.5 where the measured length is shown for the −6 dB drop calculation and the LOS 

calculation.
7-16

Table 7.5 Actual Versus Measured Values for Detection Results of Sawcut F from a 

Tutorial-Based Analysis of the 400-kHz, 45° Shear-Wave Inspection of Piping 

Segment #2, from the Near Side

Piping 

Segment 

ID#

Flaw or 

Crack ID#

Actual 

Length

Actual

Depth

Measured 

Length

−6 dB/LOS

Measured 

Depth

(%TW)

Signalto-Noise 

Ratio

Segment #2 Sawcut F 5.97 cm

(2.35 in.)

19%TW

Angled @35°

5.08 cm/6.10 cm

(2.0 in./2.4 in.)

Not 

determined

15.8 dB

This section describes the results from the LF/SAFT inspection of austenitic piping specimen 

02-24-15, providing the detection and length-sizing data as a function of examination frequency, 

incident angle, wave modality, and scan orientation relative to the introduced flaw. For a review 

of the true-state information for all flaws and sawcuts, the reader is referred to Section 5, 

Tables 5.1 and 5.2 of this report. The following sub-sections will discuss the results by piping 

segment. 

The data acquired were first qualitatively evaluated by determining the ―quality‖ of the data set 

for each frequency, angle, wave mode, and orientation to the introduced flaw. Each data set 

was reviewed and labeled as good, marginal, or poor, relative to the specific quality factor of the 

data. The quality factor was based on SNR, evidence of coupling problems in the images, 

saturation of signal responses, scattered ultrasonic energy from the austenitic weld structure, 

and the prevalence of SAAs attributed to geometrical reflectors, weldment structure, and other 

features other than introduced flaws.

After this initial qualitative assessment, the detailed analysis procedure discussed in the tutorial 

section was employed and the detection determinations, length-sizing data, and SNR

calculations were recorded in an Excel spreadsheet for subsequent comparative analysis. 

Although near-side data were acquired using both modalities, both frequencies, and all incident 

angles, this section will only focus on inspection results acquired from the far-side scan 

orientation. Near-side data were used for benchmarking purposes only, as the objectives of this 

study are directed at evaluating the detection performance of inspection methodologies for farside examinations through austenitic welds.

7.2.2 Piping Segment 1

Piping segment 1, defined as the first 120° segment of pipe spanning from the 0° datum to the 

120° point, consists of five implanted fatigue cracks and machined sawcuts. In circumferentially 

ascending order, these are sawcut A, crack A, sawcut B, crack B, and sawcut C. A discussion 

of the resultant detections and flaw characterization data is best described in tabular format. 

Table 7.6 depicts the far-side detection calls for piping segment 1 along with a grading for data 

quality and flaw characterization data as a function of incident angle and wave mode for 

250-kHz inspections. Table 7.7 depicts the same information for inspections acquired at 

400-kHz. The nomenclature used for deciphering the tables is described here. The split 

column beneath each incident angle and wave mode describes the measured −6 dB lengthsizing value of the flaw in centimeters, the length determined by using LOS in centimeters, and 
7-17

the measured SNR in decibels (dB). The column labeled with the ―Q‖ designates the preassigned quality of the data set defined earlier in the discussion. A ―G‖ represents good quality, 

an ―M‖ represents marginal quality, and a ―P‖ represents poor quality. Good quality data sets 

provided the analyst with the best opportunities for detecting the introduced flaws. Marginal 

data sets contained anomalies in the data and higher background noise levels that introduced a 

reasonable level of difficulty for interpreting the SAFT-processed data. Poor data sets resulted 

in SAFT-processed images that could rarely be used to extract useful detection, localization, or 

sizing information; and these data sets exhibited poor resolution, unusually high incidence of 

SAAs that were not attributed to introduced flaws, and generally very high background noise 

levels. Cells containing ―--" indicates no data were acquired for that particular combination of 

scan parameters, and ―ND‖ indicates that the introduced flaw was not detected.

Analysis of the far-side, 250-kHz data using the 70° L-wave modality indicated that the quality of 

the resultant images were marginal as the SNR was generally lower than anticipated. Good 

separation of signal responses from counterbore geometry is illustrated from the data, and this 

modality and frequency combination provided the capability to clearly detect three of the five 

introduced flaws in this piping segment. Sawcut B, crack B, and sawcut C were detected and 

sized in the length dimension. The computed SNR for these detectable introduced flaws ranged 

from approximately 4 dB to nearly 12 dB. From a review of the data in Table 7.6, the measured 

length-sizing data correlates very well with the true-state values for these introduced flaws. 

C-scan and B-scan end-view images as well as the individual data sheets for each detected 

flaw can be found in the Appendix D.

Analysis of the results from the 45° L-wave scans were of marginal quality due to signal 

response saturation of significant portions of the volume of interest. The total system gain is 

pre-selected based upon a quick preliminary manual scan of the largest response amplitude in 

the particular specimen being evaluated. On occasion, this manual scan procedure for 

detecting the highest amplitude signal response is not accurate, and results in a higher total 

system gain value; in turn, yielding areas of saturation that were not anticipated. The data in the 

45° L-wave scans did not provide suitable information or clarity for detections to be made for 

any of the five introduced flaws in this piping segment.
7-18

Table 7.6 250-kHz, Far-Side Detection Results for Piping Segment 1

Introduced 

Flaw

70° L 60° L 45° L 70° S 60° S 

−6 dB (cm), 

LOS (cm),

SNR (dB) “Q”

−6 dB (cm), 

LOS (cm), 

SNR (dB) “Q”

−6 dB (cm), 

LOS (cm), 

SNR (dB) “Q”

−6 dB (cm), 

LOS (cm), 

SNR (dB) “Q”

−6 dB (cm), 

LOS (cm), 

SNR (dB) “Q

Sawcut A ND M -- -- ND M -- -- -- -

Flaw A ND M -- -- ND M -- -- -- -

Sawcut B 6.35, 6.60, 11.4 M -- -- ND M -- -- -- -

Flaw B 3.05, 3.05, 5.1 M -- -- ND M -- -- -- -

Sawcut C 3.30, 3.81, 3.5 M -- -- ND M -- -- -- -

Table 7.7 400-kHz, Far-Side Detection Results for Piping Segment 1

Introduced 

Flaw

70° L 60° L 45° L 70° S 60° S 

−6 dB (cm), 

LOS (cm),

SNR (dB) “Q”

−6 dB (cm), 

LOS (cm), 

SNR (dB) “Q”

−6 dB (cm), 

LOS (cm), 

SNR (dB) “Q”

−6 dB (cm), 

LOS (cm), 

SNR (dB) “Q”

−6 dB (cm), 

LOS (cm), 

SNR (dB) “Q

Sawcut A ND G ND M 3.56, 4.57, 

6.0

G ND P ND P 

Flaw A ND G ND M ND G ND P ND P 

Sawcut B 3.81, 4.32, 8.3 G 3.05, 3.30, 

10.4

M 7.87, 8.89, 

10.6

G ND P 4.32, 5.59, 

7.2

P 

Flaw B 2.79, 2.79, 5.6 G 1.52, 2.03, 

6.4

M ND G ND P 2.79, 2.79, 

8.6

P 

Sawcut C ND G ND M ND G ND P ND P 
7-19

Analysis of the 400-kHz data using the 45°, 60°, and 70° L-wave combinations indicates that 

although the quality of the data were generally good, (marginal for the 60° data set), the SNR 

was again lower than anticipated. Some images illustrate a higher degree of clarity for detection 

of crack A and sawcut B in the unprocessed data sets, as the SAFT process at times would 

smear signal responses from indications or otherwise reduce their relative signal strength in the 

processed images. This is because the inspection probe does not employ a divergent sound 

field, but rather is a large zone-focused sound field and the flaws are all smooth specular 

reflectors. The analysis also illustrates that higher L-wave angle inspections (60° and 70°) 

result in less attenuation of the sound field and reduced signal responses from counterbore, 

weld root, and other structure that contributes to the ―noise‖ in the images. The 45° L-wave data 

resulted in the detection of sawcut A and sawcut B, while the 60° and 70° L-wave inspections 

resulted in the detection of sawcut B and flaw B. Three of the five introduced flaws went 

undetected for any given data set using the L-wave mode. The computed SNR for the detected 

introduced flaws ranged from approximately 6 dB to nearly 11 dB. In reviewing the measured 

length values for the 45° L-wave data, the −6 dB drop method provided close correlation with 

the true-state length for sawcut A, while slightly oversizing sawcut B. The 60° L-wave

inspection data provided undersized length values for both sawcut B and crack B, while the 70°

L-wave data undersized sawcut B but correlated well with the true-state value for crack B. The 

reader is directed to view any composite C-scan, B-scan end-view, or magnified/isolated views 

of SAAs in Appendix D.

The data quality for the 400-kHz shear-wave data, specifically the 60° and 70° scans, were 

poor. The unprocessed images were very grainy and exhibited a generally noisy background 

where little clarity was provided. In particular, the SAFT-processed 70° data provided no useful 

information for resolving or detecting SAAs that even remotely met the detection criteria. SAFT 

processing of the 60° data improved the image quality, enabling separation of SAAs and 

positional criteria to be used in the detection of sawcut B and crack B. The 45° shear-wave data 

exhibited good clarity, separation, low background noise levels, and was regarded as a good 

quality data set. Detection criteria were employed for the detection and sizing of three of the 

introduced flaws in this piping segment, including sawcut A, sawcut B, and crack B. The 

computed SNR for the two detected introduced flaws at 60° ranged from 7 dB to nearly 9 dB, 

while the SNR for the three detected introduced flaws at 45° ranged from 4 dB to approximately 

16 dB. A review of the −6 dB measured values for length show that the 60° data resulted in 

undersizing sawcut B and slightly under-sizing crack B, while the 45° data oversized sawcut A 

and sawcut B but correlated well with the true-state value for crack B.

From a detection performance standpoint, it is interesting to note the specific introduced flaws 

that were not detected or those that were only detected at a specific angle. From the far-side, 

flaw A went undetected throughout the evaluation. Sawcut A was only detected at the 45°

incident angle, but was detected by both longitudinal and shear wave modalities. Sawcut C was 

only detected with the 250-kHz, 70° L-wave inspection. As expected, the two deepest flaws, 

sawcut B and crack B, were detected most often across the range of angles, modalities, and 

frequencies employed.
7-20

7.2.3 Piping Segment 2

Piping segment 2, defined as the second 120° segment of pipe spanning from the 121° point to 

the 240° point, consists of four introduced flaws consisting of implanted fatigue cracks and 

machined sawcuts. In circumferentially ascending order, these introduced flaws are sawcut D, 

crack C, sawcut E, and sawcut F. The discussion of the resultant detections and flaw 

characterization data continues here based on the tabular format. Table 7.8 depicts the far-side 

detection calls for piping segment 2 along with a grading for data quality and flaw characterization data as a function of incident angle and wave-mode for 250-kHz inspections. Table 7.9

depicts the same information for inspections acquired at 400-kHz. The same nomenclature as 

described previously is used for deciphering the tables discussed here.

Analysis of the far-side, 250-kHz data using the 70° L-wave modality indicated that the quality of 

the resultant images were marginal as the SNR was again lower than anticipated. At the higher 

incident angles, particularly 60° and 70°, ultrasonic signal scatter and anomalous amplitude 

responses from component geometry and weld structure should be reduced. On occasion, 

resultant images sometimes illustrate unusually high background noise levels, higher than 

anticipated prevalence of non-flaw–related SAAs, or consistent signal responses that appear to 

result from weld structure such as counterbore or weld root. Typically, these conditions are not 

anticipated at the higher angles of inspection and often can only be attributed to the 

microstructure or attenuative characteristics of the austenitic weld.

Reasonable separation of signal responses from background noise levels can be seen from the 

data at 70°, and this modality and frequency combination provided the capability to detect and 

length-size all four of the introduced flaws in this piping segment. The computed SNR for these 

detectable introduced flaws ranged from approximately 4.5 dB to nearly 10 dB. From a 

comparative review of the data in Table 7.8 with that of the true-state values for length, it is 

shown that the −6 dB length-sizing method resulted in slightly oversizing sawcut D and 

sawcut E, and more significantly undersized crack C and sawcut F. Again, the reader is 

referenced to the C-scan and B-scan end-view images as well as the individual data sheets for 

each detected introduced flaw in Appendix D.

Analysis of the results from the 45° L-wave scans at 250 kHz were of poor quality due to higher 

background noise levels and less clarity and definition for the lower amplitude SAAs that met 

the detection criteria. This data set did not provide suitable information for detecting sawcut D, 

crack C, or sawcut E. The only detectable introduced flaw in the processed data was the SAA 

corresponding with sawcut F. The SNR for this call was approximately 10 dB and the measured 

length for this SAA undersized the true-state length.
7-21

Table 7.8 250-kHz, Far-Side Detection Results for Piping Segment 2

Introduced 

Flaw

70° L 60° L 45° L 70° S 60° S 

−6 dB (cm), 

LOS (cm),

SNR (dB) “Q”

−6 dB (cm), 

LOS (cm), 

SNR (dB) “Q”

−6 dB (cm), 

LOS (cm), 

SNR (dB) “Q”

−6 dB (cm), 

LOS (cm), 

SNR (dB) “Q”

−6 dB (cm), 

LOS (cm), 

SNR (dB) “Q

Sawcut D 5.84, 9.02, 6.0 M -- -- ND P 3.30, 3.30, 

4.6

M 2.29, 3.30, 

6.0

M 

Flaw C 2.54, 3.43, 6.0 M -- -- ND P ND M 4.32, 5.08, 

5.6

M 

Sawcut E 4.57, 5.21, 4.6 M -- -- ND P ND M ND M 

Sawcut F 4.19, 5.84, 9.5 M -- -- 4.83, 4.83, 

9.8

P ND M ND M 

Table 7.9 400-kHz, Far-Side Detection Results for Piping Segment 2

Introduced 

Flaw

70° L 60° L 45° L 70° S 60° S 

−6 dB (cm), 

LOS (cm),

SNR (dB) “Q”

−6 dB (cm), 

LOS (cm), 

SNR (dB) “Q”

−6 dB (cm), 

LOS (cm), 

SNR (dB) “Q”

−6 dB (cm), 

LOS (cm), 

SNR (dB) “Q”

−6 dB (cm), 

LOS (cm), 

SNR (dB) “Q

Sawcut D ND P ND P ND G ND P 3.81, 3.56, 

6.0

P 

Flaw C ND P ND P ND G ND P 4.83, 5.59, 

11.8

P 

Sawcut E ND P ND P ND G ND P ND P 

Sawcut F ND P ND P 4.83, 5.33, 

17.2

G ND P ND P 
7-22

Analysis of the far-side, 250-kHz data using the shear-wave modality is discussed here. The 

70° incident angle demonstrated marginal data quality due to the high amount of background 

noise and lack of SAAs meeting the detection criteria. Reasonable separation of only one SAA 

from background noise levels can be seen from the data, and this modality and frequency 

combination provided the capability to detect and length-size only sawcut D in this piping 

segment. The computed SNR for this introduced flaw was approximately 5 dB. Both the −6 dB 

and LOS length-sizing methods resulted in significantly undersized measured lengths for this 

flaw from the true-state value. The 60° incident angle also provided marginal data quality due to 

image graininess, higher than anticipated background noise levels, and a lack of SAAs meeting 

the detection criteria. However, the 60° shear wave did provide suitable information and 

separation of SAAs for the detection of both sawcut D and crack C, but did not detect sawcut E 

or sawcut F. The computed SNR for the two detections was approximately 6 dB. The 

measured length for sawcut D was significantly undersized, while the measured length for crack 

C correlated quite well with the true-state length (see Section 5, Tables 5.1 and 5.2). The 45°

incident angle provided good quality data where the SAAs were easily separated and the 

background noise levels were reasonably low relative to the SAA peak amplitudes in the 

processed data. A good degree of clarity and SAA separation were provided allowing for the 

detection of all introduced flaws in this piping segment with the exception of sawcut E. The 

measured SNR values for these SAAs ranged from 7 dB to over 13 dB. Sawcut D and sawcut F 

were both undersized in length, while both −6 dB and LOS length-sizing methods oversized 

crack C.

Analysis of the 400-kHz data using the 60° and 70° L-wave combinations indicates that the 

quality of the data at both incident angles was poor, and the SNR was again lower than 

anticipated. These data exhibited higher background noise levels and areas of signal saturation 

from material and/or weldment structure. The 60° data also exhibited a large area where the 

transducer had lost coupling, either as a result of physical de-coupling of the wedge from the 

part surface or from a malfunction of the coupling drip lines. From an analysis of these data and 

by employing the detection criteria, it is shown that neither of these scans provided suitable 

information for detection of any introduced flaws in this piping segment. The 45° L-wave 

incident angle provided good quality data and clearly allowed for the detection of one SAA. The 

background noise levels in the volume of interest were lower than anticipated, and the clarity 

and continuity of the signal responses from weldment geometry and material structure were not 

difficult to discriminate from the SAA of the introduced flaw. The computed SNR for this SAA 

was nearly 18 dB. This data set provided a detection of only one of the introduced flaws in this 

piping segment, that being sawcut F. In reviewing the measured length values for the 45°

L-wave data, both the −6 dB drop and LOS length-sizing methods resulted in undersizing the 

true-state length for sawcut F.

The data quality for the 400-kHz shear-wave data, specifically the 60° and 70° scans, were 

poor. The unprocessed images were very grainy and exhibited a generally noisy background 

where little clarity was provided. In particular, the SAFT-processed 70° data provided no useful 

information for resolving or detecting SAAs that even remotely met the detection criteria. SAFT 

processing of the 60° data improved the image quality, enabling separation of SAAs and 

positional criteria to be used in the detection of sawcut D and crack C. The 45° shear-wave 

data exhibited marginal quality as the clarity, separation, and lower background noise levels 
7-23

provided a suitable means for identifying two SAAs and applying the detection criteria for 

properly detecting and sizing sawcut D and crack C. The computed SNR for the two detected 

introduced flaws at 60° ranged from 6 dB to nearly 12 dB, while the SNR for the same two 

detected introduced flaws at 45° ranged from 7 dB to approximately 11 dB. A review of the 

−6 dB measured values for length show that the 60° data resulted in undersizing sawcut D and 

slightly undersizing crack C, while the 45° data oversized crack C and correlated quite well with 

the true-state value for sawcut D.

From a detection performance standpoint, it is again interesting to note the specific introduced 

flaws that were not detected or those that were only detected at a specific angle. From the far

side, no introduced flaw went undetected in this piping segment. Sawcut E was only detected at 

the 70° incident angle using the lower frequency of 250 kHz. The data from this piping segment 

showed that the 400-kHz L-wave modality had a very difficult time providing suitable information 

for detection of any of the introduced flaws in this segment with the exception of sawcut F being 

detected using 45°.

7.2.4 Piping Segment 3

Piping segment 3, defined as the third 120° segment of pipe spanning from the 241° point to the 

0° datum point, consists of the final four introduced flaws consisting of implanted fatigue cracks 

and machined sawcuts. In circumferentially ascending order, these are crack D, sawcut G, 

sawcut H, and crack E. The discussion of the resultant detections and flaw characterization 

data continues here based on the tabular format. Table 7.10 depicts the far-side detection calls 

for piping segment 3 along with a grading for data quality and flaw characterization data as a 

function of incident angle and wave-mode for 250-kHz inspections. Table 7.11 depicts the 

same information for inspections acquired at 400 kHz. Again, the same nomenclature as 

described previously is used for deciphering the tables discussed here.

