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Instrumented impact testing and its application to the study of acicular ferrite steels McConnell, Paul 1978-12-31

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INSTRUMENTED IMPACT TESTING AND ITS APPLICATION TO THE STUDY OF ACICULAR FERRITE STEELS  by PAUL McCONNELL B.S., Case Western Reserve University, 1970  A THESIS SUBMITTED IN PARTIAL FULFILLMENT OF THE REQUIREMENTS FOR THE DEGREE OF MASTER OF APPLIED SCIENCE  in THE FACULTY OF GRADUATE STUDIES Department of Metallurgy  We accept this thesis as conforming to the required  standard  The University of B r i t i s h Columbia January, 1978 0  Paul McConnell, 1978  In p r e s e n t i n g t h i s  thesis  in p a r t i a l  fulfilment of  the requirements f o r  an advanced degree at the U n i v e r s i t y of B r i t i s h Columbia, the L i b r a r y s h a l l make i t I  f u r t h e r agree  freely available  that permission  for  I agree  reference and  f o r e x t e n s i v e copying o f  this  that  study. thesis  f o r s c h o l a r l y purposes may be granted by the Head of my Department or by h i s of  this  representatives. thesis  It  is understood that copying or p u b l i c a t i o n  f o r f i n a n c i a l gain shall  written permission.  Department The U n i v e r s i t y o f B r i t i s h  2075  Wesbrook Place Vancouver, Canada V6T 1W5  Columbia  not be allowed without my  ABSTRACT  An instrumented impact test (IIT) machine was constructed and calibrated using s t a t i c and dynamic loading.  The theory and  fundamentals of IIT have been reviewed.  Tests were performed validity criteria.  to assess the proposed ASTM IIT  The requirements that the fracture time be  greater than 3 times the period of specimen o s c i l l a t i o n s and  1.1  times the electronic response time appear to be conservative.  The 2  data confirm that adoption of the c r i t e r i o n , B > 2.5  (K^/a ^) ,  ensures plane s t r a i n fracture, whereas the acceptance of a l i n e a r l o a d - t o - f a i l u r e condition ( i . e . , P  M A V  < P ) rv  may not be conservative  enough. For general y i e l d f a i l u r e s , crack i n i t i a t i o n was shown to occur p r i o r to the attainment of maximum load.  Thus, i n i t i a t i o n  energies calculated by assuming that crack i n i t i a t i o n occurs at the maximum load are nonconservative.  The dynamic properties of two acicular f e r r i t e p i p e l i n e steels were characterized by IIT.  The Information obtained, p a r t i -  c u l a r l y the fracture toughness parameters and the i n i t i a t i o n energies, revealed s i g n i f i c a n t inadequacies i n the toughness s p e c i f i c a t i o n s and test methods presently used by the pipeline industry.  iii  Tests performed  to assess the significance of testing,  standard Charpy V-notch specimens versus f u l l pipe wall thickness Charpys showed that lower upper shelf energies were obtained f o r the f u l l wall specimens.  However, the magnitude of the t r a n s i t i o n  and lower shelf energies and the t r a n s i t i o n temperatures were similar.  Fatigue precracked standard Charpys specimens absorbed much lower energies and had higher t r a n s i t i o n temperatures  than  did the standard specimens.  Tests were also performed  to assess the s t r a i n aging  behaviour of the two acicular f e r r i t e s t e e l s .  Strain aging the  semi-killed s t e e l resulted i n a decrease i n the propagation energy, with no change i n the magnitude of the i n i t i a t i o n energy.  For this  s t e e l , strain-aging does not increase the p o t e n t i a l for crack initiation.  Tests also revealed that s i t e s near the seam weld of  the pipe made with that semi-killed s t e e l had experienced s u f f i c i e n t pipe-forming s t r a i n and thermal energy from the welding process to exhibit s t r a i n age e f f e c t s .  The f u l l y k i l l e d acicular f e r r i t e  steel  did not s t r a i n age; i t s strength and toughness were increased upon aging.  The instrumented impact test provided fracture toughness data that correlated very well with that obtained by more conventional  iv  fracture toughness testing techniques.  The t o t a l fracture energy  from a standard Charpy test was shown to often mask the fracture toughness value of a material.  The i n i t i a t i o n energy obtained from  testing a precracked Charpy specimen accurately indicated the r e l a t i v e magnitude of the fracture toughness, however.  V  TABLE OF CONTENTS Page ABSTRACT  1  TABLE OF CONTENTS  .. .  1  v  LIST OF FIGURES  x  LIST OF TABLES  i  : x v ± ±  ACKNOWLEDGEMENTS  x  DEDICATION  i  x  x  x  1.  INTRODUCTION  1  2.  INSTRUMENTED IMPACT TESTING  3  2.1  Introduction  3  2.2  Instrumented Impact Test Machine  8  2.2.1  Machine Design  8  2.2.2  Instrumentation  H  2.2.3 . C a l i b r a t i o n 2.2.4  2.3  Test Variables  6  1  9  2.2.4.1  Drop Height  1  9  2.2.4.2  Temperature  2  2  2.2.4.3  Instrumentation Parameters  Interpretation of Load-Time Data 2.3.1  ..  1  V a l i d i t y C r i t e r i a f o r Load-Time Signals 2.3.1.1  Response Time  2.3.1.2  Signal O s c i l l a t i o n s  2  3  2  ^  2  4  2  6  2  8  vi  Page 2.3.1.3 2.3.2  Impact Energy  Data Reduction from Load-Time Curves  32  2.3.2.1  32  Energy  \  2.3.2.1.1  1  2.3.2.1.2  V e l o c i t y Reduction Correction  35  2.3.2.2  Deflection  38  2.3.2.3  Dynamic Y i e l d Strength  38  •2.3.2.4  Fracture Toughness Calculations  39  2.3.2.5  Computer Programs  40  2.3.2.6  Data Sheet  40  E f f e c t s of Test and Specimen Parameters  41  2.4.1  Significance of Test V a l i d i t y C r i t e r i a  41  2.4.1.1  I n e r t i a l Loading Effect  41  2.4.1.2  E f f e c t s of Impact V e l o c i t y  2.4.1.3  Electronic Response Time  2.4.2  2.5  33  Compliance Correction f o r I n i t i a t i o n Energy  2.4  31  .. 44 46  Specimen Parameters  50  2.4.2.1  Notch Radius  50  2.4.2.2  Notch Angle  54  2.4.2.3  Specimen Thickness  57  Crack I n i t i a t i o n  57  2.5.1  High Speed Movie Films  60  2.5.2  E l e c t r i c a l Resistance Study  62  2.5.3  Reduced Energy Tests  66  vii  Page 3.  INSTRUMENTED IMPACT STUDY OF ACICULAR FERRITIC PIPELINE STEELS  70  3.1  Acicular F e r r i t i c Steels  70  3.2  Pipeline Applications  73  3.3  Fracture Control i n Pipelines  76  \  3.4  3.5  Test Program  i  5  8  3.4.1  Steels/Pipelines  3.4.2  Metallography  3.4.3  Instrumented Impact Test Specimens  8  3.4.3.1  Specimen Preparation and Configuration..  88  3.4.3.2  Specimen Dimensions  95  8  8 7  8  3.4.3.2.1  Standard Charpy V-Notch Specimens ..  96  3.4.3.2.2  Precracked Charpy Specimens  96  3.4.3.2.3  F u l l Wall Charpy Specimens  9 7  Instrumented Impact Test Results 3.5.1  6  Absorbed Energy 3.5.1.1  98 '  1  Comparison of the Two AF Steels  3.5.1.1.1  3.5.1.1.1.3 3.5.1.1.1.4  Crack P a r a l l e l to Pipe Axis Crack P a r a l l e l to R o l l i n g Direction Crack Transverse to R o l l i n g Direction  1  Hi  Standard Charpy Data  3.5.1.1.1.1 3.5.1.1.1.2  1  1  ..  1  1  HI 120 124  Crack Transverse to Pipe Axis .. 126  viii  Page 3.5.1.1.2  F u l l Wall Charpys  3.5.1.1.2.1  Crack P a r a l l e l to Pipe Axis  3.5.1.1.2.2  Crack P a r a l l e l to R o l l i n g Direction  130  Crack Transverse to R o l l i n g Direction  132  3.5.1.1.2.3 3.5.1.1.3  Precracked Charpys  .. 128  134  3.5.1.1.3.1  Crack P a r a l l e l to Pipe Axis  3.5.1.1.3.2  Crack P a r a l l e l to R o l l i n g Direction  134  Crack Transverse to R o l l i n g Direction  138  3.5.1.1.3.3  3.5.1.2  128  .. 134  Significance of Specimen Size and Notch Acuity. 138 3.5.1.2.1  AF-1 Steel  138  3.5.1.2.1.1  Crack P a r a l l e l to Pipe Axis  3.5.1.2.1.2  Crack P a r a l l e l to R o l l i n g Direction  143  Crack Transverse to R o l l i n g Direction  144  3.5.1.2.1.3  3.5.1.2.2  AF-2 Steel  .. 138  145  3.5.1.2.2.1  Crack P a r a l l e l to Pipe Axis  3.5.1.2.2.2  Crack P a r a l l e l to R o l l i n g Direction  146  Crack Transverse to R o l l i n g Direction  146  3.5.1.2.2.3  .. 145  3.5.1.3  Conclusions of Absorbed Energy Study  147  3.5.1.4  Drop Weight Tear Test Correlations  149  ix  Page  3.6  4.  3.5.2  Dynamic Y i e l d S t r e n g t h s  155  3.5.3  Load-Time Behaviour  157  3.5.4  Fractography  166  S t r a i n Age Study  168  3.6.1  E f f e c t s o f S t r a i n i n g and S t r a i n Aging  169  3.6.2  S t r a i n Aged S i t e s i n AF-1 P i p e  176  DYNAMIC FRACTURE TOUGHNESS  183  4.1  Introduction  183  4.2  The C a l c u l a t i o n o f F r a c t u r e Toughness Parameters  from  IIT Data  187  4.2.1  L i n e a r - E l a s t i c Fractures  187  4.2.2  E l a s t i c - P l a s t i c Fractures  189  4.2.2.1  J-Integral  189  4.2.2.2  Crack Opening Displacement  193  4.2.2.3  Equivalent  194  4.2.2.4  C r i t i c a l Crack S i z e s  Energy Method  195  4.3  Dynamic F r a c t u r e Toughness of P i p e l i n e S t e e l s  196  4.4  Correlations  207  4.4.1  R e l a t i o n s h i p Between Dynamic S t r e s s F a c t o r and C r a c k I n i t i a t i o n Energy  Intensity  4.4.2  Comparisons  4.4.3  C r i t i c a l Flaw S i z e s  4.4.4  E m p i r i c a l C o r r e l a t i o n s Between K j and Other Material Properties  207  Between K j d and S t a t i c a l l y Obtained  219 I (  225  X  Page  5.  4.4.4.1  K-j-^ versus Charpy Energy  225  4.4.4.2  K  228  I d  vs Y i e l d Strength  CONCLUSION  231  5.1  Conclusions  231  5.2  Suggestions for Future Work  234  REFERENCES  237  APPENDIX A  246  APPENDIX B  247  APPENDIX C  251  APPENDIX D APPENDIX E  . •  ?  • 256 257  xi  LIST OF FIGURES  Figure Number  Page  2.1  Instrumented impact machine  2.2  (a)  2.3  9  Diagram of tup showing p o s i t i o n of s t r a i n gauges (b) Schematic of instrumented tup c i r c u i t r y and ITT components Closeup view of tup, a n v i l s , centering device, and test specimen  2.4  Instrumented impact load-time photographs (a) e l a s t i c - p l a s t i c fracture (b) l i n e a r - e l a s t i c fracture  2.5  E f f e c t of impact v e l o c i t y , V , on load-time trace (a) V = 5.46 m/s (b) V = 3.46 m/s  12 12 13  ....  25  Q  45  0  0  2.6  Effect of electronic response time, T^, on loadtime trace (a) T = 2.3 ms (c) T = 0.0729 ms (b) T = 0.719 ms (d) T = 0.0007 ms R  R  R  R  48  2.7  E l e c t r i c a l resistance study of crack growth . . . . (a) Load-time curve Scale: 500 l b / d i v . x 0.2 ms/div. (b) Potential-time curve Scale: 1 mV/div. x 0.2 ms/div.  64  3.1  AF-1 photomicrograph, 225X (a) unetched (b) etched, 2% n i t a l  89  3.2  AF-2 photomicrographs, etched 2% n i t a l , 363X  ..  90  3.3  AF-1 SEM photomicrograph (3000X) and X-ray energy analysis of inclusions  91  xii  Figure Number 3.4  3.5 3.6 3.7 3.8  Page AF-2 SEM photomicrographs and X-ray energy analysis of inclusions  92-94  (a) 480X (b) 1000X (c) 4000X AF-1 steel load-time photographs, crack parallel to pipe axis  99  AF-2 steel load-time photographs, crack parallel to pipe axis  100  AF-1 steel load-time photographs, crack parallel to rolling direction  101  AF-2 steel load-time photographs, crack parallel to rolling direction  102  3.9  AF-1 steel load-time photographs, crack transverse to rolling direction 103  3.10  AF-2 steel load-time photographs, crack transverse to rolling direction 104  3.11  AF-1 steel fracture surfaces, crack parallel to pipe axis  105  AF-2 steel fracture surfaces, crack parallel to pipe axis  106  AF-1 steel fracture surfaces, crack parallel to rolling direction  107  3.12 3.13 3.14  AF-2 steel fracture surfaces, crack parallel to rolling direction  3.15  AF-1 steel fracture surfaces, crack transverse to rolling direction  3.16  AF-2 steel fracture surfaces, crack transverse to rolling direction  3.17  AF-1 IIT absorbed energies vs temperature, standard Charpys, crack parallel to pipe axis  116  3.18  AF-2 IIT absorbed energies vs temperature, standard Charpys, crack parallel to pipe axis  116  xiii  Figure Number 3.19  3.20  3.21  3.22  3.23  3.24  3.25  3.26  Page AF-l/AF-2 IIT average absorbed energies vs temperature, standard Charpys, crack parallel to pipe axis  • • 116  AF-1 IIT absorbed energies vs temperature, standard Charpys, crack parallel to rolling direction  121  AF-2 IIT absorbed energies vs temperature, standard Charpys, crack parallel to rolling direction  121  AF-l/AF-2 IIT average absorbed energies vs temperature, standard Charpys, crack parallel to rolling direction  121  AF-1 IIT absorbed energies vs temperature, standard Charpys, crack transverse to rolling direction  125  AF-2 IIT absorbed energies vs temperature, standard Charpys, crack transverse to rolling direction  125  AF-l/AF-2 IIT average absorbed energies vs temperature, standard Charpys, crack transverse to rolling direction  125  AF-1 IIT absorbed energies vs temperature, crack transverse to pipe axis  127  3.27  AF-1 IIT absorbed energies vs temperature, f u l l wall/standard Charpys, crack parallel to pipe axis . .. 129  3.28  AF-2 IIT absorbed energies vs temperature, f u l l wall/standard Charpys, crack parallel to pipe axis . .. 129  3.29  AF-l/AF-2 IIT average absorbed energies vs temperature, f u l l wall Charpys, crack parallel to • axis • pipe  129  3.30  AF-1 IIT absorbed energies vs temperature, f u l l wall/ standard Charpys, crack parallel to rolling direction.. 131  3.31  AF-2 IIT absorbed energies vs temperature, f u l l wall/ standard Charpys, crack parallel to rolling direction.. 131  xiv  Figure Number 3.32  3.33  3.34  3.35  3.36  3.37  3.38  3.39  3.40  3.41  3.42  Page AF-l/AF-2 IIT average absorbed energies vs temperature f u l l wall Charpys, crack p a r a l l e l to r o l l i n g d i r e c t i o n  131  AF-1 IIT absorbed energies vs temperature, f u l l wall/standard Charpys, crack transverse to rolling direction  133  AF-2 IIT absorbed energies vs temperature, f u l l wall/standard Charpys, crack transverse to r o l l i n g direction  133  AF-l/AF-2 IIT average absorbed energies vs temperature, f u l l wall Charpys, crack transverse to r o l l i n g d i r e c t i o n  133  AF-1 IIT absorbed energies vs temperature, precracked/standard Charpys, crack p a r a l l e l to pipe axis  135  AF-2 IIT absorbed energies vs temperature, precracked/standard Charpys, crack p a r a l l e l to pxpe axis AF-l/AF-2 IIT average absorbed energies vs temperature, precracked Charpys, crack p a r a l l e l to pipe axis  "I 3S x  -  ,J  135  AF-1 IIT absorbed energies vs temperature, precracked/standard Charpys, crack p a r a l l e l to rolling direction  136  AF-2 IIT absorbed energies vs temperature, precracked/standard Charpys, crack p a r a l l e l to rolling direction  136  AF-l/AF-2 IIT average absorbed energies vs temperature, precracked Charpys, crack p a r a l l e l to r o l l i n g d i r e c t i o n  136  AF-1 IIT absorbed energies vs temperature, precracked/standard Charpys, crack transverse to rolling direction ..  139  XV  Figure Number 3.43  3.44  3.45  3.46  3.47  3.48  3.49 3.50  3.51  3.52  3.53  3.54  Page AF-2 IIT absorbed energies vs temperature, precracked/standard Charpys, crack transverse to rolling direction  139  AF-l/AF-2 IIT average absorbed energies vs temperature, precracked Charpys, crack transverse to r o l l i n g d i r e c t i o n .. .. ,.  139  AF-1 DWTT percent shear and IIT f u l l wall Charpy absorbed energies vs temperature, crack p a r a l l e l to pipe axis  151  AF-1 DWTT percent shear and IIT precracked Charpy absorbed energies vs temperature,crack p a r a l l e l to pipe axis  151  AF-2 DWTT percent shear and IIT f u l l wall Charpy absorbed energies vs temperature,crack p a r a l l e l to pipe axis  152  AF-2 DWTT percent shear and IIT precracked Charpy absorbed energies vs temperature, crack p a r a l l e l to pipe axis  152  AF-l/AF-2 IIT dynamic y i e l d strengths vs temperature  , . 156  AF-1 IIT maximum and general y i e l d loads vs temperature, standard Charpys, crack p a r a l l e l to pipe axis  158  AF-2 IIT maximum and general y i e l d loads vs temperature, standard Charpys, crack p a r a l l e l to pipe axis  158  AF-1 IIT maximum and general y i e l d loads vs temperature, standard Charpys, crack p a r a l l e l to r o l l i n g direction  158  AF-2 IIT maximum and general y i e l d loads vs temperature, standard Charpys, crack p a r a l l e l to r o l l i n g direction  158  AF-1 IIT maximum and general y i e l d loads vs temperature, standard Charpys, crack transverse to rolling direction  159  xvi  Figure Number 3.55  3.56  Page AF-2 IIT maximum and general yield loads vs temperature, standard Charpys, crack transverse to rolling direction  159  AF-1 IIT maximum and general yield loads vs temperature, standard Charpys, crack transverse to pipe axis  159  3.57  Schematic of variation i n general yield load, fracture load, and absorbed energy with temperature for a Charpy specimen. Effect of notch on T_ .. .. 160  4.1  AF-1 IIT dynamic fracture toughness vs temperature, crack parallel to pipe axis 197  4.2  AF-2 IIT dynamic fracture toughness vs temperature, crack parallel to pipe axis 197  4.3  AF-1 IIT dynamic fracture toughness vs temperature, crack parallel to rolling direction 198  4.4  AF-2 IIT dynamic fracture toughness vs temperature, crack parallel to rolling direction 198  4.5  AF-1 IIT dynamic fracture toughness vs temperature, crack transverse to rolling direction 199  4.6  AF-2 IIT dynamic fracture toughness vs temperature, crack transverse to rolling direction 199  4.7  ^Id^ initiation energy for acicular ferrite steel IIT Kid data meeting different validity requirements  211  AF-2 IIT dynamic J-Integral fracture toughness vs static J-Integral fracture toughness, crack parallel to rolling direction  216  4.8  v  s  4.9  Reduced pearlite steel IIT dynamic fracture toughness, crack parallel to rolling direction  .. 216  4.10  Dynamic fracture toughness vs standard Charpy energy for acicular ferrite steels  227  4.11  Dynamic fracture toughness vs dynamic yield strength for acicular ferrite steels  227  E.l  Temperature gradient in Charpy specimens taken from near AF-1 seam weld for strain age study 259  xvii  LIST OF TABLES Table  Page  2.1  Dynamic C a l i b r a t i o n Results  18  2.2  Response Times  27  2.3  Validity Criteria  32  2.4  Comparison of Valid and Invalid Data as Determined by t < 3T  43  2.5  Comparison of V a l i d and Invalid Data as Determined by  t < 1.1 T  49  R  2.6  Notch Radii Study  53  2.7  Notch Angle Study  56  2.8  High Speed Movie Film Results  61  2.9  Reduced Energy Test Results  68  3.1  Proposed Fracture Control Requirements for A r c t i c Pipelines  8 4  3.2  Steel Compositions  3.3  Standard Charpys - Average Absorbed Energies  3.4  F u l l Wall Charpys - Average Absorbed Energies .. .. H3  3.5  Precracked Charpys - Average Normalized Energies 1  3.6  D u c t i l i t y Index - Steels AF-1 and AF-2  US  3.7  E f f e c t of Notch Acuity  137  3.8  C  16-  3.9  Strain Age Study  1 7  3.10  Strain Age Sites i n Steel AF-1  1 8 3  v  100 Temperatures  87 .. .. 112  '  xviii  Table  Page  4.1  Transition Temperatures  201  4.2  Fracture Toughness and Energy Data  203  4.3  K i Values for Different V a l i d i t y C r i t e r i a  214  4.4  C r i t i c a l Crack Sizes  222  E.l  Equivalent Aging Conditions  261  I(  xix  ACKNOWLEDGEMENT S  I wish to thank the s t a f f and graduate students of the Department of Metallurgy who provided assistance throughout t h i s research, including Jim Brezden, Yvonne Chung, Orman Leszek, Jim Walker, and i n p a r t i c u l a r , my good friend V. Venkateswaran for h i s contributions i n writing the computer programs.  Bob Butters' and Ed  Klassen's e f f o r t s i n the design and construction of the test machine are commendable.  The support and advice of Mr. R.J. Cooke of the Alberta Gas Trunk Line Company, who provided much of the material tested i n this program, and the Aluminum Company of Canada, who provided f i n a n c i a l support, i s g r a t e f u l l y appreciated.  The guidance of my supervisors, Drs. E.B. Hawbolt and N.R. Risebrough,  i s acknowledged.  Bruce's constant interest i n the project  and h i s open-door p o l i c y are especially and sincerely appreciated.  To my wife, Janny, and my children, Heather and Grant, I express my utmost gratitude for their patience.  Janny's considerable  assistance during the test program and preparation of the manuscript was invaluable: she, a f t e r a l l , actually "broke" the specimens!  XX  1. INTRODUCTION  This thesis i s the culmination of a study of instrumented impact testing (IIT).  Instrumented  impact testing d i f f e r s from the  standard Charpy test i n that the load-time response of a specimen i s measured and recorded during the fracture event.  The t o t a l absorbed  energy, the area under the load-deflection curve, can be separated into two components, EI and EP.  EI i s the energy to i n i t i a t e the  crack, whereas EP i s the crack propagation energy.  In addition, the  data provide a measure of the dynamic y i e l d strength and the dynamic fracture toughness of a material.  The project objectives were: 1.  To construct, c a l i b r a t e , and render operational an  instrumented impact machine. 2.  To conduct a series of tests by which the proposed IIT  v a l i d i t y c r i t e r i a could be assessed. 3.  To conduct tests to show the advantages of IIT as  compared with standard Charpy testing.  These tests included a study  of the effects of specimen geometry and notch acuity. 4.  To demonstrate  the a p p l i c a b i l i t y of IIT by character-  i z i n g the d i r e c t i o n a l dynamic properties of two X70 acicular f e r r i t e pipeline steels.  This study included an assessment of their p o t e n t i a l  for s t r a i n age embrittlement.  - 2 -  Considerable d e t a i l on the theory and applications of IIT has been included i n this thesis to provide the necessary basis for future studies. i  A comparison of the toughness properties of the' current generation of Canadian X70 HSLA p i p e l i n e steels was  included i n this i l  study since these steels are being proposed for use i n Northern pipelines.  gas  The toughness c h a r a c t e r i s t i c s of these steels are of  prime importance,  since one of the most important design problems  i s the prevention of pipe f a i l u r e .  The work has shown that IIT i s i  p a r t i c u l a r l y valuable i n providing rapid, inexpensive, jand detailed !  dynamic fracture toughness data.  V a l i d fracture toughness values,  p a r t i c u l a r l y at high s t r a i n rates, are d i f f i c u l t to obtain by other test procedures.  '  It i s hoped that t h i s thesis w i l l provide the necessary background f o r future studies using instrumented  impact  testing.  - 3 -  2.  2.1  INSTRUMENTED IMPACT TESTING  Introduction  Instrumented impact testing (IIT) i s becoming widely  accepted  as a means to rapidly and inexpensively generate data describing the ( dynamic response of materials  2—13) '  . The American Society f o r  Testing and Materials recently devoted an entire Symposium to the (2) subject  , and are currently preparing a tentative ASTM IIT s p e c i -  f i c a t i o n to be included i n the 1978 Annual Book of ASTM Standards (14-15)^ A conventional impact testing machine (Charpy, dynamic tear, etc.) can be instrumented  by l o c a t i n g calibrated load c e l l s on the  s t r i k i n g hammer (tup) near the contact points. displacement  signal i s obtained i n place of the conventional  energy to f a i l u r e information. be made of:  1)  3)  total  From such curves, determinations can  the d i f f e r e n t i a t i o n between crack i n i t i a t i o n and  crack propagation strengths;  A load-time or load-  energies;  2)  the dynamic y i e l d and fracture  dynamic fracture toughness values, and many other  useful parameters.  For Charpy-type tests, these parameters can be  calculated by applying notch-bar three-point bending theories, with due regard for metallurgical p r i n c i p l e s .  Although c e r t a i n assumptions  must be made to permit these c a l c u l a t i o n s , meaningful, reproducible, and generally acceptable information can be generated.  - 4 -  The instrumented impact test also has a l l the advantages of the  standard Charpy test: reveals temperature t r a n s i t i o n s , low cost,  simple procedure, high s t r a i n rate, large sampling c a p a b i l i t y , established correlations with service performance.  Fracture and toughness tests generally measure either energy absorbed or c r i t i c a l loads from which design data, for example, stress intensity factors, may be derived.  The IIT, when employing a pre-  cracked Charpy specimen, y i e l d s both energy and fracture toughness data.  Several workers have made unique contributions to the development of IIT.  The e a r l i e s t references to obtaining load-deflection curves representative of the dynamic response of materials appeared i n the l a t e 1920's, although the f i r s t uses of s t r a i n gages to record the loads were not reported u n t i l t h i r t y years l a t e r ^ ^  ^ \  Augland was the f i r s t to correlate the energy results obtained from integrating the IIT load-time curve with the energy measured d i r e c t l y from the pendulum d i a l gauge of a standard Charpy machine  (18)  He i s also credited with deriving the expression which corrects the value of the area under the load-time trace to account for the  - 5 -  reduction i n hammer v e l o c i t y during impact.  Tardif and Marquis were apparently the f i r s t to suggest that the  t o t a l energy of the impact event could be separated into the (19)  energy to i n i t i a t e and the energy to propagate the crack  .  They  also proposed that the dynamic fracture toughness might be measured from IIT data. Fearnehough and Hoy used IIT data to calculate the dynamic yield strength^ ^. 2  Their paper, and that of Kobayashi, et  al^^  described i n d e t a i l the fracture process i n terms of the load-time data obtained over a range of temperatures.  In the l a t e 1960's, with the increasing interest i n fracture and fracture mechanics, papers dealing with IIT became more numerous. Commercial IIT units became available.  Many authors contributed by  reporting the dynamic fracture toughness values obtained from IIT (21-25) for a range of materials  .  Radon and Turner were the f i r s t (25)  to employ fatigue precracked Charpy specimens  .  Server and  Tetelman published a comparison between the fracture toughness data generated from f u l l size compact tension specimens, tested at various s t r a i n rates, and the fracture toughness data obtained using IIT and small precracked Charpy specimens.  The data obtained from the much  less expensive IIT (- $500) compared favourably with that generated  - 6 -  (if.)  from the standard compact tension tests (- $2 m i l l i o n )  Significant advances have been made since that time, p a r t i c u l a r l y i n assuring that the IIT data obtained i s representative (13 27-32) of the true mechanical response of the test specimen Approximately  '  half the published l i t e r a t u r e on IIT i n recent years  has been concerned with r e f i n i n g the instrumentation, c r i t i c a l l y analyzing the nature and effect of inherent signal o s c i l l a t i o n s and/or establishing v a l i d i t y c r i t e r i a by which a l l tests can be compared.  In addition, a considerable amount of research has been (2 32-33) directed to applying IIT to composites and non-metallics ' Unfortunately, almost a l l the work published p r i o r to the early 1970's must be considered suspect.  I n s u f f i c i e n t information  regarding experimental parameters i s included i n these e a r l i e r papers to ensure that the now  established v a l i d i t y c r i t e r i a were met  during  „ (34-35) testing  Data that can otherwise be obtained only at high costs (eg.,  fracture toughness test programs requiring f u l l size specimens);  or that cannot be obtained by other means (eg., including high s t r a i n rate y i e l d  dynamic data, per se,  strengths and dynamic stress i n t e n s i t y  factors) can e a s i l y be generated with IIT.  - 7 -  Precracking specimens, to simulate naturally occurring fatigue flaws, enhances the test and has been shown to provide sharper t r a n s i t i o n temperature curves and lower i n i t i a t i o n energies. Precracking i s considered essential to obtain fracture  toughness  data, i t being required to ensure that the minimum fracture toughness parameters may be measured.  IIT of precracked Charpys have  been shown to give t r a n s i t i o n temperature curves which correspond closely to those of the 5/8-in dynamic tear t e s t s ^ ' ' ' ^ 38)^ 7  3  Correlations with other tests, including the Battelle-Drop Weight Tear Test (used to determine full-thickness percent shear or the n i l - d u c t i l i t y t r a n s i t i o n temperature), have also been attempted  ..  with some success f  (7, 39-40)  Having the advantage of requiring simple, inexpensive specimen preparation while y i e l d i n g v a l i d energy, strength, and fracture toughness data, ensures that IIT w i l l become more important i n the future.  Standardization of test techniques and the e s t a b l i s h -  ment of v a l i d i t y c r i t e r i a w i l l further the acceptance and growth of IIT, thereby extending i t s application from the research laboratory  (41) to i n d u s t r i a l quality control programs  and to the general area  (42) of materials selection and evaluation  . Adoption of nonstandard  test techniques should further extend the scope of instrumented  . (43)  xmpact testing  - 8 -  2.2  Instrumented Impact Test Machine  The instrumented impact test machine used i n this study was designed, constructed, and calibrated at the Department of Metallurgy, University of B r i t i s h Columbia, and i s the only such unit i n Western Canada.  Credit f o r the design and instrumentation  go primarily to Messrs. Robert Butters and Ed Klassen, respectively. Their e f f o r t s resulted i n the construction of an extremely r e l i a b l e unit at a savings of thousands of d o l l a r s . note that,  It i s i n t e r e s t i n g to  at this writing, the unit q u a l i f i e s as the only calibrated  Charpy machine i n B r i t i s h Columbia.  2.2.1  Machine Design  The IIT machine was designed and constructed  to comply  (44) with the ASTM E 23  requirements for notched-bar impact testing,  within, of course, the l i m i t s imposed due to the machine being of drop tower design, as opposed to pendulum loading.  A photograph of  the machine i s shown i n Figure 2.1. The frame consists of a massive base plate, firmly secured to the concrete f l o o r . from the base plate.  Two 2.44 m (8 f t ) v e r t i c a l shafts extend up These shafts act as the runners f o r the s t r i k i n g  edge (tup). At the top of the two shafts, a small variable speed e l e c t r i c motor i s located which l i f t s or lowers the tup assembly.  A "drop"  Figure 2.1  Instrumented impact machine.  - 10 -  button opens a solenoid clamp, dropping the tup and i t s associated mass onto a specimen.  The f a l l of the tup assembly i s assumed to  follow the laws of gravity; the v e r t i c a l shafts being w e l l greased to minimize the effects of f r i c t i o n .  Attached to the base plate, are two shock absorbers which absorb the excess energy from the f a l l i n g tup and thereby prevent the large mass from rebounding.  The tup assembly i s that portion of the machine which provides the energy necessary to fracture a specimen.  In order to  achieve a large t o t a l impact energy, and yet a r e l a t i v e l y low impact v e l o c i t y (reasons f o r this s h a l l be discussed i n Section 2.3), the entire assembly has a mass of 45.76 kg (100.88 l b ) .  This  i s a somewhat larger mass than a t y p i c a l commercial pendulum Charpy machine.  Variations i n the mass of the s t r i k i n g tup can be obtained  by bolting to or removing massive s t e e l blocks from the sides of the tup frame.  The tup, being that portion of the machine which actually strikes the sample, i s made of t o o l s t e e l , f u l l y hardened and drawn back to 5 5 . R  c  The tup i s i d e n t i c a l i n dimension and design to that  stipulated by the ASTM for the Charpy impact t e s t , with one exception: i t has recesses i n i t s face to accommodate s t r a i n gauges which are essential to obtain load-time information.  A diagram of the tup  - I l -  l s shown i n Figure 2.2a.  The test specimens s i t on hardened tool s t e e l anvils (R 55), the size and shape of this support area also conforming c  to the ASTM s p e c i f i c a t i o n s .  These anvils can be removed to  accommodate other specimen types or test methods. i n turn, rest on larger a n v i l supports.  The a n v i l s ,  A close-up view of the  tup, specimen, and anvils i s shown i n Figure 2.3.  Specimen guides were c a r e f u l l y aligned, perpendicular to the tup face, so that placing a specimen against these supports assured that the test piece was impacted by the tup at the exact midpoint of i t s s t r i k i n g edge.  A small notch centering device, con-  s i s t i n g of a "pointer" attached to a hinged bar, was i n s t a l l e d to position the specimen so that the notch i n the specimen would l i e d i r e c t l y under the tup and midway between the a n v i l supports. No end stops were used i n positioning the specimen.  Correct  positioning and centering of the test specimen i s , of course, c r u c i a l and great care was exercised i n assuring proper alignment of the specimen guides and the hinged notch centering device.  2.2.2  Instrumentation  The essential difference between the instrumented  impact  - 12 -  8 mm rod.  ,1/4 rod  (a)  1/8" dia.  T  ~(Q B m )  Strain Gauge Recess  3/4-  Gauges not to scale  Figure 2.2  T  "  ,ension  C - compression  scale:  i  1/2" 1  (a) Diagram of tup showing position of strain gauges. (b) Schematic of instrumented tup circuitry and IIT components.  Figure 2.3  Closeup view of tup, a n v i l s , centering device, and test specimen.  - 14 -  machine and a "standard" Charpy unit i s the electronic  instrumen-  tation employed to y i e l d an analog of the dynamic load-time response of a fracturing specimen.  The electronics package consisted of an instrumented tup (load c e l l ) , a power supply, a dynamic transducer amplifier, a signal recording and display system, and a system to trigger the signal just p r i o r to the impact  event.  The tup has been recessed on both faces for protective placement of the highly sensitive semi-conductor Figures 2.2a and 2.2b  s t r a i n gauges.  show the tup design with" the position of  the s t r a i n gauges, and a schematic of the instrumentation, respectively.  Semi-conductor s t r a i n gauges (Micro-sensor Type P01-05-120) were chosen due to t h e i r high gauge factor (+110), high signal/noise r a t i o , and small s i z e (active gauge length of 1.27 mm).  As shown  i n the c i r c u i t r y diagram (Figure 2.2b),all four arms of the bridge are active gauges, which provide temperature and higher s e n s i t i v i t y .  compensating a b i l i t y  This i s an improvement over many other designs,  including commercial units.  The gauges, once placed into the tup recesses, were covered with a protective epoxy.  The gold gauge lead wires were soldered to  heavier copper wires which led through a groove within the tup.  - 15 -  Upon impact, these s t r a i n gauges sense the compressive forces on the tup and provide the signal output voltage  to the  transducer amplifier.  A Tektronix 3A10  Transducer Amplifier Module provided the  DC excitation for the s t r a i n gauges, and amplified and the output s i g n a l . 10 kHz) was  A suitable upper frequency cutoff  conditioned (generally  utilized.  The signal was  displayed on a Tektronix Type 564B o s c i l l o -  scope with a 2B67 Time Base.  Although the oscilloscope trace  could  be stored, for ease of data reduction and for permanent records, sweep was  usually photographed.  A Tektronix C27  Type 57 High Speed f i l m were used (f3).  camera and  The camera shutter  the  Polaroid was  manually opened p r i o r to and independent of dropping the tup assembly.  The oscilloscope display was tup s t r i k i n g the test specimen.  triggered just p r i o r to the  A small magnet attached to the  f a l l i n g tup assembly activated a reed switch. switch triggered the oscilloscope sweep. switch was was  Closure of t h i s  The p o s i t i o n of the reed  c r i t i c a l to ensure that the complete load-time s i g n a l  recorded on the screen of the oscilloscope.  Small gauge marks  were etched onto the drop tower frame for positioning of the trigger switch, the marks corresponding to a range of drop heights. At this writing, an electronic triggering system has been  (N.B. constructed  - 16 -  which can trigger the oscilloscope sweep and the camera system simultaneously  2.2.3  as the "drop" button i s pushed.)  Calibration  The load c e l l was  calibrated to determine the relationship  between the s t r a i n gauge voltage output and the applied load.  Initially,  a s t a t i c c a l i b r a t i o n was  pressed against a standard  made.  The tup  test specimen, the load being  was  selectively  increased by using a hydraulic jack mounted atop the tup assembly. A calibrated compression load c e l l attached was  located between the jack and a r i g i d  to an Instron machine  r e s t r a i n i n g rod.  By  activating the jack, a force of calibrated magnitude could be applied to the tup and the specimen thereby c o r r e l a t i n g the compressive load and the s t r a i n gauge response of the tup.  The voltage output from the tup s t r a i n gauge was  found to  be l i n e a r l y related to applied load from approximately 50 to 3000 lbs(222-13350 N). By adjusting the s t r a i n gauge  transducer  excitation voltage, a convenient output of 10 uV per pound of applied load was  obtained.  exceeded 3000 l b .  (N.B.  Actual loads i n the test program commonly  However, the dynamic c a l i b r a t i o n extended the  range of loads for which accurate c a l i b r a t i o n existed to over  - 17 -  8000 lb (35600 N). been redone.  Additionally,  t h i s s t a t i c c a l i b r a t i o n has  The voltage output was  again l i n e a r l y related  applied load, t h i s time to 8500 lb (37825 N), and was  since to  within 4%  of  (45) the o r i g i n a l c a l i b r a t i o n  The  ).  s t r a i n gauge signals, which are equated to load, are  the results of e l a s t i c strains. strain-rate independent, and,  E l a s t i c properties are r e l a t i v e l y  so, s t a t i c c a l i b r a t i o n should apply  to dynamic loading as w e l l ^ * ^ .  However, as Ireland suggests i n  (47) his excellent  review  desirable since:  1)  2)  s t r a i n gauges may  due  to variations  3)  the amplifier may  , dynamic c a l i b r a t i o n i s nevertheless dynamic conditions are to be monitored; have different response to dynamic loading,  i n the properties of the bonding medium;  and,  have characteristics which vary with the rate  at which the signal passes through the component.  In addition,  the  ASTM E 23 impact test s p e c i f i c a t i o n requires that a dynamic c a l i bration be made p e r i o d i c a l l y . Thus, i n order to determine i f the microvolt-load r e l a t i o n ship established from the s t a t i c c a l i b r a t i o n would indeed apply to dynamic loading conditions, a set of standardized Charpy specimens, with guaranteed values of energy, were obtained from the U.S.  Army  Materials and Mechanics Research Center, Watertown, Massachusetts,  - 18 -  the only supplier of ASTM standard Charpy specimens.  Impact tests were performed at the s p e c i f i e d temperatun of - 40°C.  The t o t a l energy to f a i l u r e was determined as describi  i n Section 2.3.2.1.  The results are shown i n Table 2.1 (a comput  printout of the results of this c a l i b r a t i o n i s given i n Appendix  Table 2.1 DYNAMIC CALIBRATION RESULTS  IIT Total Energy (ft-lb)  Guaranteed Energy(ft-lb)  Tl-0070  14.0  14.3 + 1.0  Tl-0296  13.9  14.3 ± 1.0  U3-0242  49.5  48.0 ± 2.4  U3-0786  46.9  48.0 ± 2.4  V7-0293  71.3  73.9 ± 3.7  V7-0963  72.3  73.9 ± 3.7  Sample  N.B.  1 ft-lb = 1.36 J  The IIT data compared favourably with the s p e c i f i e d energies of the standard samples.  Thus, the results indicated that the  c e l l was calibrated f o r dynamic  conditions.  load  - 19 -  It should be emphasized that i n c a l c u l a t i n g the t o t a l energy of these c a l i b r a t i o n samples, the s t a t i c c a l i b r a t i o n factor (10 yV/ 1 lb) was used.  In addition, the V7-series of samples was impacted at  a higher s t r a i n rate than were other two series, confirming that the IIT machine c a l i b r a t i o n was v a l i d f o r a range of s t r a i n rates from " s t a t i c " to impact loading.  Since i t was shown that the s t a t i c c a l i b r a t i o n was accurate under dynamic conditions, a l l subsequent checks of the c a l i b r a t i o n were done s t a t i c a l l y .  The tup assembly has a known weight (100.88 l b ) ,  and, by allowing the entire assembly to rest on a test specimen, the tup assembly could be "weighed" by reading the s t r a i n gage voltage output on the oscilloscope and employing the relationship between output voltage and load.  This was done p e r i o d i c a l l y to v e r i f y that  the machine was s t i l l calibrated.  With well over 900 impact tests  conducted, the machine never deviated from the o r i g i n a l c a l i b r a t i o n .  The time base on the oscilloscope was calibrated with a Tektronix Type 184 TimeMark Generator and found to be within the manufacturer's s p e c i f i c a t i o n .  2.2.4  Test Variables  2.2.4.1  Drop Height  The height from which the tup assembly drops determines  - 20 -  the t o t a l energy available to fracture a specimen, E , and the v e l o c i t y q  at which the tup strikes the specimen, V  o  For reasons to be discussed, E enough to assure fracture, but V  q  q  must, of course, be large  must be controlled to minimize  effects due to:  and  1)  the i n i t i a l acceleration  of the specimen  2)  the amplitude of various o s c i l l a t i o n s  3)  the limited frequency response of the electronic system.  The drop tower design allows the tup assembly to be raised to any height from approximately 0.15 m to 1.525 m (0.5 - 5.0 f t ) .  The v e l o c i t y of the tup at time of contact with the test specimen was calculated  from:  V  where,  V  q  g h  Q  o  =  (2gh )** o  =  impact v e l o c i t y  =  gravitational acceleration constant  =  drop height  (Eq. 2.1)  The corresponding t o t a l energy available upon impact was obtained from the r e l a t i o n :  - 21 -  E  where,  E  q  m  =  o  ^mV  (Eq. 2.2)  2  o  n  =  available impact energy  =  mass of tup assembly  Thus, the range of available impact v e l o c i t i e s was 1.73 to 5.47 m/s  (5.67 - 17.94 f t / s ) .  m/s  The ASTM requires that, f o r a v a l i d  Charpy test, the tup must impact the specimen at v e l o c i t i e s between 3.05 m/s  to 6.10 m/s  (10 - 20 f t / s ) .  In some instances, however, the impact v e l o c i t y of a test was less than that required by. the ASTM.  This lower impact v e l o c i t y  was sometimes necessary to decrease the amplitude of undersirable specimen o s c i l l a t i o n s and to extend the time f o r f a i l u r e which (47) avoided problems of limited electronic frequency response  .  These  problems s h a l l be discussed i n more d e t a i l i n Section 2.3.1. Although the impact v e l o c i t y was at times as low as m/s  (5.64 f t / s ) , this was not considered to be of major  1.72  consequence  when compared to the serious problems encountered i n data reduction should the v e l o c i t y be too high.  Even s t r a i n rate sensitive  require a factor of 10-100 change i n s t r a i n rate to produce  steels measurable  (13 25) changes i n mechanical properties  '  .  The lower v e l o c i t y of 1.72  m/s s t i l l yields a s t r a i n rate more than 2 x 10^ times that of a  - 22 -  conventional " s t a t i c " t e n s i l e test rate of 0.5 cm/min (0.2 in/min) and therefore can c e r t a i n l y be described as being a "dynamic" test, although i t does deviate s l i g h t l y from the ASTM Charpy test s p e c i f i c a t i o n .  These  nonstandard impact v e l o c i t i e s were necessary only f o r very low toughness materials.  The drop height f o r a given sample at a given temperature was selected so that the t o t a l available energy from the f a l l i n g tup assembly would be s u f f i c i e n t to fracture the specimen, and so that the i n i t i a t i o n energy (energy to i n i t i a t e a stable crack) would be less than a t h i r d of that t o t a l energy.  This l a t t e r r e s t r i c t i o n was important, since  i n order to apply appropriate corrections to the data from the loadtime records, the reduction i n tup v e l o c i t y must be m i n i m i z e d ' . However, care was taken not to use impact v e l o c i t i e s ( i . e . drop heights and energies) much larger than necessary.  2.2.4.2  Temperature  Instrumented impact tests were carried out over a range of temperatures from -196°C to +100°C.  A l l low temperature test samples were brought to temperature by holding them i n a l i q u i d ^ - a l c o h o l bath and were impacted within (44) f i v e seconds as prescribed i n the ASTM Standard E 23  .  Incidentally,  - 23 -  Weiss, et al*''* '' have conducted tests i n which thermocouples were 7  inplanted within the Charpy specimens and have shown that, even for samples cooled to as low as -60°C, 9 seconds out of the bath results i n less than a 2°C r i s e i n temperature.  High temperatures were achieved by placing the specimens in boiling  H 0. 2  2.2.4.3  Instrumentation  Parameters  T r i a l and error dictated the Time/Division setting and Volts/Division setting for the s t r a i n gauge transducer.  Generally,  for high toughness materials (> 70 J = 52 f t - l b ) , the maximum applied load was  on the order of 5000 lb (22250 N).  Since the s t r a i n gauge  voltage output was previously calibrated and found to be equivalent to 10 uV/lb, a setting of 10 mV/division was used and was 1000  equivalent  l b / d i v i s i o n , thus ensuring that the t o t a l impact event could  be recorded on the 10 d i v i s i o n oscilloscope screen. materials required a 5 mV/division  Low  toughness  setting.  The oscilloscope time scale for high toughness materials was  set at 0.5 ms/division since the e n t i r e Impact event took  approximately  0.005 s.  Usually, b r i t t l e samples fractured i n less  than 0.002 s, allowing a 0.2 ms/division scale to be used.  - 24 -  2.3  Interpretation of Load-Time Data  Upon testing a specimen, a photograph of the analog of the dynamic load-time response i s obtained.  Figure 2.4 i s t y p i c a l  of that response for both an e l a s t i c - p l a s t i c f a i l u r e (maximum load > general y i e l d load) and a l i n e a r - e l a s t i c f a i l u r e (maximum load < general y i e l d load).  The load-time information i s s i m i l a r to the  load-deflection curves obtained from slow bend tests of notched specimens on an Instron machine.  2.3.1  V a l i d i t y C r i t e r i a f o r Load-Time Signals  Ireland has reviewed the problems associated with obtaining (29 47) v a l i d instrumented impact data  '  . His works are based on pro-  grams which established test procedures to obtain consistent and v a l i d IIT data f o r the determination of dynamic fracture toughness , ,, . (34-35) from small specimens  parameters  Other than the obvious errors due to improper load c e l l c a l i b r a t i o n , the major sources of error i n an instrumented impact test load-time signal are:  and,  1.  inadequate electronic frequency response  2.  o s c i l l a t i o n s inherent i n the tup signal  3.  i n s u f f i c i e n t impact energy.  - 25 -  Figure 2.4  Instrumented impact load-time photographs: (a) e l a s t i c - p l a s t i c fracture (b) l i n e a r - e l a s t i c fracture.  - 26 -  2.3.1.1  Response Time  A l l electronic instrumentation has a limited frequency response, that i s , the amplitude of a signal passing through the component may be attenuated.  Most oscilloscope manufacturers  define any acceptable frequency response as that at which the signal has been attenuated by 30% (3 dB).  However, f o r instrumented  impact tests, attenuation of at most 10% i s considered acceptable which corresponds to 0.915 dB attenuation, where  dB  =  20 log(volts i n / v o l t s out)  (Eq. 2.3)  The frequency response i s more e a s i l y represented by the frequency response time, T , which i s that time required f o r a signal to r i s e to the desired amplitude (90% of the f u l l amplitude i n the case of (47) IIT).  Ireland  has pointed out that the relationship between  signal frequency and T  f o r a sine wave (which approximates an  instrumented impact load-time curve) i s : °-  35/f  .915dB  ( E  «- ' 2  4 )  This response time i s experimentally determined by superimposing a constant amplitude sine wave on the output of the s t r a i n gauge c i r c u i t .  The frequency of the sine wave i s then increased  u n t i l attenuation, of ten percent i s observed, giving f g-^SdB'  ^  e  - 27 -  response time i s then calculated from the above relationship.  The  0.915dB frequency, and, hence, the response time, i s a function of the upper frequency of the band width as set on the transducer amplifier.  This response time was determined f o r the system employed i n this work and the values are given i n Table 2.2.  Table 2.2 RESPONSE TIMES  3A10 Setting (kHz)  f  0.915dB  (  k  H  z  )  T  R  (y sec)  10  4.8  72.9  30  12  29.2  100  45  7.8  300  130  2.7  1000  500  0.7  The problem of errors due to the attenuation of the output signal can be avoided by adhering to tentative proposals of ASTM Committees test  (15 51-52) (27 29) ' ' and others ' which suggest that f o r a v a l i d  - 28 -  t where,  t  =  5-  1.1 T  (Eq. 2.5)  R  any ellapsed time to be used i n a data reduction calculation.  2.3.1.2  Signal O s c i l l a t i o n s  The second major problem i s the interpretation of the o s c i l l a t i o n s generated upon impacting a specimen. have four primary sources  These o s c i l l a t i o n s  :  1.  the true mechanical response of the specimen  2.  high frequency noise generated by the amplification system  3.  i n e r t i a l loading of the tup as a result of specimen acceleration  and  A.  low frequency o s c i l l a t i o n s caused by r e f l e c t e d stress waves and stored e l a s t i c energy.  The f i r s t i s obviously the desired response.  The electronic noise i s e s s e n t i a l l y eliminated by using the high gain (large signal/noise ratio) semiconductor s t r a i n gauges.  The t h i r d source of o s c i l l a t i o n s has been discussed i n  - 29 -  . ~ , • (13,18-19,25, depth, and from many points of view, by many authors 29-31,47,53-56)  —  .  . I t  results  from the specimen s resistance to  sudden changes i n i t s motion and i s often described as being an " i n e r t i a l loading" o s c i l l a t i o n . fluctuation  I t i s i d e n t i f i e d as the f i r s t  on the load-time trace.  The period of this o s c i l l a t i o n  i s estimated to be on the order of 30 ys for s t e e l and aluminum (29) specimens  . However, i t decays within approximately the f i r s t  two o s c i l l a t i o n s , since, as the specimen accelerates the i n e r t i a l (29 47) load decreases  '  . Thus, the loads recorded during that i n i t i a l  period of time are dominated by this i n e r t i a l loading phenomenon. The amplitude of this o s c i l l a t i o n , which can cause serious problems (53) i n data analysis, i s d i r e c t l y proportional to the impact v e l o c i t y The l a s t source of o s c i l l a t i o n s i s said to be due to a combination of reflected  stress waves and the damping of stored e l a s t i c  (13 31) ' . The period of these o s c i l l a t i o n s can be r e l i a b l y (29 35) predicted through an empirical expression ' : energy  T  where,  =  1.68S/C (W/S)^(EBC )^ Q  s  T  =  period of specimen o s c i l l a t i o n s  S  =  support span  W  =  specimen width  B  =  specimen thickness  E  =  e l a s t i c modulus  (Eq. 2.6)  - 30 -  C  q  =  speed of sound i n specimen  C  G  =  specimen  compliance  For s t e e l and aluminum Charpy specimens, x i s on the order of 33 us (approximately the same period as the i n e r t i a l o s c i l l a t i o n ) .  The  amplitude of the stress waves i s again a d i r e c t function of the impact v e l o c i t y and can cause serious data analysis errors.  In order to avoid problems associated with the amplitude of a l l these various o s c i l l a t i o n s and the period for which the i n e r t i a l o s c i l l a t i o n masks the true s i g n a l , i t has been proposed that any data to be used i n a calculation meet the requirement (27,29).  t  >  3T  (Eq. 2.7)  This requirement i s most e a s i l y met by decreasing the impact v e l o c i t y , V , thereby extending the time for fracture. q  period that the i n e r t i a l load  The  dominates i s approximately 2T, S O the  above r e s t r i c t i o n w i l l assure that the true specimen response i s not masked by contributions due to the i n e r t i a l acceleration. further advantage i n decreasing the impact v e l o c i t y i s that the amplitudes of a l l the specimen o s c i l l a t i o n s are decreased, thus improving signal analysis.  Furthermore, increasing the time to  fracture, by decreasing the impact v e l o c i t y , assures that the  A  - 31 -  electronic signal attenuation i s much less than the acceptable 10% realized by meeting the requirement i n Equation 2.5.  For some very low energy f a i l u r e s the fracture time can be quite short, so the impact v e l o c i t y may have to be lowered to below that required by ASTM E 23 to meet the above s t i p u l a t i o n s . of lowering V  q  The e f f e c t  to below the ASTM s p e c i f i c a t i o n i s not s i g n i f i c a n t , as  was discussed i n Section 2.2.4.1.  However, i n most cases, the time  to fracture obtained when using ASTM specified impact v e l o c i t i e s s a t i s f i e s the above r e s t r i c t i o n s .  2.3.1.3  Impact Energy  The t h i r d source of error i s that associated with the energy supplied to fracture the specimen.  Some calculations  used  to reduce the data obtained from an instrumented impact test r e l y on the assumption that the tup v e l o c i t y i s not reduced by more than approximately 20% so that the corresponding decrease i n i t s v e l o c i t y i s considered to be e s s e n t i a l l y l i n e a r .  To meet t h i s  requirement, a conservative s t i p u l a t i o n i s that the available impact energy, E , be greater than three times that required to q  reach maximum load, and, of course, be s u f f i c i e n t to completely fracture the specimen.  A compromise i s necessary between  i n t e n t i o n a l l y reducing the v e l o c i t y at impact, to l i m i t the amplitude of specimen o s c i l l a t i o n s and extend the fracture time, and  - 32 -  keeping that v e l o c i t y high enough to supply the energy to completely fracture the specimen with a l i n e a r decrease i n v e l o c i t y .  The tentative ASTM specifications  s h a l l require a l l of  these r e s t r i c t i o n s to be met for a load-time signal response to be accepted as being i n d i c i t i v e impact.  of the true specimen behaviour during  These requirements are summarized  i n Table 2.3.  Table 2.3 VALIDITY CRITERIA  Potential  Source of Error  C r i t e r i o n to Prevent Error  I n e r t i a l Loading Effects  t > 3T  Signal Attenuation  t > 1.1  I n s u f f i c i e n t Energy;  E > o E > o  Excessive Tup Deceleration  2.3.2  R  ^Max Load  3  E  T  Total  Data Reduction from Load-Time Curves  2.3.2.1  Energy  The t o t a l energy obtained from a standard Charpy test i s of limited value, even for comparing the r e l a t i v e toughness of  - 33 -  materials.  A high strength, b r i t t l e material may have a high crack i n i -  t i a t i o n energy though a low crack propagation energy.  A low strength,  d u c t i l e material, which may absorb the same t o t a l energy, can have a low i n i t i a t i o n energy and a high propagation energy.  The fracture  c h a r a c t e r i s t i c s must be examined i n terms of the both energy to i n i t i a t e and the energy to propagate a crack i f fracture control i s to be attempted.  2.3.2.1.1  Velocity Reduction Correction  The area under the load-time curve (which i s actually the change i n momentum or impulse) can be converted into the apparent energy f o r fracture: E  a  (Eq.  Pdt  2.8)  area under load-time curve  where,  t  =  ellapsed time from i n i t i a l contact between tup and specimen  V  =  impact v e l o c i t y  However, this i s not the true energy absorbed by the specimen since the  impact v e l o c i t y decreases from V  during the fracture event.  - 34 -  Assuming that the v e l o c i t y decrease i s l i n e a r , a correction for the apparent energy, which accounts f o r the decreasing v e l o c i t y , can be ,(18,57). derived E  where,  =  c  E (1 - E /4E ) a a o  (Eq. 2.9)  E  c  =  corrected energy  E  q  =  available energy at impact  The derivation can be found i n Appendix B.  This correction factor  often ranges as high as 10-12%.  That the v e l o c i t y indeed decreases l i n e a r l y ( i f the available energy, E , i s more than twice that absorbed) and that the magnitude o of the v e l o c i t y change i s usually small (on the order of 5%) have been experimentally demonstrated  '  .  That this correction gives the  same t o t a l energy as that conventionally observed from the dial-gauge (18 20 on a standard Charpy machine has also been amply demonstrated 46 47,49,57 60)^  j ^ ^ ^  '  c a l i b r a t i o n conducted for the present  work v e r i f i e s the correction as well. The corrected energy up to the point of maximum load i s generally considered to be the energy necessary to i n i t i a t e a stable crack which propagates through the sample.  For cleavage  '  -  .,  F  .  . ,  35 -  .  •J i  f a i l u r e s , t h i s assumption i s widely accepted  ,(19,48,59,61-62)  . For  fibrous fractures, at least f o r slow bend tests, the crack may i n i t i a t e at a load less than maximum, and i t may not begin to propagate rapidly through the specimen width u n t i l a time a f t e r maximum load has been reached^ ^ 2 1 , 6 3 ) ^ 2  For  consistency i n data analysis, i t was assumed that the  crack i n i t i a t i o n energy corresponds to the area under the load-time curve up to the point of maximum load.  For cleavage f a i l u r e s this  For 100% shear f a i l u r e s t h i s assumption, should  i s no doubt true.  i t not be s t r i c t l y v a l i d , can cause nonconservative errors i n the " i n i t i a t i o n " energy calculations estimated to be on the order of 20%.  2.3.2.1.2  Compliance Correction for I n i t i a t i o n Energy  Ireland has suggested that when the tup strikes the specimen, the available energy, E , i s reduced due to a variety q  of  factors:  A E  o  where,  "  AE E  E  ACC  =  q  ACC  =  +  E  SD  +  E  B  +  +  \ Z  (Eq  '  reduction i n impact energy e n e r  & y required to accelerate the  specimen  '  2  10)  - 36 -  =  energy r e q u i r e d the  E  B  E ^  to bend and f r a c t u r e  specimen  =  energy due t o B r i n e l l - t y p e d e f o r m a t i o n  =  energy absorbed by impact machine through vibrations  E^,  =  s t o r e d e l a s t i c energy absorbed by the machine  The energy absorbed i n deforming the specimen, E ^ , i s the d e s i r e d v a l u e . small.  E  The v a l u e s  can be d i s r e g a r d e d  o f E ^ and E ^  are u s u a l l y  quite  when the o s c i l l o s c o p e response  ALL  time i s l e s s than the specimen  f a i l u r e time ( t > 1.1  The s t o r e d e l a s t i c energy term, E ^ , machine compliance.  An a p p r e c i a b l e  T ). K  i s r e l a t e d t o the  amount o f the apparent  initiation  energy, E I , as determined from E q u a t i o n 2.9,  can be a r e s u l t o f t h i s  energy, and so, t h i s term must be e l i m i n a t e d  from the  value  of EI.  (N.B.  "corrected"  The t o t a l energy need not be c o r r e c t e d f o r  the machine compliance s i n c e the compliance i s an e l a s t i c  energy  term).  (47) I t has been shown t h a t t h i s energy term can be g i v e n by  :  - 37-  where,  P,,.„ = maximum load on load-time trace MAX = machine compliance  The machine compliance can be calculated f r o m ^ ^ :  C  T  = M C  +  where, C  C  s " GY GY d  " VGY GY  /P  /P  (Eq. 2.12)  T  =  t o t a l compliance of the syst em  Y  =  deflection at general y i e l d ( e l a s t i c l i m i t )  P  G Y  =  load at general y i e l d ( e l a s t i c l i m i t )  t  G Y  =  time at general y i e l d ( e l a s t i c l i m i t )  dg  V  Q  =  impact v e l o c i t y  C  g  =  specimen  compliance  The non-dimensional specimen compliance, C , f o r a notched threeg  point beam with Charpy dimensions has also been e s t a b l i s h e d '  .  Thus, the machine compliance, and hence, the stored e l a s t i c energy of the machine can be conveniently determined.  