Analysis of the far-side, 250-kHz data using the 70° L-wave modality indicated that the quality of 

the resultant images were marginal as the SNR was generally lower than anticipated. The 

SAFT-processed data, however, provided good separation of signal responses from background noise levels as illustrated from the data, and this modality and frequency combination 

provided the capability to clearly detect three of the four introduced flaws in this piping segment. 

Sawcut G, sawcut H, and crack E were detected and sized in the length dimension. The 

computed SNR for these detectable introduced flaws ranged from approximately 5 dB to over 

9 dB. From a review of the data in Table 7.10, the measured length-sizing data correlates very 

well with the true-state values for sawcut H and crack E, but significantly oversized sawcut G. 

C-scan and B-scan end-view images as well as the individual data sheets for each detected 

introduced flaw can be found in Appendix D.
7-24

Table 7.10 250-kHz, Far-Side Detection Results for Piping Segment 3

Introduced 

Flaw

70° L 60° L 45° L 70° S 60° S 

−6 dB (cm), 

LOS (cm),

SNR (dB) “Q”

−6 dB (cm), 

LOS (cm), 

SNR (dB) “Q”

−6 dB (cm), 

LOS (cm), 

SNR (dB) “Q”

−6 dB (cm), 

LOS (cm), 

SNR (dB) “Q”

−6 dB (cm), 

LOS (cm), 

SNR (dB) “Q

Flaw D ND M -- -- ND G -- -- -- -

Sawcut G 10.92, 11.18, 

9.3

M -- -- 7.62, 7.87, 

8.0

G -- -- -- -

Sawcut H 6.35, 6.10, 8.6 M -- -- 8.13, 8.13, 

19.0

G -- -- -- -

Flaw E 3.30, 3.30, 4.6 M -- -- ND G -- -- -- -

Table 7.11 400-kHz, Far-Side Detection Results for Piping Segment 3

Introduced 

Flaw

70° L 60° L 45° L 70° S 60° S 

−6 dB (cm), 

LOS (cm),

SNR (dB) “Q”

−6 dB (cm), 

LOS (cm), 

SNR (dB) “Q”

−6 dB (cm), 

LOS (cm), 

SNR (dB) “Q”

−6 dB (cm), 

LOS (cm), 

SNR (dB) “Q”

−6 dB (cm), 

LOS (cm), 

SNR (dB) “Q

Flaw D ND G ND G ND G ND P ND P 

Sawcut G 6.10, 6.86, 8.9 G 4.32, 3.81, 

6.9

G 5.59, 5.33, 

12.3

G ND P 5.33, 5.33, 

8.1

P 

Sawcut H 5.33, 6.35, 11.8 G 4.83, 5.08, 

8.1

G 5.84, 7.11, 

12.3

G ND P 3.05, 3.30, 

12.0

P 

Flaw E 2.79, 3.30, 13.6 G ND G ND G ND P 2.03, 4.06, 

6.4

P 
7-25

Analysis of the results from the 45° L-wave scans at 250 kHz demonstrated good quality data 

due to lower levels of background noise and good clarity and definition for the SAAs that met the 

detection criteria. Sufficient separation is evident in the SAFT-processed images for effective 

detection of two of the four introduced flaws in this piping segment. This data set did not provide 

suitable information for detecting crack D or crack E; however, application of the detection 

criteria did result in the detection of both sawcuts (G and H) from these data. The computed 

SNR for these calls ranged from approximately 8 dB to 19 dB, and the measured lengths for 

these introduced flaws were significantly oversized in comparison to the true-state lengths.

Analysis of the 400-kHz data using the 45°, 60°, and 70° L-wave combinations shows that the 

quality of the data were good, as the background noise levels were lower than anticipated. 

These data sets exhibited good clarity and provided the necessary separation of detected 

introduced flaws from other SAAs and background noise levels for detection purposes. The 

SAFT-processed images resulting from these data sets illustrate a high degree of coherence 

and smoothness for the SAAs that were identified by applying the detection criteria. The 45°

and 60° L-wave data resulted in the detection of sawcut G and sawcut H, but did not detect 

crack D or crack E. The computed SNR for the detected introduced flaws ranged from 

approximately 7 dB to over 12 dB. In reviewing the measured length values for the 45° L-wave 

data, the −6 dB drop method provided close correlation with the true-state lengths for the two 

sawcuts, while the 60° L-wave inspection data provided undersized length values for each of 

these sawcuts. The 70° L-wave data resulted in the detection of all introduced flaws in this 

piping segment with the exception of crack D. The computed SNR for the detected introduced 

flaws ranged from approximately 9 dB to nearly 14 dB. The measured length values for the 

−6 dB sizing method undersized sawcut H and crack E and oversized sawcut G, while the LOS 

method correlated well with the true-state values for sawcut H and crack E. The LOS method 

did however oversize sawcut G. Again, the reader is directed to view any of the specified 

composite C-scan, B-scan end-view, or magnified/isolated views of SAAs in Appendix D.

The data quality for the 400-kHz shear-wave data, specifically the 60° and 70° scans, were poor. 

The unprocessed images were again very grainy and exhibited a generally noisy background 

where little SAA separation or identification could be accomplished. In particular, the SAFTprocessed 70° data provided no useful information for resolving or detecting SAAs that even 

remotely met the detection criteria. SAFT processing of the 60° data improved the image quality, 

enabling separation of SAAs and positional criteria to be used in the detection of all of the introduced flaws in this piping segment with the exception of crack D. The SNR for the three 

detected introduced flaws ranged from approximately 6 dB to over 12 dB. Length-sizing 

measurements using the −6 dB method undersized the introduced flaws relative to the true-state 

length values, while the LOS method undersized sawcut G and sawcut H and oversized crack E. 

The 45° shear-wave data exhibited good quality data resulting in clear processed images, good 

separation of SAAs, and generally low background noise levels. Detection criteria were 

employed for the detection and sizing of three of the introduced flaws in this piping segment, 

including sawcut G, sawcut H, and crack E. Once again, flaw D went undetected. The computed SNR for the three detected introduced flaws at 45° ranged from over 9 dB to nearly 16 dB. 

A review of the −6 dB measured values for length show that the 45° data resulted in significant 

undersizing both sawcut G and sawcut H, while matching very well with the true-state value for 

crack E.
7-26

From a detection performance standpoint, it is again interesting to note the specific introduced 

flaws that were not detected, or those that were only detected at a specific angle. From the far 

side, crack D went undetected throughout the evaluation. The remaining introduced flaws were 

detected by combinations of both longitudinal- and shear-wave modalities. As expected, 

sawcut G and sawcut H were detected most often across the range of angles, modalities, and 

frequencies employed.

7.3 Phased Array UT: Data Analysis and Results

One of the advantages of PA data is that all information is stored digitally during each scan, 

allowing off-line processing and interpretation to be performed in an environment better suited 

for analysis than one usually encounters during data acquisition. Several image channels of the 

data can be displayed, based on the preference of the analyst. The volume-corrected B-, Cand S-scan images, along with supplemental A-scans, were used to assist our understanding of 

the data.

For the TomoView PA software used during these trials, the B-scan image is a projection of the 

sound field through the component, oriented as if looking along the direction of sound propagation, displaying a circumferential cross section of the weld (shown in the right pane of the 

analysis screen). The S-scan (shown in the upper left pane of the analysis screen), or sector

view, provides a projection of the sound fields (from initial to final angle being used; i.e., 30° to 

70° for the TRL and 40° to 70° for the TRS arrays). The black line in the S-scan shows which 

angle is currently selected. Finally, the A-scan (lower-left pane of analysis screen) represents 

the electronic responses for reflectors based on X-Y axes of time and amplitude, respectively. 

A typical analysis screen showing the scan images described above is included as Figure 7.11.

Each of the B- and S-scan images contains all data in the linear scan. Measurement, gating, 

and sound field cursors are used to provide ―slices‖ of material to view along the projected 

sound beam and discrete beam angle responses are shown in the A-scan. By manipulating 

these cursors, the analyst is able to ―walk-through‖ cross sections of the material along each 

linear scan. As Figure 7.11 was made using the longitudinal-wave array, a mode-converted 

shear-wave response is also detected. Images in the A-, B-, and S-scans appearing farther in 

time (greater than 36-mm [1.4-in.] wall thickness) are the result of mode-converted shear waves 

produced when longitudinal waves strike the inner surface, or other reflectors, in the 

component. This echo is not present in the shear-wave data.

One important analysis tool of the phased-array system is the ability to examine each angle 

individually. This allows the investigator to view the different response images in a given scan 

and discriminate between the various features in a specimen. In Figure 7.11 the system is 

displaying the results for 35°, which shows the counterbore of the weld. When the 53° scan is 

displayed, shown in Figure 7.12, the weld root becomes visible, and when one displays 65°, 

shown in Figure 7.13, the flaws become visible. This angular discrimination is a very useful tool 

when coupled with the S-scan and B-scans. Using the location of the weld root, one can 

distinguish between weld-root response and flaw responses and determine if a flaw is on the 

near side or far side of a weld. The results of all scans and details of each flaw are given in 

Appendix E.
7-27

Figure 7.11 Typical Analysis Screen with a Low Angle (35°) Selected, Showing the Weld 

Counterbore with the TRL Array

Figure 7.12 Typical Analysis Screen with a Mid-Range Angle (53°) Selected, Showing the 

Weld Root with the TRL Array
7-28

Figure 7.13 Typical Analysis Screen with a High Angle (65°) Selected, Showing the Flaws 

with the TRL Array

In Phase 1 of this study, a series of implanted cracks and sawcuts were examined from the near 

and far side of welds using both the longitudinal- and shear-wave arrays. As described in 

Section 1, the pipe specimen 02-24-15 consists of two 610-mm (24.0-in.) diameter, 36-mm 

(1.4-in.) thick pieces of ASTM A-358, Grade 304 vintage austenitic stainless steel. The 

specimen is cut into three segments to allow easier handling and scanning. For the TRL array, 

which is optimized for lower angles, the segments were scanned with the front of the array 

situated against the toe of the weld, 12.7 mm (0.5 in.) back from the weld toe, 12.7 mm (0.5 in.) 

onto the weld, and with the probe centered on the weld crown. For the TRS array, which is 

optimized for higher angles, the segments were scanned with the probe next to the weld crown, 

6 mm (0.25 in.) back, 12.7 mm (0.5 in.) back, and 19 mm (0.75 in.) back from the weld crown.

Figure 7.14 shows the best near-side and far-side B-scan images for pipe segment 1 using both 

the TRL and TRS arrays. Pipe segment 1 contains cracks A (15% through-wall) and B (43% 

through-wall), and sawcuts A (8% through-wall, straight cut), B (28% through-wall, straight cut), 

and C (7% through-wall, 35° angled cut). The term ―straight cuts‖ refers to sawcuts that were 

made 90° to the inner surface, and the 35° angled cuts follow the heat-affected zone as shown 

in Section 5, Figure 5.4. The near-side scans (left two images) show all defects very clearly 

with both arrays. Sawcut A and crack A are both detected and easily resolved as two distinct 

defects, and sawcut C is also easily detected. Sawcut A appears as two nearby indications in 

the shear wave scan. The TRL array is able to clearly detect tip signals for the sawcuts and 

cracks from the near side. The TRS array is not able to detect tip signals from the near-side.
7-29

Figure 7.14 B-scan Views for Pipe Segment 1 from the Near (left two images) and Far

Side (right two images) with the Longitudinal- and Shear-Wave Arrays

Detecting the flaws from the far side is more challenging. The far-side B-scans in Figure 7.14

(right two images) were taken at an angle that hits the shallow defects, which results in the weld 

root being visible in these images. The TRL array is able to clearly detect sawcut A but is 

challenged to detect crack A. The TRS array is able to clearly see both sawcuts and crack A, 

but the response from the 35° angled sawcut C is weaker for the TRS than for the TRL. The 

deeper sawcut B and crack B were easily and clearly detected by both arrays.

Figure 7.15 shows the best near-side and far-side B-scan images for pipe segment 2 using both 

the TRL and TRS arrays. Pipe segment 2 contains crack C (64% through-wall) and sawcuts D 

(19% through-wall, straight cut), E (12% through-wall, straight cut), and F (19% through-wall, 

35° angled cut). Each of these defects is easily detected using both arrays. The TRL array 

again shows clear tip signals on the near side of the weld, but no tips are detectable from the far 

side of the weld. The TRS array did not detect clear tip signals from the near side, however. 

Neither array was able to detect tip signals for any of the defects from the far side of the weld.
7-30

Figure 7.15 B-scan Views for Pipe Segment 2 from the Near and Far Side with the 

Longitudinal- and Shear-Wave Arrays

Figure 7.16 shows the best near-side and far-side B-scan images for pipe segment 3 using both 

the TRL and TRS arrays. Pipe segment 3 contains cracks D (19% through-wall, axiallyoriented) and E (48% through-wall), and sawcuts G (18% through-wall, 35° angled cut) and 

H (26% through-wall, 35° angled cut). Once again the TRL array is able to detect strong corner 

traps and tip signals from the near side of the weld. These deeper flaws are easily detected by 

both arrays from the far side of the weld. Neither array was able to detect the axially oriented 

crack D from the near or far side. This is to be expected as the axial crack presents virtually no 

cross section to the beam in this configuration. To detect the axial flaw, the arrays would have 

to be oriented so that the sound field would strike the flaw at near 90°. While an array could be 

designed to detect axially oriented flaws, this was not feasible during these trials.

The results for pipe segments show that while all cracks and sawcuts are visible, the shallower 

defects are somewhat more difficult to detect than the flaws with features that penetrate farther 

through-wall. The scans taken from the near side of the weld show all implanted features very 

clearly. The longitudinal-wave array shows both a strong corner trap and a clear tip signal from 

the near side. The shear-wave array shows the corner trap and a weak tip signal, if any, from 

the near side. From the far side of the weld, the longitudinal-wave array does a very good job 

detecting the 35° angled sawcuts while both arrays are able to detect the straight sawcuts and 

cracks.
7-31

Figure 7.16 B-scan Views for Pipe Segment 3 from the Near and Far Side with the 

Longitudinal- and Shear-Wave Arrays

The most challenging circumferential flaw for the longitudinal array is crack A. The response 

from crack A is weak and is difficult to distinguish from the edge of the stronger response from 

the very nearby sawcut A. The most challenging circumferential flaw for the shear-wave array 

to detect is sawcut C. This is most likely the result of the shallow depth of sawcut C 

(7% through-wall) and the 35° angle of the cut which weakens the corner trap response.

Neither array is able to detect tip signals for the defects from the far side of the weld. This 

presents a significant challenge to depth-sizing for any detected flaws. Length-sizing of flaws 

from the far side of a weld is quite accurate, however. As the UT signal from a linear defect is 

often interrupted as it passes through the weld microstructure, the 6 dB drop technique is not as 

useful as the loss of signal technique. The sawcuts and implanted cracks were length-sized 

using the 6 dB drop technique and the LOS technique. The length-sizing results for the sawcuts 

for the TRL probe are given in Table 7.12. The RMSE for far-side length sizing of the sawcuts 

from the far side of the weld was 6.3 mm (0.25 in.), well within the 19-mm (0.75-in.) acceptance 

criteria of ASME Section XI Code, Appendix VIII. The results for length sizing of the sawcuts 

with the TRS probe are given in Table 7.13. The TRS results for the 6 dB drop technique were 

consistently undersized from both the near and far side of the weld. This is the result of the 

discontinuous nature of the ultrasonic responses from the sawcuts. While the sawcuts are 

straight linear features, the ultrasonic response from them can appear as a series of tightly

spaced smaller features. The LOS technique did provide reasonably accurate lengths for the 

sawcuts, although they were also undersized. The RMSE from the far side of the weld was 

15.3 mm (0.60 in.), which is lower than the 19-mm (0.75-in.) ASME Section XI, Appendix VIII 

acceptance criteria, but is not as accurate as the TRL results. Sawcut length-sizing data is 

plotted in Figure 7.17.
7-32

Table 7.12 Sawcut Lengths in Pipe Specimen 02-24-15 as Measured by the TRL Array

Near Side Far Side

Sawcut Actual Length TRL −6 dB TRL LOS TRL −6 dB TRL LOS

A 32.8 mm 33.0 mm 42.0 mm 18.0 mm 21.4 mm

(1.29 in.) (1.30 in.) (1.65 in.) (0.71 in.) (0.84 in.)

B 65.1 mm 55.9 mm 73.7 mm 60.0 mm 60.0 mm

(2.57 in.) (2.20 in.) (2.90 in.) (2.36 in.) (2.36 in.)

C 36.3 mm 19.1 mm 49.9 mm 30.5 mm 42.3 mm

(1.43 in.) (0.75 in.) (1.96 in.) (1.20 in.) (1.67 in.)

D 54.1 mm 44.0 mm 51.1 mm 47.0 mm 61.6 mm

(2.13 in.) (1.73 in.) (2.01 in.) (1.85 in.) (2.43 in.)

E 43.7 mm 29.0 mm 53.5 mm 37.1 mm 48.7 mm

(1.72 in.) (1.14 in.) (2.11 in.) (1.46 in.) (1.92 in.)

F 59.7 mm 57.9 mm 64.4 mm 45.2 mm 65.6 mm

(2.35 in.) (2.28 in.) (2.54 in.) (1.78 in.) (2.58 in.)

G 57.3 mm 48.3 mm 67.7 mm 35.6 mm 54.0 mm

(2.25 in.) (1.90 in.) (2.66 in.) (1.40 in.) (2.13 in.)

H 75.2 mm 68.6 mm 83.8 mm 58.4 mm 75.6 mm

(2.96 in.) (2.70 in.) (3.30 in.) (2.30 in.) (2.98 in.)

RMS Error 10.2 mm 9.0 mm 12.9 mm 6.3 mm

(0.40 in.) (0.36 in.) (0.51 in.) (0.25 in.)

Table 7.13 Sawcut Lengths in Pipe Specimen 02-24-15 as Measured by the TRS Array

Near Side Far Side

Sawcut Actual Length TRS −6 dB TRS LOS TRS −6 dB TRS LOS

A 32.8 mm 13.0 mm 33.0 mm 8.0 mm 29.4 mm

(1.29 in.) (0.51 in.) (1.30 in.) (0.32 in.) (1.16 in.)

B 65.1 mm 53.0 mm 75.5 mm 61.0 mm 45.1 mm

(2.57 in.) (2.09 in.) (2.97 in.) (2.40 in.) (1.77 in.)

C 36.3 mm 24.0 mm 40.0 mm N/A 32.1 mm

(1.43 in.) (0.95 in.) (1.58 in.) (1.26 in.)

D 54.1 mm (18.0 mm 65.1 mm 35.0 mm 47.1 mm

(2.13 in.) 0.71 in.) (2.56 in.) (1.38 in.) (1.86 in.)

E 43.7 mm (29.0 mm 48.4 mm 18.0 mm 28.1 mm

(1.72 in.) 1.14 in.) (1.90 in.) (0.71 in.) (1.11 in.)

F 59.7 mm 30.0 mm 74.5 mm 25.0 mm 42.2 mm

(2.35 in.) (1.18 in.) (2.93 in.) (0.98 in.) (1.66 in.)