Therefore, the true energy to i n i t i a t e a crack (energy to maximum load) can be calculated by making corrections f o r both reduction i n tup v e l o c i t y and the effects of machine compliance:  EI  -  where,  [ E I (1 - E I / 4 E ) - h a  EI  a  =  o  C ] M  (Eq. 2.13)  i n i t i a t i o n energy as calculated from  3.  Equation 2.8.  - 38 -  2.3.2.2  Deflection  The corrections required f o r determining the specimen deflection at any time are s i m i l a r to those employed i n the energy (47-48) calculations; the p r i n c i p l e s are discussed f u l l y i n References It can be shown, however, that: d  =  t  where,  tV (1 - E /4E ) - P C o a o t M  (Eq.  d^  =  deflection at any time  t  =  ellapsed time from i n i t i a l impact  P  =  load at time of interest  =  energy as calculated from Equation 2.8  E  2.14)  3.  2.3.2.3  Dynamic Y i e l d Strength  The point on the load-time trace corresponding to the load at which the curve f i r s t deviates from l i n e a r i t y i s the general y i e l d load.  General y i e l d i s considered to occur when p l a s t i c  y i e l d i n g has spread across the entire cross-section of the specimen (for  some s t e e l s , this corresponds to the lower y i e l d l o a d ) ^ ^ ' ^ 2  Green and H u n d y ^ ^ have developed a relationship for determining the  dynamic y i e l d strength from the general y i e l d load, which f o r  three-point bending of notched specimens reduces to:  ^\  - 39 -  = PGY  L  (Eq. 2.15)  B(W - a) 1.21 2  general y i e l d load from IIT photo  where, B  =  specimen thickness  W  =  specimen width  a  =  crack length  L  =  support span  The equation has been validated for standard Charpy "V-notch" specimens by several investigators(14,20,65)^  Employment of this equation, i n  conjunction with the data obtained from an instrumented impact  test, i s  e s s e n t i a l l y the only means available f o r determining the y i e l d strength of a s t r a i n rate sensitive material at very high s t r a i n rates (although the experimentally d i f f i c u l t Hopkinson-split bar technique has been used  (67)  )•  2.3.2.4  Fracture Toughness Calculations  The s i m i l a r i t y between the instrumented impact  load-time  curve and the load-deflection curve used f o r fracture toughness determinations l e d to the application of fracture mechanics theory to IIT.  The topic of dynamic fracture toughness i s considered of  such importance to be discussed separately i n Chapter 4.  - 40 -  2.3.2.5  Computer Programs  To f a c i l i t a t e the many lengthy calculations necessary i n analyzing the instrumented impact test data, two computer programs i n FORTRAN language were written.  One, ENERGY, l i s t e d i n Appendix C, must be supplied values for the area under the load-time curve f o r energy calculations. Measuring  this area was most conveniently and accurately accomplished  with a polar planimeter.  Alternate methods of area measurement were  investigated and included employing  Simpson's Rule to integrate  the curve, using a Quantimet, and, cutting and weighing the curve area.  These tedious techniques were found to give inconsistent  r e s u l t s , with errors greater than 10%, when compared with the accurate and reproducible results obtained by measuring the area under the load-time curves of the Army c a l i b r a t i o n samples with the planimeter.  The other program, IMPACT, uses d i g i t i z e d data of the loadtime signal to f i t a polynomial to the curve, and subsequently i n t e grates that expression to determine the area under the curve.  2.3.2.6  Data Sheet  For each impact t e s t , a number of data points were obtained  - 41 -  from the load-time trace for data reduction.  An "Instrumented  Impact Test Record" data sheet was printed to provide a permanent record of the parameters for each test and to f a c i l i t a t e computer analysis.  2.4  Such a sheet i s reproduced i n Appendix D.  E f f e c t s of Test and Specimen Parameters  2.4.1  Significance of Test V a l i d i t y C r i t e r i a  2.4.1.1  I n e r t i a l Loading Effect  As described i n Section 2.3.1, ASTM tentative proposals suggest that to obtain consistent and u n i v e r s a l l y acceptable IIT data, c e r t a i n v a l i d i t y c r i t e r i a must be met.  A major problem  i n IIT i s that the o s c i l l a t i o n associated with the i n e r t i a l loading of the specimen may overshadow the true specimen response i f the time used i n any data reduction calculation i s less than the time for that i n e r t i a l o s c i l l a t i o n to decay.  To avoid such problems,  a v a l i d i t y c r i t e r i o n has been conservatively set for the times to be used i n c a l c u l a t i o n s :  t  >  (Eq. 2.7)  3T  The i n e r t i a l o s c i l l a t i o n s decay i n approximately  2x.  - 42 -  To determine i f errors i n data calculations existed when this c r i t e r i o n was not met, samples were tested at v e l o c i t i e s which inherently resulted i n general y i e l d and fracture times of less than the duration of the i n e r t i a l loading event, 2x ( i . e . , impact v e l o c i t i e s were used that were higher than that necessary to simply achieve fracture and thereby decreased the time required f o r the fracture).  Also, a l l  data generated during this program which did not meet the t >, 3T c r i t e r ion were examined.  Representative results are shown i n Table 2.4.  For those tests i n which the fracture event occurred p r i o r to 2T, i . e . , f o r tests i n which the i n e r t i a l load was considered to dominate the load-time curve, no s i g n i f i c a n t nor consistent differences i n any calculated property (e.g., absorbed energy, fracture toughness) were evident when compared with " v a l i d " tests, under i d e n t i c a l conditions, i n which the f a i l u r e times exceeded 3x.  A l l " i n v a l i d " results were within  a reasonable scatter band.  For those tests i n which the f a i l u r e times were greater than the period of i n e r t i a l loading(2x), though less than the 3T v a l i d i t y c r i t e r i o n , again no consistent deviation i n properties was evident when compared with the " v a l i d " data obtained from tests where t ^ 3T.  Although others have indicated that v i o l a t i n g this c r i t e r i o n results i n erroneus data ( p a r t i c u l a r l y fracture toughness values)  (29 68) '  Table 2.4 COMPARISON OF VALID AND INVALID DATA AS DETERMINED BY t < 3x  Specimen Code AF-l-STR-PC-07 AF-l-STR-PC-09 AF-l-STR-PC-08# RP-PC-19 RP-PC-20* RP-PC-21# AF-2-STR-PC-5P AF-2-STR-PC-5Q AF-2-STR-PC-5L* AF-2-STR-PC-5N* AF-2-STR-PC-50* AF-2-STR-PC-5M# AF-l-STR-PC-01 AF-l-STR-PC-02 AF-l-STR-PC-03* AF-l-STR-PC-11 AF-l-STR-PC-12 AF-l-STR-PC-10* AF-2-STR-PC-3L AF-2-STR-PC-3M AF-2-STR-PC-3N* AF-2-STR-PC-5J AF-2-STR-PC-5K AF-2-STR-PC-5I* *  Test Temperature (°C)  2x (ms)  - 40 - 40 - 40 - 80 - 80 - 80 -100 -100 -100 -100 -100 -100 + 20 + 20 + 20 - 60 - 60 - 60 - 40 - 40 - 40 - 80 - 80 - 80  .079 .085 .104 .085 .092 .092 .092 .092 .092 .092 .092 .104 .085 .085 .101 .082 .085 .092 .085 .092 .101 .101 .092 .092  Indicates i n v a l i d t e s t : t „ < 3T 2 o 1 ft-lb/in = 0.21 J/cnf 1 ksi-in^ = 1.1 MPa-nr v  5  < < > < < > < < < < < > < < < < < < < < < < < <  Time To General Y i e l d (ms) .133 .156 .101 .163 .103 .088 .143 .141 .104 .106 .127 .101 .137 .155 .121 .126 .141 .122 .172 .167 .143 .155 .163 .118 #  Initiation Energy (ft-lb/in )  K^(ksi-in )  (ms)  Total Energy (ft-lb/in2)  .118 .127 .156 .127 .138 .138 .138 .138 .138 .138 .138 .156 .127 .127 .152 .123 .127 .138 .127 .138 .153 .152 .138 .138  116.7 121.2 95.2 39.7 32.8 38.2 46.0 51.1 50.8 43.9 42.4 57.9 120.7 146.8 105.2 74.8 68.7 79.4 240.0 225.4 237.5 89.3 102.9 92.7  4.7 3.2 4.1 2.5 2.5 0.9 3.3 4.1 3.5 5.0 3.0 4.6 7.6 4.9 4.4 3.6 4.6 7.2 15.8 18.8 21.2 0.4 6.6 2.8  61.6 61.3 64.2 39.5 41.2 36.6 43.0 45.2 47.4 47.1 45.4 50.4 68.3 68.1 68.3 47.1 63.0 55.9 74.1 70.1 76.0 53.4 54.1 48.8  3T  > > < > < < > >  < < < < > > < > > < > >  < > >  <  Indicates i n e r t i a l load dominated:  2  t„  < 2T  - 44 -  the present work does not bear t h i s out.  This i s not to suggest,  however, that the c r i t e r i o n i s not useful; only that i n this work, the i n e r t i a l o s c i l l a t i o n s may have decayed i n a time less than 2T, and/or that the c r i t e r i o n may be quite conservative.  Adherence to  this c r i t e r i o n does not impose unreasonable r e s t r i c t i o n s i n testing specimens; merely decreasing the impact v e l o c i t y s l i g h t l y i s usually a l l that i s required to meet s p e c i f i c a t i o n s .  2.4.1.2  Effects of Impact Velocity  High impact v e l o c i t i e s not only decrease the f a i l u r e times, as just discussed, but, also increase the amplitudes of a l l the specimen o s c i l l a t i o n s .  V  should be controlled for this reason, as o  well.  To demonstrate this and the associated p o t e n t i a l for error, samples were tested at both very high impact v e l o c i t i e s and at v e l o c i t i e s which minimized the amplitudes of the specimen o s c i l l a t i o n s .  Figure 2.5a shows the effect of impacting a specimen at 5.46 m/s  (17.9 f t / s ) , which i s within the standard Charpy test v e l o c i t y  range (10-20 f t / s ) , but r e l a t i v e l y high f o r IIT. specimen was impacted at 3.46 m/s test v e l o c i t y range (Figure 2.5b).  Another i d e n t i c a l  (11.34 f t / s ) , also within the standard  - 45 -  4 J  s'.  • '  — j  k  *f (a)  1  till  +  +  +  t l  I  1 1  ]E  I  . • i 1 IM —It-  :  :  *  (b)  \ i  4  Figure 2.5  Effect of impact velocity, trace: (a) v = 5.46 m/s (b) v = 3.46 m/s D  Q  - 46 -  The time to general y i e l d , t . , was only 0.078 ms f o r the n v  specimen impacted at the higher v e l o c i t y (Figure 2.5a), which i s less than the 3T (0.099 ms) c r i t e r i o n used to assure that the i n i t i a l portion of the load-time trace i s not overshadowed by the i n e r t i a l oscillation.  Also, t h i s test did not meet the requirement that the  signal not be unduly attenuated, since t (Equation 2.5).  was less than 1.1 T  This high v e l o c i t y test must therefore be considered  i n v a l i d on these two counts.  This example shows that 1)  interpretation of the load-  time trace can be made much more d i f f i c u l t due to o s c i l l a t i o n s (compare Figures 2.5a and 2.5b);  2)  thus p o t e n t i a l for error i n  data reduction i s consequently increased; and, 3)  a test can be  rendered i n v a l i d by using an excessively high impact v e l o c i t y .  For these reasons, a l l tests were performed at v e l o c i t i e s which minimized the amplitudes of the o s c i l l a t i o n s and extended the fracture time.  2.4.1.3  Electronic Response Time  The response time of the electronic system (a function of the upper band width frequency) must be such that a s i g n a l i s displayed which has not been excessively attenuated.  - 47 -  Tests were performed to determine i f attenuated data would give erroneous r e s u l t s .  Steel specimens known to give very  reproducible results were tested under i d e n t i c a l conditions, except that the setting of the upper band width frequency was varied.  The  response time, T , was correspondingly increased (refer to Table 2.2). K  Therefore, the fracture event, i n some cases, occurred i n a time much less than the response time of the e l e c t r o n i c system, and the s i g n a l was attenuated by more than 10% ( i . e . t < 1.1 T ) . D  (N.B. The same  results may have been obtained i f the impact v e l o c i t y were unduly increased and the response time kept constant.  However, the attendant  increase i n the amplitudes of the o s c i l l a t i o n s would confuse the comparisons of the effects of varying the response time r e l a t i v e to the fracture time).  Results of this series of tests are given i n Table 2.5. The corresponding impact photographs are shown i n Figures 2.6a - d.  The data show that inadequate system response times result in signals that have been grossly attenuated and thus y i e l d inaccurate results.  The accurate values are those of the t o t a l l y u n f i l t e r e d  test with the 1MHz setting and corresponding 0.0007 ms response time (Figure 2.6d).  The 0.3 kHz and 1 kHz band width settings, which  give response times of 2.3 ms and 0.714 ms, respectively, yielded data with greatly extended fracture times (time to maximum load),  Figure 2.6 Effect of electronic response time, T , on load-time trace: R  (a) (b)  T = 2.3 ms T = 0.714 ms R R  (c) T = 0.0729 ms (d) T = 0.0007 ms R  R  Table 2.5 COMPARISON OF VALID AND INVALID DATA AS DETERMINED BY t < 1.1 T„  Response Time(ms)  2.3  AF-1-SLP-TR1*  Stress-Intensity Factor j_ (ksi - i n c h )  Time To General Y i e l d (ms)  Time To Maximum Load (ms)  General Yield Load (lb)  Maximum Load (lb)  Total Energy (ft-lb)  .334  .661  1761  2287  22.2  104.9  2  AF-1-SLP-TR2 *  .714  .231  .435  2481  3134  26.0  109.0  AF-1-SLP-TR3  .0729  .169  .328  3388  3821  24.5  138.0  AF-1-SLP-TR5  .0007  .160  .345  3507  3955  25.0  141.7  *  Indicates Invalid Test For a V a l i d Test T^/TL, > 1.1  - 50 -  though attenuated loads (Figure 2.6a, b). The fracture  toughness  i  parameters were also seriously attenuated.  I t i s interesting to  note that the t o t a l absorbed energy was not affected by attenuation, however. These results are i n agreement with those of Hoover who studied Borsic-aluminum  composites.  Note that f i l t e r i n g the signal somewhat by using a 10 kHz setting (Figure 2.6c) results i n v a l i d , accurate data with the advantage that the amplitudes of the superfluous specimen are  oscillations  greatly suppressed.  2.4.2  Specimen Parameters  2.4.2.1  Notch Radius  It was found to be extremely d i f f i c u l t to cut large numbers of Charpy notches with the accurate 0.25 mm± 0.025mm standard notch (44) radii  .  Specimens received from outside sources and samples  produced within the Department commonly deviated from this standard radius. The effect of any notch i s 1)  to r a i s e the e f f e c t i v e  s t r a i n rate below the notch root, which implies f o r bcc materials,  - 51 -  that the y i e l d stress increases; 2)  to concentrate p l a s t i c s t r a i n  and raise the y i e l d strength a d d i t i o n a l l y by s t r a i n hardening; and 3)  to introduce a t r i a x i a l stress state at the notch root.  The  result i s to raise t e n s i l e stress levels below the notch and to raise the d u c t i l e - b r i t t l e t r a n s i t i o n temperature^^,69)^  Decreasing  the notch root radius, as i n precracking accentuates those effects by increasing the p l a s t i c stress concentration factor; decreasing the stress l e v e l required to achieve the maximum degree of stress i n t e n s i f i c a t i o n ; and, by decreasing the p l a s t i c zone size required to maximize the degree of t r i a x i a l i t y ^ .  A series of preliminary tests were performed on Charpy specimens with nonstandard notches to determine the effect of t h i s test variable on IIT results.  Steel specimens with notch r a d i i from 0.16 mm to 0.33 mm were tested.  Charpy samples having fatigue cracks at the notch  root were also tested.  A l l other specimen dimensions conformed  to the ASTM E 23 requirements.  The results of the study are presented i n Table 2.6.  The  data show that the propagation energy was not affected by variations i n the notch radius.  However, the results of the tests conducted  at +20°C indicate that the i n i t i a t i o n energy was affected by decreases  - 52 -  in the radius.  The fatigue precracked notches, having a very sharp  radius, yielded extremely low i n i t i a t i o n energy values.  For notches  with r a d i i i n the range from 0.25 mm to 0.16 mm, there i s l i t t l e difference i n the i n i t i a t i o n energy.  The r e l a t i v e l y blunt 0.33 mm  notches gave the highest i n i t i a t i o n energies.  The results obtained at temperatures  of -20°C and below  (transition region) show that specimens with the 0.18 mm radius notch have a higher i n i t i a t i o n energy than those with the standard 0.25 mm radius notch.  Ciampi and coworkers  also found that  specimens with a 0.12 mm radius notch often had higher i n i t i a t i o n energies than those with the 0.25 mm notch i n this temperature This i s unexpected  on a t h e o r e t i c a l b a s i s ^ \  range.  Apparently, the  variations i n the notch toughness due to changes i n the notch radius (at least f o r this s t e e l i n the limited range of from 0.25 mm to 0.18 mm) are not so great as to overshadow either the inherent scatter found i n toughness data or the bimodal behaviour of Charpy , , • (69,71) data i n the t r a n s i t i o n region  Results of the c r i t i c a l crack opening displacement, COD, data are also indicated i n Table 2.6.  The c r i t i c a l COD can be  defined as the amount of i n e l a s t i c stretching of the material immediately ahead of the crack t i p at the moment of crack i n i t i a t i o n (72) Discussion on the calculation of this parameter from IIT  Table 2.6 NOTCH RADII STUDY  0.33 T( C)  0.25  mm  0 . 18 mm  mm  0.16  0 mm  mm  EP/A  EI/A  COD  EP/A  EI/A  COD  EP/A  EI/A  COD  EP/A  EI/A  COD  EP/A  123  52  6.6  124  40  5.6  125  38  5.4  123  35  4.8  - 20  124  29  4.1  123  43  - 40  107  21  3.1  107  33  - 50  93  17  2.8  - 60  74  14  2.0  + 20  44  - 70  19  3.1  EI/A  COD  125  7  1.3  5.3  112  7  1.4  4.2  107  4  1.4  85  5  1.2  67  4  1.1  60  4  1.0  72  33  4.3  48  15  2.2  44  15  2.4  A l l values are averages of several tests EP/A  =  2 Propagation Energy/Unit Area ( f t - l b / i n )  EI/A  =  Initiation  o  Energy/Unit Area ( f t - l b / i n )  3  COD = C r i t i c a l Crack Opening Displacement ( i n x 10 ) A l l tests performed on Steel AF-1 with crack running p a r a l l e l  to r o l l i n g d i r e c t i o n .  - 54 -  data i s deferred u n t i l Chapter 4.  The trends are very s i m i l a r to those noted f o r the i n i t i a t i o n energy.  In comparing COD data f o r the fatigue precracked versus the  notched samples tested at +20°C, i t i s evident that the crack opening deflection decreases with decreasing notch radius.  The r e l a t i v e l y  blunt 0.33 mm notch resulted i n the highest displacements p r i o r to fracture i n i t i a t i o n .  The specimens with the 0.25 mm, 0.18 mm, and  0.16 mm notch r a d i i did not show s i g n i f i c a n t differences i n COD.  Again, i n the t r a n s i t i o n temperature range, the 0.18 mm notch radius specimens experienced more deflection p r i o r to crack i n i t i a t i o n than did the standard  specimens.  The ASTM E 23 s p e c i f i c a t i o n f o r the Charpy notch radius i s 0.25 mm ± 0.025 mm.  The results of this preliminary study have shown  that though i t i s d i f f i c u l t to machine specimens to such a tolerance, specimens with notch r a d i i which deviate only s l i g h t l y from the standard s t i l l y i e l d data within the expected scatter band of the material.  2.4.2.2  Notch Angle  Charpy specimens having a 60° notch angle as opposed to  - 55 -  the specified 45° ± 1° standard were tested, a l l other specimen dimensions  adhering to the ASTM standard.  The results of the study  are shown i n Table 2.7.  The observed increases i n both the i n i t i a t i o n and the propagation energies of the 60° notched materials (versus the 45° notch) conform to the t h e o r e t i c a l expectations.  On a t h e o r e t i c a l basis, increasing the notch angle (as with increasing the root radius) should have the effect of decreasing the maximum value of the p l a s t i c stress i n t e n s i f i c a t i o n factor.  The  p l a s t i c stress i n t e n s i f i c a t i o n factor i s defined as the r a t i o of the maximum t e n s i l e stress existing below a notch to the t e n s i l e y i e l d stress of an unnotched s p e c i m e n ^ .  Lowering the maximum value of  this factor decreases the magnitude of the t e n s i l e stresses i n the p l a s t i c zone ahead of a notch f o r a given applied stress and thereby increases the measured d u c t i l i t y and toughness manifested by increases i n both the propagation and i n i t i a t i o n energies.  Thus, the greater  the notch angle, the less the constraint at the notch root, and the greater the notch toughness of the specimen.  The maximum possible  value of the stress concentration factor below the notch i s 2.82 f o r a 0° notch and 1.0 f o r an unnotched b a r ^ \ 60°  The values f o r 45° and  notches are 2.18 and 2.05, respectively, which are not s i g n i f -  icantly different.  Nevertheless, measurable increases i n toughness  - 56 -  Table 2.7 NOTCH ANGLE STUDY  AF-l-Crack P a r a l l e l RD 60° Notch T(°C) Initiation Energy*  Propagation Energy*  Initiation Energy*  Propagation Energy*  6.0 5.8 5.6 5.3 4.2 3.5 1.3  18.0 17.9 19.2 17.2 17.2 12.9 4.7  5.6 4.6 5.7 4.4 3.2 3.2 1.1  15.9 15.6 18.5 15.3 13.8 11.5 4.9  + 100 + 20 0 - 20 - 40 - 60 - 80  T(°C)  AF-l-Crack Transverse RD 60° Notch  AF-l-Crack Transverse RD 45° Notch  Initiation Energy*  Propagation Energy*  Initiation Energy*  34.3 37.3 30.9 26.1 20.1 18.3 1.2  69.3 71.7 85.3 92.9 74.2 59.5 8.4  + 100 + 20 0 - 20 - 40 - 60 - 80  *  in ft-lb  A l l values are averages of several tests 1 ft-lb  AF-l-Crack P a r a l l e l RD 45° Notch  =  1.36 J  30.8 30.0 22.3 21.4 18.9 14.5 1.2  Propagation Energy* 64.9 72.3 82.4 84.4 69.0 55.8 9.1  - 57 -  were observed f o r the 60° notched  specimens.  The results of this study are important i n that 1)  they  show the usefullness of instrumenting an impact test for revealing differences i n the dynamic response of d i f f e r e n t types of specimens, and 2)  they underline the importance of adhering to the notch angle  requirement i n the ASTM E 23 s p e c i f i c a t i o n .  2.4.2.3  Specimen Thickness  Tests were performed to determine the effect of specimen thickness on the IIT r e s u l t s .  Since pipeline steels were employed  i n this study, these results are included i n Chapter 3.  2.5  Crack I n i t i a t i o n  An important assumption i n the analysis of the load-time data from an instrumented impact test i s that the area under the curve to the point of maximum load i s a direct measure of the energy required to i n i t i a t e the crack. parameter,  This i n i t i a t i o n energy  EI, i s not only used to describe the crack i n i t i a t i o n  event, but also i s used to calculate fracture toughness  parameters,  such as the J - i n t e g r a l and i t s associated s t r e s s - i n t e n s i t y factor, K. T  In addition, the assumed relationship between the peak load  - 58 -  and crack i n i t i a t i o n i s used i n establishing c r i t i c a l deflections for crack opening displacement  calculations and c r i t i c a l loads for other  fracture toughness values.  It has been reported  '^  that for work hardenable  notched three-point bend specimens, tested under slow s t r a i n rate conditions, the idealized crack initiation/propagation behaviour i s as follows:. Several microscopic cracks i n i t i a l l y appear, e s s e n t i a l l y simultaneously, at mid-width, on the tension side of the specimen being loaded.  Edge e f f e c t s and unconstrained p l a s t i c i t y i n the center of  the specimen account for t h i s .  These small s u b c r i t i c a l cracks  eventually j o i n together into a much wider and deeper crack, r e s u l t i n g in a "thumbnail" appearance on the fracture surface.  This i s known  to occur at a point on the load-deflection trace beyond general y i e l d but prior to maximum load, the exact location on that curve being a function of specimen size, composition, and s t r a i n rate.  The depth  of the i n i t i a t i n g crack at this stage remains e s s e n t i a l l y constant up to the point of maximum load; whereas, i t s width extends l a t e r a l l y , as the regions near the edges of the specimen begin to form microcracks which combine with the central crack.  The crack eventually  reaches the sides of the test specimen at the point of maximum load. Beyond maximum load  the crack propagates  specimen with an attendant loss i n load.  through the width of the The mode of propagation  at this stage, whether i t be cleavage, fibrous, or a combination of  - 59 -  the two, determines  the magnitude of the propagation energy.  Iyer and M i c l o t ^ " ^ reported that for non-work hardening materials, however, no s u b c r i t i c a l crack growth occurs i n the post y i e l d region prior to reaching maximum load.  Crack extension was  always accompanied by a drop i n load.  For those specimens which fracture p r i o r to general y i e l d (i.e.  l i n e a r - e l a s t i c f a i l u r e s ) or which fracture e n t i r e l y by cleavage,  crack i n i t i a t i o n i s believed to occur at the maximum load as the crack front  , .1 . , , • • t.. , (20-21,62) pops i n straight across the entire specimen thickness  These descriptions of the cracking process have been drawn from studies done under slow bend conditions.  I t i s reasonable to  assume that the general crack formation process i s the same under (19-20) impact loading conditions  ; the only difference being the precise  position between general y i e l d and maximum load at which the crack i n i t i a t e s - the crack should s t i l l reach f u l l specimen thickness at maximum load. Various instrumented impact tests were performed to determine 1)  i f , i n f a c t , during fibrous or mixed mode fracture the crack  extends to f u l l specimen thickness at the maximum load; and, 2) i f the precise point on the load-time curve at which a crack i n i t i a t e s  - 60 -  could be established under impact loading conditions.  These t e s t s , to study the relationship between peak load and fracture, included:  1)  High speed movie films to record the  surface evidence of the fracture event; tance technique;  2.5.1  and, 3)  2)  an e l e c t r i c a l r e s i s -  "reduced energy" tests.  High Speed Movie Films  The high speed f i l m technique employed a Hycam movie camera, capable of up to 10,000 frames per second, to examine the surface of the test specimen during impact.  Comparison of the  specimen surface on each frame of the high speed f i l m with the corresponding IIT load-time trace was used to determine the actual load at which a surface crack appears.  Due to the r e s t r a i n t s imposed by the high i n t e n s i t y l i g h t i n g required for high speed filming and the coordination required i n triggering the impact machine and the high speed camera, tests were possible only at ambient temperatures (30° - 35°C). Thus, only fibrous f a i l u r e s could be studied.  Charpy specimens of an acicular f e r r i t e s t e e l were notched so that the crack would propagate transverse to the r o l l i n g d i r e c t i o n .  - 61 -  This p a r t i c u l a r s t e e l (designated AF-1) exhibits r e l a t i v e l y high impact energies when cracking occurs i n t h i s d i r e c t i o n .  Table 2.8 summarizes the r e s u l t s of this study.  Table 2.8 HIGH SPEED MOVIE FILM RESULTS  Specimen Code  Film Speed At Impact  Time to Maximum Load From IIT Photo  Time of F i r s t Observation of Surface Crack on High Speed Movie Film  AF-1-49  4625 f t / s  0.633 ms  0.649 < t << 0.845 ms  AF-1-47  5250 f t / s  0.727 ms  0.762 < t «  0.952 ms  These r e s u l t s indicate that under impact loading conditions the crack does indeed appear on the surface at a time that i s approximately equivalent to that required to a t t a i n the maximum load, i n agreement with slow bend test r e s u l t s .  The time uncertainty shown  i n the tabulated data i s associated with the time ellapsed between i n d i v i d u a l frames of the movie f i l m .  These times determined from  the high speed films do f a l l on the maximum load "plateau" of the load-time trace.  The difference i n i n i t i a t i o n energy between that  - 62 -  obtained using the maximum load from the load-time trace versus that corresponding to the median time obtained from the high speed f i l m tests i s equivalent to approximately 6-7 J ; the magnitude of this difference i s similar to other errors inherent i n analyzing the load-time data.  2.5.2  E l e c t r i c a l Resistance Study  (76) Mclntyre and P r i e s t  have described the application of  the e l e c t r i c a l resistance technique to study crack growth.  A con-  stant current i s passed through a notched specimen, a c e r t a i n potential difference existing between the current leads placed at the ends of the specimen.  As a crack i n i t i a t e s and extends  from  the notch during loading, this p o t e n t i a l difference w i l l  suddenly  increase due to the increase i n the path of resistance.  This change  i n the p o t e n t i a l drop across the sample can be monitored using an oscilloscope.  By comparing the time at which the p o t e n t i a l difference  i n i t i a l l y increases with the time f o r attaining the peak load on the load-time trace, the relationship between peak load and crack i n i t i a t i o n can be assessed.  This method offers the advantage that  the crack formation process can be monitored impact  throughout the entire  event. Stranded copper wire:current leads were spot welded to the  - 63 -  ends of standard Charpy specimens (the material was i d e n t i c a l to that employed i n the high speed movie study).  A constant  of 20 amps was supplied across the test specimen.  current  Great care was  taken to insulate the entire system, p a r t i c u l a r l y by separating the specimen from the anvils with a t h i n sheet of i n s u l a t i n g material.  Typical r e s u l t s are shown i n Figure 2.7. The p o t e n t i a l difference across the specimen generally increased r a p i d l y , and unexplainably, at the instant of impact, then dropped to a value below that of the i n i t i a l p o t e n t i a l i n a time span within the e l a s t i c region on the corresponding load-time trace.  The p o t e n t i a l difference  then, usually, rose r a p i d l y i n the time range between that with general y i e l d i n g and the maximum load.  associated  The rapid r i s e i n  potential apparent after y i e l d i n g was thought to be associated with the crack i n i t i a t i o n and consequent increase i n the path of e l e c t r i c a l resistance.  Late stages of the fracture event generally  correlated  well with a rapid increase i n the p o t e n t i a l across the specimen.  However, i n some tests the p o t e n t i a l difference decreased rather than increased  as the specimen fractured.  In other t e s t s ,  the i n i t i a l increase i n p o t e n t i a l difference corresponded to a point beyond the maximum load.  Such inconsistencies cast doubt as to the  r e l i a b i l i t y of the technique.  - 64 -  Figure 2.7  Electrical resistance study of crack growth. (a) load-time curve. Scale: 500 lb/div x 0.2ms/div (b) potential-time curve. Scale: 1 mV/div x 0.2ms/div.  - 65 -  The following are possible reasons f o r these test problems:  1.  Although Mclntyre and P r i e s t were successful i n  monitoring crack growth i n Charpy specimens, they employed a well insulated system and slow bend tests.  The present work, under  impact loading conditions, necessarily involved the f a l l i n g tup assembly, a massive block of s t e e l . f i e l d may  Thus, this moving magnetic  have generated e l e c t r i c f i e l d s which d r a s t i c a l l y influenced  the test r e s u l t s . 2.  The i n i t i a l crack front of these specimens i s curved.  The resistance method produces an output which i s proportional to the average crack length between the mid-section the sample.  and the edges of  Thus, the resistance change i n the i n i t i a l stages of  crack formation are small and may  have been overshadowed by  the  factors described i n 1. 3.  During the i n i t i a l stages of cracking, p r i o r to  extensive bending of the specimen, rough surfaces on the opposing fracture faces may thereby reducing  have interconnected  and caused short c i r c u i t i n g ,  the magnitude of the p o t e n t i a l drop or giving  erroneous r e s u l t s altogether.  If the system could be better insulated, this technique could be useful i n monitoring crack growth under impact loading  - 66 -  conditions to establish the exact point on the load-time curve at which fracture  2.5.3  initiates.  Reduced Energy Tests  A series of s t e e l specimens were subjected to a range of impact energies varying from a magnitude i n excess of that required to i n i t i a t e the crack to energies less than that required f o r initiation.  This " i n i t i a t i o n " energy (energy to maximum load) was  determined from previous tests of the same material and found to be 50.6 ± 4.1 J (37.3 ± 3.0 f t - l b ) .  S p e c i f i c a l l y , specimens i n t h i s study were impacted at one of the following l e v e l s of available energy, E : q  1)  energy of less than the lowest value of the i n i t i a t i o n including scatter; i . e . , less than 46.5 J (34.3 f t - l b ) ; impact energy within  An impact  energy, 2) an  the i n i t i a t i o n range, i . e . 46.5 to 54.7 J  (34.4 - 40.3 f t - l b ) ; or, 3)  an.E value greater than the highest o  known value of the i n i t i a t i o n energy (> 54.7 J ) .  After the reduced energy was imparted to each  specimen,  the specimens were heat tinted to oxidize any r e s u l t i n g crack surfaces and subsequently fractured to reveal the extent of crack propagation.  - 67 -  The results of these tests, presented i n Table 2.9, supported, at least qualitatively, the description of crack formation previously discussed: crack Initiation apparently occurs near mid-center of the specimen, prior to maximum load; the crack extends laterally to f u l l specimen thickness at maximum load; and, thereafter, the f u l l width crack propagates through the specimen.  Specimens 1 and 2, impacted with available energies of less than 46.5 J, both showed very slight evidence of crack initiation.  These cracks were extremely short (< 1 mm) and did not  extend across the samples.  Specimen 4, impacted at an energy of 51.2 J (within the "initiation" range) exhibited a crack which had extended to the sides of the specimen and showed some f u l l width propagation. Whereas, specimen 3, impacted at 45.6 J, just less than the "initiation" energy, displayed a crack which had not quite extended across the test sample.  The behaviour shown in specimen 4 i s comparable to that found using slow bend t e s t i n g ^  3  and agrees with the results  of the high speed movie films which showed that the crack f i r s t extends to the sides of the specimen at maximum load.  - 68 Table 2.9 REDUCED ENERGY TEST RESULTS  Specimen  Impact Energy (J)  Photograph  Energy to maximum load ( " i n i t i a t i o n " energy) Predetermined to be 50.6 ± 4.1 J (37.3 ± 3.0 f t - l b ) . A l l specimens AF-1 s t e e l notched transverse to r o l l i n g Tests a l l at +20°C.  direction  - 69 -  Specimens 5 and 6, impacted with energies exceeding the i n i t i a t i o n energy, exhibited cracks which had extended across the sample and then propagated approximately halfway through the remaining ligament.  In summary then, these tests show that for shear type f a i l u r e s resulting from impact loading: p r i o r to maximum load, and 2)  1)  crack i n i t i a t i o n occurs  that the crack extends to the sides  of the sample at the point of maximum load.  Calculations of  i n i t i a t i o n energy, among others, depend upon the assumption crack i n i t i a t i o n starts at maximum load.  that  Since no r e l i a b l e  technique exists to establish the precise point of crack i n i t i a t i o n under high s t r a i n rate conditions, a l l calculations i n this study were made assuming that the maximum load i s equivalent to the point of crack i n i t i a t i o n .  I t i s recognized that some values determined  employing this assumption may be non-conservative (unless fracture was e n t i r e l y cleavage or occurred before general y i e l d i n g ) .  This  remains one of the major areas i n the f i e l d of instrumented  impact  * . ^ • • * « further .!. i (14,34,72,75) testing requiring work  - 70 -  3. INSTRUMENTED IMPACT STUDY OF ACICULAR FERRITIC PIPELINE STEELS  3.1  Acicular F e r r i t l c  The r e l a t i v e l y new  Steels  high strength,  low a l l o y (HSLA) a c i c u l a r  f e r r i t i c steels have become an important class of s t r u c t u r a l materials due to t h e i r low cost per unit of strength, high toughness, and good formability and weldability. 70 to 80 k s i (480-550 MPa), 115  Material having a y i e l d strength  of  a Charpy upper shelf energy of well over  f t - l b (155 J ) , and a Drop Weight Tear Test 50% shear fracture  appearance temperature of about -45°C i s now  The this new  innovative  available.  production techniques employed to produce  generation of steels have been reviewed i n several  publications  (77-86)  Acicular f e r r i t e (AF) i s defined as a highly substructured, non-equiaxed f e r r i t e that forms on continuous cooling by a transformation involving both d i f f u s i o n and shear.  The  transformation  temperature i s s l i g h t l y higher than that of upper b a i n i t e . also distinguished  AF i s  from b a i n i t e i n that only a very small amount of  carbide i s present due to the limited amount of carbon available i n such steels(78)^  - 71 -  Strengthening i s achieved through several independent mechanisms.  The AF i s inherently f i n e grained (ASTM 12-14)  and has a high d i s l o c a t i o n density. 0.07 w/o)  The niobium addition (0.04 -  provides a d d i t i o n a l strengthening by p r e c i p i t a t i n g as a  niobium carbonitride.  The very low carbon additions (less than 0.07 w/o)  and the  exceptionally f i n e grain size of AF contribute to i t s high toughness, a property not generally associated with high strength materials. Higher carbon l e v e l s and the consequent formation of carbides r e s u l t i n higher t r a n s i t i o n temperatures and lower shelf energies.  The  low carbon l e v e l has the added advantage that both w e l d a b i l i t y and formability are markedly improved.  A minimum carbon l e v e l of 0.01 (78)  0.02 w/o  i s desirable to f a c i l i t a t e p r e c i p i t a t i o n strengthening  The addition of molybdenum (0.25 - 0.50 w/o), manganese (1.50 - 2.25 w/o),  and to a lesser extent, niobium, suppresses  the  a u s t e n i t e - f e r r i t e transformation temperature to below 700°C which increases the nucleation time required to form polygonal f e r r i t e ; the alternative acicular f e r r i t e microstructure i s thereby  allowed  (79) to form upon cooling  That the AF transformation occurs at a r e l a t i v e l y temperature i s s i g n i f i c a n t , i n that:  1)  the decreased  low  solubility  -  - 72 -  of Nb i n f e r r i t e promotes the formation of Nb(C,N) at a slower rate and  therefore produces a f i n e r , more homogeneous p r e c i p i t a t e ; the (1Q\  s t e e l ages only s l i g h t l y  ; and,  2)  the low transformation  temperature also contributes to the formation of the very f i n e f e r r i t e ( 81^ grain sizes.  Both effects improve strength and toughness  S t r i c t process control during the hot r o l l i n g of these HSLA steels i s essential to achieve the very f i n e f e r r i t i c grain size and the desired balance of toughness and strength necessary f o r c r i t i c a l applications  such as pipelines.  A reduced slab reheating temperature (approximately 1150°C) ensures that the austenite i s f i n e grained prior to hot r o l l i n g . During the i n i t i a l r o l l i n g stages (- 1150° - 980°C) the s t e e l i s heavily deformed and i t r e c r y s t a l l i z e s repeatedly, further r e f i n i n g the y - g r a i n /on  size  . There i s some Nb(C,N) p r e c i p i t a t i o n i n the austenite at  these temperatures which tends to retard austenite r e c r y s t a l l i z a t i o n and  grain growth^^'^"*"'^^ ^6) . h i s i s another b e n e f i c i a l e f f e c t of t  the niobium addition  (Al, V, and T i , f o r instance, retard ( 82") but not r e c r y s t a l l i z a t i o n )  grain growth  Below 980°C, where r e c r y s t a l l i z a t i o n of the austenite ceases, f i n i s h r o l l i n g takes place (980°-800°C).  At the lower end of t h i s  temperature range, heavy deformation of the f i n e grained y imparts a  - 73 -  heavily dislocated structure and elongates the grains, thereby providing more s i t e s f o r the subsequent nucleation and growth of a fine grained f e r r i t e .  The heavy deformation of the y-phase also suppresses  the y-a transformation t e m p e r a t u r e .  The minimum r o l l i n g  temperature  i s controlled to ensure that no deformation of the f e r r i t e phase occurs, as this would be detrimental to the toughness of the f i n a l product.  In general, decreasing the slab reheat and f i n i s h r o l l i n g temperature  (to l i m i t y r e c r y s t a l l i z a t i o n and grain growth) and increas-  ing the degree of deformation i n the l a t e r o l l i n g stages (to enhance substructure strengthening and to provide more f e r r i t e nucleation s i t e s ) results i n more refined f e r r i t e grains thereby improving the strength j . * (81-82,86) and toughness of the s t e e l  Very low sulfur levels and/or additions of rare earths, f o r s u l f i d e i n c l u s i o n shape control, are desirable to assure adequate toughness and to reduce the anisotropy of the toughness and d u c t i l i t y „. (78,85) properties  Acicular f e r r i t e steels are often k i l l e d , sometimes with (78) aluminum  since s i l i c o n can impair impact resistance  3.2  Pipeline Applications  Vast resources of recoverable o i l and gas exist; nearly  - 74 -  20% of the gas fields are believed to be situated in the distant offshore Arctic regions of Alaska and Canada and perhaps another 50% in Siberia a l o n e ^ \ 87  The social and economic pressures to retrieve these resources are enormous. The Alyeska trans-Alaska o i l pipeline and the proposed Alcan/Foothills gas line are notable engineering projects which extend the limits of pipeline technology.  Such pipelines are being built at incredible costs  through  an extremely hostile, yet fragile, environment - the Alyeska pipeline was recently completed at a cost of over $9 b i l l i o n .  Thus, to ensure  the integrity of these lines, imperative for economic and ecological reasons, stringent demands must be made on the materials of construction, fabrication techniques, and test procedures  From a design standpoint, economics dictate larger diameter lines operating at higher pressures to maximize throughput and thereby lower the operating costs over the l i f e of the l i n e ^ ^ .  These factors  necessitate the employment of higher strength materials and/or greater wall thicknesses since the maximum hoop stresses, a , in a pipeline cai H  be determined  from^^: a  H  =  Pd/2t  (Eq. 3.1)  - 75 -  where,  P  =  operating pressure  d  =  pipe diameter  t  =  pipe wall thickness  There are l i m i t s , however, to the pipe wall thickness due to: 1)  r e s t r i c t i o n s imposed by m i l l f a c i l i t i e s , 2)  requirement 3)  the toughness  of a pipeline (which i s also a function of thickness),  d i f f i c u l t i e s i n retaining high strength and toughness i n very  thick plate, and 4) additional d i f f i c u l t i e s i n welding and f i e l d inspection.  The new generation of A r c t i c pipelines have been or are proposed to be constructed using HSLA acicular f e r r i t e s t e e l s .  Their  higher strengths per unit of cost and weight allow reduction i n pipe wall thickness and t o t a l pipeline weight and consequent savings i n i n i t i a l cost, transportation, and f i e l d handling, while permitting the use of higher operating pressures.  Important, too, i s the fact  that these steels have very high toughness thereby s i g n i f i c a n t l y reducing the potential f o r f a i l u r e s i n the pipelines.  Additional advantages of the AF s t e e l to the p i p e l i n e industry i s the fact that they do not exhibit discontinuous y i e l d i n g and they have higher work hardening rates than conventional f e r r i t e p e a r l i t e pipeline steels.  Thus, they o f f s e t the y i e l d strength losses,  observed when testing flattened t e n s i l e specimens or when forming pipe,  - 76 -  known to r e s u l t from the Bauschinger effect(80,84,86,91)^  This i s  p a r t i c u l a r l y important f o r s p i r a l welded pipe which i s not cold expanded after forming.  (N.B. A Canadian producer plans to cold  expand i t s s p i r a l welded pipe i n the near future - a unique innovation to take advantage of t h i s feature of AF s t e e l s .  They s h a l l r e a l i z e  a net increase i n the strength of the pipe, r e l a t i v e to that of the (92) controlled r o l l e d plate)  The Canadian metallurgical community has pioneered and continues to be a leader i n the production and use of AF steels f o r pipeline applications.  The f i r s t commercial application of such  s t e e l was a 130 km section of 107 cm (42 in) diameter, 9.4 mm (0.370 in) wall thickness s p i r a l welded gas p i p e l i n e produced by The Interp r o v i n c i a l Steel and Pipe Corporation, Ltd., Regina, Saskatchewan (IPSCO) i n the l a t e 1960's.  The Steel Company of Canada  (Stelco)  has recently begun to market HSLA acicular f e r r i t e steels suitable for A r c t i c pipeline applications.  Pipeline manufactured by both  companies w i l l l i k e l y be used i n the Alcan/Foothills gas pipeline project.  3.3  Fracture  Control i n Pipelines  Gas pipeline failures, pose a p a r t i c u l a r l y serious problem because the v e l o c i t y of a propagating crack may be greater than the decompression rate of the gas.  Thus, the crack front would remain i n  - 77 -  a high pressure region, a condition which could lead to long catas-  (88) trophic f a i l u r e s : one of up to 12 km has been documented  .  The  crack i n i t i a t i o n and propagation resistance of the p i p e l i n e s t e e l i s therefore of importance i n designing gas l i n e s . In o i l pipelines, decompression i s rapid, and cracks do not propagate great distances.  However, the environmental damage  resulting from a cracked o i l l i n e and the high costs involved i n repairing such a leak i n remote locations requires that great  im-  portance be placed on preventing crack i n i t i a t i o n .  Studies by the B a t t e l l e Memorial I n s t i t u t e , sponsored by the American Gas Association, have been ongoing for the past twenty years to delineate the causes and c r i t e r i a for the prevention and arrest of b r i t t l e and d u c t i l e pipeline f a i l u r e s .  Conclusions from  this program have been incorporated into most pipeline strength and toughness s p e c i f i c a t i o n s throughout the world, including those of the Canadian National Energy Board, the Canadian Standards Association (Z 245.1), the American Petroleum Institute s p e c i f i c a t i o n s (API  5LX  (92—93)  and 5LS), and v i r t u a l l y every p i p e l i n e company.  The basic fracture control philosophy inherent to a l l of these standards i s : 1)  to prevent b r i t t l e f a i l u r e s by assuring that  the pipeline operates above the material's d u c t i l e - t o - b r i t t l e t r a n s i tion temperature;  2)  to prevent d u c t i l e crack i n i t i a t i o n by specifying  - 78 -  a minimum toughness f o r a pipe operating at a s p e c i f i c stress l e v e l ; and  3) to control d u c t i l e crack propagation by specifying some (88 93—95)  average toughness that w i l l assure s e l f - a r r e s t  '  c r i t e r i a must be met at some specified minimum design  . These temperature.  For the Alcan/Foothills gas l i n e the most severe design is-18°C  ( 9 2 )  temperature  .  That the f u l l scale pipe fracture behaviour s h a l l be d u c t i l e and, therefore, that b r i t t l e fracture s h a l l be prevented i s ensured, according to the B a t t e l l e s t u d i e s  s  i f the fracture (39)  appearance of a Battelle-Drop Weight Tear Test specimen 85% or more shear when tested at the minimum operating  exhibits temperature.  Typical pipeline s p e c i f i c a t i o n s therefore require that of a l l the material tested ( i . e . , 50% of the heats per every 10 miles of pipe shipped), the average percent shear exhibited by the DWTT specimens be greater than 85%. However, any given heat can be accepted i f (92) 60% shear i s obtained i n the DWTT . The B a t t e l l e studies also generated an empirical formula which relates Charpy upper shelf energies to the c r i t i c a l defect size necessary for d u c t i l e crack i n i t i a t i o n under s t a t i c loading (95-97)  conditions.  This equation i s geometry dependent  „  2 : 1 T  OC  <J_ 2  "  Intsec (^Ma /2cr )] P c  (Eq. 3.2)  - 79 -  where,  K  =  12C E/A  C  =  Charpy upper shelf energy ( f t - l b ) = ET  A  =  Area of Charpy specimen ligament (- 0.124  2c  =  Length of sharp through-wall flaw (in)  a  =  Flow stress (- y i e l d stress + 10 k s i ) ( p s i )  =  " F o l i a s correction" (= f(pipe radius and wall  2  M  2 in )  thickness)) E  =  E l a s t i c Modulus (psi)  0 P  =  F a i l u r e stress (psi)  This equation i s widely used i n the pipeline industry to predict allowable defect sizes for the prevention of d u c t i l e crack (94-95) initiation  .  However, i t should be noted that this equation  i s applicable only for the range of temperatures over which the Charpy upper shelf energy.exists.  Futhermore, i t does not account  for i n i t i a t i o n resulting from dynamic loading or as a r e s u l t of unusual stress states and/or geometries. These l a t t e r conditions could be created by impacts from machinery or seismic action, or from bending stresses and buckling due to frost heave and d i f f e r e n t i a l (92) settlement which could cause p l a s t i c i n s t a b i l i t y At very high levels of toughness a material becomes e s s e n t i a l l y "flow stress dependent", that i s the f a i l u r e stress predicted from Equation 3.2 becomes dependent only upon crack length, y i e l d strength, and pipe geometry  - 80 -  For design purposes, pipeline companies t y p i c a l l y calculate the size of the toughness independent c r i t i c a l crack and then determine, from Equation 3.2, the minimum Charpy upper shelf energy necessary to prevent the i n i t i a t i o n of a crack 95% the size of that flow (92) stress dependent defect  This c r i t i c a l defect size must be large enough to be observed as a leak during hydraulic proof testing, or to be detectable (95) by nondestructive test techniques  . Proof testing of the p i p e l i n e  i s generally conducted at 1.05 times the specified minimum y i e l d stress (92) (SMYS) to assure that no c r i t i c a l size defects exist  Equation 3.2 predicts f o r a 42-in (107 cm) diameter, 0.540-in (13.7 mm) w a l l , X70 p i p e l i n e , operating at a design factor of 0.8, that a minimum Charpy energy of 51 f t - l b  (69 J) i s needed to prevent the  i n i t i a t i o n of the 95% flow stress dependent c r i t i c a l size defect of 5.8-in (15 cm).  This toughness l e v e l i s being specified as the a l l (92)  heat minimum value f o r the Alcan/Foothills p i p e l i n e  . However,  current proposals f o r that project c a l l for X70 pipe of either 48-in (122 cm) or 54-in (137 cm) outside diameter, 0.540-in wall thickness. Such pipe would require 55 f t - l b  (75 J) and 60 f t - l b  (81 J) minimum  Charpy toughness to assure i n i t i a t i o n prevention of the 95% flow stress dependent defect. minimum, however.  Current pipeline s p e c i f i c a t i o n s remain at 50 f t - l b  - 81 -  Another empirical r e l a t i o n s h i p generated by B a t t e l l e from f u l l scale burst tests simulating actual p i p e l i n e operating  conditions  has established the toughness required of a s t e e l to arrest a f a s t running du£tile_£rack.  For.certain grades of s t e e l , pipe geometries  y i e l d strengths, and stress l e v e l s , an empirical formula has been derived which predicts the minimum Charpy'energy, C^, required f o r d u c t i l e fracture a r r e s t :  C  =  V  where,  2  1/3  0.0873a„ (Rt) ' A ( f t - l b )  (Eq. 3.3)  H  a  =  operating stress (ksi)  R  =  pipe radius (in)  t  =  pipe wall thickness (in)  A  =  area of Charpy specimen ligament (^ 0.124 i n ) 2  The operating stress l e v e l , a„, i s t y p i c a l l y 0.8 SMYS. H  Although the f i n a l dimensions of the Alcan/Foothills pipel i n e has not been set, a current proposal i s for a 48-in (122 cm) diameter, 0.540-in (13.7 mm) wall thickness pipe operated at a design factor of 0.8 of the SMYS of 70 k s i (483 MPa).  For t h i s material  and the operating conditions chosen, Equation 3.3 requires that the pipeline s t e e l have 79 f t - l b  (107 J) Charpy upper shelf energy at  the minimum operating temperature. (108 J) i s , i n f a c t ,  An energy l e v e l of 80 f t - l b  often used i n p i p e l i n e s t e e l s p e c i f i c a t i o n s as  - 82 -  an a l l heat average toughness value.  This "average" toughness  philosophy is based upon the hypothesis that although a l l sections of the pipeline may not have sufficient toughness to arrest a propagating crack, the crack would eventually run into a section with high toughness and would thereby be arrested.  Unfortunately, the empirical relationship has not always correlated well with results from larger diameter (over 42-in), (94_95 98) higher strength (over X65 grade), and heavier wall pipe  '  In fact, for the newer, controlled-rolled AF steels " i t i s impossible [from this relationship] to accurately specify the toughness requirements" for ductile fracture a r r e s t U n t i l relationships similar to that given in Equation 3.3 are established for the AF steels and/or the larger pipelines (through full-scale burst testing), employment of empirical correlations should be used with caution.  Nevertheless,  pipeline companies continue to apply such correlations to establish (92 minimum toughness levels for crack arrest in line pipe materials  '  94-95) Pipeline manufacturers when testing pipeline material to determine i f i t meets the touhgness specifications, test specimens in which the fracture path is representative of the longitudinal axis of the pipe; the maximum operating stresses in a pipeline are the  - 83 -  hoop stresses, a , which tend to open cracks along the pipe axis.* n  The Battelle-Drop  Weight Tear Test and the standard Charpy  impact test are specified for specimens oriented for fracture along the pipe axis d i r e c t i o n .  Pipeline s t e e l s , p a r t i c u l a r l y i f not treated with rare earths and/or desulfurized,are  known to exhibit anisotropy i n  mechanical properties, especially toughness.  Thus, although a s t e e l  may be s p i r a l welded and thereby exhibit a maximum toughness along the pipe axis, the properties of the pipe at small angles to the pipe axis may be s i g n i f i c a n t l y below the specified minimums; yet the  pipe stresses would s t i l l be a s i g n i f i c a n t f r a c t i o n of the  maximum hoop stress.  However, as yet no s p e c i f i c a t i o n s require  tests or minimum toughnesses for directions other than the pipe axis.  The fracture control philosophy used i n the p i p e l i n e i n dustry i s one which s t r i v e s for prevention of both fracture i n i t i a tion and propagation.  Table 3.1 summarizes the current  fracture  control proposals for A r c t i c pipelines.  *  Weld zones are also tested, but t h i s area of investigation i s not a topic of discussion i n t h i s work.  Table 3.1 PROPOSED FRACTURE CONTROL REQUIREMENTS FOR ARCTIC PIPELINES  CURRENT REQUIREMENTS PIPE  WALL  MINIMUM  DIAMETER  THICKNESS  YIELD  PRESSURE  Pipe  DWTT  STRENGTH Toughness (in)  (in)  (ksi)  42  0.540  70  1440  48  0.540  70  1260  54  0.540  70  1120  48  0.720  70  1680  (psi)  Tests done only i n pipe axis o r i e n t a t i o n Minimum design temperature, -18°C Maximum operating stress, 56 k s i * **  From Equation 3.2 From Equation 3.3  BASIS FOR REQUIREMENTS  OPERATING  % Shear  (ft-lb)  50 f t - l b minimum any heat 80 f t - l b average a l l heats  95% Flow Stress Dependent* Critical Flaw Size (in)  C  v  for I n i t i a t i o n Prevention (ft-lb)  C  v  for Crack Arrest** (ft-lb)  60% minimum  5.8  51"  76  6.2  55  79  85%  6.6  60  82  average  7.2  65  88  - 85 -  To prevent b r i t t l e fracture, an average of 85% shear i s required on the DWTT conducted at the minimum design temperature. However, this test i s primarily a measure of the propagation mode as revealed  by the associated  percent shear on the fracture  Ductile f a i l u r e s are controlled only through empirical which are based upon the Charpy upper shelf energy.  surfaces.  equations  The standard  Charpy t e s t , i n i t s e l f , reveals nothing of the crack i n i t i a t i o n process.  Both i t and the B-DWTT employ blunt notches and are  therefore,representative  not,  of the most severe defect, for example, a  weld or fatigue crack.  3.4  Test Program  In t h i s series of tests the response of two a c i c u l a r f e r r i t e HSLA pipeline steels to dynamic loading by IIT was determined. The comparison study i s important for the following reasons: 1.  Both steels are to be employed i n the proposed Alcan/  F o o t h i l l s gas pipeline and, steels must be characterized  therefore, the dynamic response of the to assess the effect of i n service  loading. 2, concern.  For pipelines, fracture control i s a high p r i o r i t y  At present only two tests are routinely conducted on  pipeline materials  to assess t h e i r fracture resistance - the Drop  Weight Tear Test and the standard Charpy impact t e s t .  Neither test  - 86 -  provides a measure of crack i n i t i a t i o n . n o r distinguishes between crack i n i t i a t i o n and crack propagation. To ensure adequate f r a c ture resistance, pipeline companies employ empirical established  correlations  for lower strength, nonacicular pipeline s t e e l s , the  relationships being v a l i d only at temperatures corresponding to  the  Charpy upper shelf energy. 3)  An IIT study w i l l provide i n i t i a t i o n and propagation  energy data which can then be used to establish more fundamental fracture toughness correlations. because i t i s a r e l a t i v e l y new required  IIT lacks general acceptance  test and correlations are  for f u l l scale behaviour.  contribute control  The  to a more meaningful and  still  Analysis of the IIT data  may  sophisticated basis for fracture  and/or point out the a p p l i c a b i l i t y of IIT as a rapid, i n -  expensive test for quality assurance purposes.  3.4.1  Steels/Pipelines  Sections of pipe from production heats were supplied two Canadian producers for the test program. were 42-in (107  cm)  wall thickness,  and were rated as X70  Both pipe products  outside diameter, with a 0.540-in (13.7  strength of 70 k s i ) .  by  mm)  grade s t e e l (minimum y i e l d  These HSLA steels had been controlled r o l l e d  to achieve a f i n e grained acicular f e r r i t e microstructure. pipe sections had been s p i r a l welded.  Both  - 87 -  The chemical compositions of these steels are given i n Table 3.2.  Table 3.2 STEEL COMPOSITIONS  C  Mo  Mn  Nb  Si  P  Cu  Ni  Cr  Sn  Ti  .023  .012  .24  .10  .04  .02  -  .006  .006  .037  .027  .068  .005  S  Al  -  AF-1  .05 1.93 .26 .063  .03  AF-2  .06 1.82 .45 .05  .26 .045  .002  Ce  .034  A l l values i n weight percent  From Table 3.2 i t can be seen that the s t e e l designated AF-2 was f u l l y k i l l e d and rare earth treated. for the s t e e l AF-1.  This was not the case  The sulfur content of the AF-1 s t e e l was also  comparatively high.  3.4.2  Metallography  The structure of both steels was examined to determine the sulphide inclusion shape and the grain size.  Both steels had a very f i n e non-equiaxed f e r r i t i c grain structure, their low carbon contents precluding the p o s s i b i l i t y of  - 88 -  visible carbide formation. acicular ferrite.  This microstructure is typical of that of  The grain sizes, determined using the Heyn-Inter(  (99) cept m e t h o d , were ASTM 13.25 v  for steel AF-2, and ASTM 12.7 for  the AF-1 material. The high sulfur AF-1 steel had not been rare earth treated for inclusion shape control and therefore exhibited numerous long sulfide "stringers".  Figure 3.1 shows the directional inclusions  and very fine grained AF microstructure of this steel.  The structure of the AF-2 steel is shown in Figure 3.2, only globular nondirectional inclusions being visible.  The inclusions in both AF steels were examined using the analytical capability of the scanning electron microscope.  The  results showed that the AF-1 steel contained only MnS inclusions, as shown in Figure 3.3, whereas the AF-2 steel contained inclusions of several different compositions, not a l l of which were sulfides (Figures 3.4a - c).  3.4.3  Instrumented Impact Test Specimens  3.4.3.1  Specimen Preparation and Configuration  Large sections were i n i t i a l l y flame cut from the pipe. Test specimens were then saw cut such that the notch in each test  - 8*1 -  Figure 3.1  AF-1 photomicrographs, 225X (a) unetched (b) etched, 2% n i t a l  - 90  Figure 3.2  -  AF-2 photomicrographs, etched 2% n i t a l , 363X.  - 91 -  Figure 3.3  AF-1 SEM photomicrograph (3000X) and X-ray energy analysis of inclusions.  - 92 -  Figure 3.4a  AF-2 SEM photomicrograph (480X) and X-ray energy analysis of inclusions.  - 93 -  Figure 3.4b  AF-2 SEM photomicrograph (1000X) and X-ray energy analysis of inclusions.  Figure 3.4c  AF-2 SEM photomicrograph (4000X) and X-ray energy analysis of inclusions.  - 95 -  specimen was a minimum of 15 cm from any weld or flame cut edge.  The specimens were notched through the pipe wall thickness and cut so that the cracks would follow a path: pipe axis;  2)  1)  p a r a l l e l to the  p a r a l l e l to the r o l l i n g d i r e c t i o n ; and, 3)  to the r o l l i n g d i r e c t i o n .  transverse  The r o l l i n g d i r e c t i o n was at an angle of  63° to the pipe axis f o r the AF-1 pipe and 45° to the pipe axis f o r the AF-2 pipe.  3.4.3.2  Specimen Dimensions  Three types of specimens were tested i n order to assess the behaviour of d i f f e r e n t defects and thicknesses.  This broader  range of test conditions allowed better characterization of the dynamic properties of the s t e e l s .  For each of the orientations  discussed i n the previous Section, standard Charpy  V-notch speci-  mens, precracked Charpy specimens, and " f u l l w a l l " Charpys were prepared.  Thus, f o r each s t e e l , a minimum of nine series of  d u c t i l e - t o - b r i t t l e t r a n s i t i o n curves were established using the instrumented impact test equipment. performed at each temperature.  A minimum of three tests were  In addition, for the AF-1 s t e e l  only, standard Charpy specimens with notches cut transverse to the pipe axis were also tested.  - 96 -  3.4.3.2.1  Standard Charpy V-Notch Specimens  Standard Charpy V-notch specimens were considered the reference samples.  The Charpy test i s widely used i n the pipeline  industry, fracture control being based upon correlations and r e lationships which employ Charpy test data.  The standard Charpy  specimen i s 10 mm x 10 mm x 55 mm, with a 2 mm deep 45° V-notch (44) with a root radius of 0.25 mm  3.4.3.2.2  Precracked Charpy Specimens  Standard Charpy V-notch specimens were fatigue precracked p r i o r to testing.  The sharp fatigue crack provides the most severe  notch acuity and i s better adapted f o r fracture toughness measurements.  Samples were precracked using a Dynatup Precracker following tentative ASTM s p e c i f i c a t i o n s for precracking small specimens.  Both surfaces perpendicular to the V-notch were  polished p r i o r to fatigue precracking so that crack growth could be observed.  The ASTM requirements are that the crack front be  r e l a t i v e l y straight ( i . e . , not fa "thumbnail"); that the plane of symmetry between the notch and the precrack be less than 10°; that the cracks have a minimum length (1.3 mm) to avoid the p l a s t i c  - 97 -  region ahead of the machined notch; that the t o t a l crack length (notch plus precrack) to sample width r a t i o (a/w) and 0.55;  and,  l i e between  0.35  that the stress i n t e n s i t y factor associated with the  fatigue precracking, K^, be s t r i c t l y controlled and  sufficiently  small to avoid p l a s t i c deformation at the crack t i p ( i n general, must not exceed 60 percent of the s t r e s s - i n t e n s i t y value determined i n subsequent t e s t i n g ) . below 25 k s i - i n  For the AF s t e e l s ,  was  necessarily kept  (27.5 MPa-m ) and the specimens were precracked i n  approximately 15-20,000 cycles.  A l l these precracking c r i t e r i a are  e s s e n t i a l l y the same as those required for large fracture toughness specimens  (100)  .  After each instrumented impact test the fracture surface was  examined with a t r a v e l l i n g microscope to determine the  precrack  length and to assure that the fatigue crack c h a r a c t e r i s t i c s met  the  ASTM requirements.  3.4.3.2.3  F u l l Wall Charpy Specimens  It i s well established that sub-size specimens can have t r a n s i t i o n temperatures, toughness, and fracture modes which d i f f e r from those of a f u l l s i z e component  102)^  ^  s  ^  e  thickness of  a specimen increases, the degree of constraint also increases u n t i l plane-strain conditions e x i s t ^ \  This increasing constraint r e s u l t s  - 98 -  i n a decrease i n both the fracture stress and toughness, to a l i m i t i n g plane-strain value.  The extent to which variations i n  thickness a f f e c t crack i n i t i a t i o n energy or fracture toughness i s , of course, of great i n t e r e s t .  To ensure that toughness and fracture  data i s conservative for design applications, i t i s necessary that the test specimens duplicate the thicknesses of the actual service components to maximize constraint and thereby minimize  the  toughness behaviour.  To test f u l l pipe wall thickness specimens, Charpy-style samples were prepared i n which the thickness (dimension across the notch) was that of the pipe wall (nominally 13.7 mm).  The other  dimensions, including the notch size and acuity, were i d e n t i c a l to the standard Charpy specimen.  3.5  Instrumented  Impact Test Results  Figures 3.5 - 3.10 over a range of temperatures,  show representative load-time traces, for the AF-1 and AF-2  each specimen type and sample orientation.  s t e e l s , and for  Corresponding  surfaces of the test specimens are shown i n Figures 3.11 -  fracture 3.16.  Figure 3.5  AF-1 s t e e l load-time curves, crack p a r a l l e l to pipe axis.  - (00 -  Figure 3.6  AF-2 s t e e l load-time curves, crack p a r a l l e l to pipe axis.  F i g u r e 3.7  AF-1 s t e e l load-time c u r v e s , c r a c k p a r a l l e l to r o l l i n g d i r e c t i o n .  -  STANDARD  innn LB/DIV , n.? ns/mv  Figure 3.8  101 -  PRECRACKED  5 0 0 u:/niv x 0 . 2 I-S/DIV  1 * 0 LE/3IV » 0 . 2 RS/D1V  AF-2 steel load-time curves, crack p a r a l l e l to r o l l i n g direction.  -  Figure 3.9  103  -  AF-1 steel load-time curves, crack transverse to r o l l i n g d i r e c t i o n .  - 104 -  Figure 3.10  AF-2 s t e e l load-time curves, k transverse to r o l l i n g d i r e c t i o n .  - 105 -  Figure 3.11  AF-1 steel fracture surfaces, crack parallel to pipe axis.  - 106 -  standard  Figure 3.12  precracked  full-wall  AF-2 steel fracture surfaces, crack parallel to pipe axis.  - 107 -  standard  Figure 3.13  precracked  full-wall  AF-1 s t e e l fracture surfaces, crack p a r a l l e l to r o l l i n g d i r e c t i o n .  - 108 -  standard  F i g u r e 3.14  precracked  full-wall  AF-2 s t e e l f r a c t u r e s u r f a c e s , c r a c k p a r a l l e l to r o l l i n g d i r e c t i o n .  - 109 -  standard  Figure 3.15  precracked  full-wall  AF-1 s t e e l fracture surfaces, crack transverse to r o l l i n g d i r e c t i o n .  - no  standard Figure 3.16  -  precracked  f u l l wall  AF-2 s t e e l fracture surfaces, crack transverse to rolling direction.  - Ill -  3.5.1  Absorbed Energy  The t o t a l energy absorbed by each specimen has been d i f f e r e n t i a t e d into two components,the crack i n i t i a t i o n energy (energy to maximum load) and the crack propagation energy (postmaximum load energy).  Energy values are l i s t e d i n Tables 3.3 - 3.5  for each specimen type; these values represent the averages of many tests.  The absorbed energy data are presented graphically  i n Figures 3.17 - 3.44. For the f u l l w a l l and precracked Charpy data, the energy values have been normalized by d i v i d i n g the measured absorbed energy values by the ligament area of the specimen p r i o r to impact.  The average absorbed energy values of the  corresponding standard Charpys have been included to f a c i l i t a t e comparisons.  3.5.1.1  Comparison of the Two AF Steels  3.5.1.1.1  3.5.1.1.1.1  Standard Charpy Data  Crack P a r a l l e l to Pipe Axis  The standard Charpy tests with the fracture path p a r a l l e l to the pipe axis i s i d e n t i c a l to that required by the p i p e l i n e specifications.  This i s considered the most important orientation  Table 3.3 STANDARD CHARPYS - AVERAGE ABSORBED ENERGIES (EI + EP = ET)  P a r a l l e l to Pipe Axis T(°C) AF-1 +100 + 20 0 - 10 -20 -30 -40 -50 -60 -70 -80 -100 -196  33+89=121 30+98=128 25+78=102 19+77=96 20+75=95 13+64=76 19+56=75 10+38=48 2+7=9 2+5=7 1+3=4 0+1=1  P a r a l l e l to R o l l i n g Direction  AF-2  AF-1  32+56=88 34+50=84  7+20=27 5+15=20 6+19=24  21+41=62 20+37=58 15+26=42 10+21=31 5+13=18 2+8=10 2+5=7 2+3=5  4+16=20 4+11=17 4+13=17 2+12=14 3+11=14 2+5=7 1+5=6 1+3=4 0+1=1  AF-2  24+37=61 22+36=58 17+34=50 14+32=45 13+26=39 9+24=34 7+21=27 1+15=16 1+5=6  A l l values given i n f t - l b 1 f t - l b = 1.4 J  Transverse to R o l l i n g Direction AF-1  AF-2  31+65=95 30+72=102 22+83=105  36+83=119  28+84=112 20+73=93 19+69=88 13+53=66 9+46=55 1+12=13 1+4=5 0+1=1  39+63=102 33+61=94 30+55=85 22+53=75 23+44=67 19+44=63 12+29=41 5+3=8  Transverse to Pipe Axis AF-1 6+17=23 5+20=25 6+25=30 6+19=25 3+13=16 3+8=11 1+6=7 1+3=4 0+1=1  Table 3.4 FULL WALL CHARPYS - AVERAGE ABSORBED ENERGIES (EI + EP = ET)  T(°C)  AF-1 + 20 0 - 20 - 30 - 40 - 50 - 60 - 70 - 80 -100  P a r a l l e l to Rolling Direction  P a r a l l e l to Pipe Axis  26+85=111 _  24+87=111 —  20+91=110 20+79=99 13+68=81 7+53=60 2+9=12 2+3=5  AF-1  AF-2  8+16=24  39+75=114 30+71=102 29+69=98 28+50=78 23+47=70 10+28=39 8+35=43 5+25=30 3+26=28 1+4=5  -  6+15=21  -  6+13=19  -  3+11=14 2+10=12 2+6=8 —  A l l values given i n f t - l b 1 ft-lb  =  1.4 J  AF-2 21+42=63 22+43=65 16+44=60 -  -  18+38=56 15+35=50 10+33=43 3+23=26 4+23=27 2+6=8  Transverse to Rolling Direction AF-1 36+91=127  24+97=121  -  23+102=125 19+81=93 12+81=93 15+68=83 5+29=34  AF-2 46+86=132 36+81=117 30+79=109 —  32+73=105 26+69=95 25+61=86 19+58=77 12+42=54 5+38=43  Table 3.5 PRECRACKED CHARPYS - AVERAGE 'NORMALIZED ENERGIES (EI - ET)  P a r a l l e l to Pipe Axis T(°C) + 20 0 - 20 - 30 - 40 - 50 - 60 - 70 - 80 -100  AF-1 82 - 442 100 71 47 9 11  -  472 420 343 311 232  4 - 120  P a r a l l e l to R o l l i n g Direction  AF-2  AF-1  AF-2  AF-1  AF-2  92 - 408 76 - 367 45 - 311 45 - 305 12 - 221 9 - 183 8 - 125 5 - 120 4 - 78 5 - 22  7 - 128  94 - 483  7 - 114  55 - 287 39 - 276 26 - 254  106 - 490 97 - 429 82 - 406  4 - 119 5 - 91 4-72 4-59 6 - 43 4-22  17 - 234 13 - 191 5 - 142 7 - 114 4-95 4-49  A l l values given i n f t - l b / i n ^ For a standard Charpy specimen 1 f t - l b represents = 8 ft-lb/in  2  Transverse to R o i l i n g Direction  =  8 ft-lb/in  1.7 J/cm  2  100 - 478 99 - 502 50 - 423 22 - 383 7 - 245 5 - 178 2 - 132 3-24  76 - 328 37 - 329 14 - 248 9 - 208 6-136 3-46  - 115 -  because f u l l scale tests indicate that f a i l u r e s do propagate along the  pipe axis.  The maximum operating stress i n pipelines i s the  hoop stress which tends to open such cracks.  The differences i n the impact resistance of the two AF steels are s t r i k i n g , as shown i n Figures 3.17-3.19.  In terms of  their meeting the suggested standards outlined i n Table 3.1 at the -18°C s p e c i f i c a t i o n temperature, the AF-2 s t e e l does not meet the 80 f t - l b (108 J) average energy c r i t e r i o n , having only 62.3 f t - l b (84 J) at -20°C (Figure 3.18).  Similar r e s u l t s from the pipe  manufacturer confirm this lack of toughness for this heat.  (N.B.  This does not imply that a l l such pipe w i l l not meet the 80 f t - l b specification.  The proposed s p e c i f i c a t i o n requires only that the  average value f o r a l l heats be 80 f t - l b ; on this basis the AF-2 (92) s t e e l has been shown to thus qualify as A r c t i c grade pipe)  The AF-1 s t e e l e a s i l y met the 80 f t - l b requirement, having an average toughness of 95.7 f t - l b (130 J) at -20°C (Figure 3.17). However, the data did exhibit considerable scatter.  The standard  deviations for the t o t a l absorbed energy ranged from a maximum of 39.4 f t - l b the  (53 J) i n the t r a n s i t i o n region, to 22.7 f t - l b  upper shelf temperatures.  (31 J) at  Scatter i n the t r a n s i t i o n region,  however, i s often observed i n Charpy testing.  - 116 ^  220h  AF-1-absorbed energy , crack parallel toj pipe axis j #  • Tjtal energy  1  • Propagation energy  ;  o Initiation energy  n  »i  • I •  Figure 3.17  : ii • u  lOOh  20 H  oL  -  ire) -  20  .100  20  I  180  -t--  100  A F - 2 - a b s o i b e d energy i standard charpy crack parallel to ' pipe axis J • Total energy • .1 o Propagation energy  Figure 3.18  160  *|  o_ JnitigjMjn_enera.y  -HO S 40  \Z0  .100  ~ ? 0 ^ 2 0 T(°C)  -60  T  1 1 1 • AF-1/AF-2-energy (average values) standard charpy crack parallel to pipe axis r • total energy AF-1 \ • propagation I » initiation , o total AF-2 \ o propagation ' I * > initiation  •  •  Figure 3.19  •  .  *  8  •  •  *  1 o[  0  O •  •  1  •  I 1 1— 1  .  a  1.  o  * Q  o Q  8  * A  • *  • * A  •  0  : ' *  -  — i  • 1  • •  •  •V  1  -100  -60  T(°C)  -20  -20  - 117 -  In a c l a s s i c paper, Crussard and c o w o r k e r s h a v e shown that this i s due to an effect.termed "bimodal behaviour".  In the  t r a n s i t i o n region, two mechanisms of fracture (cleavage and fibrous) may be operative at the same temperature.  The magnitude of the  absorbed energies therefore f a l l into two d i s t i n c t groups:  one of  high energy, void formation and coalescence being the dominant fracture mechanism; or a low energy group c h a r a c t e r i s t i c of cleavage dominating the fracture event.  This bimodal behaviour i n the  t r a n s i t i o n region was often observed i n the AF-1 s t e e l , though not i n the AF-2 material.  It should be emphasized that the AF-1 s t e e l revealed a high degree of scatter at a l l temperatures, not j u s t i n the t r a n s i t i o n region.  This may be due to the greater number of inclusions present  (0.023 w/o S i n AF-1; 0.006 w/o S i n AF-2) which are well known to be deleterious to impact resistance.  Although the AF-1 s t e e l exhibited higher t o t a l energies than did the AF-2 s t e e l (to -70°C), the AF-2 s t e e l had e s s e n t i a l l y equivalent i n i t i a t i o n energies (refer to Table 3.3 and Figure 3.19). Thus, the AF-2 s t e e l requires an equivalent energy to i n i t i a t e a crack while crack propagation i s much more d i f f i c u l t i n the AF-1 steel.  This implies that the matrix of AF-1 had a higher work  hardening rate and hence void coalescence was more d i f f i c u l t ^ ' ^ " ^ .  - 118 -  The higher work hardening rate must counteract the p o s s i b i l i t y of a lower energy due to fibrous fracture associated with the higher inclusion content of the AF-1.  However, elongated inclusions  aligned normal to the crack t i p require a larger p l a s t i c zone size to "envelop" them before the fracture s t r a i n can be attained.  Thus,  the rate of void growth from such inclusions i s lower than from inclusions aligned p a r a l l e l to the crack t i p ^ " ^ * ^ .  In addition, the  advancing crack can propagate i n the transverse d i r e c t i o n i n AF-1 upon reaching a band of inclusions, and thereby e f f e c t i v e l y blunt the crack (5 9) tip  '  . The fracture surface of the AF-1 standard Charpy specimen at  +20°C (Figure 3.11)  i s i n d i c a t i v e of such behaviour.  The r e l a t i v e d u c t i l i t i e s of the two steels can be compared by examining the " d u c t i l i t y index", DI.  This i s the r a t i o of pro-  pagation energy to i n i t i a t i o n energy  :  DI  =  EP/EI  (Eq.  3.4)  Low indices imply a " b r i t t l e " material behaviour since most of the energy Is absorbed e l a s t i c a l l y .  Typical values range from 0.4 for  E-glass to 61.5 for laminate composites, with steels having values generally between 2 to 20, depending on temperature, microstructure, etc<  (33,48,104)  m  Table 3.6 l i s t s d u c t i l i t y indices for both steels f o r each specimen geometry tested, at selected temperatures. s t e e l , independent  For the AF-1  of crack orientation, the propagation energy i s ,  i n general, a s i g n i f i c a n t l y higher proportion of the t o t a l energy  Table 3.6 DUCTILITY INDEX - STEEL AF-1  P a r a l l e l to Pipe Axis  P a r a l l e l to Rolling Direction  T(°C)  Transverse to R o l l i n g Direction  Standard  Full Wall  Precracked  Standard  Full Wall  Precracked  Standard  Full Wall  Precracked  +20  2.7  3.2  4.4  3.3  2.2  17.2  2.4  2.7  4.2  -20  4.0  3.7  3.7  3.5  2.3  14.9  3.1  4.0  3.8  -40  5.0  4.5  6.4  3.8  2.2  29.1  3.6  4.4  7.4  -60  3.8  5.4  20.9  3.3  3.2  16.5  3.9  6.6  33.5  -80  3.2  8.0  23.3  4.6  3.6  6.4  9.9  6.4  38.1  .  DUCTILITY INDEX - STEEL AF-2 +20  1.8  1.9  3.4  1.6  2.0  4.3  2.3  1.9  3.6  -20  1.9  2.3  6.0  2.0  2.7  8.8  1.6  2.6  3.8  -40  1.7  2.0  17.2  2.3  2.1  12.5  1.8  2.2  3.3  -60  2.5  4.1  14.4  2.6  3.2  26.5  1.9  3.0  17.0  -80  2.2  9.4  19.4  2.2  6.2  26.1  1.6  3.6  19.8  - 120 -  to f a i l u r e as compared to the AF-2  The 50 f t - l b  steel.  (68 J) t r a n s i t i o n temperature  s t e e l was -59°C versus -37°C for AF-2  for the AF-1  (Figure 3.19).  I t should be  noted that the AF-1 material retained a 50 f t - l b propagation  energy  down to -55°C.  Below -70°C the absorbed energy values of the two  steels  were e s s e n t i a l l y equivalent.  3.5.1.1.1.2  Crack P a r a l l e l To R o l l i n g Direction  The absorbed energies of specimens of both steels oriented with the crack path i n the r o l l i n g d i r e c t i o n are shown i n Figures 3.20 -  3.22.  In terms of the potential A r c t i c gas pipeline s p e c i f i c a t i o n s outlined i n Table 3.1 neither s t e e l meets the 80 f t - l b  (108 J) average  t o t a l energy c r i t e r i o n at -18°C, nor does the AF-1 material meet the 50 f t - l b  (68 J) minimum (at -20°C, the AF-2  average; AF-1 only 20.2 f t - l b  s t e e l had 50.4  ft-lb  (68 J)  (27 J ) ) . However, at this time, toughness  s p e c i f i c a t i o n s require testing only i n the pipe axis orientation.  This raises disturbing questions regarding pipeline toughness s p e c i f i c a t i o n s and their usefulness i n preventing f a i l u r e s .  The  s p e c i f i c a t i o n s are written to ensure high toughness along the pipe  - Ul -  100  AF-l-obsorbed energy standard charpy crack parallel to rolling direction • total energy o propagation energy o initiation energy  Figure 3.20  H60 .  60  7. » «  5 £3 1  i t  100  80  60  20  40  0 • TCC)  „  20  40  50  80  100  J80 AF-2-absorbed energy standard charpy crack parallel to rolling direction  100  • Total  energy  • Propagation energy  Figure 3.21  o Initiation energy  r 60f  20r  -20  .100  -20 TCC)  1  1  1  1  11  r  AF-l/AF-2-energy(overage vjjlues) standard charpy crack parallel to rolling direction • total a propagation . initiation  { {  o total • propagation a initiotion  180  Figure 3.22 2  80h  ,60  l  J40  40f  420 20 ^  -100  "^60  TCC)  -ET  »  - 122 -  axis since the maximum crack opening hoop stress i s operative i n that d i r e c t i o n . However, the magnitude of the minimum stresses i n a p i p e l i n e are only 1/3 AF-1  operating  that maximum hoop stress. The  s t e e l exhibits a f i v e f o l d decrease i n toughness at -20°C on  changing the test d i r e c t i o n from the crack p a r a l l e l to the pipe axis to the crack p a r a l l e l to the r o l l i n g d i r e c t i o n ( i . e . , f t - l b (130 J) to 20.2 AF-1  f t - l b (27 J ) ) .  95.7  The r o l l i n g d i r e c t i o n i n the  pipe i s 63° from the pipe axis and hence the operating  are greater than the minimum. effectiveness of preventing  stresses  This casts serious doubt upon the  crack i n i t i a t i o n when the weakest d i r e c -  tion i s not included i n the toughness test s p e c i f i c a t i o n s .  In addition, the AF-1,  when tested with the crack p a r a l l e l  to the r o l l i n g d i r e c t i o n , exhibits an extremely low crack  initiation  energy of approximately 5 f t - l b (7 J) for the entire test temperature range from +100°C to -60°C (refer to Table 3.3 and Figure 3.20). This suggests that a defect having even a r e l a t i v e l y blunt notch radius may  e a s i l y i n i t i a t e i n this lower toughness d i r e c t i o n due to  any sudden damage from pipe-laying equipment or buckling as a r e s u l t of frost heave. also.  High r e s i d u a l stresses could contribute to i n i t i a t i o n ,  It seems quite possible that a c r i t i c a l s i z e crack may  created which could propagate i n t h i s low energy d i r e c t i o n .  then be Even i f  the crack was not s u f f i c i e n t to cause a long running f a i l u r e , i t s t i l l represents a l o c a l i z e d crack requiring r e p a i r .  - 123 "  The pipeline industry should address i t s e l f to those p o s s i b i l i t i e s by: orientations;  and,  1)  requiring minimum toughness values i n a l l 2)  establishing s p e c i f i c a t i o n s based upon  i n i t i a t i o n energies f o r crack i n i t i a t i o n prevention.  In the r o l l i n g d i r e c t i o n , the AF-2  s t e e l was  significantly  superior to the AF-1 material, both i n i n i t i a t i o n and propagation energy.  No doubt the numerous elongated MnS  " s t r i n g e r s " i n the  AF-1 s t e e l provided low energy crack i n i t i a t i o n and propagation paths i n the r o l l i n g d i r e c t i o n (Figures 3.1 and 3.3), which could explain the lack of scatter i n the test data for the AF-1 material i n this orientation.  However, the AF-2 in this direction.  s t e e l also exhibited i t s lowest energies  Since this i s a low sulfur rare earth treated  s t e e l , the reduced energy i s probably due to the r o l l i n g texture and alignment of grain boundaries developed during controlled r o l l i n g .  In addition, as Figure 3.4b  indicates, the spherical inclusions  present i n AF-2 were aligned along the r o l l i n g d i r e c t i o n .  Crack  propagation, through the mechanism of void coalescence, i s therefore easier for the AF-2  s t e e l i n the r o l l i n g d i r e c t i o n (compare +20°C  data for AF-2 i n Table 3.3).  In comparing the general shape of the t r a n s i t i o n curves  - 124 -  for both s t e e l s , the AF-2 s t e e l , f o r a l l orientations, showed a continuous decrease i n i n i t i a t i o n and propagation energy with decreasing temperature.  The AF-1 s t e e l exhibited a more sudden  change from a high energy upper shelf region to the low energy values.  3.5.1.1.1.3  Crack Transverse to R o l l i n g Direction  The data obtained from testing both steels with the crack running transverse to the r o l l i n g d i r e c t i o n are shown i n Figures 3.23 - 3.25.  The two steels showed similar t o t a l energy values over the t o t a l range of test temperatures.  The AF-2 i n i t i a t i o n energy  was considerably higher than the EI of the AF-1 material, whereas the propagation energy of AF-1 was higher- than that of AF-2 (refer to Table 3.3).  As noted i n the previously discussed data, the AF-1 s t e e l displayed more scatter and showed a " c l a s s i c " upper shelf and sharp energy t r a n s i t i o n temperature at -80°C (Figure 3.23).  The AF-2  s t e e l showed less scatter and a continuous decrease i n energy with decreasing temperature (Figure 3.24).  - 125 -  "T A F - l - o b s o r b e d energy standard charpy c r a c k t r a n s v e r s e toi rolling direction i 180  Figure 3.23  •  Total  o  Propagation  o  Initiation  J  energy energy  energy  1  : :  g  . A-  8 » -100  TCC)  1  Figure 3.24  1  "T ] — 1  A F - 2 - a b s o r b e d energy standard charpy c r a c k t r a n s v e r s e to rolling direction  . I  • Total  |  1  energy  -  d Propagation  -  o Initiation  energy  —  |IOOj  • • •• • •u  •  • •  9 1 • j  0  o 1 o |  8  *  •  As  O 9  °  9 8  •  o o  -a  ] 1 1  °  8  o.  -100  ,!  40<  I  1  20  TCC) 0  120 A F - l / A F - 2 - e n e r g y (averoge standard charpy crack transverse to rolling direction  { AF-2] 120  . total . propagation . initiation , o total opropagation I .initiation  % • 80  Figure 3.25  |  UJ 60|  r •° B  80  1  •  •  o>  B ••  —  60  120  !  1• j -•— i ••  •  20  •  I  energy  -  <  1  100  60 i  • 80  • 40' 20 20!  -20 TCC)  • 20  - 126 -  Both steels met the 80 f t - l b  (108 J) average/50 f t - l b (68  J) minimum -18°C c r i t e r i a ; AF-2 had 102.0 f t - l b 112.1 f t - l b  (138 J) and AF-1  (152 J) average t o t a l energies at -20°C.  t r a n s i t i o n temperatures  The 50 f t - l b  were similar f o r both s t e e l s : -73°C f o r  AF-1; -77°C for AF-2.  3.5.1.1.1.4  Crack Transverse to Pipe Axis  Only the AF-1 s t e e l was examined i n t h i s d i r e c t i o n , the results being shown i n Figure 3.26.  Very low energies, 25.3 f t - l b obtained.  This orientation  (34 J) at -20°C, were  i s only 27° from the r o l l i n g  direction  which i s also the d i r e c t i o n i n which the s u l f i d e inclusions l i e . The observed low energies are thought to be due to cracks following the path of these "low energy stringers".  The high energy  orientations  observed for the AF-1 s t e e l were oriented at 63° (crack p a r a l l e l to the pipe axis) and 90° (crack transverse to the r o l l i n g to the r o l l i n g d i r e c t i o n , both orientations the path of the s u l f i d e  direction)  lying at a high angle from  stringers.  The toughness of the AF-1 s t e e l i n t h i s d i r e c t i o n and i n the r o l l i n g d i r e c t i o n was very poor, exhibiting  low values of propagation  energy and extremely low values of i n i t i a t i o n energy.  Since the pipe  - 127 -  100 80  I I  AF-1 -absorbed energy standard charpy crack tranverse to pipe axis  ' [ | |  • total energy  I  • propagation energy  I  o initiation energy  [  „ i  .  "|60  i  X60  480  i  s>  140^  UJ  UJ  -o40| JQ  20  JO <  20| 0L  -100  Figure  -60  3.26  -20  t ( o c )  .20  AF-1 absorbed energies, standard Charpy, crack transverse to pipe axis.  - 128 -  operating stresses i n these directions are s i g n i f i c a n t f r a c t i o n s of the maximum operating hoop stress (> 0.33 o^) some toughness requirement should be established.  3.5.1.1.2  F u l l Wall Charpys  A v a l i d c r i t i c i s m of the standard Charpy specimen i s that i t cannot predict the toughness of thicker materials.  More important,  though, the standard Charpy test i s nonconservative, since toughness decreases with increasing thickness would better represent  , F u l l pipe wall Charpys  the f u l l - s c a l e behaviour of a p i p e l i n e because  the constraint across the notch simulates  the service conditions.  By  testing such specimens for pipeline applications, i t was hoped that correlations between the f u l l wall Charpy data and the f u l l wall Battelle-Drop Weight Tear Test could be generated, and that the adequacy of the standard  Charpy specimen to represent  full-size  behaviour could be ascertained.  3.5.1.1.2.1  Crack P a r a l l e l to Pipe Axis  The standard Charpy test samples showed that the AF-1 material had a much higher t o t a l absorbed energy with the i n i t i a t i o n energy being comparable for both s t e e l s .  - Ill -  AF-1 "absorbed energy full pipe wall thickness charpy crack parallel to pipe axis • .•full wall • Total energy {. ovg  _  s t d  rofull wall 1. ,  .Initiation..  Q v g  s t d  c  h  h a r p >  a  r  p  y  460C*t  Figure 3.27  6 60  i• .100  -60  !  T  (  „  c  p-  .20  -20  )  —1  1  1  AF-2-absorbed energy full pipe wall thickness charpy crack parallel to pipe axis r • full wall Total energy ( . charpy s  Initiation »  t  •  d  j o full wall [ charpy a  a  v  g  s  t  d  t  • •  •  t  •  joo  Figure 3.28  •  1 1 20  -  • • • •  1 &  0  -100  0  e  • •  -  •  • •  200" B  •  0 D  8  6  1  i  0  • 0  ? O ] -60 TCC]  -20  1  AF-l/AF-2-energy(average values full wall charpy crack parallel to pipe axis total propagation initiation  AF-1 <  {  o total • • propagation * initiation  I20h  180  Figure 3.29  -60  -20 TCC)  - 130 -  The f u l l wall data shows similar behaviour.  However, at  +20°C, the t o t a l energies were e s s e n t i a l l y equal (Figures 3.29), the AF-1 s t e e l had 111 f t - l b the  (155 J ) ; and  AF-2 material exhibted a higher i n i t i a t i o n energy (39 f t - l b  as compared to 26 f t - l b not  (150 J) and AF-2 had 114 f t - l b  (35 J) for AF-1.  (53 J ) )  Since the AF-2 s t e e l did  exhibit an upper shelf, as did the AF-1 s t e e l  which retained  upper shelf energies to -40°C, the AF-2 t o t a l energy decreased progressively to values less than that of AF-1 to -70°C. The higher toughness of the AF-1 was related to i t s higher propagation energy.  There was l i t t l e scatter i n the data for either s t e e l , nor was bimodal behaviour observed (Figures 3.27 - 3.28).  3.5.1.1.2.2  Crack P a r a l l e l to R o l l i n g D i r e c t i o n  As with the standard Charpy comparison, the f u l l wall AF-2 specimens were tougher, requiring considerably more energy than those of the AF-1 s t e e l (Figures 3.30 - 3.32).  At no temperature was the  toughness of the AF-1 s t e e l comparable to that of the AF-2.  The upper shelf i n i t i a t i o n energy of the AF-1 s t e e l was constant at 6 f t - l b  (8 J) down to -40°C, compared with approximately  20 f t - l b (27 J) EI for AF-2 to. -40°C. the  The t o t a l absorbed energy for  f u l l wall AF-1 Charpy was only 21.4 f t - l b  (29 J) at -20°C; the  « 131. -  AF-1  -Absorbed energy full pipe wall thickness charpy crack parallel to rolling direction  Total energy/* full-wall I • avg. std. std. charpy  240  ° full-wall • avg std. charpy  Initiation  200^  Figure 3.30  160  j c  120  |  UJ  I I  80  »  I 10  40  AF-2  '  9 TOO  '  i  l  l  1  -80  -60  40 TCC)  -ZO  0  1  1  :  1  1—  T  1—l  2  —1  0  — I  - absorbed energy full pipe wall thickness charpy crack parallel to rolling direction  Total  energy  full wall •• avg. std. charpy  Initiation  "  r° full wall ° avg. std. charpy  500  • •  L  80  •  Figure 3.31  •  •  60  i 1  •  300  200  • • -  •  0  e  „  .  o  •  o  o  •  e  g  0  o  100  o  0  o e  -100  •  -60  1  1  1  •  _l  • 20  -20 T  1  1  • • •  •  20 1 •  1  CC)  1  1  1  1  A F - l / A F - 2 - e n e r g y (average values)  AF-1  AF-2  full wall charpy crack parallel to rolling • • total • propagation • * initiation • o to-tal • propagation • * initiation  direction  •  -  Figure 3.32  o  - 60 T  o o  -  o  •  •  o •  •  o O  a  •  o D  t  • t  •  • *  . * * S 8  •  •  •  1  -  •  -  A  »  i -20  i -60 TCC)  ,  •  uj  a  <  - 132 -  average t o t a l energy f o r the AF-2 specimens was 60.0 f t - l b  (81 J)  (refer to Table 3.4).  