G 57.3 mm 15.0 mm 66.0 mm 13.0 mm 38.7 mm

(2.25 in.) (0.59 in.) (2.60 in.) (0.51 in.) (1.52 in.)

H 75.2 mm N/A 79.0 mm N/A 53.0 mm

(2.96 in.) (3.11 in.) (2.09 in.)

RMS Error 26.4 mm 8.5 mm 28.3 mm 15.3 mm

(1.04 in.) (0.33 in.) (1.12 in.) (0.60 in.)
7-33

Figure 7.17 Sawcut Lengths Measured Using the 6 dB Drop and Loss of Signal 

Techniques with the TRL and TRS Probes from the Near and Far Side of the 

Weld Plotted Versus the Actual Sawcut Lengths

Tables 7.14 and 7.15 show the results for the length sizing of the implanted thermal fatigue 

cracks for the TRL and TRS probes, respectively. This data is plotted in Figure 7.18. The LOS 

technique performed much better than the 6 dB drop technique for both probes from the far side 

of the weld. This is again the result of the uneven amplitude along the scanning direction of the 

ultrasonic responses from the implanted cracks. Both the TRL and TRS were able to accurately 

length size the defects from the far side; the TRL had an RMSE of 6.3 mm (0.25 in.) and the 

TRS array had an RMSE of 8.9 mm (0.35 in.). It should be noted that the best scans were 

taken with the front of the probe either against the toe of the weld or slightly back from the toe of 

the weld. These scans could easily be taken with the weld crown left in the as-welded position 

and do not require the weld crown to be machined off.
7-34

Table 7.14 Thermal Fatigue Crack Lengths in Pipe Specimen 02-24-15 as Measured by 

the TRL Array

Near Side Far Side

TFC Actual Length TRL −6 dB TRL LOS TRL −6 dB TRL LOS

A 10.7 mm 12.0 mm 16.2 mm 16.0 mm 17.3 mm

(0.42 in.) (0.47 in.) (0.64 in.) (0.63 in.) (0.68 in.)

B 30.5 mm 27.9 mm 27.6 mm 21.0 mm 27.1 mm

(1.20 in.) (1.10 in.) (1.09 in.) (0.83 in.) (1.07 in.)

C 43.6 mm 22.1 mm 50.1 mm 16.0 mm 53.1 mm

(1.72 in.) (0.87 in.) (1.97 in.) (0.63 in.) (2.09 in.)

E 33.8 mm 17.8 mm 43.5 mm 17.8 mm 37.4 mm

(1.33 in.) (0.70 in.) (1.71 in.) (0.70 in.) (1.47 in.)

RMS Error 13.5 mm 6.6 mm 16.8 mm 6.3 mm

(0.53 in.) (0.26 in.) (0.66 in.) (0.25 in.)

Table 7.15 Thermal Fatigue Crack Lengths in Pipe Specimen 02-24-15 as Measured by

the TRS Array

Near Side Far Side

TFC Actual Length TRS −6 dB TRS LOS TRS −6 dB TRS LOS

A 10.7 mm 12.0 mm 20.1 mm 6.0 mm 20.9 mm

(0.42 in.) (0.47 in.) (0.79 in.) (0.24 in.) (0.82 in.)

B 30.5 mm 14.0 mm 40.0 mm 7.4 mm 42.3 mm

(1.20 in.) (0.55 in.) (1.58 in.) (0.29 in.) (1.67 in.)

C 43.6 mm 7.0 mm 60.1 mm 21.0 mm 45.0 mm

(1.72 in.) (0.28 in.) (2.37 in.) (0.83 in.) (1.77 in.)

E 33.8 mm 30.0 mm 74.5 mm 25.0 mm 42.2 mm

(1.33 in.) (1.18 in.) (2.93 in.) (0.98 in.) (1.66 in.)

RMS Error 20.1 23.0 mm 16.9 mm 8.9 mm

(0.79 in.) (0.90 in.) (0.67 in.) (0.35 in.)
7-35

Figure 7.18 Thermal Fatigue Crack Lengths Measured Using the 6 dB Drop and Loss of 

Signal Techniques with the TRL and TRS Probes from the Near and Far Side 

of the Weld Plotted Versus the Actual Flaw Lengths

8-1

8 Phase 2 Evaluation

The Phase 2 evaluation focused on examination of three full-circumferential welds to determine 

the effects of weld microstructure on far-side inspections. Austenitic pipe specimen 3C-022 was 

used for this assessment. Work conducted in Phase 2 again included all three ultrasonic NDE 

techniques—conventional-UT, low-frequency/SAFT-UT, and phased-array UT. As described in 

Section 5, pipe specimen 3C-022 is made of Schedule 80 ASTM 358 Grade 304 wrought 

stainless steel and consists of four 610-mm (24.0-in.) diameter pipe sections joined by three 

welds: (1) weld horizontally with air backing, (2) welded vertically with air backing, and 

(3) welded horizontally with water backing. A 10% through-wall, 360° notch is machined into the 

heat-affected zone of welds one and three to provide a consistent standard reflector. Weld two 

has three 10% through-wall EDM notches and three implanted cracks of 5%, 10%, and 15% 

through-wall depth in the HAZ. See Appendix F for all conventional-UT data and images on 

specimen 3C-022.

8.1 Conventional UT: Data Analysis and Results

8.1.1 Weld One

Weld number one, as described in Section 5, was made in the horizontal position in a fieldsimulated manner with air backing. The longitudinal seam weld was at the top (12 o’clock 

position) of the pipe during welding. The 10% notch in weld one was scanned using 1.5-MHz, 

60° and 70° shear waves from the far side of the weld. The probes were raster-scanned in the 

x-direction (perpendicular to the weld) over the pipe specimen over a distance of 152 mm 

(6.0 in.) and indexed (stepped) at 1.3-mm (0.05-in.) increments for a total of 1966 mm (77.4 in.) 

around the outside circumference of the pipe specimen, which produced a total of 1548 lines 

(strokes) per data set. These scans were time-intensive.

Figure 8.1 shows C-scans for the 60° and 70° scans for a representative section of weld one. 

The 10% notch in weld number one was easily detected by both the 60° and 70° shear waves. 

The signal-to-noise ratios were very similar—the SNR for the 60° shear waves is 19 dB and the 

SNR for the 70° shear waves is 20 dB. Responses for the 10% notch varied by approximately 

+6 dB in amplitude along the length of the notch at both angles. There was no measurable 

difference in the amplitude of the response from the 10% through-wall notch related to the 

overall position on the weld. The amplitude response from near the 12 o’clock position that was 

welded from above is virtually identical to the amplitude response from the areas near the 3 and 

6 o’clock positions on the weld. There is no correlation between the welding orientation and the 

locations of the high-amplitude responses in the 60° and 70° scans.
8-2

Figure 8.1 C-scan Views for a Representative Area of Weld Number One Acquired with 

1.5-MHz 60° and 70° Shear Waves from the Far Side. The 10% notch 

response is marked by arrows and varies throughout the length of each scan 

by 6 dB. 

8.1.2 Weld Two

Weld two was fabricated in the horizontal position with air backing. Weld number two was 

scanned from the near side with the 60° shear-wave probe and from the far side with the 60° 

and 70° shear-wave probes. The probes were raster-scanned over the pipe specimen where 

the x-direction scan length was 152 mm (6.0 in.) and y-direction increment step was 1.3 mm 

(0.05 in.). The total circumferential path traveled (with the inclusion of overlapped part path) 

was 2118 mm (83.4 in.) around the outside circumference of the pipe specimen for a total of 

1668 lines per data set. The 152 mm (6.0 in.) of overlap in each scan provides two built-in 

fiduciaries in the scans as the longitudinal seam appears as a ―dead zone‖ in the scans. Again, 

these scans were time-intensive.

Figure 8.2 shows C-scan images of a representative section of the 10% though-wall EDM notch 

in weld two from the 60° near-side scan, the 60° far-side scan, and the 70° far-side scans. The 

near-side scan provides the sharpest image of the EDM notch and has the best SNR of 27 dB. 

The 60° and 70° far-side scans both have SNRs of 24 dB. The amplitude along the 60° far-side 

scan varies by 6 dB, but the amplitude of the response in the 70° scan varies by as much as 

10 dB in some sections.
8-3

Figure 8.2 C-scan Views for a Representative Area of Weld Number Two Acquired with 

1.5-MHz, 60° and 70° Shear Waves. The 10% notch response is marked by 

arrows in each image and varies by as much as 10 dB in the 70° data.

Figures 8.3 through 8.5 show ultrasonic typical responses using 60° and 70° shear-wave probes 

from three implanted cracks in weld two. Figure 8.3 shows the response from crack 4 (5% 

through-wall), Figure 8.4 shows the response from crack 5 (10% though-wall), and Figure 8.5

shows the response from crack 6 (15% though-wall). Crack 4 was only detectable from the 

near side. Neither the 60° nor 70° shear-wave probes were able to detect the 5% though-wall 

crack from the far side of the weld. All three scans were able to detect crack 5. The strongest 

response for crack 5 comes from the near side, and the 60° far-side scan has a better SNR than 

the 70° far-side scan. As with the phased-array scans, the amplitude of the response from 

crack 5 is roughly 6 dB smaller than the response from the peak responses of the 10% EDM 

notch. This is to be expected as the crack is tighter than the EDM notch, the crack is much 

shorter than the EDM notch, and the crack is at a 35° angle and provides a weaker corner trap 

than the perpendicular EDM notch. Crack 6 is easily detected by all three scans. No tip 

responses are visible from any of the cracks, from the near or far side. This makes accurate 

depth-sizing of the cracks virtually impossible. The cracks were length-sized using the 6 dB 

drop technique and the loss of signal (LOS) technique. Table 8.1 shows the results for the 

length-sizing of the three cracks. Both the 6 dB drop and LOS technique oversized each crack 

in all scans.
8-4

Figure 8.3 C-scan Views of Crack 4 in Weld Number Two Acquired with 1.5-MHz, 60° and 

70° Shear Waves. The crack was not detected in the two far side data sets.

Figure 8.4 C-scan Views of Crack 5 in Weld Number Two Acquired with 1.5-MHz, 60° and 

70° Shear Waves
8-5

Figure 8.5 C-scan Views of Crack 6 in Weld Number Two Acquired with 1.5-MHz, 60° and 

70° Shear Waves

Table 8.1 Thermal Fatigue Crack Lengths in Weld Two of Pipe Specimen 3C-022 as 

Measured by the 60° and 70° Conventional-UT Probes

60° Near 60° Far 70° Far

Crack Actual Length 6 dB LOS 6 dB LOS 6 dB LOS

4 3.6 mm 19 mm 50.8 mm No detection No detection

(0.14 in.) (0.75 in.) (2.00 in.)

5 7.1 mm 6.4 mm 31.8 mm 10.2 mm 16.5 mm 10.2 mm 17.8 mm

(0.28 in.) (0.25 in.) (1.25 in.) (0.40 in.) (0.65 in.) (0.40 in.) (0.70 in.)

6 10.7 mm 11.4 mm 15.2 mm 12.7 mm 20.3 mm 11.4 mm 20.3 mm

(0.42 in.) (0.45 in.) (0.60 in.) (0.50 in.) (0.80 in.) (0.45 in.) (0.80 in.)

RMS Error 8.9 mm 30.9 mm 2.6 mm 9.5 mm 2.2 mm 10.2 mm

(0.35 in.) (1.22 in.) (0.10 in.) (0.37 in.) (0.09 in.) (0.40 in.)

8.1.3 Weld Three

Weld number three was made in the horizontal position in a field-like manner with water 

backing. The longitudinal seam weld was at the top (12 o’clock position) of the pipe during 

welding. The probes were again raster-scanned over the pipe specimen, with an x-direction 

scan stroke length of 152 mm (6.0 in.) and a total y-direction (circumferential) indexing length of 

2118 mm (83.4 in.) around the outside circumference of the pipe specimen. The probes were 

indexed in 1.3-mm (0.05-in.) increments, for a total of 1668 lines per data set. As in the 

previous scans, these scans were also initiated at a distance of 76 mm (3.0 in.) before the 

centerline of the longitudinal seam weld, indexed 360° around the pipe, and then continued 
8-6

indexing an additional 152 mm (6.0 in.) to 76 mm (3.0 in.) beyond the longitudinal weld 

centerline to obtain sufficient overlap. These scans were also quite time-intensive.

Figure 8.6 shows C-scans for the 60° and 70° scans for a representative section of weld three. 

The 10% notch in weld number three was easily detected by both the 60° and 70° shear waves. 

The SNRs for the different scans of weld number three were very similar—the SNR for the 60° 

shear waves is 8 and the SNR for the 70° shear waves is 5. The amplitude of the responses

from the 10% notch varied by 6 dB along the length of the notch for both angles. There was no 

measurable difference in the amplitude of the response related to the overall position on the 

weld. The amplitude response from near the 12 o’clock position that was welded from above is 

virtually identical to the amplitude response from the areas near the 3 and 6 o’clock positions on 

the weld. As with weld one, there was no correlation between the locations of the highamplitude responses in the 60° and 70° scans. 

Figure 8.6 C-scan Views for a Representative Area of Weld Number Three Acquired with 

1.5-MHz, 60° and 70° Shear Waves from the Far Side. The 10% notch 

response is marked by an arrow in each image and varies by 6 dB along the 

length of the notch at both angles 

8.2 Low-Frequency/SAFT: Data Analysis and Results

Each of the three welds was scanned using 400-kHz and 250-kHz probes. The welds were 

examined using 45° shear and longitudinal waves, 60° longitudinal waves, and 70° longitudinal 

waves at both frequencies. The data were stored electronically and analyzed using SAFT. The 

SAFT procedure used on the vertical pipe section was the same as used for processing the 

Phase 1 pipe segments. The SAFT angle used for the processing was 12° and the beam entry 

diameter was 12.7 mm (0.5 in.). No skipping was used and a dynamic range factor of one was 
8-7

used. Scans obtained using LF/SAFT proved very time-intensive with regard to both data 

collection and SAFT post-processing of the data.

8.2.1 Weld One

Weld number one, as described in Section 5, was made in the horizontal position in a field-like 

manner with air backing. The longitudinal seam weld was at the top of the pipe during welding. 

The 10% notch in weld one was scanned using 400-kHz 45° shear and longitudinal waves and 

60° longitudinal waves.

Each scan angle was able to detect the 10% notch. The 45° shear wave sound field was more 

suitable at consistently detecting the notch with the lowest variability (between 3 dB and 6 dB) in 

the reflected amplitude along the circumferential extent of the notch. The 45° longitudinal scan 

had the poorest response from the notch, with the amplitude of the response dropping 6 dB for 

several centimeters at a time along the notch. The 45° shear-wave scan provides the best 

width sizing for the notch, with the 6.35 mm (0.25 in.) wide notch appearing to be 15.2 mm 

(0.6 in.) in width using the 6 dB drop method. In both of the longitudinal scans, the notch 

appears to be much wider than it is while the shear-wave response is more tightly defined. The 

45° longitudinal scan shows the notch to have a width of 22.9 mm (0.9 in.) to 48.3 mm (1.9 in.). 

The 60° longitudinal scan shows a width ranging from 45.7 mm (1.8 in.) to 55.9 mm (2.2 in.). 

Width sizing of the narrow notch really provides an indication of the probe beam size rather than 

the notch width. As the probe sweeps past the flaw the beam width is mapped out. For this 

inspection, the data demonstrates that the 45° shear probe has a tighter beam than the 45°

longitudinal probe. 

Figure 8.7 shows representative C-scans from the three angles used to scan the notch (each 

taken from the same 38.1-cm [15.0-in.] section of pipe). It is interesting to note that there is 

weak, if any, correlation between the regions of high and low amplitude in these scans. The 

weld metal absorbs or redirects the 45° shear and longitudinal and 60° longitudinal waves 

uniquely. There was no strong correlation with the overall position on the weld with the 

amplitude of the responses from the different angles and modes for the 400-kHz transducers. 

The amplitude of the responses from the 10% notch was similar at the region near the 

12 o’clock position as the region near the 6 o’clock position.

The 10% notch in weld one was also scanned using 250-kHz probes with 45° and 70° 

longitudinal waves. The 250-kHz scans had a poorer response from the notch, with the 45° 

having as much as a 10 dB drop along the scan and showing the 6.4-mm (0.25-in.) wide notch 

to be over 50.8-mm (2.0-in.) wide in places. The response amplitude using 70° longitudinal 

waves had a lower variability, with regions dropping as many as 6 dB, and showing the notch to 

be 25.4–50.8-mm (1.0–2.0-in.) wide. As with the 400-kHz scans, the 250-kHz 45° and 70° 

scans show no correlation with specific regions of the pipe allowing the sound to penetrate and 

which regions absorb or redirect sound. The weld metal appears to have a different profile for 

the two angles. Figure 8.8 shows representative C-scans from the three angles used to scan 

the notch (each taken from the same 38.1-cm [15.0-in.] section of pipe).
8-8

Figure 8.7 C-scan Views for a Representative Area of Weld Number One Acquired with 

400-kHz, 45° S-waves, 45° L-waves, and 60° L-waves from the Far Side. The 

10% notch response is best at 45° shear with the least amount of signal loss 

and worst at 45° longitudinal. 
8-9

Figure 8.8 C-scan Views for a Representative Area of Weld Number One Acquired with 

250-kHz, 45° L-waves and 70° L-waves from the Far Side. The 45 L-wave 

response has as much as a 10 dB variation along the imaged length of the 

10% notch. 

As with the 400-kHz transducers, the 250-kHz scans showed no correlation with the overall 

position on the weld and with the amplitude of the responses from the different angles. The 

amplitude of the responses from the 10% notch was similar at the 12 o’clock position as at the 

6 o’clock position for the 250-kHz scans.

Finally, the weld metal appears to have a different ultrasonic profile depending entirely on the 

inspection frequency. Weld one was scanned with 45° longitudinal waves at both 400 kHz and 

250 kHz, and these two scans show no correlation as to where the weld allows sound to 

penetrate and where it interferes with the response from the 10% notch. Figure 8.9 shows the 

two 45° longitudinal-wave scans. This limited correlation between the different scans show that 

the weld metal microstructure is very complex and the ability of sound to pass through the weld

metal is a strong function of the angle and frequency of the sound.
8-10

Figure 8.9 Same Section of Weld Number One Scanned Using 45° Longitudinal Waves 

at Both 400 kHz and 250 kHz from the Far Side. There is limited correlation 

between the two frequencies in where sound penetrates and where it drops 

out.

8.2.2 Weld Two

Weld two was welded in the vertical position with air backing. It has three EDM notches with 

depths of 10% through-wall and three implanted cracks, one with a depth of 5% through-wall, 

one with a depth of 10% through-wall, and one with a depth of 15% through-wall. The flaws 

were scanned from the far side of the weld using 400-kHz 45° shear waves, and 45°, 60°, and 

70° longitudinal waves. Each of the scans was able to detect the 10% EDM notches easily. 

The amplitude of the response from the four scans was very similar in sensitivity. All four scans 

showed a 6 dB variation along the EDM notches and the signal responses from the very thin 

EDM notches are shown to slightly smear, covering a larger axial extent at lower frequencies. 