L i t t l e scatter and no bimodal behaviour was evident i n either s t e e l .  3.5.1.1.2.3  Crack Transverse to Rolling D i r e c t i o n  The t o t a l energy of the two steels was similar over the entire temperature range (Figures 3.33 - 3.35).  The AF-2 s t e e l  exhibited only a marginally higher energy than the AF-1 material at +20°C, but was less at lower temperatures.  As noted with the standard Charpy specimens, the i n i t i a t i o n energies of the AF-2 s t e e l were comparable to, but s l i g h t l y higher than those of the AF-1 s t e e l ; whereas, the propagation energies of the AF-1 were higher.  These differences were most noticeable at -40°C,  where the AF-1 and AF-2 propagation energies were 102 f t - l b and 73 f t - l b were 23 f t - l b  (138 J)  (99 J ) , respectively; but the i n i t i a t i o n energy values (31 J) f o r the AF-1 and 32 f t - l b  (43 J) f o r AF-2.  The AF-1 s t e e l exhibited bimodal behaviour from -50° through -80°C i n this orientation (Figure 3.33).  - 133 -  180  AF-1 - absorbed energy full pipe wall thickness charpy crack transverse to rolling direction ' • „.„i e n e r g y ( full wall T Total ^ ^ .• i  initiation ..  Figure 3.33  {  0,u  0  <  a  J  t  d  c t ) a r p >  " „ avg. std.charpy  '  •  1600_  J0O 4400:  !  8 9  8  -100  i  ft (959)  AF-2-absorbed energy full pipe wall thickness charpy crack transverse to rolling direction  H800  I  _ . , r • full wall Total energy | . Q v f l  s  t  d  c h o r p y  l a avg std. charpy  Figure 3.34  •  I  <  §60  -60  1 160 H•  p  , TCC)  1  -20  1 —i  A F - l / values) A F - 2 -energy (average full wall charpy crack transverse to rolling direction °  1  •  * 112 7) O  ° -  m •  •  o m  m  0  0  . total  •  Figure 3.35  -  •  AF-1- • propagation A initiation  °  o total  a  AF-2- • propagation ° initiation  o  °  a  •  • -  *  • A  A  *  *  A  20 A  *  A  •  -100  -60  -20 T(°C)  -  *  •20  -  - 134 3.5.1.1.3  3.5.1.1.3.1  Precracked Charpys  Crack P a r a l l e l to Pipe Axis  The AF-1 s t e e l shows higher t o t a l energies at a l l test temperatures (Figures 3.36 - 3.38) and superior i n i t i a t i o n energies for  the temperature range of -20° to -40°C.  For the f u l l wall and  standard Charpys, the i n i t i a t i o n energy of the AF-2 s t e e l was generally equivalent or higher than that of the AF-1 material over the  entire temperature range studied (compare Table 3.3 - 3.5). This  indicates that the AF-2 s t e e l may be more susceptible to crack i n i t i a t i o n as the notch acuity increases.  At -40°C, both steels exhibit sharp i n i t i a t i o n energy 2 transitions and lower shelf energies of 5-10 f t - l b / i n  2 (1-2 J/cm ).  There was very l i t t l e data scatter nor was bimodal behaviour observed.  3.5.1.1.3.2  Crack P a r a l l e l to R o l l i n g Direction  The AF-2 s t e e l required more energy f o r fracture than the AF-1 s t e e l i n t h i s d i r e c t i o n (Figures 3.39 - 3.41).  This same effect  was observed for the standard and f u l l wall specimens.  Fatigue precracking the samples s i g n i f i c a n t l y reduced the  - 135 -  AF-1-absorbed energy precracked charpy crack parallel to pipe axis , • precracked Total energy { . „ o v g  Initiation  s  t  0  c  o  r  p  y  r o precracked L o avg. std. charpy  -moo  140  Figure 3.36  •J600§  loo w UcJ •£60  •  -100  9 - 9 -60  -20  Tra  AF-2-absorbed energy precracked charpy crack parallel to pipe axis 140  Toto, energy  { avg. I ^ 'std. c cnorpy h o precracked • ovg. std. charpy  100  Figure 3.37 S60  20  1  0  8  " a H 8  D  S  1 '»  - °TCC) 6  80  AF-1/AF-2-energy (average values) precracked charpy crack parallel to pipe axis • total A F - i {: r * initiation rototal *- \ .-initiation •  -  2  °  .20  -440 - 420  - 380  2  Figure 3.38  - 340  • 300  60  - 260^ - 220  i  i 40 - 180  |  UJ  - 140  J  o  - 100 <  •420  -•o TCC)  - 136 -  AF-1 - obsorbed energy precrocked chorpy crack parallel to rolling direction Totol energy  precrocked chorpy  {*  avg. std. chorpy  e "  Figure 3.39  precracked charpy avg. std. chorpy  50  J240 5  g 40| « 1  c UJ  •o 301  0  1  I 20 10  • 40  Bfla - 2 0  0  *20 T (°C)  AF-2 - obsorbed energy precracked charpy crock parallel to rolling directic _ , . Tota  Figure 3.40  energy * „ " 7  . ... Initiation 3  /• 1 " r° {„ °  L  L  precrocked .. . avg. std. chorpy precrocked ovg. std. charpy u  80  H300  H200  i -100  -60  -20 T PC)  -1— A F - l / A F - 2 - e n e r g y (overage precracked crock  pardllel  values)  charpy to rolling  direction  total  AF-1 .-* initiation AF-2  ° total  Figure 3.41  J  40  +40  H60  _1 » » * -60  i  , , TPC)  i  -20  i  ±-  £  g  - 137 -  i n i t i a t i o n energies: the AF-1 specimens, at a l l temperatures, exhibited an i n i t i a t i o n energy of less than 1 f t - l b (1.4 J ) ; whereas at -20°C, the AF-2 precracked Charpys required only 3 f t - l b (4 J) for crack i n i t i a t i o n .  The effect of notch acuity i n this d i r e c t i o n can be seen by examining Table 3.7.  The i n i t i a t i o n energy of AF-2 i s more  Table 3.7 EFFECT OF NOTCH ACUITY  Standard Notch  +20°C  -20°C  Fatigue Precracked Notch  AF-1  AF-2  AF-1  AF-2  20  61  16  36  EI  5  24  0.9  ET  20  50  14  EI  4  17  0.9  ET  :  7 32 3  A l l values i n f t - l b Precracked values determined by multiplying normalized energies by area of standard Charpy ligament.  sensitive to the increased notch acuity.  The i n i t i a t i o n energy of  the AF-1 i s very low f o r both notch conditions at a l l temperatures.  - 138 -  3.5.1.1.3.3  Crack Transverse to R o l l i n g D i r e c t i o n  The two steels i n t h i s orientation, for a l l specimen types, yielded similar test r e s u l t s .  For the precracked Charpys,  the t o t a l energies were approximately equal f o r both s t e e l s , although the AF-1 s t e e l maintained an upper shelf energy to below -20°C and exhibited a sharp t r a n s i t i o n at -60°C.  The energy of  the AF-2 s t e e l gradually decreased with decreasing temperature (Figure 3.43).  The i n i t i a t i o n energy of the AF-2 s t e e l was more s e n s i t i v e to the presence of the sharper fatigue crack than was AF-1 (compare i n i t i a t i o n energies on Figures 3.42 and 3.43).  3.5.1.2  Significance of Specimen Size and Notch Acuity  3.5.1.2.1  3.5.1.2.1.1  AF-1 Steel  Crack P a r a l l e l to Pipe Axis  Figures 3.27 and 3.36 show the IIT absorbed energy r e s u l t s obtained from the f u l l wall and precracked specimens versus the standard Charpys, respectively,, for the AF-1 s t e e l .  - 139 -  I80T 800  I AF-1-absorbed energy precracked charpy crack transverse to rolling direction -J60O^ • precracked Total energy{ * J i avg. std. chorpj Initiation ii |j. o precracked I ° avg std. char pi 400 =  Figure 3.42  60  r •{zoo*  20 • 0.  Figure 3.43  140  i  -20  »20  TCC)  AF-2 - absorbed energy precracked charpy crack tranverse to rolling direction • precrocked Total energy{, avg. std. charpy • o precracked Initiation >< { avg. std. chorpy •  2(999) B  3  1001  • ! * v 60  < 20  _S_&_ -100  -60  -20  • 20  T(°C)  AF-1/AF-2-enerov (average values) precrocked chorpy crack transverse to rolling direction > total A F - I { : ; initiation  80 A M  F r  2  #  f o total l . initiation  Figure 3.44  HI40  « >- 6 0 t  TCC)  -20  .20  - 140 -  It should be remembered that the data from the AF-1  steel  exhibited s i g n i f i c a n t scatter; the average values also exhibit similar scatter (refer to Figure 3.17).  The minimum standard  deviation was approximately 20 f t - l b (27 J) or more.  Bimodal  behaviour was also observed i n the t r a n s i t i o n region for the standard specimens, but not f o r the precracked or f u l l wall Charpys.  In comparing the f u l l wall and standard specimens (Figure 3.27), the standard Charpys show a much higher t o t a l absorbed (977 f t - l b / i n  2  (205 J/cm  2  ) for the standard specimen, 651  energy  ft-lb/in  2  2 (137 J/cm  ) for the f u l l wall at +20°C).  This behaviour can be  explained i n that the greater thickness of the f u l l wall specimen provides greater constraint across the notch, thereby lowering the toughness. the absorbed  However, as the temperature decreased to below -40°C,  energies of the two specimen types became e s s e n t i a l l y  equivalent.  j  The i n i t i a t i o n energies were nearly equal at -20°C and below.  The f u l l wall Charpys exhibited a sharp energy t r a n s i t i o n at about -80°C, whereas the standard size specimens had a t o t a l energy t r a n s i t i o n 10°C higher. , O t h e r s h a v e  observed that  increasing the thickness of a Charpy specimen increases the t r a n s i t i o n temperature, however.  i  i  ' .  - 141 -  Thus, for the AF-1 s t e e l i n t h i s orientation, the standard Charpy specimen i s not representative of the f u l l size behaviour, except at very low temperatures; the standard Charpy data i s nonconservative.  For p i p e l i n e applications, adoption of a f u l l size  specimen would provide more meaningful data and would be easier to prepare.  A comparison of the precracked and standard Charpy data i s shown i n Figure 3.36.  specimen  As expected, the introduction of a  fatigue precrack s i g n i f i c a n t l y reduces the crack i n i t i a t i o n energy.  It i s i n t e r e s t i n g to note the e f f e c t that notch acuity has on the fracture process and the corresponding t o t a l absorbed energy.  By multiplying the -20°C precracked specimen normalized  energy (absorbed energy per unit area) by the area of a standard Charpy ligament (0.124 i n ) , a value of 58.6 f t - l b (79 J) t o t a l 2  energy and 12.4 f t - l b (17 J) i n i t i a t i o n energy would be obtained. The standard Charpys, at that same temperature, absorbed 96 f t - l b (130 J) t o t a l and 19 f t - l b (26 J) i n i t i a t i o n .  Thus, the presence  of the sharp- fatigue crack s i g n i f i c a n t l y reduces the energy to i n i t i a t e and the energy to propagate a crack, as shown i n Figure 3.36.  The e f f e c t of the precrack can be further demonstrated by comparing the r e l a t i v e amount of energy absorbed i n fracture propagat i o n and i n i t i a t i o n through an examination of the d u c t i l i t y indices  - 142 -  (DI) i n Table 3.6.  That Table shows that for the AF-1 s t e e l , tested  with the crack p a r a l l e l to the pipe axis at temperatures  from +20°C  to -40°C, the DI f o r the precracked specimens i s only s l i g h t l y higher than that of the standard blunt notched Charpys.  Thus,  although the fatigue flaw requires a much lower crack i n i t i a t i o n energy than the standard Charpy notch, the propagation energy i s also greatly reduced  (though not to the same extent).  As a notch  lengthens and sharpens during the fracture event, the s t r a i n concentrated near i t s t i p increases and the point of maximum stress i n t e n s i f i c a t i o n moves back towards the crack t i p .  The stress l e v e l  immediately ahead of the crack consequently increases and the crack accelerates.  Therefore, the presence of a sharper notch i n the  sample w i l l f a c i l i t a t e the crack i n i t i a t i o n and subsequently the propagation process by causing crack acceleration and subsequent s t r a i n hardening e a r l i e r i n the fracture event ^ \  At temperatures  of -60°C and -80°C, the AF-1 precracked  specimens exhibited very high d u c t i l i t y indices (> 20) indicating that the propagation energy was a s i g n i f i c a n t  proportion of the  t o t a l energy absorbed i n the crack process; i . e . , very l i t t l e  energy  was required f o r crack i n i t i a t i o n .  The t r a n s i t i o n behaviour of the t o t a l energy f o r the precracked specimens was better defined than that of the standard Charpys.  I I i  !  - 143 -  j  The precracked Charpy i n i t i a t i o n energy also showed a marked trans i t i o n between -40° and -50°C, decreasing from an upper shelf i n i t i a t i o n energy of approximately 10 f t - l b (14 J) to a lower shelf value of less t h a n l f t - l b (1.4 J ) . The t r a n s i t i o n i n the standard Charpy i n i t i a t i o n curve was not so sharp, exhibited b i modal behaviour, and did not reach the lower shelf u n t i l -70°C (Figure 3.36).  This observed behaviour i l l u s t r a t e s the c l a s s i c  effect of a sharp flaw i n a structure: the t r a n s i t i o n  temperature  curve i s shifted to higher temperatures and the magnitude of the upper shelf energy i s s i g n i f i c a n t l y reduced.  i 3.5.1.2.1.2  Crack P a r a l l e l to R o l l i n g Direction  The r e s u l t s from the standard, f u l l w a l l , and fatigue precracked Charpy specimens were not s i g n i f i c a n t l y d i f f e r e n t , probably due to the low magnitudes of the energies involved i n cracking  I along the b r i t t l e MnS  inclusions (Figures 3.30 and 3.39).  i Although the standard Charpy specimens did have marginally higher t o t a l energies, the differences between the standard and  full  wall specimens were small, being approximately 3-4 f t - l b (4-5 J) at +20°C.  The i n i t i a t i o n energies of those specimen types were v i r t u a l l y i  equal at a l l temperatures.  Thus, for this orientation, a standard  Charpy adequately represents the f u l l thickness impact behaviour.  -. 144 -  A l l three specimens of the AF-1 s t e e l , when tested with the  crack running p a r a l l e l to the r o l l i n g d i r e c t i o n , showed almost  no i n i t i a t i o n energy t r a n s i t i o n ; values of EI ranged from 8 f t - l b (11 J) to 2 f t - l b  (3 J) f o r the f u l l wall Charpys, whereas the pre-  cracked Charpys had only a constant low magnitude i n i t i a t i o n energy of less than 1 f t - l b studied.  (1.4 J) over the entire temperature range  These extremely low values of the i n i t i a t i o n energy f o r  the AF-1 s t e e l i n t h i s orientation must be emphasized.  The precracked d u c t i l i t y indices were quite high (Table 3.6), indicating that crack i n i t i a t i o n was an i n s i g n i f i c a n t component of the t o t a l absorbed energy.  In f a c t , the normalized pro-  pagation energies f o r the precracked specimens were e s s e n t i a l l y equivalent to that of the standard Charpys f o r t h i s orientation (Figure 3.39).  3.5.1.2.1.3  Crack Transverse to R o l l i n g D i r e c t i o n  The f u l l wall and standard specimens i n this orientation gave similar t o t a l and i n i t i a t i o n energy r e s u l t s (Figure 3.33). Both specimens exhibited bimodal behaviour i n the -60° to -80°C range. A s i g n i f i c a n t decrease i n the t o t a l and i n i t i a t i o n energies occurred for both specimens between -70° and -80°C.  - 145 -  As with the previous orientations, the standard Charpy specimens with the crack running transverse to the r o l l i n g d i r e c t i o n had much higher energies than the precracked samples (Figure 3.42). Only at -80°C d i d the two energy curves coincide.  The precracked  i n i t i a t i o n energy curve exhibited a sharp t r a n s i t i o n between -50° and -60°C, the energy decreasing to less than 1 f t - l b (1.4 J ) . The corresponding t r a n s i t i o n for the standard specimens was not as d i s t i n c t , and occurred over a lower temperature range of -60° to -80°C.  3.5.1.2.2  3.5.1.2.2.1  AF-2 Steel  Crack P a r a l l e l to Pipe Axis  The r e s u l t s of the AF-2 f u l l wall thickness Charpy tests are compared with those of the standard Charpys i n Figure 3.28. The increased constraint at the root of the notch i n the f u l l wall specimens caused a reduction i n the toughness of the AF-1 samples (Figure 3.27), but did not decrease the energy of the AF-2 specimens i n this direction.  The f u l l wall Charpys showed similar normalized t o t a l  and i n i t i a t i o n energies to those obtained from standard Charpys down to -60°C.  At lower temperatures, the propagation energy, and hence  the t o t a l energy of the f u l l wall specimens was greater. s i t i o n behaviour showed no thickness e f f e c t .  The tran-  Thus, the standard  - 146 -  Charpy adequately describes the f u l l wall impact behaviour of the AF-2 s t e e l .  This i s i n d i r e c t contrast to the comparison of r e s u l t s  made on the AF-1 s t e e l .  The data i n Figure 3.37 shows that precracking greatly reduced the i n i t i a t i o n energy and increased the t r a n s i t i o n temperature by approximately 30°C.  3.5.1.2.2.2  The standard Figure 3.31.  Crack P a r a l l e l to R o l l i n g D i r e c t i o n  and f u l l wall Charpy r e s u l t s are compared i n  I t can be seen that for the AF-2 s t e e l , the standard  Charpys adequately describe the f u l l s i z e behaviour f o r temperatures below -20°C.  The precracked AF-2 specimens exhibited a continuously decreasing  i n i t i a t i o n energy with decreasing  temperature, with l e s s  than 10 f t - l b (14 J) required below -40°C (Figure 3.40).  3.5.1.2.2.3  Crack Transverse to R o l l i n g D i r e c t i o n  The differences between the standard Charpy and the f u l l wall Charpy are more d i s t i n c t i n this orientation for the AF-2 material (Figure 3.34). standard  The normalized t o t a l energies of the  specimen are higher f o r temperatures down to -70°C.  The  - 147 -  i n i t i a t i o n energies of the standard Charpys were higher at a l l temperatures.  The f u l l wall Charpys exhibited bimodal behaviour  between -80° and -100°C, unusual for the f u l l wall specimens i n this study.  The differences between the standard and the precracked specimens were also more pronounced than for the other orientations (Figure 3.43).  The normalized i n i t i a t i o n energy of the standard  Charpys was approximately equal to the t o t a l normalized energy of the  precracked specimens.  This points out the importance of basing  fracture control s p e c i f i c a t i o n s on the worst possible defect and on the  i n i t i a t i o n energy, rather than t o t a l energy, since the t r a n s i t i o n  temperatures and energy l e v e l s of the two are quite d i f f e r e n t .  3.5.1.3  Conclusions of Absorbed Energy Study  The AF-1 s t e e l exhibited a higher degree of anisotropy than did the AF-2 material.  Very low toughness was observed for the  AF-1 i n the orientations for which the crack followed the r o l l i n g d i r e c t i o n or was transverse to the pipe axis.  The AF-1  initiation  energies i n these directions were extremely low, even for the r e l a t i v e l y blunt notched standard Charpy specimens.  - 148 -  The i n i t i a t i o n energy of the AF-2 s t e e l was a greater portion of the t o t a l energy than was the EI for the AF-1 material, i n a l l d i r e c t i o n s , as shown by the d u c t i l i t y indices i n Table 3.6. Although the AF-1 s t e e l often showed higher t o t a l energies than AF-2, i t also exhibited lower i n i t i a t i o n energies.  It i s suggested that pipeline s p e c i f i c a t i o n s require testing and minimum toughnesses i n a l l directions of the pipe. Further tests are required to establish a s p e c i f i c a t i o n for an acceptable i n i t i a t i o n energy to better ensure protection against crack i n i t i a t i o n .  The AF-1 s t e e l had more scatter i n the data and often exhibited bimodal behaviour i n the t r a n s i t i o n region.  The energy t r a n s i t i o n curves f o r AF-1 showed a c l a s s i c upper and lower shelf connected by r e l a t i v e l y sharp t r a n s i t i o n regions. The energy of the AF-2 s t e e l decreased continuously with decreasing temperature.  The shape of the curves for the propagation and i n i t i a t i o n energy components of the t o t a l energy also showed t h i s t r a n s i t i o n behaviour.  This suggests that the i n i t i a t i o n energy may have s i g -  nificance i n terms of a t r a n s i t i o n temperature approach.  - 149 -  Both steels met the 80/50 f t - l b (108/68 J) energy c r i t e r i a at -18°C for the standard specimens i n which the crack followed the pipe axis.  However, due to the very low toughnesses of the AF-1 s t e e l  in other orientations, the adequacy of t h i s fracture control s p e c i f i cation i s questioned.  Fatigue precracked specimens greatly reduced the absorbed energies and yielded higher t r a n s i t i o n temperatures than f o r the standard Charpys.  F u l l wall Charpys generally displayed lower upper  shelf energies, although the energy values of the standard and f u l l wall specimens were similar i n the t r a n s i t i o n and lower shelf regions. The standard Charpy specimen often gave nonconservative r e s u l t s at the  pipeline s p e c i f i c a t i o n temperature of -18°C; i t i s suggested that  a f u l l wall Charpy be adopted for routine testing of p i p e l i n e s t e e l .  3.5.1.4  Drop Weight Tear Test Correlations  The pipeline industry employs two tests to ensure the toughness of the steels used i n p i p e l i n e s : the standard Charpy impact (39) test and the Battelle-Drop Weight Tear Test  The DWTT i s used to define the percent shear on the fracture surfaces of a f u l l - t h i c k n e s s test specimen.  Absorbed energy data i s  also obtained, although i t i s not employed i n pipeline s p e c i f i c a t i o n s .  - 150 -  The specimen i s provided a pressed notch, the flank angle and root radius being indentical with that of the standard Charpy specimen. in that:  The DWTT specimen d i f f e r s from that of the Charpy test 1) i t i s a f u l l plate thickness specimen to ensure maxi-  mum constraint;  2) the pressed notch provides a b r i t t l e crack  i n i t i a t i o n s i t e ; and 3) the dimensions are such that the propagation stage dominates the fracture event (76 x 305 mm).  The two nonstandard  specimens i n t h i s study have features  i n common with the DWTT specimen:  the f u l l wall Charpy specimens  have the same thickness (dimension across the notch); and the precracked Charpys require a low i n i t i a t i o n energy making the propagation stage the major component of the fracture process.  In addition, the instrumented impact test provides i n f o r mation regarding the crack propagation stage which may be associated with the percent shear.  Both pipe manufacturers provided DWTT data f o r the steels used i n this study.  Only data f o r the crack path following the  longitudinal axis of the pipe could be obtained since that i s the only specified test d i r e c t i o n .  Figures 3.45-3.48 compare the data from the DWTT with that obtained from the IIT of the f u l l wall and precracked Charpys of the  -  —I  1  1  151-  n  1  1  AF-I - Drop weight tear test vs full wall charpy crack parallel to pipe axis • % shear. DWTT{ I energy i total | initiation, propagation  {  Full wall energy  110+25  B  I  0>  „j70f 50-  5;_8o-  4000-  -O D  -15  c aj  J-100  F 5000OJ c 0)  -8-  —  # 90'^20  I  21  1-80  3000-  o  o  o  Q  c  HO  30i h5 10•0  •  JL  _L  -100  -60  T(°C)  -20  60 A -60  2000H  40 H1-40 o  1000-  20-L  20  0 *20  Figure 3.45  T  T  T  AF-1 - Drop weight tear test vs precracked charpy crack parallel to pipe axis • % shear DWTT{ ° energy • total Precracked energy A initiation o propagation  {  <4  H00  ;  ?400-T  "O  80 a?>  S300c OJ  I200-  OJ  -  (-60 -40-  l00-f-20  •a 4 0 0 0 o in o 3000-  "I 300- -80 5  2000-  -40 °  1000-  a  0  -100  -60  T(°C)  -20  Figure 3.46  Hoo  «5000-  c  • 20  &  c o  1-60°. ! 200i K  I00H h20  i AF-2 - Drop weight tear test VS| full wall charpy crack parallel to pipe axis • o Full wall energy 1 * ^ • r  D W T T - % shear total initiation propagation  80 H mat  -40  120-  o  S 60"  h-30  c c  ?80H c  CD  c o  H  §40a.  -5 60--20 o  o  40 H Ho  • e  20-  20H -60  -100  T(°C)  -20  ,20  Figure 3.47  T  T  T  AF-2 - Drop weight tear test vs precracked charpy crack parallel to pipe axis j r • total Precracked energy \ initiation A  L o propagation  I  a D W T T - % shear I  -100  -60  T(°C)  Figure 3.48  -20  .20  - 153 -  AF-1 and AF-2 s t e e l s , respectively.  D i s t i n c t s i m i l a r i t i e s exist between the IIT and the DWTT data for the AF-1 s t e e l .  For the f u l l wall specimens, at the -18°C  s p e c i f i c a t i o n temperature, where the s p e c i f i c a t i o n requires an average of 85% shear, a l l the energy components, t o t a l , i n i t i a t i o n , and propagation,were a t upper shelf conditions as was the percent shear from the DWTT (Figure 3.45). IIT  Furthermore, the shapes of the  energy t r a n s i t i o n curves and their i n i t i a l deviation from upper  shelf values closely matched that of the percent shear t r a n s i t i o n curve of the DWTT, although the DWTT data exhibited a sharper transition.  The DWTT absorbed energy data exhibited a similar energy t r a n s i t i o n to those of the f u l l wall Charpy specimens. did  However, i t  not maintain a constant upper shelf energy even though approxi-  mately 100% shear was reported for temperatures down to -40°C.  The precracked Charpy data also showed close s i m i l a r i t i e s with that of the AF-1 DWTT (Figure 3.46).  In comparing the precracked  Charpy data with the B-DWTT data, a l l three energy components of the precracked specimens were at the upper shelf or peak energy condition at  the -18°C s p e c i f i c a t i o n temperature.  Both the t o t a l energy and  the propagation energy decreased from their upper shelf values at  - 154 -  approximately the same temperature as did the DWTT percent shear, although their t r a n s i t i o n s were not as steep.  The precracked  i n i t i a t i o n energy t r a n s i t i o n was sharp but the t r a n s i t i o n  temperature  was higher than the DWTT t r a n s i t i o n by approximately 15°C.  Although more data correlations are c e r t a i n l y necessary, this work does indicate that an empirical r e l a t i o n s h i p may exist for the AF-1 s t e e l between a f u l l w a l l Charpy or a precracked f u l l wall Charpy and the Drop Weight Tear Test.  I t i s possible that the f u l l  wall Charpy test could provide the industry with a single instrumented impact test that would measure the propagation behaviour i n terms of percent shear and absorbed energy while s t i l l providing a measure of the i n i t i a t i o n energy.  Such a test would have time and cost saving  advantages for quality assurance purposes.  No simple correlations existed between the DWTT r e s u l t s and the f u l l wall and precracked Charpy data of the AF-2 3.47 - 3.48).  The AF-2  s t e e l (Figures  s t e e l energy decreased continuously with  decreasing temperature, whereas the DWTT percent shear curve did have an upper shelf which remained constant with decreasing temperature to approximately -30°C where i t exhibited a very sharp d u c t i l e - b r i t t l e transition.  - 155 -  3.5.2  Dynamic Y i e l d Strengths  Dynamic y i e l d strengths were obtained for a l l  specimens  from the load-time traces as described i n the preceeding Chapter. F i g . 3.49 shows the average values of the dynamic y i e l d strengths versus temperature for each s t e e l i n each orientation, as determined from the standard Charpy tests.  This property was very reproducible.  The y i e l d strength of: a bcc material increases with: 1.  decreasing temperature since the P e i e r l ' s stress i s  a strong inverse function of temperature; 2.  and  with increasing s t r a i n rate since the density and  v e l o c i t y of moving dislocations i s proportional to s t r a i n  rate^.  The s t r a i n rate effect can be seen i n t h i s study by examining the data at +20°C.  The dynamic y i e l d strengths of the s t e e l s , for the  orientations examined, ranged from 100-120 k s i (690 - 828 MPa), whereas the " s t a t i c " y i e l d strengths of these s t e e l s , at t h i s same temperature and for the same orientations ranged from 72-79 k s i (497 - 545 MPa).  The s t r a i n rates imparted by the IIT were approximately the same at a l l temperatures.  Thus, the increase i n dynamic y i e l d strengths  with decreasing temperature, observed on Figure 3.49, can be attributed primarily to the temperature dependence of the y i e l d stress.  At the  Dynamic  rv>  Yield  -£>  O ~i  CD  O 1  1  Strength  n  O 1  r  oo o  (MPa)  o o T  —J—  I  •fo • o  >  >  >  -n i  I  ro  a-£t> «o  • > + • • • o Q O  =  T3 Q - «  :  ro  -»  Q  = Q  —  +>B  o •  +•»•  Q-  P +> •  o>  O Q  O  a Q Q  •  I  o -  -i Q IT  • jbJKJ*  I  3  Q  CD — • ^ _? ~i 3 ° < —« ° < / > Q CD 05 < — < Q «< & 91«/)CD CD — . —CD </> o O a> O CL o -o — —^ o — —• T3  +t>a  • o»  —  r  - = ~0  CD 3  CD  a X  uQ  +> ¥ | o 3  (£>  -  0  3T  —•  a ^ w = a> o O O 3 o  = CD  3  JL  ro O  a.  CD  o  Dynamic  Yield  O O Strength (ksi)  - 9SI -  -J^ O  - 157 -  lower temperatures, the y i e l d strengths increased to 140 k s i (966 MPa) .  The dynamic y i e l d strength data was employed i n calculations used to v e r i f y dynamic fracture toughness v a l i d i t y and should be used to estimate the y i e l d strength of pipe sections subjected to dynamic loading.  3.5.3  Load-Time Behaviour  Figures 3.50 to 3.56 show the maximum and general y i e l d loads and the time to r e a l i z e the maximum load ("crack i n i t i a t i o n " ) f o r each orientation of the two steels as a function of temperature.  These data  were obtained from standard Charpy specimen t e s t s ; the load-time traces are shown i n Figures 3.5 - 3.10.  Several investigators have proposed theories which permit detailed analyses of such load-temperature diagrams i n terms of the mechanisms of deformation and fracture  behaviour21,64,105)^  Diesburg^^ has described the load/time behaviour of a c i c u l a r f e r r i t i c steels.  At low temperatures, cleavage fracture takes place at a load less than that required f o r general y i e l d i n g - l i n e a r - e l a s t i c f a i l u r e s occur.  - 158'  1  1  1  1  1  1  •  » I  1  —•  AF-1 -Load/Time standard charpy crack parallel to pipe axis • maximum load atime to maximum load ogeneral yield load • percent shear fracture  1  1  1  1  1  A F - 2 - Load/Time crack parallel to pipe axis i maximum load • time to maximum load > general yield load • percent shear fracture  . t  e • t • i  S30H  E i-  w  100-  80-j  20 80--.8  60-  ! i •I • • • •  6 0 - -.6 4040-  -.4 20-  20  -100  -60  -60  -20  TCC).  Figure 3.50  1  —i  i  Figure 3.51  1  1  1  A F - I - Load/Time crack parallel to rolling direction . maximum load o general yield load • time to maximum load • percent shear fracture  r i  i •| . . •  40 f  a •  :  .  o  X  2 3CH  E 100-  iO  80-I-.8 60  6  «  1  -tOO  -60  -20 TCC)  Figure 3.52  *  ;  9  8  8  a  8  -14 S  TO J  1  - 3  -  -10  -  •• •t' ••  -4  •  •  4  2 0 +.2  8  •  -8 40  •  1 •  -18  '  1  . maximum load o general yield load • time to maximum lood  -22  • f ' i "•I  -1  A F - 2 - Load/Time crack parallel to rolling direction  +20  1-16  -20 TCC)  1  1 '  i  i  i -60  i TCC)  1  r -20  Figure 3.53  |  1 •  1  1  AF-1 -Load/Time standard charpy crack transverse to rolling direction  ' i i r— 1 1 AF-2-Lood/Time Crack transverse to rolling direction • maximum lood ° general yield lood • time to maximum load  • maximum load • time to max. load o general yield load • percent shear fracture  50H  • t I  40  : i  • • .  o  - 4  100  + 0 |  Ho  I  80-  I *  60-  I .  40-  -100  -60  -20  .20  TCC)  -100  Figure 3.54  -80  -60  -40 TCC)  Figure 3.55  A F - t - L o a d / T i m e standard charpy crock transverse to pipe axis • maximum load •time to maximum load ogeneral yield lood percent shear fracture a  22  •  *:8 I  D o  a  Q  100--1.0 80--.8 60--  r-6  4020+2  -100  -60  TCC)  -20  Figure 3.56  • 20  - 160  D O  max  If)  i_  00  Charpy  cn i—  Energy  CD C  UJ  \  \ \ V) If) CO  \  s  \  V  i_  s  -4— CO  s  N N  *  - oy  Test  Figure 3.57  Temperature  Schematic of variation of general yield load, fracture load, and absorbed energy with temperature. Effect of notch on  Since the effective yield strength of notched steel specimens i s an inverse function of temperature, the extent of plastic deformation required to raise the tensile stress at the root of the notch to that required for cleavage fracture increases with t e m p e r a t u r e :  - 161 -  a yy  ( = K a ) = a o p y '  f o r cleavage  £ f  x  a  ^  (Eq. 3.5)  e \ M  fracture  J J  where,  :  m a x  =  maximum t e n s i l e stress below notch  = =  cleavage fracture stress (- constant) y i e l d strength, below notch rj /a* (by d e f i n i t i o n ) = p l a s t i c yy y stress concentration factor  yy  CJ£* 0"^ K op  =  m a X  The loads required for cleavage f a i l u r e therefore also increase with increasing temperature  since:  o* so, as but, K  =  , t T, m  =  f(l/T)  ,„ + K ap  r  max for a yy  f ( p l a s t i c zone size)  (Eq. 3.6)  = c*rf=  f(applied load)  In this study, i n c i d e n t a l l y , at temperatures as low as -100°C, considerable p l a s t i c deformation at the crack t i p was evidently required for the cleavage f a i l u r e s observed, since the load required for f a i l u r e (and, hence, the p l a s t i c zone size required) was found to be much less i n tests conducted at -196°C (approximately 2000 l b (8900 N) at -196°C Versus loads on the order of 3500 l b (15575 N) at -100°C).  - 162 -  In Charpy specimens, the stress concentration f a c t o r ,  K ,  reaches a maximum, ^ax (= 2.18), at some temperature less than T (point A Figure 3.57).  This temperature has been  experimentally  determined to be that at which the applied load, P, equals 0.8 P At this point, work hardening i s required to r a i s e the t e n s i l e max * * stresses below the notch to equal the cleavage stress (K a < a.). y andf Above the temperature at which the general y i e lop d load maximum (fracture) load are equal, called the b r i t t l e n e s s t r a n s i t i o n temperature, T^, the fracture mode becomes a combination of fibrous tearing (ductile) and cleavage f r a c t u r e ^ \  The fracture load necessarily increases with temperature beyond T„ due to a relaxation of the t r i a x i a l stress state (K D  ):  o"p  the s t r a i n needed to produce the work hardening required to r a i s e c r  m  a  yy  X  to a i s so large that the p l a s t i c constraint i s decreased f r  plane stress conditions are approached.  A peak i n the maximum load curve i s eventually reached at some temperature above T^. This peak temperature, T^, i s termed the d u c t i l i t y t r a n s i t i o n temperature.  I t corresponds to the point where  the s t r a i n (and thus the load) required to i n i t i a t e cleavage fracture i s so large that i t exceeds that required for the i n i t i a t i o n of fibrous tearing.  Beyond that temperature, the fracture load decreases  with increasing temperature.  - 163 -  This "knee" i n the load-temperature curve (T^_) i s associated with the temperature at which fracture i s i n i t i a t e d s o l e l y by fibrous t e a r i n g .  Other researchers have i d e n t i f i e d this point, f o r  acicular f e r r i t i c s t e e l s , as corresponding to the "C (7,107).  100" temperature  This temperature i s defined as the lowest temperature at  which the fracture surfaces of a Charpy specimen exhibit 100% shear, i . e . , the lowest temperature at which fracture i n i t i a t e s and propagates in an e n t i r e l y d u c t i l e manner, and i s often used i n p i p e l i n e s p e c i f i cations.  For nonacicular s t e e l s , fibrous cracks do i n i t i a t e at the d u c t i l i t y t r a n s i t i o n temperature, T^-, the temperature associated with the peak load on the load-temperature d i a g r a m .  However, at high  rates of s t r a i n , the stress f i e l d ahead of the advancing d u c t i l e crack can.cause large increases i n the d i s l o c a t i o n density.  This may result  i n cleavage fracture i n i t i a t i n g ahead of the advancing d u c t i l e crack t i p and a mixed mode f a i l u r e would be apparent on the fracture surface even at temperatures above T_^.  This has been observed for polygonal  f e r r i t i c structures(20,64)^ such behaviour i s manifested on the loadtime trace by a sudden drop i n the load at some point beyond that of the maximum load.  An acicular f e r r i t e s t e e l i s already highly dislocated and consequently the d i s l o c a t i o n b u i l d up ahead of an advancing crack w i l l  - 164 -  not be s i g n i f i c a n t and thus no change i n fracture mode i s observed. The fracture resistance at the onset of crack i n i t i a t i o n i s therefore the same as the fracture resistance ahead of the propagating crack. The implication i s that i f cleavage fracture does not occur early i n the fracture process i n AF s t e e l s , i t w i l l not occur during the propagation stage.  Therefore, for the AF s t e e l s , the peak load tempera-  ture, T„, i s also associated with C N  100^\ v  This suggests that the temperature at which the i n i t i a t i o n energy f i r s t deviates from i t s upper shelf value may also be associated with the  100 temperature, since that i n i t i a l decrease i n EI may  s i g n i f y the t r a n s i t i o n from fibrous to cleavage i n i t i a t i o n .  If this  i s true, then that i n i t i a t i o n energy t r a n s i t i o n temperature may  be  used to protect against cleavage f a i l u r e s .  The percent shear on the fracture surfaces of standard Charpy specimens of the steels tested i n this work was supplied by the s t e e l manufacturers. 3.52,  3.54, and  These data have also been plotted on Figures 3.50-  3.56.  Table 3.8 presents the C  100 temperatures as determined from  the fracture surfaces and from the peak i n the IIT load-temperature curves.  In addition, the i n i t i a t i o n energy t r a n s i t i o n  temperature  ( i n i t i a l deviation from upper shelf value) for the corresponding standard Charpys i s given.  - 165 -  Table 3.8 C 100 TEMPERATURES v  Peak on „ ,„ . Fracture Load-Temperature . _ Appearance Curve, T T  T  EI . . Transition m  AF-l-Crack P a r a l l e l to PA  -21°C  >+22°C  +20°C  AF-l-Crack P a r a l l e l to RD  -40°C  -40°C  -40°C  AF-l-Crack Transverse to PA  -60°C  -51°C  -60°C  AF-l-Crack Transverse to RD  -20°C  -51°C  -20°C  AF-2-Crack P a r a l l e l to PA  -30°C  .  >+22°C  0°C  -,-„_,„ DWTT . „, 85% Shear  o r a  -51°C  -30°C  The correlations between the temperatures presented i n Table 3.8 are inconsistent.  For those specimens which did not exhibit  " s p l i t t i n g " on the fracture surfaces (to be discussed i n the next Section), a l l three temperatures were i n agreement (AF-1 specimens with cracks p a r a l l e l  to r o l l i n g d i r e c t i o n and transverse to pipe a x i s ) .  However, when s p l i t t i n g was observed (see Figures 3.11 -3.16) no corr e l a t i o n s could be made.  I t i s extremely d i f f i c u l t to determine the  percent shear on the fracture surfaces of heavily c o n t r o l - r o l l e d AF steels which exhibit s p l i t t i n g by direct-examination; the r e l i a b i l i t y  - 166 -  of such measurements i s i n question T  N  . However, the peak load,  (C 100) on the load-temperature curves (Figures 3.50 - 3.56) v  was not always well defined either.  More correlating data i s  required.  In addition, i t has been suggested that the B-DWTT 85% shear temperature should l i e between the peak temperature, T^, and the b r i t t l e n e s s t r a n s i t i o n temperature, IL (P = P„„) D max GY long as  . As  i s less than the pipeline s p e c i f i c a t i o n temperature for  85% shear (-18%C), a Drop Weight Tear Test may not be required should IIT be employed to evaluate p i p e l i n e materials.  However, the v a l i d i t y  of this suggestion could not be established, as Table 3.8 indicates.  3.5.4  Fractography  Although no fractography study was made i n this thesis, several unique c h a r a c t e r i s t i c s of the fracture surfaces were noted.  The fracture surfaces of both the acicular f e r r i t e steels exhibited i r r e g u l a r i t i e s known as " s p l i t s " .  These appear as sharp,  deep, quasi-cleavage fractures normal to the fracture face and p a r a l l e l to the plane of the plate (see Figures 3.11 - 3.16).  Splitting i s  commonly observed on the fracture surfaces of f u l l - s c a l e tests of pipe made of AF steels and on impact specimens tested i n the upper  - 167 -  shelf and transition temperature range.  The effect of splitting on  the absorbed energy i s not fully established ^'"^^^ HO) ^  Some workers have claimed that rolling AF steels below the Ar^ temperature i s a necessary prerequisite for s p l i t t i n g ^ ^ ^ ^ . However, the AF-1 steel was finish rolled at approximately 800°C, whereas AF-2 was finish rolled at about 760°C, both temperatures (86 lf)8)  being above the 700°C Ar.. temperature  '  . Others have suggested  that even though HSLA AF steels are rolled above Ar^, at very low finish rolling temperatures where essentially no y-recrystallization can occur, the elongation of the y-grains i s severe. The mechanical anisotropy thereby introduced may be the cause of splitting  .  This anisotropy is increased with decreased rolling temperatures. The presence of Nb(C,N), which retards the y-recrystallization, effects the degree of that splitting 0.063 w/o Nb; AF-2 contained 0.05  110)^  steel contained  w/o.  Killed steels have been said to have a greater tendency to split than do semi-killed steels^"^^ . However, this was not observed in this study:  the semi-killed AF-1 steel (0.03 w/o Si) had a slightly  greater tendency to split than did the killed AF-2 material (0.26 w/o Si, 0.045 w/o A l ) .  - 168 -  Diesburg  suggests that no d i s t i n c t upper shelf energy  plateau exists i n that temperature range where s p l i t t i n g occurs, even though the fracture remains 100% d u c t i l e .  Instead, the d u c t i l e  fracture energy decreases with decreasing temperature and a sloping shelf i s observed. the AF-2 s t e e l .  Indeed, sloping energy curves were observed for  However, the AF-1 s t e e l , which had the s l i g h t l y  greater tendency to s p l i t , had a d i s t i n c t upper shelf plateau (compare Figures 3.23 - 3.24 and 3.15 - 3.16).  The s p l i t t i n g phenomena evidently  i s a r e s u l t of a complex  interaction of composition and processing variables.  A complete  understanding of the causes and effects of s p l i t t i n g i n AF pipeline steels i s s t i l l to be resolved.  3.6  Strain Age Study  Pipeline s t e e l s p e c i f i c a t i o n s c a l l f o r high strengths and toughnesses i n the a s - r o l l e d and the as-formed product.  Pipe and  f i t t i n g s are subjected to p l a s t i c s t r a i n i n g , after s p e c i f i c a t i o n testing, p a r t i c u l a r l y during f i e l d bending.  Subsequent g i r t h welding  then provides the potential f o r s t r a i n aging.  The p o t e n t i a l f o r  s t r a i n aging also exists i n areas adjacent to seam welds since the i n i t i a l cold pipe forming operations impart p r i o r s t r a i n to the s t e e l .  - 169 -  An IIT study was conducted  to determine the effects of  straining and the subsequent aging on the dynamic properties of the AF-1 and AF-2 pipeline  steels.  This study was conducted  i n two parts.  F i r s t , a characteri-  zation of the effects of straining and s t r a i n aging of the two steels was made.  Second, f o r the AF-1 s t e e l , impact  specimens taken from  near the seam weld were tested to determine i f s t r a i n aging had occurred within the pipe.  3.6.1  Effects of Straining and Strain Aging  The effects of straining and subsequent aging on the IIT properties were examined i n the AF-1 s t e e l i n two orientations: 1)  f o r cracks running p a r a l l e l to the pipe axis;  running p a r a l l e l to the r o l l i n g d i r e c t i o n .  and  2)  f o r cracks  The AF-2 s t e e l , which had  shown much less anisotropy, was tested only with cracks running p a r a l l e l to the pipe axis.  To introduce a constant amount of s t r a i n into the test materials, large t e n s i l e bars were cut from the pipe.  The reduced section of these  bars was approximately 28 cm long and at least 55 mm wide to permit the cutting of a standard Charpy specimen.  The t e n s i l e bars were cut from  the pipe so that the straining d i r e c t i o n was p a r a l l e l to the Charpy crack  - 170 -  path.  The grip areas were pressed f l a t , although the gauge length  retained the original pipe curvature.  Gauge marks were carefully  scribed every 13 mm along the reduced section to allow determination of the actual strain after testing.  The bars were plastically strained 3-5% i n uniaxial tension on a 100,000 kg tensile machine. This strain level approximates the combined maximum strain involved i n fabrication and installation of pipe^^'"^"  i t should be emphasized, however, that this  operation provided strain i n excess of that already introduced due to the pipe forming operations.  After straining, the bars were stored i n dry ice until Charpy specimens could be cut from the gauge section. Upon cutting the Charpy blanks from the strained bars, half the specimens were placed in stainless steel bags and aged in an air furnace for one hour at 275°C. This time and temperature was chosen to optimize the expected effects of strain aging  . Charpy specimens were  then machined from the blanks, notched through the pipe thickness, and stored in dry ice until tested.  In a l l cases, control specimens taken from the pipe adjacent to the position from which the tensile bar was cut were f i r s t tested to establish the properties of the cold formed pipe. An  - 171 -  instrumented impact test temperature equivalent to the transition temperature of the as received pipe was chosen; i f straining and strain aging produced any measurable change in dynamic properties, the magnitude of those changes would therefore be expected to be large.  The results of these tests are presented in Table 3.9. A measure of the ductility of the specimens was made by:  1) cal-  culating the ductility index, DI; and 2) taking the difference between the time to maximum load, t .,,, and the time to general MAX w  yield load (elastic limit), t  r v  , from the IIT load-time traces.  This time should be directly related to the amount of strain occurring prior to plastic instability. The shift in transition temperature was estimated by assuming that the strain and strain aging did not change the shape of the energy transition curves. A measured energy could then be associated with a specific temperature on the control specimen energy curve; the difference between this temperature and the test -  temperature was considered to be the shift i n the transition temperature.  In general, the data show that the toughness and ductility of the AF-1 steel was reduced by straining and strain aging as i t s  Table 3.9 STRAIN AGE STUDY  Material  T (°C)  ET  EI  EP DI  G  A  yd  (ft - l b )  MAX ~ GY  t  t  TT S h i f t  (ms)  (°C)  (ksi) AF-2-TR-CN AF-2-TR-S AF-2-TR-SA  -60° -60° -60°  35.0 46.4 64.2  10.0 13.5 16.6  24.2 32.9 47.7  2.4 2.4 2.9  127.1 137.3 145.7  .247 .309 .353  AF-1-TR-CN AF-1-TR-S AF-1 TR-SA  -60° -60° -60°  10.9 11.1 . 8.4  1.7 2.1 1.3  9.2 9.0 7.1  5.4 4.3 5.5  >120.1 109.8 >120.7  .117 .055 .008  AF-1-TP-CN AF-1-TP-S AF-1-TP-SA  -30° -30° -30°  94.4 89.1 81.0  18.3 19.0 17.3  76.1 70.1 63.7  4.2 3.7 3.7  122.2 135.1 139.0  .434 .336 .333  AF-1-TP-CN AF-1-TP-S AF-1-TP-SA  -40° -40° -40°  84.5 76.9 65.7  14.4 15.0 13.6  70.1 61.9 52.1  4.9 4.1 3.8  129.0 132.9 >138.2  .277 .264 .210  TR: TP: CN: S: SA:  Crack p a r a l l e l to r o l l i n g d i r e c t i o n Crack p a r a l l e l to pipe axis Control specimens from pipe 3-5% p l a s t i c s t r a i n strained and 1 hour at 275°C  A l l values are averages of several  tests.  -30° -91° _ 0° + 4 _ "  + 5° +12° _ + 7° +14°  94.2 92.7 95.6 91.2 93.8 94.0 92.4 94.5 96.5 92.4 94.5 96.5  - 173 -  strength and hardness increased.  It i s significant to note that  the strength, toughness, and ductility of the AF-2 steel a l l i n creased upon straining and subsequent aging.  Straining of the AF-1 steel produced the classic effects expected of a cold worked material:  hardness and yield strength  increased and the ductility decreased (as indicated by the observed decrease in t_^^ - t ^ and decrease in DI) as the dislocation density i n c r e a s e d .  Consequently, the impact resistance was  reduced as manifested by a decrease in the absorbed energy and a shift in the transition temperature to a higher temperature.  Subsequent strain aging of the AF-1 steel produced a further increase in strength and hardness, and a decrease in the ductility and the impact resistance.  Others have observed a  similar increase i n the transition temperature in semi-killed AF, steels, but without the loss of absorbed  e n e r g y .  It was not surprising that the AF-1 steel showed these effects.  This steel was semi-killed (0.03 w/o Si), contained few  nitride formers other than Nb, and thus, no doubt contained a high free nitrogen c o n t e n t ,  i t i s well established that strain  aging results from free i n t e r s t i t i a l s , particularly carbon and nitrogen, diffusing to dislocations and locking them; acicular  - 174 -  • • , (111,113,115-116) ferrite steels are no exception  One significant observation from the AF-1 data, however, was that the decrease in total absorbed energy resulting from straining and strain' aging was due primarily to a decrease in the crack propagation energy.  The initiation energy was only marginally  affected by either straining or strain aging.  This observation is  significant to the pipeline industry in particular.  Although the  total toughness of the AF-1 steel i s adversely affected by straining (as could occur from field bending or frost heave) and strain aging (as could result from welding prestrained pipe), the initiation energy of the steel i s not reduced.  Therefore, though the AF-1  steel i s susceptible to strain age embrittlement, i t does not increase the potential for crack initiation.  One possible explanation for this important observation is that at the temperatures at which the AF-1 steel was tested (-30° and -40°C), the primary fracture mode for cracking parallel to the pipe axis was ductile.  Figure 3.50 indicates that s i g n i f i -  cant plastic flow was associated with the fracture event.  The  initiation of ductile failure i s known to occur by void formation at inclusions or precipitates, either by interface separation or particle cracking ^'^""^ . The ductile crack propagates through the matrix as these voids coalesce between particles.  The initiation  - 175 -  step in the fibrous crack process i s therefore a function of the strength of the inclusions and/or the strength of the inclusion/ matrix interface.  The propagation stage of ductile fracture i s  dependent upon the matrix properties (strength, ductility). ing and  Strain-  strain aging of the AF-1 steel was shown to increase the  yield strength and decrease the ductility, the combination of which results in lower toughness.  It i s suggested that the matrix  properties are primarly affected by these factors; the properties of the inclusions remaining relatively unaffected. This would explain why the crack initiation energy was unaffected, whereas the crack propagation energy was reduced with straining and strain aging.  The AF-2 steel gave much different results (Table 3.9). Not only did the strength increase with straining and aging, but so did the ductility.  As this i s a highly killed steel containing  several nitride formers (Nb, A l , T i , and La), a minimum concentration of free interstitials should be available to lock dislocations and ,  __  „  '  ,  ,(113-114)  therefore no strain aging effects would be expected  . How-  ever, the magnitude of the improvement in impact resistance with straining and aging was significant, the total energy increasing from 35 ft-lb to 64 ft-lb (47 - 87 J) with a 91°C improvement in transition temperature.  - 176 -  The AF-2  steel was probably underaged in the as rolled  (79) condition ^^"^.  and aging may have enhanced the Nb(C,N) precipitation  This increased precipitation could increase the ductility  by retarding the dislocation motion and thus increasing the work hardening rate.  The prior strain could have optimized the effective  precipitate size as dislocation loops formed around the fine underaged precipitates.  This factor, plus the increased dislocation  density, may have caused a significant increase in the work hardening rate, and hence, forced necking to occur at a higher plastic strain thereby increasing the observed ductilities. In summary, the behaviour of the AF-1 straining and aging was markedly different.  and AF-2  The AF-1  steels with  steel exhibited  the classic effects of cold work and strain aging, while the ductility and toughness properties of the AF-2  steel were significantly improved  by straining and aging.  3.6.2  Strain Aged Sites in AF-1 Pipe  Having established  that the AF-1  steel was susceptible  to  strain aging and recognizing the fact that the pipe forming operation imparts a degree of strain to the material (approximately 2-3%)(H-1,118)^ a series of tests were performed to assess the presence of strain aged  - 177 -  embrittlement in the AF-1 pipe. The obvious location for a strain aged structure would be adjacent to the seam weld; this material would have been strained during the pipe forming operation and subsequently aged from the heat associated with the seam welding process.  Although the properties of the weld bead and the heat-  affected-zone (HAZ) have been documented(^,119)^  n Q  polished  work has been done to identify the strain aged site outside the HAZ in welded AF pipe, though the need for such a study has been . .(92,95) recognized  Using welding parameters supplied by the pipe manufacturer, the peak temperatures versus distance from the HAZ were established (the pertinent calculations are included in Appendix E). At a position approximately 15 mm (0.6-in) from the weld fusion boundary a peak temperature of approximately 337°C was realized.  Using cooling  rate equations, i t was further determined that this region of the pipe should experience temperatures optimum for strain aging for approximately 30 seconds (337° to 2 8 5 ° C ) d u r i n g the two-pass spiral welding process.  To check the r e l i a b i l i t y of a reported activation energy equation for strain aging HSLA steels from a region well away from the seam weld.  Charpy blanks were cut Using the relationship  between time and temperature, several equivalent strain aging times  - 178 -  and temperatures were established.  The Charpy blanks were aged at  s p e c i f i c times and temperatures chosen to simulate the s t r a i n aging conditions predicted for the p o s i t i o n 15 mm  from the edge of the  weld fusion boundary.  It should be noted that Rashid's equation:  log t / t 1  2  =  7500 [ 1 / ^ - 1/T ] 2  T  1  < T  2  (Eq. 3.7)  indicates that the HSLA steels w i l l not s t r a i n age at room temperatures.*  After aging the blanks, standard Charpy specimens were prepared and notched so that the crack would propagate p a r a l l e l to the pipe axis.  Charpy control specimens with no aging treatment, only  pipe forming s t r a i n , were also prepared.  In addition, Charpy specimens were also cut so that the structure below the notch was approximately 15 mm from the weld fusion boundary of the seam weld and had experienced the 337°C - 30 second  *  An optimum s t r a i n age condition i s 1 hour at 275°C = 548°K; the equivalent s t r a i n age time at room temperature (298°K) i s : log t^l t,  =  7500[l/298 - 1/548]  =  3 x 10^" hours !  - 179 -  aging treatment.  In actual f a c t , the structure below the notch had  experienced a temperature gradient ranging from 285°C d i r e c t l y below the notch (19 mm from the weld fusion boundary) to 421°C at the opposite face (11 mm from the weld fusion boundary). are  given i n Appendix E.  Calculations  It was recognized that these Charpy  specimens taken from near the seam weld a c t u a l l y experienced a range of aging conditions which could only be approximated by the a r t i f i c a l l y aged specimens. 1.  Nevertheless, should:  the a r t i f i c a l l y aged Charpys y i e l d similar results  upon impact testing - these r e s u l t s having d i f f e r e n t values from the  IIT data of the nonaged control samples; and, 2.  should the r e s u l t s compare reasonably w e l l with those  obtained from the "seam weld" specimens, then i t could be concluded that the area near the seam weld had indeed experienced s t r a i n aging and that Equation 3.7 i s applicable to the acicular f e r r i t e steels studied.  Instrumented impact tests were conducted at -30°C, this being a temperature near the upper shelf t r a n s i t i o n region of the control specimens.  Table 3.10 gives the r e s u l t s of this study.  The data shows: 1.  that those samples a r i t i c a l l y aged showed d e f i n i t e  signs of s t r a i n aging - their y i e l d strengths and hardness values  - 180 -  increased and their absorbed energies decreased r e l a t i v e to the values observed for the control specimens.  A simple aging or  stress r e l i e v i n g treatment, i . e . , one f o r which no p r i o r s t r a i n had been imparted to the specimens, should not have reduced the impact resistance(85,111)^  T  hi  s  indicates that the pipe forming  strains were s u f f i c i e n t to cause s t r a i n aging. 2. the  As i n the previous phase of the s t r a i n age study, only  propagation energy was affected by s t r a i n aging; the i n i t i a t i o n  energies reported i n Table 3.10 are constant f o r each aging condition. 3.  Equation 3.7 was  equivalent time/temperature  shown to be accurate i n determining  s t r a i n aging conditions. The fact that  the specimens aged 1 minute at 337°C gave a somewhat higher y i e l d strength can be r a t i o n a l i z e d i n that Rashid's equation actually predicts that 1 minute at 316°C or 1/2 minute at 330°C would give the equivalent s t r a i n age e f f e c t s as the other three aging conditions. 4.  The samples removed from adjacent to the seam weld  showed a marked decrease i n absorbed energy and an increase i n y i e l d strength and hardness; both are d e f i n i t e indications that the region had been s t r a i n aged.  In f a c t , the properties near the seam weld  indicated that the area had been s t r a i n aged to a greater extent than that estimated from the peak temperature and cooling rate calculations. . However, those calculations were considered conservative.  - 181 -  Table 3.10 STRAIN AGE SITES IN STEEL AF-1  ET Age Treatment  EI  EP  G  R  A  B  yd (ksi)  (ft - lb)  As Formed Pipe  94.4  18.3  76.1  92.4  122.2  1 hr @ 244°C  85.5  18.2  67.2  93.3  127.9  15 min.@ 266°C  84.5  18.1  66.4  93.2  127.6  5 min. @ 289°C  84.0  19.0  65.0  95.0  129.2  1 min. @ 337°C  83.4  19.2  64.2  93.9  134.0  15 mm from seam weld  78.5  18.1  60.4  97.3  134.4  Grain Size  A l l notches parallel to pipe axis A l l tests at -30°C A l l values are averages of many tests.  ASTM  12.7  12.4  - 182 -  A microstructural examination of the specimens taken from adjacent  to the seam weld of the pipe indicated that no apparent  structural modification could be associated with the observed change i n properties, the microstructure and grain s i z e being i d e n t i c a l to that found i n the control speciemens.  It i s important to note that the p o t e n t i a l for crack t i a t i o n (EI) was  ini-  not increased by s t r a i n aging, even near the weld.  Thus, although s i t e s do_ exist which have been s t r a i n aged i n the AF-1  pipe, the e f f e c t of the s t r a i n aging i s r e l a t i v e l y small, the  t o t a l energy s t i l l meeting the toughness s p e c i f i c a t i o n s (78.5 (106 J) at  -30°C).  ft-lb  - 183 -  4. DYNAMIC FRACTURE TOUGHNESS  4.1  Introduction  A fundamental principle of fracture mechanics i s that the stress f i e l d ahead of a crack can be characterized by a single parameter, K, the stress intensity factor.  The magnitude of K i s  directly related to the crack size: K a o(ir a)  2  where,  (Eq. 4.1)  a = applied stress a  = sharp flaw size  thus providing the design engineer with a means of relating the defect size and allowable stress.  For:a particular combination of stress and defect size, the stress intensity factor reaches a c r i t i c a l value, K^, where unstable crack growth occurs.  This c r i t i c a l value i s described as  the "fracture toughness" and i s a basic property of a material.  This relationship i s significant i n that i t allows considerable f l e x i b i l i t y i n design for fracture control.  Trade-  offs in material selection (K^), design stress ( c ) , and allowable  - 184 -  flaw sizes (a), as determined by NDT flaw detection c a p a b i l i t y , can be made i n a quantitative manner.  The c r i t i c a l stress i n t e n s i t y factor decreases to a minimum value as the thickness of a plate increases  to a condition  of maximum constraint where t r i a x i a l stresses exist at the t i p of the notch.  A condition of plane s t r a i n then exists since p l a s t i c defor-  mation i n the d i r e c t i o n p a r a l l e l to the crack front (through-thickness) i s r e s t r i c t e d .  