This effect is reduced as the frequency is increased to 400 kHz (see Figure 8.9) and a more 

representative signal response is shown for axial extent of the notch in the images. Unlike in 

weld one, there was some correlation in the response amplitude from the EDM notches and the 
8-11

position along the pipe. The highest amplitude responses for each scan were found at the 

beginning and end of the EDM notches. None of the scans were able to detect any of the three 

implanted flaws. As discussed in Section 6, the dimensions of the sound field (in particular the 

dimensions of the overlap or cross-over point between transmit and receive transducers) and 

the wavelength play a key role in the resolution capability of the inspection approach. At these

low frequencies, resolution of small flaws is limited.

Figure 8.10 shows C-scans for each 400-kHz scan from the same representative region of the 

pipe. The region chosen shows the end of the second EDM notch, the region with the 15% 

through-wall crack (two thirds of the way down in each figure), and the beginning of the third 

EDM notch. In each C-scan the 10% EDM notches are clearly discernable as is the gap 

between them but the implanted flaw is not detected in the gap.

Figure 8.10 C-scan Views for a Representative Area of Weld Number Two Acquired with 

400-kHz, 45° S-waves, 45° L-waves, 60° L-waves, and 70° L-waves. The 10% 

notches are detected but not the 15% implanted crack.

The EDM notches and cracks were also scanned using 250-kHz probes with 45° shear waves 

and 45° and 70° longitudinal waves. The results for the 250-kHz scans show that the amplitude 

of the three scans varies by 6 dB along the length of the EDM notches. As with the 400-kHz 

scans, none of the implanted flaws were detected with the 250-kHz scans. Figure 8.11 shows 

C-scans for each 400-kHz scan from the same representative region of the pipe. The EDM 

notches are discernable in all three scans, but the 45° longitudinal method does not resolve the 

gap between the EDM notches very clearly and the implanted 15% flaw is not detected.
8-12

Figure 8.11 C-scan Views for a Representative Area of Weld Number Two Acquired with 

250-kHz, 45° S-waves, 45° L-waves, and 70° L-waves. The 10% notches are 

detected but not the 15% implanted crack.

8.2.3 Weld Three

Weld number three, as described in Section 5, was made in the horizontal position in a field-like 

manner with water backing. The longitudinal seam weld was at the top of the pipe during 

welding. As with weld one, there is a 10% through-wall notch machined into the heat-affected 

zone of the weld. The 10% notch was scanned at 400 kHz from the far side of the weld with 

45° shear waves, and 45°, 60°, and 70° longitudinal waves. The 45° shear waves, and 45° and 

60° longitudinal waves, were able to clearly and consistently detect the 10% through-wall notch. 

The amplitude of the signal varies by 6 dB along the length of the notch for these angles. The 

70° longitudinal wave scan was very uneven, however, and has regions in which the amplitude 

drops to the noise level. C-scans from a representative region of the four scans are shown in 

Figure 8.12. 
8-13

Figure 8.12 C-scan Views for a Representative Area of Weld Number Three Acquired with 

400-kHz, 45° S-waves, 45° L-waves, 60° L-waves, and 70° L-waves. The 10% 

notch responses as viewed from the far side are noted with arrow in each 

image. 

Weld number three was also scanned using 250-kHz transducers. The 10% notch was 

examined using 45° and 70° longitudinal waves. The 45° scan was very poor. The amplitude of 

the response from the 10% notch was very uneven and often fell to the level of the noise. The 

70° scan was much more consistent. The amplitude of the response from the 10% through-wall 

notch was always less than −6 dB and was usually less than −4 dB. As with weld number one, 

the 400-kHz and 250-kHz scans showed no correlation with the overall position on the weld with 

the amplitude of the responses from the different angles. Figure 8.13 shows C-scans from a 

representative region of the two scans. The uneven nature of the 45° scan is visible.
8-14

Figure 8.13 C-scan Views for a Representative Area of Weld Number Three Acquired with 

250-kHz, 45° L-waves and 70° L-waves. The 10% notch responses as viewed 

from the far side are noted with arrow in each image. 

8.3 Phased Array UT: Data Analysis and Results

For the Phase 2 PA-UT assessments of specimen 3C-022, each weld was scanned using both 

TRL and TRS phased arrays. Unlike pipe specimen 02-24-15, this pipe specimen was not cut 

into individual sections. The welds were scanned with the front edge of the arrays directly 

against the weld toe and with the front edge of the arrays 6.4 mm (0.25 in.), 12.7 mm (0.5 in.), 

and 19 mm (0.75 in.) back from the weld toe. The scans were initiated at a distance of 51 mm 

(2.0 in.) back from the longitudinal seam (L-seam) weld, and traveled 360° around the pipe, 

continuing until the longitudinal seam was scanned again. By running over the L-seam twice, a 

beginning and ending fiduciary is put into the data. This locator works well, as the L-seam 

appears as a dead zone in the scan. Welds one and three were scanned using both the TRL 

and TRS arrays and were only scanned from the far side. Weld two was scanned from both the 

near and far side using both arrays.
8-15

8.3.1 Weld One

The 10% through-wall notch in the HAZ of weld one was examined using the TRL and TRS 

arrays. A sample region of the ultrasonic response of the 10% notch is shown in Figure 8.14. 

The dead zone near the top of both scans is caused by the longitudinal seam weld in the pipe. 

To allow for consistency, both figures are shown with the gain set so that the peak noise level 

was approximately 20% full screen height (FSH).

Figure 8.14 B-scan Views for the Ultrasonic Response for the 10% Notch in a Portion of 

Weld One from the Far Side with the Longitudinal- and Shear-Wave Arrays. 

Arrows Mark the Notch Response.

The scans showed that while there was strong short-range variability in the strength of the 

ultrasonic notch response, there was no overall ―structure‖ to the interference. The overall 

amplitude response for the notch located near the top (at the 12 o’clock position and welded 

downhand from above) is essentially the same as the notch response from both sides and the 

bottom (at the 6 o’clock position and welded from underneath). The amount of sound coming 

back through the weld along its length varies strongly, however. The amplitude of the ultrasonic 

response from the 10% through-wall notch was not constant and steady, but instead increased 

and decreased strongly along the scanning direction. Both the TRL and TRS showed ―cold 

spots‖ in the responses from the 10% notch that were as much as 20 dB lower than the peak 

response. These cold spots were between 1 and 5 mm (0.04 and 0.20 in.) long. Overall, the 

TRL array has a better SNR than the TRS array.
8-16

8.3.2 Weld Two

Weld number two contains three 10% through-wall EDM notches and three implanted thermal 

fatigue cracks with depths of 5%, 10%, and 15% through-wall. As described in Section 1, this 

weld was made in the vertical position with air backing. The EDM notches and cracks provided 

both standard reflectors and realistic cracks to test the longitudinal- and shear-wave methods to 

penetrate and detect flaws from the far side of the weld.

Weld number two was examined from both the near and far side. Figure 8.15 shows the 

response from the 10% EDM notch for the TRL array with the gain in both scans set to 44 dB. 

Using this gain, the peak signal amplitude in the near-side scan is 100% FSH with a low value 

of 57% FSH along the notch response. A clear corner response is visible in the scan. The farside scan shows a similarly intense response, but the ultrasonic response from the notch is less 

discrete than from the near side. No tip signal responses can be called in either scan. Using a 

gain of 44 dB, the peak response from the far side is 100% FSH and the lowest signal along the 

response is 40% FSH. The results show that the 2-MHz longitudinal waves are able to pass 

through the weld metal without a large loss of amplitude, but with varied levels of sound field 

redirection. Typical A-scans, shown in Figure 8.16, show the differences in the nature of the 

signal coming back from the EDM notch. The far side response is much broader than the near 

side response. 

The shear-wave response from the EDM notch was quite different than the longitudinal-wave 

response. The shear waves were severely attenuated but appear to be less strongly redirected 

by the weld metal. Figure 8.17 shows a typical response for the EDM notch scanned from the 

near and far side using the TRS array. In both scans, the gain is set to 23 dB. (Much less gain 

is needed for the TRS array as it has a larger aperture and there is no mode conversion 

compared to the TRL array.) From the near side, the maximum response is 100% FSH and the 

minimum response along the EDM notch is 78% FSH; and from the far side, the maximum 

response is 30% FSH with a minimum of 5% FSH. This represents an amplitude drop of 10 to 

20 dB through the weld. The EDM notch response does, however, look like the response from 

a linear feature and shows that the shear-wave beam is not redirected as severely as the 

longitudinal waves. The A-scans, shown in Figure 8.18, show that the largest effect on the 

beam as it passes through the weld metal is the reduction in amplitude.

The three implanted thermal fatigue cracks are all detectable from the far side using both the 

TRL and TRS arrays. The 5% through-wall crack has the smallest response, as would be 

expected. The 15% through-wall crack is the easiest to detect. No tip signals were detected 

from the far side by either array for any of the flaws. Figures 8.19 and 8.20 show the ultrasonic 

responses from the far side for the three implanted cracks using the TRL and TRS arrays, 

respectively. In each figure the crack is in the center and the linear indications above and below 

the crack result from the 10% through-wall EDM notches. It is interesting to note that crack 5 

(10% through-wall crack) returns a response that is 6 dB smaller than the 10% through-wall 

EDM notch for both the shear and longitudinal probes. Crack 5 is 7.1-mm (0.28-in.) long, which 

is larger than two wavelengths for the longitudinal-wave probe and more than four wavelengths 

for the shear-wave probe. Also, the cracks are at a 35° angle while the notch is perpendicular 

to the pipe wall, so there will be associated differences in the strength of the reflection. The 
8-17

6 dB difference in the amplitude of the responses from the crack and the EDM notch are 

consistent with previous PNNL studies (Schuster et al. 2010).

Figure 8.15 B-scan Views for the Ultrasonic Response from the EDM Notch in Weld Two 

from the Near and Far Side with the Longitudinal-Wave Array. Arrows Mark 

the Notch Response.
8-18

Figure 8.16 A-scan Views for the Ultrasonic Response from the EDM Notch in Weld Two 

from the Near and Far Side with the Longitudinal-Wave Array

Figure 8.17 B-scan Views for the Ultrasonic Response from the EDM Notch in Weld Two 

from the Near and Far Side with the Shear-Wave Array. Arrows Mark the 

Notch Response.
8-19

Figure 8.18 A-scan Views for the Ultrasonic Response from the EDM Notch in Weld Two 

from the Near and Far Side with the Shear-Wave Array

Figure 8.19 B-scan Views for the Ultrasonic Response from the Three Implanted Cracks 

in Weld Two from the Far Side with the Longitudinal-Wave Array.
8-20

Figure 8.20 B-scan Views for the Ultrasonic Response from the Three Implanted Cracks 

in Weld Two from the Far Side with the Shear-Wave Array

While the lack of tip signals is a barrier to depth-sizing the cracks, length-sizing is possible. The 

flaws were consistently oversized by both the TRL and TRS arrays, but the values were well 

within the ASME Code acceptance criterion of 19-mm (0.75-in.) RMSE. Table 8.2 shows the 

results of the length sizing using the 6 dB drop and LOS-sizing methods.

Table 8.2 Thermal Fatigue Crack Lengths in Weld Two of Pipe Specimen 3C-022 as 

Measured by the TRL and TRS Arrays

Far Side

TRL Results TRS Results

Crack Actual Length TRS −6 dB TRS LOS TRS −6 dB TRS LOS

4 3.6 mm 23.6 mm 25.4 mm 16.0 mm 18.0 mm

(0.14 in.) (0.93 in.) (1.00 in.) (0.63 in.) (0.71 in.)

5 7.1 mm 9.0 mm 22.0 mm 15.0 mm 24.0 mm

(0.28 in.) (0.35 in.) (0.87 in.) (0.59 in.) (0.95 in.)

6 10.7 mm 4.0 mm 7.6 mm 9.0 mm 14.5 mm

(0.42 in.) (0.16 in.) (0.30 in.) (0.35 in.) (0.57 in.)

RMS Error 12.2 mm 15.7 mm 8.6 mm 13.0 mm

(0.48 in.) (0.62 in.) (0.34 in.) (0.51 in.)

8.3.3 Weld Three

The 10% through-wall notch in the HAZ of weld three (water-backed) was examined using the 

TRL and TRS arrays. A sample region of the ultrasonic response of the 10% notch is shown in 

Figure 8.21. The dead zone near the top of both scans is caused by the L-seam weld in the 
8-21

pipe. As in weld one, both figures are shown with the gain set so that the maximum noise level 

was 20% FSH.

As in weld one, there was no correlation with the overall azimuthal position on the quality of the 

ultrasonic response; that is, the ultrasonic response from the 10% notch at the 12 o’clock 

position is essentially the same as the ultrasonic response at the 6 o’clock position. Again, as in 

the air-backed weld one, the ultrasonic response can randomly vary strongly over the weld 

length, and the response from the 10% notch appears as a string of high-intensity points 

connected by lower-intensity regions. The TRS array had a higher SNR compared to the TRL 

array.

Figure 8.21 B-scan Views for the Ultrasonic Response for the 10% Notch in Weld Three 

from the Far Side with the Longitudinal- and Shear-Wave Arrays. Arrows 

Mark the Notch Response.

9-1

9 Discussion

The technical discussion presented in this section is focused on efforts conducted for the 

assessment of Phase 1 and Phase 2 evaluations, where conventional-UT, low-frequency/SAFTUT and phased array-UT were assessed on specimens 02-24-15 and 3C-022.

9.1 Conventional UT – Phase 1

The Phase 1 results show that the 60° and 70° 1.5-MHz probes used together can detect and 

length-size flaws through welds in wrought austenitic piping. All of the circumferential flaws 

were detected, and the subsequent length-sizing was accurate enough to meet the ASME 

Section XI, Appendix VIII requirements. It was not possible to depth-size the flaws from the far

side of the welds, as there were no tip signals detected. Depth-sizing is complicated by the high 

noise levels associated with through-weld inspections, as the noise can in some circumstances 

mimic a tip signal and provide false depth information.

The automated scanning, digital data collection, and electronic post-scan analysis provided 

improved detection and more accurate length-sizing results than the manual ultrasonic tests 

performed in 2002. In the 2002 manual scans, crack A was considered undetectable and 

crack E was only detectable with medium confidence. The 70° shear-wave automated scan can 

clearly detect crack A, and crack E is clearly detectable by both the 60° and 70° automated 

scans. The automated scans are able to detect implanted cracks as small as 5% through-wall 

from the near side of the weld and cracks as small as 10% through-wall from the far side. The 

main drawback of the automated scanning is the very long scan times. Under laboratory

conditions, one circumferential scan of a 610-mm (24.0-in.) diameter pipe takes over four hours. 

If one wants to scan a pipe using multiple angles, this becomes prohibitive for inspections in the 

field. Also, the 60° probe would not be able to insonify the region of interest on the far side of 

the weld unless the weld crown has been removed so as to be flush with the pipe.

9.2 Conventional UT – Phase 2

The Phase 2 work showed that the welding process had little effect on the ultrasonic response 

from a 10% through-wall notch on the far side of the weld. The signal-to-noise ratio for weld 

one was slightly better than the SNR for weld three. The amplitude of the response from the 

10% through-wall notch varies along the scans. There was no effect in the azimuthal position 

along the weld on the ultrasonic response of the 10% notch through the welds. The amplitude 

response from the 10% through-wall notch near the 12 o’clock positions is virtually identical to 

the responses from the 3 and 6 o’clock positions, despite the different positions during welding. 

These results strongly suggest that the Phase 1 results on a shop-welded pipe specimen are

valid for field-welded specimens.

One method for potentially improving the conventional scanning would be the use of 

longitudinal-wave probes for possible tip-diffracted signals. While the shear waves were very 

effective at detecting the corner trap responses without any confusing mode-converted signals, 

the shear-wave probes were not able to provide any tip signals.
9-2

9.3 Low-Frequency/SAFT-UT – Phase 1

The low-frequency/SAFT inspection technique uses lower frequencies to employ longer

wavelengths and thereby improve penetration and reduce the attenuative effects typically 

attributed to microstructure and weld geometry. In the case of 304 wrought stainless steel 

piping components, it is the austenitic weldment that contributes to reducing the inspection 

effectiveness as the weld structure is attenuative to standard inspection frequencies. Coupled 

with the requirement to reliably and consistently detect cracks from the far side of the weld, 

these inspections are often difficult to conduct with standard inspection techniques used in 

industry. In addition, typical weld crowns quite often preclude the use of inspection angles less 

than 60°, limiting the inspections that can be conducted and thereby reducing the capability to 

employ effective inspection protocols from the far side. In this evaluation, the weld crowns were 

ground smooth and the inspection methodologies were provided opportunities for the use of 

shallow angles from 30° through 70° in both the longitudinal- and shear-wave modalities. 

Although the LF/SAFT inspection technique takes advantage of lower frequencies for improved 

penetration and reduction of microstructural and geometric attenuation, the performance of the 

technique is also reduced as the wavelength is too long and the sound-field dimensions too 

large for detection of smaller flaws. By employing advanced signal processing in the form of the 

SAFT post-processing methodology, some degree of resolution can be maintained while using 

lower frequencies.

As expected, many of the smaller introduced flaws were not consistently detected, and some 

introduced flaws were not detected at all with any combination of employed frequency, wavemode, or incident angle. In large part, this is attributed to the resolution issues described here;

however, physical size of the transducers and key sound-field parameters also play a role in the 

detection system’s ability to detect flaws in these materials. The inspection technique has been 

primarily developed for thicker section, primary loop cast stainless steel components, 5.8-cm

(2.3-in.) thick and greater, where the sound path lengths typically range from 8.9 cm to 18.6 cm

(3.5 to 7.3 in.). In this study, the transducers and associated wedges were modified and 

fabricated to conform to the smaller diameter and smaller wall thickness dimensions, 3.6-cm

(1.4-in.) thick, of the wrought piping components evaluated here. Typical sound path lengths in 

this study ranged from 5.0 cm to 10.4 cm (2.0 to 4.1 in.). The ―footprint‖ or contacting surface 

area of the transducer/wedge configuration against the part surface was still rather large for 

these components, as coupling, smooth translation of the transducer over the part surface, and 

overall sound field dimensions limited the effectiveness of the technique for detection of smaller 

flaws. Table 9.1 shows the detection results for the LF/SAFT inspection technique as a function 

of wave-mode and incident angle for 250-kHz far-side inspections of each piping segment. 

―D‖ represents a detected flaw, ―ND‖ represents not detected, and ―--" means that the specific 

scan combination was not conducted.

Table 9.1 illustrates the detection performance for the inspection technique at 250 kHz, and 

indicates that the 70° longitudinal-wave scans were generally successful in detecting the vast 

majority of the introduced flaws in all three piping segments. This particular scan combination 

demonstrated an optimal balance between penetration capabilities, resolution, and reduction of 

signal clutter and noise due to the microstructure and austenitic weld material. This frequencyangle-modality combination proved to be the most effective and consistent at detecting the 
9-3

introduced flaws provided for this study. Table 9.2 shows the detection results for the LF/SAFT 

inspection technique as a function of wave-mode and incident angle for 400-kHz far-side 

inspections of each piping segment. Again, the same nomenclature is used for Table 9.2. As 

the frequency increases to 400 kHz, the data shows a reduction in resolution and detection 

capabilities. In the longitudinal mode, the 45° 250-kHz scan combination did not provide useful 

detection data even though the weld crowns were ground smooth. This is primarily due to a 

large portion of the energy being mode-converted. Because the corner-trap signal responses 

peak at 45°, the poor detection performance at this angle may also be attributed to the fact that 

the main lobe of the sound field at 45° must propagate through the thickest portion of the weld 

volume, thereby encountering more scattering and attenuation from the austenitic weld structure

Table 9.1 Summary of Detection Results for the 250-kHz Low-Frequency/SAFT 

Inspection Technique

Introduced Flaw 70°L 60°L 45°L 70°S 60°S 45°S

Sawcut A ND -- ND -- -- --

Flaw A ND -- ND -- -- --

Sawcut B D -- ND -- -- --

Flaw B D -- ND -- -- --

Sawcut C D -- ND -- -- --

Sawcut D D -- ND D D D

Flaw C D -- ND ND D D

Sawcut E D -- ND ND ND ND

Sawcut F D -- D ND ND D

Flaw D ND -- ND -- -- --

Sawcut G D -- D -- -- --

Sawcut H D -- D -- -- --

Flaw E D -- ND -- -- --

Table 9.2 Summary of Detection Results for the 400-kHz Low-Frequency/SAFT 

Inspection Technique

Introduced Flaw 70°L 60°L 45°L 70°S 60°S 45°S

Sawcut A ND ND D ND ND D

Flaw A ND ND ND ND ND ND

Sawcut B D D D ND D D

Flaw B D D ND ND D D

Sawcut C ND ND ND ND ND ND

Sawcut D ND ND ND ND D D

Flaw C ND ND ND ND D D

Sawcut E ND ND ND ND ND ND

Sawcut F ND ND D ND ND ND

Flaw D ND ND ND ND ND ND

Sawcut G D D D ND D D

Sawcut H D D D ND D D

Flaw E D ND ND ND D D
9-4

than the 70° sound field. The higher angle of incidence directs the sound field toward the lower 

sections of the weld, where the distance across the austenitic weld material is shorter. This, in 

turn, means that the sound field encounters less weld material and attenuation effects are 

reduced when employing larger incident angles.

At 400 kHz, the overall detection performance is less effective than the lower-frequency 

(250-kHz) inspections, and although both conventional and phased-array technologies should 

provide strong performance results, the reason for the reduced 400-kHz performance is not well 

understood. The longitudinal-wave modality is generally not as effective as the shear-wave 

results especially when compared to the 45° and 60° incident angles. Table 9.2 illustrates the 

trends in reduced detection performance. The 70° shear-wave inspection mode provided no 

useful information for detection of any introduced flaws in the piping segments. It is seen from 

the data provided in Appendix D that the general background noise levels increased at 400 kHz, 

adding to the reduced capability for resolving SAAs and detecting introduced flaws from the 

data. The significant decrease in detection performance (and general increase in background 

noise) associated with a relatively small increase in inspection frequency (from 250 kHz to 

400 kHz) is not well understood. As noted in this study, both the conventional method and the 

phased-array technique are operating at much higher frequencies, and both of these methods 

are shown to exhibit more effective detection performance using the same wave modality. The 

45° and 60° shear-wave results indicate promising detection performance; however, sawcuts C, 

E, and F were not detected, revealing the inconsistency so prevalent in the 400-kHz data.

It should also be noted that to effectively meet project milestones within the constraints of scope 

and budget, the number of scan combinations were optimized and reduced from initial plans to 

conduct the entire range of scan-angle-frequency-mode combinations. In reviewing the 

250-kHz data, the acquisition of 45° and 60° shear-wave scans for piping segments 1 and 3 

would definitely have provided additional insight into the detection performance results for the 

lower-frequency inspections.

Regarding length-sizing accuracy, the LF/SAFT inspection technique again is limited by the 

resolution issues described earlier. The longer wavelengths will not provide the resolution for 

accurate sizing as only the deepest portions of the introduced flaws will be detected and the

size of the wavelength relative to the flaw size tends to smear the actual SAA in the SAFTprocessed images, thereby making sizing of small flaws difficult to accurately determine. 

Table 9.3 summarizes the −6 dB drop length-sizing results for the LF/SAFT inspection 

technique as a function of incident angle and wave mode for 250-kHz data relative to the truestate values for each introduced flaw. Dimensions are in centimeters. Again, an ―ND‖ 

represents not detected, and ―--" means that the specific scan combination was not conducted.

Table 9.4 summarizes the LOS length-sizing results for the LF/SAFT inspection technique as a 

function of incident angle and wave mode for 250-kHz data relative to the true-state values for 

each introduced flaw. Dimensions are in centimeters. The nomenclature used is the same as 

in Table 9.3.

In general, the data shown in Table 9.3 show well-correlated measured length values for the 

70° L-wave inspection mode with the exception of crack C and sawcut G.
9-5

Table 9.3 Summary of Length-Sizing Results for the 250-kHz Low-Frequency/SAFT 

Inspection Technique Using the −6 dB Drop Method

Introduced 

Flaw

TrueState 

Length 

(cm)

70°L 

Measured 

Length 

(cm)

60°L 

Measured 

Length 

(cm)

45°L 

Measured 

Length 

(cm)

70°S 

Measured 

Length 

(cm)

60°S 

Measured 

Length 

(cm)

45°S 

Measured 

Length 

(cm)

Sawcut A 3.28 ND -- ND -- -- --

Flaw A 1.07 ND -- ND -- -- --

Sawcut B 6.52 6.35 -- ND -- -- --

Flaw B 3.05 3.05 -- ND -- -- --

Sawcut C 3.63 3.30 -- ND -- -- --

Sawcut D 5.41 5.84 -- ND 3.30 2.29 4.06

Flaw C 4.36 2.54 -- ND ND 4.32 5.97

Sawcut E 4.37 4.57 -- ND ND ND ND

Sawcut F 5.97 4.19 -- 4.83 ND ND 3.43

Flaw D 1.33 ND -- ND -- -- --

Sawcut G 5.73 10.9 -- 7.62 -- -- --

Sawcut H 6.84 6.35 -- 8.13 -- -- --

Flaw E 3.38 3.30 -- ND -- -- --

RMS Error 1.84 N/A 1.48 2.11 2.21 1.90

Table 9.4 Summary of Length-Sizing Results for the 250-kHz Low-Frequency/SAFT 

Inspection Technique Using the Loss-of-Signal Method

Introduced 

Flaw

TrueState 

Length 

(cm)

70°L 

Measured 

Length 

(cm)

60°L 

Measured 

Length 

(cm)

45°L 

Measured 

Length 

(cm)

70°S 

Measured 

Length 

(cm)

60°S 

Measured 

Length 

(cm)

45°S 

Measured 

Length 

(cm)

Sawcut A 3.28 ND -- ND -- -- --

Flaw A 1.07 ND -- ND -- -- --

Sawcut B 6.52 6.60 -- ND -- -- --

Flaw B 3.05 3.05 -- ND -- -- --

Sawcut C 3.63 3.81 -- ND -- -- --

Sawcut D 5.41 9.02 -- ND 3.30 3.30 4.06

Flaw C 4.36 3.43 -- ND ND 5.08 6.48

Sawcut E 4.37 5.21 -- ND ND ND ND

Sawcut F 5.97 5.84 -- 4.83 ND ND 4.70

Flaw D 1.33 ND -- ND -- -- --

Sawcut G 5.73 11.2 -- 7.87 -- -- --

Sawcut H 6.84 6.10 -- 8.13 -- -- --

Flaw E 3.38 3.30 -- ND -- -- --

RMS Error 2.12 n/a 1.59 2.11 1.58 1.63
9-6

Table 9.5 summarizes the −6 dB drop length-sizing results for the LF/SAFT inspection 

technique as a function of incident angle and wave mode for 400-kHz data relative to the truestate values for each introduced flaw. Dimensions are again in centimeters and the 

nomenclature remains the same as in previous tables.

Table 9.5 Summary of Length-Sizing Results for the 400-kHz Low-Frequency/SAFT 

Inspection Technique Using the −6 dB Drop Method

Introduced 

Flaw

TrueState 

Length 

(cm)

70°L 

Measured 

Length 

(cm)

60°L 

Measured 

Length 

(cm)

45°L 

Measured 

Length 

(cm)

70°S 

Measured 

Length 

(cm)

60°S 

Measured 

Length 

(cm)

45°S 

Measured 

Length 

(cm)

Sawcut A 3.28 ND ND 3.56 ND ND 4.32

Flaw A 1.07 ND ND ND ND ND ND

Sawcut B 6.52 3.81 3.05 7.87 ND 4.32 5.33

Flaw B 3.05 2.79 1.52 ND ND 2.79 3.30

Sawcut C 3.63 ND ND ND ND ND ND

Sawcut D 5.41 ND ND ND ND 3.81 5.33

Flaw C 4.36 ND ND ND ND 4.83 5.08

Sawcut E 4.37 ND ND ND ND ND ND

Sawcut F 5.97 ND ND 4.83 ND ND ND

Flaw D 1.33 ND ND ND ND ND ND

Sawcut G 5.73 6.10 4.32 5.59 ND 5.33 1.78

Sawcut H 6.84 5.33 4.83 5.84 ND 3.05 4.83

Flaw E 3.38 2.79 ND ND ND 2.03 3.30

RMS Error 1.43 2.26 0.92 n/a 1.85 1.69

Table 9.6 summarizes the LOS length-sizing results for the LF/SAFT inspection technique as a 

function of incident angle and wave mode for 400-kHz data relative to the true-state values for 

each introduced flaw. Dimensions are again in centimeters and the nomenclature remains the 

same as in previous tables.

A review of Table 9.5 shows that no trend is evident, the length-sizing results are less accurate,

and vary in both oversizing and undersizing of the true-state lengths in apparently random 

fashion. From the −6 dB drop length-sizing method, the data indicates that the 45° shear-wave 

inspection was more accurate (in a relative sense) in determining the introduced flaw lengths 

than the other modes at 400 kHz. Table 9.6 depicting the LOS measured length values also 

shows no trends regarding oversizing or undersizing the true-state lengths. In general, the LOS 

method is less accurate than the −6 dB drop method from a review of the 400-kHz data.
9-7

Table 9.6 Summary of Length-Sizing Results for the 400-kHz Low-Frequency/SAFT 

Inspection Technique Using the Loss-of-Signal Method

Introduced 

Flaw

TrueState 

Length 

(cm)

70°L 

Measured 

Length 

(cm)

60°L 

Measured 

Length 

(cm)

45°L 

Measured 

Length 

(cm)

70°S 

Measured 

Length 

(cm)

60°S 

Measured 

Length 

(cm)

45°S 

Measured 

Length 

(cm)

Sawcut A 3.28 ND ND 4.57 ND ND 3.81

Flaw A 1.07 ND ND ND ND ND ND

Sawcut B 6.52 4.32 3.30 8.89 ND 5.59 5.84

Flaw B 3.05 2.79 2.03 ND ND 2.79 3.05

Sawcut C 3.63 ND ND ND ND ND ND

Sawcut D 5.41 ND ND ND ND 3.56 5.59

Flaw C 4.36 ND ND ND ND 5.59 4.83

Sawcut E 4.37 ND ND ND ND ND ND

Sawcut F 5.97 ND ND 5.33 ND ND ND

Flaw D 1.33 ND ND ND ND ND ND

Sawcut G 5.73 6.86 3.81 5.33 ND 5.33 2.03

Sawcut H 6.84 6.35 5.08 7.11 ND 3.30 5.33

Flaw E 3.38 3.30 ND ND ND 4.06 3.30

RMS Error 1.13 2.13 1.26 n/a 1.65 1.46

In summary, the detection performance results indicate that the 70° longitudinal-wave 

inspections at 250 kHz provide the most meaningful data and result in the best detection results 

of any of the scan combinations evaluated here. Crack A and crack D were by far the most 

difficult introduced flaws to detect, as neither introduced flaw was detected using any of the 

scan combinations in this study. These two thermal fatigue cracks were the smallest in the 

depth dimension. Regarding length-sizing, the 400-kHz inspections provided more accurate 

results (of those introduced flaws that were detected) in comparison to the true-state lengths, in 

particular for the 45° inspection angle. The LOS length-sizing method provided an RMSE of 

3.4 mm (0.13 in.) for the 45° shear-wave inspections of thermal fatigue cracks, and proved to be 

the most accurate of the scan combinations. The −6 dB length-sizing method provided an 

RMSE of 9.2 mm (0.36 in.) for the 45° L-wave inspections for sawcuts. From the work reported 

here, it is shown that lower frequencies generally are not appropriate for the inspection problem 

in these piping components as the sound field is too large and the wavelength too long for 

effective detection and sizing of smaller cracks. With a range of longer wavelengths spanning 

0.79 cm (0.31 in.) to 2.32 cm (0.91 in.) and the large zone-focal dimensions of the sound field 

(discussed in Section 6) employed in this study, the resolution is generally too low relative to the 

size of the introduced flaws. The application of the LF/SAFT inspection technique does,

however, shed light on the effectiveness of lower examination frequencies and provides a 

general lower bound for the use of longer wavelengths for flaw detection from the far side.

9.4 Low-Frequency/SAFT-UT – Phase 2

The Phase 2 results show that the 400-kHz and 250-kHz probes are able to detect the 

10% through-wall notch from the far side of the weld using a variety of angles and with both 

shear and longitudinal waves. The scans along welds one and three show that the amplitude of 
9-8

the sound returning from the 10% through-wall notch is not strongly affected by the position of 

the pipe during the welding process.

None of the scans were able to detect the three implanted flaws in weld number two. The long 

wavelengths, 7.5 mm (0.3 in.) for the shear waves and 15 mm (0.6 in.) for the longitudinal 

waves, are larger than the 3.5-mm (0.14-in.) depth of the EDM notches, but are able to detect 

the EDM notches as they are very long and essentially fill the sound field with a continuous 

reflector. The implanted cracks are much shorter in length, with the 15% crack being only 

10.7-mm (0.42-in.) wide. With the shear-wave probe, which has the smaller wavelength, the 

largest implanted flaw is less than half a wavelength deep and only slightly more than one 

wavelength long. While this is within the theoretical limit (1⁄2 of a wavelength) for resolving a 

defect, the defects are not detectable from the far side of the weld. The longitudinal waves are 

at the far limit of being able to detect the 15% through-wall crack, as this crack is much less 

than 1⁄2 of a wavelength deep and just over 1⁄2 of a wavelength long.

The different welding techniques did not strongly affect the amount of sound returning from the 

10% through-wall notch. Table 9.7 shows the gains used when scanning the pipe specimen. 

The gains were chosen to provide a near-saturation response from the 10% through-wall notch 

and represent a rough measure of how well the weld transmits the ultrasound. Analysis of total 

system gain values indicate that the air-backed weld appears to be the most efficient at 

transmitting longitudinal waves. The water-backed weld appears to absorb 3–4 dB of the 

longitudinal waves when compared to the air-backed weld. The opposite was true for the shear 

waves, in which the water-backed weld absorbed 3 dB less sound than the air-backed weld.

Table 9.7 Gain (in dB) Settings Used for Scanning to Reach Near-Saturation for 10% 

Notch

400 kHz

Air-Backed Middle Weld Water-Backed

45° Shear 33 35 30

45° Longitudinal 29 31 33

60° Longitudinal 29 30 33

70° Longitudinal Na 29 30

250 kHz

Air-Backed Middle Weld Water-Backed

45° Longitudinal 17 17 21

70° Longitudinal 17 16 20

The true benefit of using the LF/SAFT inspection technique is the composite approach taken 

when all incident angles, frequencies, and wave modes can be evaluated together to take a 

comprehensive approach to the inspection and use SAA redundancy to capture the true value in 

the detection and discrimination capabilities of this technique. In this study, each scan 

combination was evaluated on its own merit, and in reviewing the individual detection 

performance results, the 250-kHz, 70° L-wave scans offered the most reliable and consistent 
9-9

combination of inspection parameters for effectively inspecting the piping segments from the far

side. These results echo the conventional-UT inspection results, and also show that the 

addition of the 45° incident angle provides a suitable means for accurate length-sizing for using

such a low-frequency regime. The LF/SAFT technique as employed in this study does not 

exploit the advantages of the SAFT signal processing algorithms to the fullest, as the zonefocused, low-frequency transducers provide a spatially wide but non-divergent sound field. For 

a SAFT-based inspection technique to hold any promise for addressing this inspection problem, 

a truly divergent beam (using a smaller transducer) and examination frequencies near 1–2 MHz 

might provide a more effective means for far-side detection and sizing of cracks through 

austenitic weld material. The SAFT signal processing enhancements are truly realized in three 

dimensions and the resultant focusing effects might very well provide a reliable and effective 

means for detection and sizing.

To provide the optimal conditions for employing and evaluating a SAFT-based inspection 

technique, a modeling and simulation effort is recommended. Modeling plays a key role in 

providing guidance for optimizing inspection frequencies, incident angles, and other important 

inspection parameters, and should be an integral part of strategic planning design of future 

experimental evaluations.

9.5 Phased-Array UT – Phase 1

The Phase 1 results show that the longitudinal-wave and the shear-wave arrays are both 

capable of detecting and length-sizing defects from the far side of a weld in wrought austenitic 

stainless steel piping. All circumferential flaws were detected, although crack A (15% throughwall) was difficult to discern from the weld-root response. For the TRL array, this is most likely 

due to the limited flaw face available for L-wave reflection, and the flaw face proximity (spatially 

and temporally) with weld-root responses. Because no corner-trap or tip-diffracted responses 

were evident from far-side flaws, only spectral reflectors can be used for detection. The TRS 

array was more effective in detecting the corner-trap response for crack A from the far side of 

the weld than the TRL array. This is the result of the better corner-trap response from shear 

waves and the larger aperture of the shear-wave array, allowing for a more tightly focused beam 

at the higher angles. Neither array was able to accurately depth-size the defects from the far

side of the weld as crack-tip signals could not be detected. Also, a LOS technique may not be 

feasible, as many defects have ultrasonic responses that go beyond the highest angles used by 

the probe and the sound fields produced by the probes are larger than the smaller flaws. For an 

example of this, see the sector scan in Section 7, Figure 7.13.

The results from Phase 1 show that PA technology provides a viable alternative to conventional 

ultrasonic techniques for detecting service-induced flaws (cracking) on both the near and far

side of austenitic piping welds. While this technology offers essentially the same sound fields 

as could be obtained with monolithic element (conventional) transducers, the ability to 

interrogate the weld with varied angular sound fields in near real time drastically reduces 

acquisition time on the component. During this study, full linear data scans of approximately 

508 mm (20.0 in.) of circumferential weld were inspected in less than one minute for each 

segment of specimen 02-24-15. This included discrete data sets swept through angles of 30° to 

70°, at 1° increments for the TRL array and 40° to 70° for the TRS array. If one attempted to 
9-10

duplicate this volumetric scanning conventionally using raster scanning, multiple set-ups and 

scans for the same length of piping weld would exceed several hours and did, in fact, exceed 

several hours for a single transducer in our conventional-UT trials on the same pipe specimens. 

Also, the phased-array measurements can be made with the weld crowns in the as-welded 

condition, while the 60° and possibly the 70° conventional raster scanning would require the 

weld crowns to be machined off to acquire similar far-side weld data.

Another asset of PA technology is evident during data review. Analyses occur off-line, with the 

analyst having options to simultaneously display varied images of raw, or merged, data sets. 

Software features such as gating, sound field, and measurement cursors, as well as traditional 

time and amplitude presentations, enable one to initially screen the data for optimum responses 

(as related to angles of sound propagation and cross sections of the weld) and to target 

suspected flaw regions. This iterative operation is performed by moving the respective cursor 

though the sector scan (S-scan) image(s) based on a given sound-beam angle, then walking 

through the B- or C-scan image(s) to produce A-scans for that "slice" of material. Once the 

suspect regions have been identified, the analyst may call up different images and preferred 

ultrasonic angles to aid interpretation of these regions. Conventional ultrasonic imaging and 

analysis systems do not generally provide these capabilities for data fusion and manipulation. 

However, the PA analysis system flexibility and large sets (about 50 megabytes each) for each 

linear scan performed during the Phase 1 trials indicate that thorough analyses of field data for 

many welds may require significantly more time than data acquisition.

9.6 Phased Array-UT – Phase 2

The results from Phase 2 show that the ultrasonic response from a continuous feature on the far

side of a weld can have variable amplitude along the feature’s length. This effect may cause 

the feature to appear as a series of shorter responses. While there was a strong centimeter-bycentimeter variability in the strength of the returning signal, the overall nature of the ultrasonic 

response from the 10% notches did not vary with the general azimuthal position along the fieldlike welds. In all welds of pipe specimen 3C-022, the ultrasonic response from the 10% 

through-wall notch was very similar in terms of positional variation. The results also show that 

the 1.4-mm (0.06-in.) wavelength shear waves are more strongly attenuated as they pass 

through the weld metal than the longitudinal waves with a 3.1-mm (0.12-in.) wavelength. While 

the longitudinal waves were redirected significantly as they passed through the weld, they were 

not strongly attenuated by the weld. The flaws in weld number two were detectable with both 

longitudinal- and shear-wave arrays from the far side of the weld.

The variation in amplitude response along the 10% notches may help explain the inability of the 

two arrays to detect tip signals from the far sides of the welds. While the longitudinal waves 

were better able to penetrate the weld metal, both the TRL and TRS arrays showed large 

variability in the response from a continuous 10% through-wall notch. This effect can be seen in 

the Phase 1 results as well.

There was little difference in the ultrasonic responses from standard reflectors with the different 

welding techniques used on the Phase 2 piping specimen. There was no evidence that 

suggests that the ―down-hand‖ weld used in the Phase 1 pipe specimen is significantly different 
9-11

from the field-like welds in the Phase-2 pipe specimen. Therefore, Phase 2 results strongly 

corroborate Phase 1 findings and show validity for field-welded piping.

There are several ways to improve flaw detectability and sizing from the far side of a weld in 

wrought austenitic materials. A larger-aperture longitudinal-wave probe should be better able to 

resolve the far-side defects, as it would have a more strongly focused beam at the higher 

angles. In addition, a lower-frequency longitudinal-wave probe would be better able to 

penetrate the weld metal. With the weld crowns removed, such a probe may be able to better 

depth-size the cracks by the loss-of-signal technique even in the absence of a tip signal. A 

lower-frequency longitudinal-wave array would also suffer less beam redirection in the material.

10-1

10 Longitudinal-Wave Sound-Field Assessments 

and Simulations

The evaluation of ultrasonic fields emitted from longitudinal-wave phased-array probes provides 

insights on the behavior of sound one might expect in a material undergoing volumetric 

examination. Computer simulations of these same sound fields in parallel with the actual 

characterizations allows one to verify that the focal laws employed for a given phased-array 

probe will produce adequate beam steering with sufficient field densities and provide focusing at 

the desired depth within the material.

Beam simulations and sound-field characterizations were performed for a 1.5-MHz, 1.5-D matrix 

phased-array longitudinal-wave probe on an austenitic stainless steel coupon at several incident 

angles and focal depths. These simulations were performed along with sound-field characterizations using commercial modeling software to verify and validate the phased-array focal laws. 

Focal laws were generated using the wrought austenitic piping and Rexolite wedge material 

properties, and manufacturer-supplied phased-array parameters (e.g., frequency, number of 

elements, element spacing, etc.) in conjunction with the specified wave mode, focal depth, 

sample volume, and incident angle(s).

The objective of the beam mapping exercise is to evaluate the effectiveness and coherence of 

phased-array sound-field propagation through wrought stainless steel (WSS) components with 

austenitic welds. This evaluation was conducted as a function of discrete incident angle, focal 

depth in the material, and probe position relative to the weld centerline. The sound fields were 

imaged through both parent material, WSS, and through parent material plus austenitic weld 

material, and subsequently analyzed visually for degradation effects (scattering, attenuation, 

field partitioning, etc.) and by measuring signal-to-noise and sound field dimensions (in both 

X- and Y-axes) at the −6 dB and −10 dB points. Longitudinal-wave sound fields produced by 

the PA probe were evaluated through a 36-mm (1.4-in.) thick WSS coupon at several discrete 

angles. The coupon was extracted from a larger WSS piping component with an austenitic weld

(a small coupon from specimen 3C-022 was used). One end of the coupon is comprised of an 

austenitic weld that was cut along the weld fusion line while the bulk of the coupon is comprised 

of WSS parent material. Both ends of the coupon were cut at a 60° angle, representative of a 

single-V counterbore, and then machined to accommodate sound-field characterizations. This 

provided the capability to obtain sound-field maps of identical inspection parameters through 

both WSS (alone), and through WSS plus the austenitic weld material. Figure 10.1 illustrates 

different views of the WSS coupon used in this study.
10-2

Figure 10.1 WSS Coupon Used for Sound Field Mapping. Left: Top view of WSS coupon 

with positional markings for probe position of various sound-field scans. 

Right: Side view of WSS coupon showing the austenitic weld and WSS 

material.

A hybrid immersion-scanning configuration where the water level in the scanning tank was filled 

so that the receiver pin-transducer was completely immersed while only the elements of the 

transmit PA probe were immersed. This was necessary in order to protect the electrical 

connections in the PA probe, which was not designed for full immersion. Sound fields 

generated by the transmitting PA probe were mapped at different incident angles and focal 

depths. The calculated exit point on the PA probe-wedge configuration was positioned at 

theoretical locations from the inspection ends of the coupon to achieve focusing at the plane of 

interest at focal depths of 15 mm (0.59 in.) and 25 mm (0.98 in.) for the various incident angles. 

See Table 10.1 for a matrix of the sound-field evaluations performed. Figure 10.2 shows a 

schematic setup for incident angle 70° with a focal depth of 15 mm (0.59 in.). For data 

acquisition, the assembly is rotated to position the 60° cut end in a horizontal plane, parallel to 

the scanning bridge.

Table 10.1 Sound Field Mapping Parameter Matrix for Data Acquired from Both Ends of 

the Coupon

Focal Depth Incident Angle

Distance Back from 

Coupon Edge

15 mm (0.59 in.) 70° 57 mm (2.24 in.)

25 mm (0.98 in.) 50° 50 mm (1.97 in.)

25 mm (0.98 in.) 60° 64 mm (2.52 in.)
10-3

Figure 10.2 Side View Schematic Showing Orientations of PA Probe, Specimen, and 

Pinducer

Prior to scanning, the WSS coupon was fitted with a PA probe and then mounted in a fixture 

that allowed free positioning (rotation) of the coupon and the probe assembly (Figure 10.3). The 

coupon was leveled relative to the X-Y scanning bridge (C-scan plane). The picture on the left 

in Figure 10.3 shows the transmit/receive longitudinal probe mounted on a Rexolite wedge and 

secured to the specimen using latex banding. Note that for this mapping study, only the 

transmit side of the probe was active and was operated in this mode via the ZETEC Tomoscan 

III phased-array system (Figure 10.4). Using ZETEC’s UltraVision 1.2R4 software package to 

control the Tomoscan III, the probe was pulsed with an 180V, 330-nsec square-wave signal 

(optimized for 1.5 MHz) with the desired output being a longitudinal-wave signal focused at the 

coupon face of interest.

A Valpey-Fisher pinducer was chosen to be the receiver in this study. It has a 1.5-mm 

(0.059-in.) diameter active aperture and a flat bandwidth of DC to 20 MHz in the receive mode. 

This type of miniature transducer was originally designed for shock-wave applications. More 

recently, newer applications have been identified where these probes are used as acoustic 

emission receivers and for sound-field characterization of ultrasonic transducers. Valpey-Fisher 

has developed these custom pinducers by incorporating standard NDT fabrication techniques, 

such as quarter-wave matching layers and dense backings, to dramatically increase bandwidth 

and sensitivity.
10-4

Figure 10.3 Left: Phased Array Probe Mounted and Secured on Austenitic WSS Coupon;

Right: Coupon Positioning Fixture for Sound Field Mapping

Figure 10.4 Phased-Array Data Acquisition System

Operating in a through-transmission configuration, the broadband pinducer acted as the 

receiver and was raster-scanned under the end of the coupon being characterized 

(Figure 10.5). Raster scanning set up involved using the PNNL-developed SAFT acquisition 

software to control the XYZR scanning bridge. The pinducer was attached via epoxy and zip 
10-5

ties to the end of a search tube and positioned normal to and under the face of the coupon to be 

scanned. The offset distance to the specimen (water path) was approximately 2 mm (0.08 in.). 

A quick and coarse XY raster scan with resolution of 2.54 mm (0.1 in.) in both axes was 

acquired so that adjustments could be made if necessary to place the beam response in the 

center of the C-scan view. Additionally, the maximum response in the image was adjusted to 

avoid saturation. The final raster scan resolution was set at 0.25 mm (0.01 in.) in the scan 

direction and 0.25 mm (0.01 in.) in the index direction over a total area of 25.4 mm × 25.4 mm 

(1.0 in. × 1.0 in.). The selected scanning resolution and area is sufficient to capture and show 

the resulting sound fields at a high degree of precision because at 1.5 MHz the wavelength in 

this material is on the order of 4 mm (0.15 in.). At this resolution, the mapping exercise will 

show minor changes/distortions in the shape and intensity of the sound fields.

Figure 10.5 Pinducer (highlighted by red arrow) Positioned Just Beneath a Coupon End 

During Beam Mapping. Side view through Plexiglas wall of automated 

immersion scanning tank.

The received signals from the pinducer were routed to a Ritec ultrasonic receive amplifier, 

where the signals were filtered and conditioned prior to digital sampling. Because the frequency 

to be mapped was 1.5 MHz, the receiver was set to filter out contributions below 100 kHz and 

above 3 MHz. This filtering reduces unwanted signals thus increasing the SNR of the beam 

image. The signal output from the Ritec receiver was routed to the SAFT data acquisition 

system where the analog signals were digitized (8-bit) and C-scan (X-Y axis) sound-field plots 

were generated and analyzed.
10-6

10.1 Results

The computer-generated beam simulations and the empirical sound-field maps resulting from 

specimen scans can be found in Figures 10.6–10.11. Figures 10.6, 10.8, and 10.10 display 

beam simulations with the B-scan side views on the left and B-scan end views on the right. An 

indication of focal length and position of maximum intensity are seen in the side views. The end 

views indicate the beam shape and size in the lateral – depth (y-z) plane. Figures 10.7, 10.9,

and 10.11 display images acquired from scanning the end of the specimen at the 60° slant, as 

shown in Figure 10.2. The images in the left of these figures are from the WSS side of the 

specimen while the images on the right side are from the WSS plus weld material, and display

sound field distortion in shape and coherence. Note that the data obtained from the actual 

specimen was along the 60° slant of the end while the beam simulation software is essentially 

on a 90° slant (vertical). The simulation display software only allows viewing of results along the 

three orthogonal axes. It should also be noted that the beam simulation only allows isotropic 

media.

Both the simulated beam images and the specimen data images were analyzed for size in the 

vertical and horizontal directions. Table 10.2 displays the sizing results from the beam 

simulations while Tables 10.3 and 10.4 provide measured information from the specimen 

images. The beams were sized at the −6 dB and −10 dB levels. Tables 10.3 and 10.4

additionally provide signal-to-noise ratios.

Figure 10.6 L-Wave Sound Beam Simulation at 50°, 25-mm (0.98-in.) Depth Focus; Side 

View on the Left and End View on the Right
10-7

Figure 10.7 Data Images at 50°, 25-mm (0.98-in.) Depth Focus, 19 dB, 50 mm (1.9 in.) Back 

from WSS Side on the Left and from the Weld Side on the Right

Figure 10.8 L-Wave Sound Beam Simulation at 60°, 25-mm (0.98-in.) Depth Focus; Side 

View on the Left and End View on the Right
10-8

Figure 10.9 Data Images at 60°, 25-mm (0.98-in.) Depth Focus, 27 dB, 64 mm (2.5 in.) Back 

from WSS Side on the Left and from the Weld Side on the Right

Figure 10.10 L-Wave Sound Beam Simulation at 70°, 15-mm (0.59-in.) Depth Focus; Side 

View on the Left and End View on the Right
10-9

Figure 10.11 Data Images at 70°, 15-mm (0.59-in.) Depth Focus, 19 dB, 57 mm (2.2 in.) 

Back from WSS Side on the Left and from the Weld Side on the Right

Table 10.2 Simulated Beam Size

10 dB Size 6 dB Size

Angle Focus

Distance 

Back X Y X Y

70° 15 mm

(0.59 in.)

57 mm

(2.24 in.)

11.7 mm 

(0.46 in.)

14.9 mm 

(0.59 in.)

9.1 mm 

(0.36 in.)

11.5 mm 

(0.45 in.)

50° 25 mm

(0.98 in.)

50 mm

(1.97 in.)

11.6 mm 

(0.46 in.)

11.6 mm 

(0.46 in.)

9.1 mm 

(0.36 in.)

10.6 mm 

(0.42 in.)

60° 25 mm

(0.98 in.)

64 mm

(2.52 in.)

13.0 mm 

(0.51 in.)

13.6 mm 

(0.54 in.)

10.1 mm 

(0.40 in.)

10.6 mm 

(0.42 in.)

Table 10.3 Wrought Stainless Steel Coupon End, Longitudinal-Wave Data

Angle Focus

Distance 

Back

Gain 

(dB)

Average 

Noise 

(Counts)

Peak 

Signal 

(Counts)

Signal 

to Noise

10 dB Size 6 dB Size 

X Y X Y

70° 15 mm

(0.59 in.)

57 mm

(2.24 in.)

19 5.92 39 6.59 19.1 mm

(0.75 in.)

21.3 mm

(0.84 in.)

14.7 mm

(0.58 in.)

15.5 mm

(0.61 in.)

50° 25 mm

(0.98 in.)

50 mm

(1.97 in.)

19 3.70 61 16.49 13.5 mm

(0.53 in.)

20.1 mm

(0.79 in.)

9.1 mm

(0.36 in.)

14.5 mm

(0.57 in.)

60° 25 mm

(0.98 in.)

64 mm

(2.52 in.)

27 12.80 116 9.06 17.8 mm

(0.70 in.)

22.9 mm

(0.90 in.)

12.7 mm

(0.50 in.)

16.0 mm

(0.63 in.)
10-10

Table 10.4 Coarse-Grain (Weld Side) Coupon End, Longitudinal-Wave Data

Angle Focus

Distance

Back

Gain 

(dB)

Average 

Noise 

(Counts)

Peak 

Signal 

(Counts)

Signal 

to Noise

10 dB Size 6 dB Size 

X Y X Y

70° 15 mm

(0.59 in.)

57 mm

(2.24 in.)

19 4.84 30 6.20 12.7 mm

(0.50 in.)

17.3 mm

(0.68 in.)

9.4 mm

(0.37 in.)

13.0 mm

(0.51 in.)

50° 25 mm

(0.98 in.)

50 mm

(1.97 in.)

19 4.08 44 10.78 15.2 mm

(0.60 in.)

20.3 mm

(0.80 in.)

7.9 mm

(0.31 in.)

14.5 mm

(0.57)

60° 25 mm

(0.98 in.)

64 mm

(2.52 in.)

27 11.80 92 7.80 16.3 mm

(0.64 in.)

20.8 mm

(0.82 in.)

11.9 mm

(0.47 in.)

16.0 mm

(0.63 in.)

The approximate sound-field dimensions based on the X and Y measurements in Tables 10.3

and 10.4 are plotted in Figure 10.12 at the −6 dB sizing level to assist in visualizing the relative 

change in sound fields as observed from the WSS side and from the WSS-plus-weld side of the 

specimen. At 50° and 60° the shapes are similar in the Y direction and slightly smaller in the 

X direction when passing through the WSS and weld material. At 70° there is a greater 

difference seen. In the X direction, the beam is 36% smaller after passing through the WSS and 

weld material and in the Y direction the beam is 16% smaller. At the higher angle (70°) the 

sound path is longer and therefore the beam has more opportunity for distortion which the data 

show. These are idealized plots because they do not show the image distortion, skewing, and 

amplitude variation that are evident in the actual specimen images.

Table 10.5 shows the difference in peak-signal response for the longitudinal-wave modality 

between the base material (WSS) and the weld. For example, at the 50° angle with a 25-mm 

(0.98-in.) focal depth, the peak weld response is 3 dB lower than the base material response. In 

all of the longitudinal data, the weld response is lower by 2–3 dB than the base material, which 

is a reasonable expectation. This reduction in peak response is likely due to the fact that the 

weld region is comprised of a different material than the base material thus creating boundary 

conditions that can attenuate or redirect sound from the desired target area.
10-11

Figure 10.12 Sound-Field Sizing Comparison at a −6 dB Level Showing WSS Results 

with Black Ovals and WSS + Weld Results with Magenta Ovals. Data are 

displayed for 50°, 60°, and 70° longitudinal waves

Table 10.5 Comparison of Peak Signal Response from the Weld to the Base Material

Angle Focus

Weld-to-Base Metal 

Response (dB)

70° 15 mm (0.59 in.) −2.4

50° 25 mm (0.98 in.) −3.0

60° 25 mm (0.98 in.) −2.1
10-12

10.2 Theoretical Analysis of Longitudinal-Wave Propagation through 

Austenitic Weld

To understand resulting complex wave fields as ultrasonic beams propagate through parent 

material (WSS) and then through austenitic weld material, a simplified mathematical model has 

been developed. Figure 10.13 shows a simplified representation of the propagation media 

under consideration. The austenitic stainless steel weld region is modeled as truncated 

V-shaped region and illustrated by solid heavy lines. For the sake of additional simplicity, the 

weld region is divided into a finite number of segments having planar interfaces, each having its 

own crystallographic axis of rotational symmetry typical of austenitic stainless having columnar 

grains. The dotted lines appearing in the schematic diagram represents the alignment of these 

crystallographic directions ([001] axis). More accurate and detailed segmentation of the weld 

region will be used in future calculations, when ultrasonic beams will also be more correctly 

approximated by finite Gaussian beams. The origin of the coordinate system (XOZ) is placed at 

the intersection of phased-array centerline and the surface of the WSS section of the specimen.

Figure 10.13 Schematic Diagram of Wrought Stainless Steel – Austenitic Stainless Steel 

Composite Media

In this report, we present a simplistic mathematical model of the ultrasonic beam propagation 

through the medium in question. Specifically, we begin by replacing the ultrasonic beam by a 

ray along its axis. At each interface between two generally anisotropic regions, mentioned in 

the previous paragraph, the beam is reflected and transmitted. Generation of various modes 

(L-, SV-, SH-waves) at each interface will be duly accounted for, as described below.

The weak-contrast reflection transmission coefficients at interfaces between two generally 

anisotropic media are given by

             1 1 2 2 3 3 4 4 5 5 6 6 R R , R R , R R ,R R ,R R ,R 1 R .

(10.1)

 

  



  



  

 

 

i i ijkl i j k l

i i i

e E c e p E P

R

2 v P p

(10.2)
10-13

where α = 1,2,3,4,5 for a non-degenerate incident wave and α = 1,2,3,4 for a degenerate 

incident wave. In the aforementioned equations, cijlk, and ρ represent elastic moduli and density 

at the incident side of the plane interface; 

c c , ijkl ijkl     

represent elastic moduli and density 

at the transmitted side of the plane interface; Pi, ni, Ei denote the slowness vector, normal to the 

interface, and unit polarization vector of the incident plane wave; 

   p , e ,R , 1, 2, 3 i i  

denote the 

slowness vectors and the unit polarization vectors of the reflected waves and the corresponding 

displacement reflection coefficients; 

   p , e ,R , 4, 5, 6 i i  

denote the slowness vectors and the 

unit polarization vectors of the transmitted waves and the corresponding displacement 

transmission coefficients.

The weak-contrast displacement transmission coefficient of the unconverted wave is given by

 



   

5

6

i i

1

R 1 e E R

for non-degenerate incident wave and 

 



   

4

5

i i

1

R 1 e E R .

These expressions for reflection-transmission coefficients also apply to evanescent-generated 

waves. Snell’s law for the generated waves can be expressed as

  p P n a , 1,2,3,4,5,6. i i i    

(10.3)

The normal slowness vector discontinuities can be calculated as the relevant solutions of the

sixth-order equations obtained by inserting Snell’s law into Christoffel equations.

To carry out the calculations, we have chosen the following physical properties of the base and 

the weld materials having same crystal elastic constants (Table 10.6). Table 10.7 shows the 

summary of computational predictions of beam propagation through weld at two different 

incidence angles. In both cases, beam exits the weld at a different location than the 

desired/designed focal point. Even though our formalism can track mode-converted beams, we 

present only the beam path for incident longitudinal wave. However, we do show that at the 

incident face of the weld, the beam splits into two beams – L-wave and SV-wave. Because of 

the two-dimensional nature of our model, no conversion to the SH-wave mode occurs.

More detailed calculations will be made in the future where (1) more realistic beam cross

sections will be used, (2) scattering attenuation will be included, and (3) macroscopic 

predictions of forward-scattered grain-induced noise accommodated. The last of these future 

enhancements of our computational strategy is important to predict signal-to-noise ratio.

Table 10.6 Material Properties of Base and Weld Metal

Material C11 (GPa) C12 (GPa) C44 (GPa) ρ (kg/m3) Orientation

WSS 219.2 136.8 109.2 7860 Random

Weld 219.2 136.8 109.2 7860 [001]
10-14

Table 10.7 Summary of Computational Results

Angle Focus

Point of Exit:

L-wave (x, z)

Point of Ideal Exit:

L-wave (x, z)

Trans. Coef.:

L-wave

(Entry Face)

Trans. Coef.:

SV-wave

(Entry Face)

70° 15 mm

(0.59 in.)

(41.35 mm, 17.57 mm)

(1.75 in., 0.69 in.)

(41.21 mm, 15.0 mm)

(1.62 in., 0.59 in.)

0.9621 0.1103

60° 25 mm

(0.98 in.)

(42.11 mm, 27.06 mm)

(1.66 in., 1.07 in.)

(43.30 mm, 25.0 mm)

(1.70 in., 0.98 in.)

1.1588 -0.1164

10.3 Discussion and Conclusions

From a review of the data provided in Tables 10.3 and 10.4 and the sound-field plots in 

Figures 10.6–10.11, the effectiveness and coherence of sound-field propagation can be 

compared in both parent material and austenitic weld material as a function of the various 

inspection parameters evaluated in this study. Because this comparative evaluation is aimed at 

a number of inspection parameters, it is important to discuss the relationship between wave 

speed (acoustic velocity), inspection frequency, and wavelength. The longitudinal wave velocity 

l v

in wrought stainless steel is 5790 m/s (0.228 in./μs). From the relationship in Eq. (10.4), 

  lv / f

(10.4)

where f is the inspection frequency (Hz) and λ is the wavelength, the wavelength for a frequency 

of 1.5 MHz was calculated to be 3.9 mm (0.15 in.) in WSS for longitudinal waves. A review of 

the phased-array probe beam simulations in Figures 10.6, 10.8, and 10.10 provide some 

baseline information regarding sound-field shape, contour, directionality, orientation, and 

intensity; however, the modeling software does not simulate the effects of material 

inhomogeneities or anisotropy (due to grain boundaries or austenitic weld structure). Thus, 

these results are what would be expected in an isotropic, fine-grained, homogeneous material 

structure and were used as baseline information for the sound fields propagating in the WSS 

parent material.

The phased-array probe used in this evaluation was constructed from 1.5-D arrays of small 

rectangular elements. The active area (aperture) for the 1.5-MHz TRL probe was 35 mm ×

17.5 mm (1.3 in. × 0.69 in.). This essentially results in a 2:1 aperture ratio. The effective 

aperture defines the focusing capabilities, directional characteristics, and overall shape of the 

primary sound-field lobe (when using all elements to beam form). In a non-scattering, 

homogeneous, and isotropic material structure, the primary sound-field lobes of this transducer 

configuration are expected to yield asymmetric distributions of acoustic energy, which is evident 

from a review of the dimensional data in Tables 10.2 through 10.4.

With an understanding of what the sound-field distributions should look like, it is quite apparent 

that the austenitic weld does not significantly contribute to degrading the coherence and 

structure of the longitudinal-wave sound field. It is observed that the compression-wave 

modality actually shows little degradation as the sound field propagates through the weld 
10-15

material from the far side. In this modality, particle oscillations occur in the same direction as 

the direction of the traveling sound field. The sound-field plots provided in Figures 10.7, 10.9, 

and 10.11 clearly show that the longitudinal-wave mode provides good penetration and the 

sound field maintains ample coherence and signal strength as it propagates through the 

austenitic weld from the far side. Only at the highest angles of insonification (70°) coupled with 

a shallow focal depth (15 mm [0.59 in.]) do the longitudinal waves start to exhibit slight 

degradation to the field shape, coherence, and structure, as is evident by a review of 

Figure 10.11. This is to be expected because for the larger incident angle, the ultrasonic beams 

strike each columnar zone interface of the weld with smaller incidence angle and thus suffers 

from larger scattering attenuation and generate greater forward-scattered signals. More 

detailed analysis of the signal-to-noise ratio at various angles of incidence will be carried out in 

the future.

The results of this evaluation show that longitudinal (compression-wave) phased array is an 

effective means for inspection of austenitic welds from the far side. When coupled with the 

appropriate wavelength relative to the size of the grain boundaries and weld structure, this 

method can provide a very effective inspection.

11-1

11 Phase 3 Evaluation

The Phase 3 evaluation incorporated lessons learned from Phases 1 and 2 of this study, and 

focused on employing phased-array-UT techniques only for examination of field-removed piping 

segments with serviced-induced IGSCC. These specimens (discussed in Section 5) were 

provided by the EPRI NDE Center, Charlotte, North Carolina.

Flaw data from the EPRI specimens were analyzed for detection with the three PA probes 

described in Section 6. The detection criteria are discussed here. Given the true state of a 

specimen, if a clear response that was distinct from geometry and possessing a good signal-tonoise ratio (SNR) was found at any of the skew angles in the region noted as flawed, a ―Yes‖ 

result was recorded for the flaw. If a weak signal that was discernable but not clearly separate 

from geometry or close to the noise levels was found in the region noted as flawed, a ―Marginal‖ 

result was recorded. If no signal was found in the region noted as flawed, then a ―No‖ result 

was recorded.

Figure 11.1 shows data from a region classified as ―Yes‖ in the detection results. Arrows point 

to the flaw signals in the three views. The top view shows an ID geometrical signal on the left 

edge that runs nearly the entire length of the scan. The flaw signal is farthest to the right in the 

top view with other material reflectors in the middle. The signal of interest is gated in the side 

view using red and blue vertical lines and these gated results are shown in the end view. 

Arrows in the end view point to the flaw.

Figure 11.2 shows data from a region that was categorized as ―Marginal‖ in detection. There is 

a signal in the region of interest that is slightly higher than the background but it does not have 

much length associated with it compared to surrounding area signals. Another signal in the 

data (end view) with similar amplitude stands out as a possible flaw but this region was not 

shown to be flawed. This could potentially lead to a false call in this area. False calls were not 

tracked in this analysis but contributed to the Marginal result as opposed to Yes detected calls.

Figure 11.3 shows data from a region that was classified as ―No‖ in detection. There is no 

signal in the flaw region that stands out from the surrounding area. A lack of signal might be 

evident in the region of interest in the end view but this was not used as one of the flaw criteria.
11-2

Figure 11.1 Example of a “Yes” Call

Figure 11.2 Example of a “Marginal” Call
11-3

Figure 11.3 Example of a “No” Call

The practice specimen data were taken to PNNL and analyzed shortly after data acquisition. 

Some of the images and detection calls are shown in Appendix G from both the near and far 

side. The data are presented from the downstream side first for a given specimen and then 

from the upstream side for the three PA probes. The secure data were later analyzed at the 

EPRI site. Data images and information on detection calls relative to particular specimens were 

not allowed to leave the EPRI testing room. Only approved summary information is included in 

this report. Results from the aggregate sets of data are summarized in Figure 11.4, with data 

plotted to show the percent detection (Yes, No, or Marginal) for all flaws with a particular PA 

probe. The ―Yes‖ calls are represented with blue boxes, the ―Marginal‖ calls with green 

triangles, and the ―No‖ calls with brown circles. Secure data are shown with dotted line 

connections and practice specimen data with solid lines. Results for the three probes from the 

near side and far side are provided. In most of the data sets, the far-side ―Yes‖ call is lower 

than the near-side ―Yes‖ call, as expected. However, from the secure set at 1.5-MHz TRL, it 

can be seen that the near-side and far-side ―Yes‖ detected calls are the same at approximately 

80%. Another interesting difference is shown for the 1.5-MHz shear wave on the secure data 

set, with far-side out-performing the near-side detections. Other differences in results from the 

secure and practice data sets may also be due to experience of the data analysts in 

interpretation, or in the specimens themselves. In general, the Yes calls fall in the 50% to 80% 

range for the far-side inspection.
11-4

Figure 11.4 Detection Calls on the Secure (S) and Practice (P) Data Sets as a Function of 

PA Probe and Near- and Far-Side Inspection

The secure and practice data results were combined to give a larger sample population and to 

protect the integrity of the secure set. A summary of the results are listed in Table 11.1. This 

combined data set represents a number of service-induced cracks in a variety of pipe and 

component configurations joined with an austenitic weld. The depth of the cracks were not 

known nor measured from the data as tip-diffracted signals were not sought out. The lengths of 

the cracks were variable, as reported by EPRI true-state information. This study focused on 

flaw detection, so crack lengths were not measured. In general, cracks on the small end of the 

observed length range could be confused with material noise, and cracks on the long end of the 

range could be confused with part geometrical reflectors such as counterbore or weld root. 

Medium to long indications whether from a flaw or geometry could also exhibit signal drop out 

due to sound-field redirection, scattering, or attenuation. These issues make signal 

discrimination on service-induced IGSCC more challenging.

Table 11.1 IGSCC Detection Summary

1.5-MHz TRL 2.0-MHz TRL 2.0-MHz TRS

Access Yes No Marginal Yes No Marginal Yes No Marginal

Near Side 73.5 14.7 11.8 78.8 18.2 3.0 70.6 26.5 2.9

Far Side 67.6 14.7 17.6 69.0 20.7 10.3 51.5 36.4 12.1

The detection summary shows that near-side inspection using methods similar to those 

described in this report will be superior with approximately 70% to 80% of the flaws being 

detected. On a far-side inspection using similar methods, the detection rate drops to 

approximately 50% for shear waves and is just under 70% for the longitudinal-wave modality.
11-5

Flaw detection generally involves two parts. One part of the detection process relates to signal 

amplitude. How high above the background noise level is the signal of interest? A 

measurement used for this process is the signal-to-noise ratio (SNR). A SNR of at least 2:1, or 

6 dB, is commonly desired, but current analytical techniques allow lower values with most 

procedures requiring analyses of flaw responses down to the material noise level. The SNR for 

these data were calculated from the peak signal responses from ―Yes‖ detected flaw responses 

and the average noise level around the flaws. Noise values were taken at the same part path 

as the signal responses. SNR values from the different detected cracks were averaged with 

results shown in Figure 11.5 for each probe and for near- and far-side access. Results indicate 

that the SNR values are good; all exceeding 10 dB (signal is 3.2 times the noise level). The 

higher near-side SNR for the TRS probe is possibly due to the lack of mode conversion taking 

place. From the far side, the SNRs are similar for each of the three probes for a detected flaw. 

With SNRs being adequate in these data sets, the challenge in flaw detection then becomes 

one of signal discrimination.

Figure 11.5 Signal-to-Noise Ratios for “Yes” Detected Cracks

The second major variable for flaw detection is signal discrimination—is the ultrasonic response 

being generated from a geometrical condition, material noise, or a flaw? Ideally, one would 

determine the weld area profile along the pipe axis using a contour gauge on the OD surface 

and UT thickness readings to exhibit the approximate ID contour. These contour measurements were not made during this exercise. In fact, the presence of weld crowns on the fieldremoved specimens prohibited the collection of normal-beam ultrasonic measurements to 

accurately map the ID surface in the weld region. If the weld area is diagramed, indications 

from subsequent specimen evaluations can be mapped to the originating location in the 

specimen profile. Signals mapped to the counter bore, weld root, etc. can be classified as 

geometrical indications. Often geometrical signals have significant length and amplitude 

associated with them. A strong counterbore signal provides a reference point for all other 

signals whether the weld area profile is known or not. As shown in previous sections of this 

report, an example of a known configuration for an austenitic stainless steel piping specimen 
11-6

typical of that found in a primary coolant BWR recirculation system, was fabricated with sawcuts 

and implanted thermal fatigue cracks. The approximate weld cross section is shown in 

Figure 11.6. PA data of a flaw acquired from the far side of the weld is shown in Figure 11.7. 

The sector side view on the left of Figure 11.7 shows, from left to right in the image, a 

geometrical signal from the inner-diameter mismatch step, a geometrical signal from the weld 

root, and the flaw signal (as if UT data was acquired from the OD surface on the right side of the 

weld shown schematically in Figure 11.6). 

Data from the EPRI specimens were generally not as straightforward and were cluttered with 

many signals to consider. Figure 11.8 shows near-side data from a practice specimen. The 

geometrical signal from the far side of the weld is clearly seen and can be used as a reference. 

The side view shows three main reflectors between the red and blue horizontal lines. This 

region is gated and shown in the top view. The far-side geometrical signal is identified and all 

other signals referenced to it. Notice that the far-side geometry produces a strong signal, but 

with signal dropout evident. The flaw signal is sandwiched between two geometrical signals in 

the side and top views and the near-side geometry signal does not extend the full length of the 

image. The vertical red and blue lines in the side view are used to gate out the flaw signal with 

results shown in the end view. The flaw signal is at the bottom of the end view and is boxed 

with the magenta lines. Other signals are detected at that same axial position but are lower in 

amplitude. For this reason, the flaw was considered ―Yes‖ detected. In general, for data 

analyses, a narrow vertical gate such as that shown in the side view is moved across the image 

while looking at the end view for signals of interest. In the PNNL analyses, analysts considered 

whether there was a signal in the area documented as flawed that was separable from other 

signals and of sufficient amplitude, when making these calls. All images in this report are from 

merged data, but unmerged data were also reviewed during the flaw detection analyses.

Figure 11.6 Cross-Sectional Schematic Showing the Weld Geometry and Location of an 

Implanted Crack in the Weld Heat Affected Zone
11-7

Figure 11.7 Far Side Data from a Specimen with the Cross Section Shown in Figure 11.6. 

Geometrical signals from the inner-diameter mismatch step and weld root 

are noted. The flaw signal is also identified.

Figure 11.8 Data Images Showing a Flaw Signal Separated from Geometrical Indications

The combined data indicate that near-side detection of IGSCC provides for a 71%–79% 

detection rate when using the applied phased-array methods and scanning techniques. 

However, far-side detection rates drop, with shear-wave detection from the far side representing 

the lower extreme at 52%, as compared to the longitudinal-wave detections at 68% or 69%. 
11-8

Earlier work (Anderson et al. 2003; Anderson et al. 2006; Anderson et al. 2008), based on 

machined-notch responses and TFCs, showed that a longitudinal sound field passed through 

the weld with a minimal amplitude loss but sound-field redirection was observed. The shearwave sound field (having a shorter wavelength) was severely attenuated (primarily by a 

scattering mechanism) by the weld material, but showed less redirection and beam skewing. At 

the frequencies typically applied, shear waves do not transmit through austenitic weld 

microstructures as efficiently as longitudinal waves, therefore resulting in lower detection rates. 

Another, albeit undocumented, contribution to the detection rates involves the location of the 

flaws on the actual pipe segments. If IGSCC flaws were very close to the edges of the 

specimens, the optimized PA probe footprints may have precluded a proper insonification near 

these edge areas. This problem would not be an issue in the field, where access to the full 

circumferential weld length is available.

PNNL was supplied with flaw true-state information, so analyses should not be considered 

representative of blind tests or field inspections. The investigation was focused on assessing 

current PA ultrasonic techniques for austenitic welds. In this regard, the detection results are a 

best-effort, given equipment and inspection conditions; that is, the arrays were not specifically 

optimized for the weld conditions encountered, and, more importantly, the weld crowns were not 

removed, which would have provided greater access for each of the PA modalities and probes 

applied. Typically, the detection process involved first identifying geometrical reflectors based 

on axial position and circumferential extent. A strong ID-geometrical signal that had a large 

circumferential extent provided a good reference point. The geometrical reflectors were then 

removed from further consideration and the remaining signals evaluated based on the true-state 

knowledge. If an indication was present in the location of a known flaw and the signal was 

separable from other signals by amplitude or length, for example, the indication was called a 

detected flaw. The data, whether from the near side or the far side, were generally cluttered 

with many signals to consider. The smaller footprint of the 2.0-MHz TRL probe provided better 

access to some flaws, and in a few cases, allowed the only access. Its sound field, however, 

was not as well focused or coherent as the 1.5-MHz TRL probe. Far-side detection rates were 

similar with both of these probes, but it is believed that detection improvement would be 

obtained if the weld crowns were removed. This would ensure full insonification of the far-side 

area of interest. The results might then be augmented with a 1.0-MHz PA evaluation that would 

be less effected by attenuation, scattering, and beam skewing. At this lower frequency, the 

probe footprint (due the required aperture for proper focusing) would be larger, so weld crown 

removal would be required for a meaningful inspection. The shear-wave inspection did not add 

value to the longitudinal-wave results of this study. While it is recognized that shear-mode nearside inspections are an acceptable industry practice, this limited study did not show any 

improvement in the shear-mode results (even on the near side of the welds) as compared to the 

longitudinal-mode results.

In order to fully understand UT flaw detection capabilities from the far side of austenitic welds, 

weld crown removal would be required. Signal discrimination could be improved by more 

effectively peaking of ultrasonic signals where weld crowns presently limit this function. More 

accurate specimen profiling in the critical weld area, presently not possible with the weld crowns 

in place, would also improve the data interpretation. It is also possible that flaw-tip signals could 

be detected in a linear or raster PA scan if closer access to the flaw could be achieved.
12-1

12 Summary and Conclusions

The ultrasonic examination of austenitic coarse-grained structures continues to be a challenge 

to the industry during inservice inspection of nuclear power plant components. This is 

especially true for detection and characterization of service-induced flaws that may initiate at 

welds where access for examination may be limited to a single side of the weld. This limited 

access is generally caused by outside surface geometries encountered; for example, pipe-tovalve, nozzle-to-pipe, or elbow-to-pipe configurations, but examination access may also be 

affected by rough as-built surface conditions of the weld crown. For these reasons, this study 

was conducted to begin to ascertain the effectiveness of ultrasonic technologies for reliable

examination of austenitic piping welds when flaws are manifested on the far side of these welds.

In the case of wrought stainless steel piping components, it is the austenitic weldment and asbuilt conditions that adversely impact the inspection effectiveness. Coupled with the need to 

reliably and consistently detect cracks from the far side of the weld, manual ultrasonic 

inspection techniques used in industry do not achieve the desired detection reliability. In this 

study, spatially-encoded data sets using conventional inspection techniques, phased-array 

techniques, and low-frequency/SAFT inspection techniques were employed to examine a set of 

introduced flaws in piping specimens containing austenitic welds when inspecting from the far

side. The piping specimens developed for the first two phases of research included a 610 mm 

(24 in.) OD, with a wall thickness of 36 mm (1.4 in.). The phase three specimens were salvaged 

from piping that had been in service at commercial operating plants, and were nominally

711 mm (28.0 in.) in OD, with thickness ranging from 20–41-mm (0.8–1.6 in.). The frequencies 

chosen for this work are in the range of those used for examining piping welds in the field. A 

significant amount of data were acquired for each UT technique and extensive analyses 

performed on this large database from which the following conclusions and recommendations 

are made.

It has been shown that the far-side detection of inner-diameter surface-breaking flaws is 

possible with precise data acquisition and rigorous analysis methods. In addition, acceptable 

length-sizing of the flaws has also been demonstrated. However, because no diffracted flaw tip 

signals could be observed with any of the UT methods applied, accurate through-wall depthsizing may not be feasible. Figure 12.1 provides a graphic representation of the detection 

capabilities for each of the technologies employed.

The phased-array technology provided the best overall detection performance, followed by 

conventional UT. Flaws greater than 20% through-wall were detected by at least one technique 

for each of the methods used; however, very small flaws (<10% through-wall) were only 

consistently detected by phased array. Note the through-wall range of 11–20% included a 

19% axially-oriented flaw that was not detected by any of the methods. This was expected, as 

the techniques used by all methods were aimed at more structurally significant circumferentiallyoriented flaws because only scans perpendicular to the weld were performed. Figure 12.2

compares images from the implanted and machined flaws for the best parameters of each 

method applied during this study. The capability to length-size flaws on the far side of welds is 

summarized by the comparison in Table 12.1.
12-2

Figure 12.1 Far-Side Detection Results for All Techniques Applied on All Thermal Fatigue 

Cracks and Machined Sawcuts and Notches as a Function of Through-Wall 

Flaw Size

Table 12.1 Far-Side Length-Sizing Comparison for All Methods Applied

Thermal Fatigue 

(RMSE)

Best 

Technique

Sawcuts

(RMSE)

Best 

Technique

Conventional 2.2 mm 70° −6 dB 4.0 mm 70° LOS

LF/SAFT 3.4 mm 400 kHz, 

45°S - LOS

9.2 mm 400 kHz, 

45°L −6 dB

Phased Array 6.3 mm TRL LOS 6.3 mm TRL LOS

The conventional-UT method set a high performance benchmark for detection and length-sizing 

of flaws during these laboratory tests. This is somewhat surprising, based on poorer results 

obtained during previous round-robin trials (Heasler and Doctor 1996). Only the smallest 

reflectors could not consistently be detected with conventional transducers, and the best RMSE 

was obtained by the conventional method. The reader is cautioned that, while the conventional 

transducers represent those typically used during manual performance demonstrations, all data
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Figure 12.2 Images of Flaws Comparing Results from Conventional-UT, Low-Frequency/ 

SAFT and Phased Array-UT Methods
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were acquired with automated scanners, and digitally stored for off-line image analyses. 

Therefore, the increased accuracy of scans and ability to thoroughly evaluate the large amount 

of data acquired during these tests significantly contributed to the success of the conventional 

UT applied. It is not believed that manual techniques, interpreted in real-time, would perform at 

this level. The down-side of automated conventional examinations is that weld crown removal 

may be needed to facilitate raster scans, and these scans can typically require long durations 

(on the order of hours for large-diameter piping). In addition, multiple scans at various 

inspection angles and propagation modes are indicated to ensure a high probability of detection.

A summary review of the low-frequency/SAFT inspection results indicate that the detection 

performance of the 70° longitudinal-wave inspections at 250 kHz provided the most meaningful 

data and the best detection results of any of the LF/SAFT scan combinations evaluated here. 

Crack A and crack D were by far the most difficult introduced flaws to detect, as neither 

introduced flaw was detected using any of the scan combinations for this inspection technique. 

With regard to length-sizing, the 400-kHz inspections provided more accurate results (of those 

introduced flaws that were detected) in comparison to the true-state lengths, in particular for the 

45° inspection angle. The LOS length-sizing method provided an RMSE of 3.4 mm (0.13 in.) for 

the 45° shear-wave inspections of thermal fatigue cracks, and proved to be the most accurate of 

the scan combinations. The −6 dB length-sizing method provided an RMSE of 9.2 mm

(0.36 in.) for the 45° L-wave inspections for sawcuts, as shown in Table 12.1. From the work 

reported here, it is shown that lower frequencies generally are not appropriate for reliable and 

consistent detection of small-dimension flaws in these piping components as the sound field is 

too large and the wavelength too long for effective detection and sizing of smaller cracks. With 

a range of longer wavelengths spanning 7.9 mm (0.31 in.) to 23.2 mm (0.91 in.) and the large 

zone-focal dimensions of the sound field (discussed in Section 6) employed in this study, the 

resolution is generally too low relative to the size of the introduced flaws as a 10% flaw is 

3.8 mm (0.15 in.) in through-wall size. The zone-focused transducer may not be appropriate in 

this application, as the non-divergent beam results in a reduction in SAFT focusing in the 

processed images. The application of the LF/SAFT inspection technique does, however, shed 

light on the effectiveness of lower examination frequencies and provides a general lower bound 

for the use of longer wavelengths for flaw detection from the far side.

As reported above, the phased-array method provided the best detection results with the 

highest number of flaws detected. The capability of phased array to generate a broad range of 

discrete, incrementally steered beams with both shear and longitudinal waves increases the 

probability that an optimum incidence angle will occur for all, including small, spectral reflectors. 

For the austenitic welds and implanted/machined flaws examined in this study, good signal-tonoise ratios were obtained by both longitudinal- and shear-wave arrays, and far-side lengthsizing was well within the acceptable range (RMSE of 19 mm or 0.75 in.) required by the ASME 

Code Section XI, Appendix VIII, Supplement 2. Data sampling and digitization rates during 

continuous line scans slightly increase the RMSE for length-sizing of circumferential flaws, but 

this error, as stated above, is well within acceptable tolerances. In addition, acquisition rates on 

the order of inches per second are easily achievable and off-the-shelf analysis software is 

available to simplify off-line interpretation. There has been some hesitancy in the industry to 

use phased-array probes due to the initial costs for instrumentation and the expense of 
12-5

individual arrays. However, these costs are expected to be outweighed by the savings achieved 

with shorter scan times and the increased performance level of the method.

It is important to note that, as with any UT technique applied, acoustic physics must be 

optimized to produce the quality of inspection results observed during these laboratory tests. 

For this reason, modeling of the sound fields significantly contributed to the selection of 

appropriate UT parameters. This is especially true when using phased arrays to achieve 

optimum performance while controlling fabrication costs, as the arrays are much more 

expensive than conventional transducers.

During Phase 1 of this work, it was shown that far-side detection and length-sizing of planar 

reflectors, including implanted thermal fatigue cracks and small sawcuts, could be adequately 

performed. It should be noted that these results may vary for service-induced flaws such as 

IGSCC or primary water stress corrosion cracking (PWSCC), which tend to exhibit more 

branching and faceting than the planar flaws used in this study. In general, these planar flaws 

are expected to produce strong specular reflections that can be more readily detected, whereas 

branched and faceted flaws may return noncontiguous (or partitioned) sound fields, which can 

be redirected at multiple angles, thus more difficult to discriminate from other acoustic noise 

sources. The results were further validated in Phase 2 on piping fabricated to simulate fieldweld microstructures. Prior to the laboratory trials, it was believed that the anisotropic nature of 

field-weld microstructures could vary azimuthally, and subsequently, affect the transmitted 

acoustic field with respect to circumferential weld position. It was shown that, while small 

(approximately 6 dB) through-weld variations in acoustic beam intensities were evident along 

the weld length, this effect was random, and no overall influence based on welding position was 

seen to affect transmission efficiency. In other words, no relation of welding position to acoustic 

transmission quality could be ascertained.

After initial studies (Phases 1 and 2) on inspection of machined flaws and implanted thermal 

fatigue cracks through austenitic stainless steel piping welds showed favorable results, PNNL 

was invited to examine specimens containing service-induced IGSCC at the EPRI NDE Center. 

Phased-array ultrasonic data were acquired on a series of specimens with three different probes 

and two different inspection modalities. The specimens were divided into two groups—a 

practice set and a secure set. Full access to the specimens was given for the practice set and 

the data were taken to PNNL for flaw detection analysis. During data analyses, the flaw true 

state on these practice specimens was known. The secure data were acquired with only OD 

access to the specimens. Secure data analyses took place at the EPRI NDE Center with a 

proctor in place to ensure specimen security. True-state information was also provided for the 

secure data set.

The practice and secure data results were combined to maintain integrity of the secure set and 

showed that TRL inspection techniques produced a far-side detection rate of approximately 

69%. The TRS results were lower at 52% and this data showed much more scattering of the 

beam in passing through the austenitic weld material. SNRs in both longitudinal- and shearwave responses were good for detected flaw signals, suggesting that signal discrimination 

remains the most challenging aspect of flaw detection.
12-6

The 2.0-MHz TRL probe, while not producing as well defined a beam as the 1.5-MHz TRL 

probe, had a smaller footprint and therefore allowed better access to some of the far-side area 

of interest. This was an advantage in tight geometrical configurations. Lower-frequency probes 

are less susceptible to beam steering, scattering, and attenuation, but require a larger probe 

design footprint and therefore make weld crown removal necessary for a far-side inspection. 

Additional justification for weld crown removal includes enhanced signal discrimination by 

allowing more effective peaking of the ultrasonic signal and better inner diameter profiling of the 

specimen in the complex weld region. Access afforded by weld crown removal may also allow 

tip signals to become evident for enhanced flaw detection and depth-sizing.

This work was systematically conducted to evaluate the basic inspection performance of 

conventional-UT, low-frequency/SAFT-UT, and phased-array UT methods by iteratively 

evaluating notches, sawcuts, machined reflectors, fabricated thermal fatigue cracks, and finally, 

service-induced, field-removed IGSCC. As the industry progresses towards developing an 

effective and reliable inspection approach for far-side examinations in components with 

austenitic welds, insights gained from the subject study that may prove useful are: modeling 

could be employed to determine optimal inspection parameters for improved phased-array UT 

probe and inspection protocol design; modeling of the inspection process, weld microstructures, 

physical acoustic propagation phenomena, and other critical aspects of the inspection could 

also be employed; and metallographic studies could be performed on existing specimens and 

results would be interpreted and correlated with the findings of the modeling and measurement 

activities.
13-1

13 References

Anderson MT, SL Crawford, SE Cumblidge, AA Diaz and SR Doctor. 2008. "A Comparison of 

Ultrasonic Flaw Responses as Observed through Austenitic Stainless Steel Piping Welds." In 

Proceedings of the Sixth International Conference on NDE in Relation to Structural Integrity for 

Nuclear and Pressurised Components, EUR 23356 EN-2008, pp. 798–806. October 8–10, 

2007, Budapest, Hungary. eds: M Bieth and J Whittle. European Communities.

Anderson MT, SE Cumblidge and SR Doctor. 2003. "Applying Ultrasonic Phased Array 

Technology to Examine Austenitic Coarse-Grained Structures for Light Water Reactor Piping." 

In Third EPRI Phased Array Inspection Seminar. June 9–11, 2003, Seattle, Washington.

Anderson MT, AA Diaz, SE Cumblidge and SR Doctor. 2006. "Capabilities of Ultrasonic 

Techniques for the Far-Side Examination of Austenitic Stainless Steel Piping Welds." In Fifth 

International Conference on NDE in Relation to Structural Integrity for Nuclear and Pressurised 

Components, May 10–12, 2006, San Diego, California. eds: M Bieth and J Whittle. European 

Commission Joint Research Centre.

Doctor SR, GJ Schuster, LD Reid and TE Hall. 1996. Real-Time 3-D SAFT-UT Evaluation and 

Validation. NUREG/CR-6344, PNNL-10571, U.S. Nuclear Regulatory Commission, 

Washington, D.C.

Gruber GJ, RL Edwards and PD Watson. 1995. "Fabrication of Performance Demonstration 

Initiative Specimens with Controlled Flaws." In 13th International Conference on NDE in the 

Nuclear and Pressure Vessel Industries. May 1995, Kyoto, Japan.

Heasler PG and SR Doctor. 1996. Piping Inspection Round Robin. NUREG/CR-5068, 

PNL-10475, U.S. Nuclear Regulatory Commission, Washington, D.C.

Lemaitre P. 1994. Report on the Evaluation of the Inspection Results of the Wrought-toWrought PISC III Assemblies No. 31, 32, 33, 34, 35 and 36. PISC III Report No. 33, Programme 

for Inspection of Steel Components, Joint Research Centre, EEC, Petten, The Netherlands.

MacDonald D, M Dennis, J Landrum and GP Selby. 2000. Phased Array UT Performance on 

Dissimilar Metal Welds: Interim Report. TE-1000116, EPRI Nondestructive Evaluation Center, 

Electric Power Research Institute, Charlotte, North Carolina.

Moysan J, A Apfel, G Corneloup and B Chassignole. 2003. "Modeling the Grain Orientation of 

Austenitic Stainless Steel Multipass Welds to Improve Ultrasonic Assessment of Structural 

Integrity." International Journal of Pressure Vessels and Piping 80:77–85.

Poguet J, J Marguet, F Pichonnat and L Chupin. 2001. "Phased Array Technology: Concepts, 

Probes and Applications." In 3rd International Conference of NDE in Relation to Structural 

Integrity for Nuclear and Pressurized Components. November 14–16, 2001, Seville, Spain.

Schuster GJ, SL Crawford, AA Diaz, PG Heasler and SR Doctor. 2010. Reliability of Ultrasonic 

In-Service Inspection of Welds in Reactor Internals of Boiling Water Reactors. PNNL-14561, 

U.S. Nuclear Regulatory Commission, Washington, D.C.

Appendix A

Piping Specimen Fabrication Photographs

A-1

Appendix A

Piping Specimen Fabrication Photographs

Figure A.1 Austenitic Pipe Segments Shipped from PNNL for Flaw Implantation

Figure A.2 Liquid Penetrant Inspection of Weld Preparation Surfaces
A-2

Figure A.3 Implant Locations

Figure A.4 Preparing Weld Bevel Area for Flaw Implantation
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Figure A.5 Flaw Coupons
A-4

Figure A.6 Quality Assurance of Implantation Coupons

Figure A.7 Welding Set-up During Flaw Implantation
A-5

Figure A.8 Upstream Pipe Segment Ready for Joining to Downstream Segment After 

Flaw Implantation is Completed

Figure A.9 Root Pass of Full Penetration Weld Near Implanted Flaw
A-6

Figure A.10 Welding Root Pass with MIG

Figure A.11 Inter-Pass Filler Welding of Full Penetration Girth Weld
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Figure A.12 Completed Girth Weld Prior to Axially Sectioning the Pipe

Appendix B

Enhanced Far-Side Analysis Using a

Synthetic Aperture Focusing Technique

B-1

Appendix B

Enhanced Far-Side Analysis Using a

Synthetic Aperture Focusing Technique

B.1 Introduction

This appendix illustrates the value of using SAFT processing of the raw ultrasonic data and is 

typical of the improvements for all of the specimens.

B.2 Part Inspected

The raw data was taken on pipe segment 2 from the downstream side of the weldment. This 

segment included Flaw C, which was on the far side of the weld (upstream side). Due to 

physical constraints, only the steeper, 60° angles were used for this inspection, and far-side 

access due to the weld crown limited insonification of the weld region. In this case, the 60° 

sound path was only adequate to detect a part of the flaw’s through-wall extent.

B.3 Data Analysis

Analysis was performed on a typical, industry standard 2.0 GHz dual-processor computer 

workstation. The processing time varies depending on the size of the data file and the size of 

the synthetic aperture. The wider the SAFT processing aperture angle, the more summations 

are required and the longer the processing. Processing of the 60° shear, 190-MB data file from 

pipe segment 2 with a 12° synthetic aperture angle took approximately 14 hours. A B-scan end 

view image comparison of the raw data and the processed data are provided as Figures B.1 

and B.2.

B.4 Results

The raw data contained a significant amount of "noise" at and above the 6 dB level. This is 

evident in Figure B.1 as the greens, yellows, and reds indicate signal response in the upper end 

of the dynamic range. Using a typical 6 dB signal drop for sizing would make this image very 

difficult to interpret for detection and sizing of defects. Figure B.2 demonstrates how the 

12° SAFT processing has removed much of the incoherent noise that would make analysis of 

the raw data difficult. Additionally, Flaw C has been enhanced by the processing and is 

indicated in red. The image represents the entire range of data collected, and it has not been 

envelope detected. The location of the pipe specimen ID is at 34.3 mm (1.35 in.) in these 

images.
B-2

Figure B.1 Pipe Segment 2, B-scan End View, Raw Data

Figure B.2 Pipe Segment 2, Raw Data from Above, After SAFT Processing with a 12° 

Synthetic Aperture, B-scan End View
Appendix C

Conventional UT Data and Analysis Images

C-1

Appendix C

Conventional UT Data and Analysis Images

All scales are in inches

End view images are given in each figure as well as length and depth sizing information and 

noise characteristics. The data includes the 60 and 70 degree shear modality. 

Appendix C Content:

Sawcut A, 8% 1.5 and 2.25 MHz Near side

Flaw A, 15% 1.5 and 2.25 MHz Near side

Sawcut B, 28% 1.5 and 2.25 MHz Near side

Flaw B, 43% 1.5 and 2.25 MHz Near side 

Sawcut C, 7% 1.5 and 2.25 MHz Near side 

Sawcut A, 8% 1.5 and 2.25 MHz Far side 

Flaw A, 15% 1.5 and 2.25 MHz Far side

Sawcut B, 28% 1.5 and 2.25 MHz Far side

Flaw B, 43% 1.5 and 2.25 MHz Far side 

Sawcut C, 7% 1.5 and 2.25 MHz Far side 

Sawcut D, 19% 1.5 MHz Near and Far side 

Flaw C, 64% 1.5 MHz Near and Far side 

Sawcut E, 12% 1.5 MHz Near and Far side 

Sawcut F, 19% 1.5 MHz Near and Far side 

Sawcut G, 18% 1.5 MHz Near and Far side 

Sawcut H, 26% 1.5 MHz Near and Far side 

Flaw E, 48% 1.5 MHz Near and Far side 
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