When t e n s i l e stresses are applied across the  notch, fracture occurs by the crack surfaces being displaced normal to themselves (Mode I ) .  This minimum plane s t r a i n value for Mode I  type fracture i s designated K^.  Most s t r u c t u r a l steels exhibit such a high fracture toughness that for the available s t r u c t u r a l thicknesses, the K value cannot be measured. calculate K (^0) ^ Ic T  g  The l i n e a r - e l a s t i c analysis used to  i } . i d a t e d when i n s u f f i c i e n t specimen nva  thickness results i n general y i e l d i n g and the formation of large p l a s t i c zones ahead of the crack t i p .  E l a s t i c - p l a s t i c analyses  have extended the fracture mechanics concepts to account f o r such behaviour  The value of K i s determined f o r quasi-static conic T  n  d i t i o n s , that i s , at s t r a i n rates of approximately 10 "Vs; t h i s  - 185 -  i s equivalent to a stress i n t e n s i t y rate, K, of approximately 10 k s i - i n / s , where K Is the r a t i o of K j 2  fracture.  c  to the time required for  For s t r a i n - r a t e sensitive materials, increasing the  loading rate to that corresponding to an impact t e s t , i . e . , approxi,5 mately 10/s(K - 10 k s i - i n / s ) , causes a decrease i n the plane s t r a i n (122-123) fracture toughness to a minimum value  .  This value i s  c a l l e d the dynamic fracture toughness, K-.^., and i s generally the most conservative value of a material's fracture toughness at a given temperature. For s t r u c t u r a l s t e e l s , at temperatures where cleavage f a i l u r e s occur, the s t a t i c and dynamic values of the plane s t r a i n fracture toughness are approximately e q u i v a l e n t 1 2 5 ) ^  The s t r a i n rate s e n s i t i v i t y of fracture toughness i s explained by the increase i n the y i e l d strength with increasing loading rate (as with decreasing temperature). Increases i n the y i e l d stress imply a higher l e v e l of t e n s i l e stress i n the p l a s t i c zone ahead of the crack and hence both a higher density of voids and easier void coalesence.  The energy required for the d u c t i l e  crack process i s thereby lowered.  Consequently,with  s t r a i n rate, the fracture toughness d e c r e a s e s ^ \  increasing  For cleavage  fracture i n i t i a t i o n , however, since the cleavage strength i s r e l a t i v e l y i n s e n s i t i v e to changes i n the s t r a i n rate or temperature,  - 186 -  K ~ K lc - Id  (  K  1  2  6  )  K  Interestingly, i t has been established that for some high strength titanium a l l o y s and high strength s t e e l s (a > 145 k s i ) , increasing the s t r a i n rate increases both the y i e l d strength and the fracture toughness ( i . e . ,  > K ). I c  Although the reasons are not  f u l l y understood, the e f f e c t i s thought to be due to adiabatic heating i n front of the crack t i p ; the l o c a l i z e d heating increases the energy required to deform the associated p l a s t i c zone by causing a r e l a t i v e decrease i n the t e n s i l e properties ^^"^.  The dynamic fracture toughness i s useful for design purposes when:  1)  conservative estimates of the fracture toughness  are desired - as i s the case i n the nuclear power industry, or  2)  dynamic loading conditions are expected i n service.  Since large s i z e specimens may be required to achieve plane s t r a i n conditions, the cost of machining the specimens and the i n t r i c a t e test procedures required have kept fracture toughness testing from being used i n other than laboratory settings.  Instrumented impact t e s t i n g using precracked Charpy specimens i s currently receiving considerable attention as a r e l a t i v e l y simple means of generating v a l i d fracture toughness values  - 187 -  from fracture occurring i n both the l i n e a r - e l a s t i c and e l a s t i c p l a s t i c regimes.  IIT has the advantage of being a rapid, inex-  pensive technique that employs small, e a s i l y machined test specimens. The impending standardization of IIT and the use of precracked Charpy specimens  (15,51-52,127)  should encourage a wider acceptance  of t h i s test approach f o r obtaining fracture mechanics data.  4.2  The Calculation of Fracture Toughness Parameters from IIT Data  4.2.1  L i n e a r - E l a s t i c Fractures  For precracked  Charpy specimens i n which the fracture  i n i t i a t e s p r i o r to general y i e l d i n g , i . e . , when the maximum load, MAX' *  P  S  -*-  ess t  n  a  n  t  n  e  general y i e l d load, FQ » Y  a  s  i  the stress i n t e n s i t y factor can be calculated from  K. Id  where,  6 Y Ma  n  Figure 2.4b,  (128) :  2  B  specimen thickness  W  specimen width  a  crack length (notch plus precrack)  M  applied moment at fracture  Y  f(L/¥, a/w)  (Eq. 4.2)  - 188 -  For three-point bend specimens (eg., Charpy specimens):  M  where,  MAX  =  L  P  =  L  M  (E<  **  4 < 3 )  specimen support span  and:  Y  =  1.93 - 3.07(a/w) + 14.53(a/w) - 25.11(a/w) 2  +  3  25.8(a/w)  (Eq. 4.4)  4  For the calculated value of K^^ to be "valid", that i s , for the value to represent plane strain conditions, the ASTM E 399 standard^^0) stipulates that:  B, a, (W - a) ^ 2.5(K la )  2  T  (Eq. 4.5)  I c y  where,  = yield strength at the test temperature and loading rate (129)  However, Tetelman, et a l  , have indicated that the central 90  percent of the Charpy specimen thickness i s i n plane strain so long as: B  >>  1.6(K la ) Ic y T  2  (Eq. 4.6)  - 189 -  However, both Equations 4.5 and 4.6 were established for statically obtained fracture toughness parameters. from which K  The expressions  is derived may not be s t r i c t l y valid for dynamic (72)  loading conditions  Ireland has r e p o r t e d t h a t the only validity requirement for K , in the tentative ASTM standard for instrumented impact Id testing shall be that fracture occur before general yielding, i.e.:  P MAX  <  P  (Eq. 4.7) GY  The size requirements of Equations 4.5 and 4.6 have been reported . . i j,, (27,35,130) to be too conservative for dynamic loading conditions (Of course, certain other criteria must be met in precracking the Charpy s p e c i m e n s ) .  The criterion outlined in Equation 4.7 was employed in this work in assessing the validity of the K-_^ measurememts.  4.2.2  Elastic-Plastic Fractures  4.2.2.1  J-Integral  The J-Integral approach to general yielding fracture  - 190 -  /-|Q*1  mechanics  1 Q/ \  characterizes the stress-strain conditions exis-  ting near the crack tip in an elastic-plastic solid.  The J-Integral  is calculated by taking the load-displacement records from the same material for two different crack lengths and determining the change (9 132) in potential energy for an incremental crack length change  '  ,  i.e.:  J  -  ^ >  <E . 4.8) q  The J-Integral is also described as being the crack extension force per unit length of crack front, or the general fracture energy release rate per unit area.  For elastic conditions, J = G, (72)  the elastic strain energy release rate  The ASTM i s presently preparing a proposed standard for J-Integral testing which is to supersede the plane strain fracture (72) toughness standard, ASTM E 399  . The main advantage i n employing  elastic-plastic toughness parameters for design i s that valid test specimens can be one tenth to one hundredth the size of those required by E 399.  Such a test shall have application for both linear-elastic  and elastic-plastic failure conditions. In practice, the calculation of the c r i t i c a l J-Integral value i s dependent upon making the appropriate measurements to the (133) point of crack initiation and calculating J  from the expression  - 191 -  J  where,  lc  A  BCW  =  =  2A - a)  (Ec  *' ' 4  9 )  area under the load-deflection curve to the point of crack initiation  Several techniques are used to determine the point of crack i n i t i a tion i n slow bend J-Integral t e s t s , including ultrasonics and elect r i c a l resistance.The common, though tedious method i s to load several specimens to various deflections and to determine the extent of crack growth i n each specimen by heat t i n t i n g .  A plot  of the applied J-Integral versus the extent of cracking i s made and the data i s extrapolated to give  J  I  c  ^  7  2  back to the crack i n i t i a t i o n condition  \  For IIT, crack i n i t i a t i o n i s assumed to occur at the point of maximum load.  Thus, the c r i t i c a l dynamic J-Integral, J-j-^* i s  determined by substituting the i n i t i a t i o n energy, EI, for A i n Equation 4.9. From t h i s expression,, the e l a s t i c - p l a s t i c stress i n t e n s i t y factor can then be calculated:  K  where,  E  =  =  (EJ^)^  e l a s t i c modulus  (Eq. 4.10)  - 192 -  This calculation i s necessarily nonconservative when crack initiation precedes the point of maximum load. Elaborate test techniques for more accurately determining the point of i n i t i a l crack extension i n a Charpy specimen during instrumented impact testing are being developed ^ ^ ^ . Such procedures have been shown to provide a better 3  estimate of the J-Integral when considerable yielding transpires prior to maximum load. However, even when assuming that  in-  dicates the point of crack initiation, good correlations between the J-Integral stress intensity factor determined from IIT and valid values for steels have been obtained^"^.  The ASTM Committee concerned with standardizing this (127) technique for IIT has suggested that for fibrous crack initiation, (27) the specimen thickness requirement be : B  >  — a  where,  a  =  (Eq.4.11)  F  flow stress  = average of yield stress  F  and ultimate stress - a +10 ksi y and, for cleavage initiation:  B  >  50 J / o I d  F  (Eq.4.12)  - 193 -  4.2.2.2  Crack Opening Displacement  This method of determining the elastic-plastic fracture toughness relies on a knowledge of the strains at the crack tip at point of f a i l u r e i  n  applying instrumented impact test infor-  mation to determine a COD value i t i s again assumed that crack i n i tiation occurs at the point of maximum load.  The specimen deflection  at the point of crack initiation, d^, can then be calculated as described in Section 2.3.2.2.  Once the initiation deflection has been established, a calculation of the crack opening displacement, COD, may be made, which for the Charpy-geometry can be expressed as^^ :  COD  where,  r  =  =  2.54[r(W - a)]d  (Eq. 4.13)  ±  rotational ratio  There i s s t i l l controversy over the computational methods (72) for establishing a COD value  . In particular, the value to be  assigned to the rotational ratio, which i s a measure of the hinge distance below the crack tip, remains to be settled. Values for (23 48 105 137-139) r ranging from 0.20 to 0.50 have been cited for V"»'» »» 0  L U J  J  I I T  In this work, for sake of computational consistency, a commonly  J /  - 194 -  accepted value of 0.33  Once the COD release rate, G^, may  G  selected for the r o t a t i o n a l r a t i o .  has been calculated, the s t r a i n energy be determined  =  d  was  :  (C0D)(a )  (Eq.  4.14)  (G E)  (Eq.  4.15)  yd  (9) from which  : K  C 0 D  =  z  d  It should be noted that when fracture occurs p r i o r to general y i e l d , the dynamic y i e l d strength cannot be determined from IIT.  Therefore, the COD  stress i n t e n s i t y factor can be calculated  only for e l a s t i c - p l a s t i c fractures.  Vitek and Chell  report that post y i e l d fracture tough-  ness calculations for f a s t fracture are best based upon the  J-Integral  c r i t e r i o n , whereas for time dependent f a i l u r e (slow crack growth) the COD  method i s more suitable.  4.2.2.3  Equivalent  The equivalent  Energy Method  energy method for c a l c u l a t i n g the stress  i n t e n s i t y factor assumes that had a s u f f i c i e n t l y large sample been  - 195 -  employed, fracture would have occurred p r i o r to general y i e l d , under plane s t r a i n conditions, at any energy corresponding to that of the i n i t i a t i o n energy of the smaller test s p e c i m e n ^ ^ . 42  The load at  which the larger specimen would have fractured i s established by extrapolating the l i n e a r slope of the e l a s t i c region of the smaller specimen's load-time curve so that the area under this  extrapolated  curve equals that corresponding to the energy to maximum load f o r the small specimen.  This "equivalent energy" load i s then used to  calculate the stress i n t e n s i t y factor, K-.--, as per the equation f o r l i n e a r - e l a s t i c fracture (Equation 4.2).  Although Robinson and Tetelman^"'" ^ have c r i t i c i z e d this 4  method as being far too nonconservative, others claim that the values obtained  for K^-j from IIT correlate better with large specimen K__  c  values than do the IIT K  J  4.2.2.4  parameters  (14)  C r i t i c a l Crack Sizes  Many functions r e l a t i n g the fracture toughness, applied stress, and flaw size have been determined for a v a r i e t y of speci_ (144-146) men configurations  For p i p e l i n e applications, a through-wall flaw i s considered to be the most severe defect.  The c r i t i c a l length of such  - 196 -  a flaw, i . e . , the length at which unstable crack propagation occurs, can be determined from:  K  where,  =  Ic  =  a  f ( ™ f  (Eq-  2  4.16)  applied stress to i n i t i a t e unstable crack growth  2a  4.3  =  sharp flaw length  Dynamic Fracture Toughness of Pipeline Steels  The fracture toughness parameters described i n the previous Section were calculated using data obtained from the precracked Charpy IIT load-time traces.  These data are graphically presented i n Figures  4.1 to 4.6 f o r each orientation of the AF-1 and AF-2 steels.  The  J-Integral v a l i d i t y requirements were applied to the data and those values of Kj not meeting these c r i t e r i a have been separated with a dashed l i n e i n each Figure.  It can be seen that for both steels there was a sharp temperature  t r a n s i t i o n i n the J-Integral dynamic fracture toughness  i n a l l test directions except i n the r o l l i n g d i r e c t i o n ; the t r a n s i t i o n was much sharper than the corresponding energy t r a n s i t i o n s f o r the same precracked specimens.  At the bottom shelf of these t r a n s i t i o n  curves, K , data correlated well with the K  T  values.  The COD stress  - 197 -  1  1  1  1  1  1  r  AF-1 - Dynamic fracture toughness crack parallel to pipe axis  • Kl-lniigrol o  K„  a  Kcoo  •  1 Kj volid  •  *  I :  -100  J_  -60  T(°C)  -20  •20  Figure 4.1 1  1  1  1  1  1  1  1  1  AF-2-Dynamic fracture toughness crack parallel to pipe axis • K  -  0  °  -  a  K  ld  •  K oo C  -  Kj  valid  *  .  • •  •  • !  -  •  o  g  8 1  1  -1010  1  i  o .8 1  -60  Figure 4.2  B  8  -  8  1  TCC)  1  -20  1  1  •20  - 198 -  80 £  b  80  2 •  e  o  60  •  g  c .c o>  •&60 3  3 o  o  40  z  D  2 40 AF-1-Dynamic fracture toughness crack parallel to rolling direction  o E o c  E o  • K.J-lntegrol 20  20  o K,  1  0  1  1  1  1  -20  60  -100  1  1 20  1  I  T(°C)  Figure 4 . 3 1  1  1  1  1  AF-2 - Dynamic fracture toughness crack parallel to rolling direction  "E  s. 5  •  K  0  o Kid C  *  K  c -C  coo  cn 3  3 O  o  *  o  200  6  -  o c >« Q  •:  100  0  • •  • 1 • « a K j valid!  I i  * i -100  • 1  i  • 8  Figure 4.4  E o  HI00 m  1 -20  T(°C)  o  a  i  -60  200 o  *20  -  199 -  T  T  T  AF-1 - Dynamic fracture toughness crack transverseto rolling direction  •  *Vi(rtegrol  O l d K  0> c  300  COD  x  f300|  c -C a> 3 O  3»  4> w 3 O O  •"'1  Kj valid j — — "  200 200h  o o o  £ o c >>  1  o c  1100  100 r•  •  2o  t  s « § 0«  X -60  -100  Q  -20  *20  Jo  TCC)  Figure  4.5  AF-2 - Dynamic fracture toughness crack transverse to rolling direction D 0_  • ,J-lntegral K  o K Id  in  a> c  * K COD  1300 » a> c  #300f  o>  o  o  Kj valid id  •fc 200f  J200 J  -  o o  u E o  e D C  iioo Q  100  8• • -100  X  -60  T(°C) Figure  4.6  X -20  •20  Jo  - 200 -  i n t e n s i t y factors lay along the upper shelf.  This fracture  toughness  t r a n s i t i o n coincided with the onset of fracture p r i o r to general y i e l d i n g ; i t i s assumed that plane s t r a i n conditions existed across much of the specimen thickness and cleavage fracture was the dominant  .  .  _  (7,124)  mode of fracture r  Table 4.1 l i s t s the t r a n s i t i o n temperatures f o r the fracture toughness curves (from Figures 4.1 - 4.6), the precracked Charpy i n i t i a tion energy curves (from Figures 3.36 - 3.44), the Drop Weight Tear Test percent shear (Figures 3.45 and 3.47), and the standard Charpy t o t a l energy curves (Figures 3.17 - 3.25).  I t i s s i g n i f i c a n t that the t r a n s i t i o n  temperatures are equivalent f o r a l l but the standard Charpy t o t a l energy curves.  The t r a n s i t i o n f o r those curves occurred, over a wide range of  temperatures and at a lower temperature than the sharp transitions of the fracture toughness, precracked i n i t i a t i o n energy, and DWTT percent shear.  Barsom and R o l f e ^ ^ have shown that at the low end of the 4  fracture toughness t r a n s i t i o n temperature curve, the mode of i n i t i a l crack extension i s cleavage. i s d u c t i l e tearing.  At the upper end, the i n i t i a t i o n mode  In the t r a n s i t i o n region, both modes occur.  Since  the c r i t i c a l fracture toughness describes a crack i n i t i a t i o n event, i t i s not surprising that the precracked Charpy i n i t i a t i o n energy curve has the same general shape and t r a n s i t i o n temperature as the fracture  Table 4.1 TRANSITION TEMPERATURES  K  AF-1 AF-2  DWTT  EI  Ci•ack Traiasverse Rcj l l i n g D i r e c t i o n  Crack P a r a l l e l to Rolling Direction  Crack P a r a l l e l to Pipe Axis std  pc  K  EI pc  std  K  EI pc  ET std  -40  -40  -45  -60  lower shelf  lower shelf  -60  -50  -50  -70  -30  -30  -30  *  -50  -50  -85  -55  -50  -80  A l l temperatures i n °C *  No sharp  transition.  - 202 -  toughness.  It i s i n t e r e s t i n g , however, that the fracture toughness transitions for the material tested i n the pipe axis orientation coincided w e l l with the Battelle-Drop Weight Tear Test percent shear t r a n s i t i o n (Table 4.1).  The fracture toughness t r a n s i t i o n i s believed  to be due to a change i n the microscopic i n i t i a t i o n mode; the DWTT i s primarily a measure of the propagation mode.  In general, the values of K  were greater than those of  •J  K  r n n  .  The c a l c u l a t i o n of both parameters i s based upon the assumption  that the crack i n i t i a t e s at the peak load.  Thus, when considerable  general y i e l d i n g occurs prior to the attainment of maximum load, these values tend to be nonconservative.  However, the Kj c a l c u l a t i o n  employs the area under the load-time curve to the point of maximum load, whereas that of load.  Thus, K  p  requires only the d e f l e c t i o n to maximum  i s a more conservative measure of a material's  dynamic e l a s t i c - p l a s t i c fracture toughness than i s K  T>  In general, the r e p r o d u c i b i l i t y of a l l the fracture toughness data was excellent; standard deviations of less than 10% were obtained, which i s within the expected scatter band for s t a t i c K  data^  1 4 3  \  - 203 -  The AF-1 and AF-2  s t e e l exhibited comparable dynamic  i  fracture toughness when tested with the crack running p a r a l l e l to the pipe axis (Figures 4.1 - 4.2).  However, the AF-2  s t e e l showed  a marked t r a n s i t i o n at -30°C, 10°C higher than that of the AF-1 material.  Data from Tables 3.3 and 3.5 have been included i n Table 4.2 and indicate that f o r the pipe axis o r i e n t a t i o n , the AF-1  steel  absorbed far more energy i n a standard Charpy test than did the AF-2 s t e e l , i n the temperature as shown i n Table 4.2, equivalent.  range from -40° to -60°C.  their fracture toughness values are v i r t u a l l y  Notice also, though, that the precracked Charpy  Table  4.2  FRACTURE TOUGHNESS AND  Material  However,  Orientation  ENERGY DATA  ET . , std  K , Id  EI pc  AF-1  P a r a l l e l Pipe Axis  75 f t - l b  60 k s i - i n  2  9 ft-lb/in  AF-2  P a r a l l e l Pipe Axis  31 f t - l b  55 k s i - i n  2  9 ft-lb/in  AF-1  Parallel Rolling Direction  14 f t - l b  62 k s i - i n ^  A l l data for -50°C  5 ft-lb/in  2  - 204 -  initiation energies were equal for both steels i n this orientation. Since the c r i t i c a l fracture toughness describes a crack initiation event, i t would be expected that the two materials have similar K^.^ values i f their crack initiation energies were similar.  A standard  Charpy test would mask such v i t a l information, however.  Although  overall the AF-1 steel had greater total absorbed Charpy energies than did the AF-2 steel, precracked initiation energies of the two materials were generally similar.  This equivalence of crack i n i t i a -  tion energy was manifested i n the very similar fracture toughnesses of the two steels.  It should be noted that the AF-2 steel precracked Charpy data showed some linear-elastic fractures at -20°C.  At that tem-  1-  perature, that steel's average K _ was 162 k s i - i n though the mean K  2  I'  (178 MPa-in ), 2  for the AF-1 steel was 207 ksi-in (228 MPa-in ). 2  2  Using Equation 4.16, this represents a variance i n the c r i t i c a l through-wall flaw size of over 3 inches (7.6 cm) for a pipeline operating at a stress of 56 ksi (386 MPa)(5.3 in[13.5 cm] for AF-2; 8.7 in [22.1 cm] for the AF-1 steel).  Both flaw sizes are quite  large, however, and could be detected as leaks prior to unstable crack propagation^. The biggest differences between the fracture toughness values of the two steels were observed for tests in which the crack  - 205 -  was parallel to the rolling direction (Figures 4.3 and 4.4). A l though neither steel showed a sharp transition, the AF-2 steel exhibited a higher fracture toughness; valid K values for temJ  peratures from +20°C to -100°C were 175 to 57 k s i - i n  2  (193-62 MPa-  in ). The Kj values for the AF-1 steel in that same temperature range were only 72 to 54 k s i - i n  2  (79-59 MPa-m ). 2  The absorbed Charpy energy of the AF-1 steel in this direction, particularly the initiation energy, was essentially at lower shelf over the complete temperature range.  It is likely that  the MnS inclusions, aligned along the rolling direction during hot  rolling, are responsible for the relatively low fracture  toughness values observed. The AF-2 steel, treated with rare earths, maintained a high toughness even in the rolling direction. This detrimental effect of elongated inclusions on fracture toughness has been observed by others  ,  At -20°C, assuming a failure stress of 56 ksi (386 MPa) and an average K  C 0 D  of 128.5 k s i - i n  2  (141 MPa-in ), the AF-2 steel 2  in the rolling direction had a c r i t i c a l through-wall defect size of 3.4-in (8.5 cm). The AF-1 steel, using a mean 68 k s i - i n (2.3 cm).  2  value of  (75 MPa-in ) had a c r i t i c a l flaw size of only 0.9-in 2  - 206 -  It i s known that strain age embrittlement can occur near the welds in pipelines and that this could reduce the fracture toughness.  In addition, residual stresses in these regions can  attain a stress level equal to that of the yield strength of 70 ksi (483 MPa).  Under such conditions, the c r i t i c a l flaw size for the  AF-1 pipe in the rolling direction would be less than 0.6-in (1.5 cm); this estimate does not account for the loss in fracture toughness that might be associated with the strain age embrittlement.  Specifying a minimum toughness of 50 ft-lb (68 J) in the pipe axis orientation is equivalent to ensuring a minimum c r i t i c a l flaw size of approximately 6-in (15.2 cm) i n this direction (Table (95) 3.1)  . Since the c r i t i c a l flaw size i s approximately 1/10 of  this value for a crack running parallel to the rolling direction, this direction must be included in any pipeline specification imposed to restrict fracture initiation. It should be noted that for the AF-1 steel, a significantly higher total absorbed energy i s required in the pipe axis orientation as compared to parallel to the rolling direction.  However, the  fracture toughness data i s similar in both directions from - 50°C and below, as shown in Table 4.2 (also refer to Table 3.3 and Figures 4.1 and 4.3).  The initiation energies for the precracked  Charpy specimens for these two orientations are equivalent, in this  - 207 -  temperature range, thus explaining this apparent discrepancy (Tables 4.2 and 3.5).  Both steels showed equivalent fracture toughness in specimens oriented with the crack transverse to the rolling direction (Figures 4.5 and 4.6).  The absorbed Charpy energy values were also  quite similar (Table 3.3).  The AF-1 steel showed a transition from  -40° to-50°C, whereas the AF-2 material's fracture toughness transition was more gradual, occurring over the range from -40°C to -60°C.  Within these transition ranges both linear-elastic and  elastic-plastic failures were observed.  At -20°C, employing K and 206 ksi-in  values of 188 k s i - i n  2  for AF-2  for AF-1 (207 and 227 MPa-in ) and an applied stress  of 56 ksi (386 MPa), the c r i t i c a l through-wall flaw sizes would be 7.2-in (18.3 cm) and 8.6-in (21.8 cm), respectively, both quite large.  4.4  Correlations  4.4.1  Relationship Between Dynamic Stress Intensity Factor and Crack Initiation Energy  - 208 -  A fundamental relationship for plane strain fracture i s (5,9).  =  K. Tc  (Eq. 4.17)  [ E G / ( l - v )] 2  H  Ic  The term G^ i s defined as the c r i t i c a l elastic energy release rate c  and can be described as being the work required to initiate unstable fracture at the tip of a flaw.  Through Equation 4.17 i t can be seen  that the stress intensity and energy approaches to fracture toughness are equivalent.  In instrumented impact testing, when fracture occurs prior to general yielding, unstable crack growth begins at the point of (27) maximum load  . The parameter G^j can therefore be associated  with the energy to maximum load, i.e., the crack initiation energy, EI per unit area.  Thus, should plane strain conditions exist, Equa-  tion 4.17 can be modified such that: 1 (1 - v ) 2  (Eq. 4.18)  elastic modulus  where, E  ligament area of Charpy specimen  A  Koppenaal  EI A  (59)  in an IIT study showed a correlation between  - 209 -  and the crack initiation energy where, of course, both terms A under A , , ' strain _ . rates.. ^ were measured equivalent  Others (60,72-73,149)  have attempted similar correlations, with and without IIT, by assuming the total energy absorbed by the precracked specimens could be related to K^  (or K^)  c  K  2  / E  through the relationship:  1  =  E|  2(1 -  i C  v )  ( E q >  4 > 1 9 )  A  These correlations were not based on sound principles for one or more of the following reasons: 1.  K^  c  is defined by conditions existing at point of crack  initiation, whereas the measurement of ET/A  involves the total fracture  process; 2.  the factor of 1/2 in Equation 4.19 was explained by  (149) Ronald  on the basis of two surfaces being created at fracture,  although that fact was accounted for in the basic definition of G  T  Ic  that Ronald ^"^^  and others  showed a correlation between Ic  (59) and ET/2A was  fortuitous, since i t was later shown  materials they studied 3.  (high strength Ti-alloys)  in some instances, K  T  that for the  ET/2A - EI/A;  was measured under slow bend  ic conditions but the data for ET/A was obtained at an impact loading rate; no correlation should be expected for strain-rate sensitive materials.  ;  - 210 -  A l l the K.^ data for which fracture occurrred prior to general yielding has been plotted against the corresponding precracked Charpy crack initiation energies in Figure 4.7.  Not a l l of the IC^ values calculated in this study could be considered valid according to the ASTM E 399 c r i t e r i o n ^ " ^ ^ which requires a minimum thickness for plane strain conditions (Equation 4.5), nor for the more liberal restriction cited in Equation 4.6. Those data that did meet those respective plane strain criteria have been so distinguished in Figure 4.7.  It i s significant that for the data obeying the most 2 conservative validity criterion, B > 2.5 (K^/o^) , excellent agreement exist between the theoretical and the measured relationship 2 between EI and (K^/E).  The initiation energy thus appears to be  a reasonable estimate of  for those tests adhering to that require-  ment . The specimens meeting the stipulation that their thickness 2 be greater than 1.6 (K^/cr^)  displayed somewhat more scatter, but  their data s t i l l agreed well with the theoretical line. However, those specimens meeting only the  <  PQ  criterion, which i s to be employed in the tentative ASTM IIT  Y  T  1  1  1  \—  EI/A  1  1  1  1  r  (in-lb/in ) 2  2 Figure 4.7  Kjd /E vs i n i t i a t i o n energy f o r a c i c u l a r f e r r i t e IIT K , data meeting d i f f e r e n t v a l i d i t y  requirements.  - 212 -  standard, show the greatest scatter, and l i t t l e , i f any, relationship to the theoretical line. This means that P„,„ < P_,, i s an MAX GY r  insufficient criteria for ensuring plane strain conditions.  Apparently for such specimens, the initiation energy as measured from the IIT load-time trace includes factors other than those strictly associated with  Possibly, energy losses due  to subcritical crack growth (doutbful for cleavage fractures occurring prior to general yielding) or plastic indentation at the contact points during impact (Brinell energy, E^, in Section 2.3.2.1.2) are a  significant.  It i s also possible that the limited ligament depth  below the crack i n a Charpy-size specimen may preclude a true measurement of a material's fracture resistance.  The ASTM E 399  validity criterion (Equation 4.5) stipulates a minimum crack size(a) and ligament depth (W-a), in addition to thickness (B), so that the stress field ahead of the crack approximates that i n a linear-elastic (72) body  . This crack length and ligament depth requirement i s not met  by the small Charpy specimens. It should also be emphasized that the definition of G ^ i s the change i n strain energy, U, with a change i n crack length (du/da) (9) The initiation energy, however, has been normalized by dividing by the total precracked Charpy ligament area, i.e., EI/A, since no accurate measure of the ligament depth associated with the i n i t i a l crack extension i s possible.  The initiation energy EI may indeed be  - 213 -  equivalent to U, but the precracked Charpy ligament area, A, i s definitely not equal to da.  Thus EI/A would be expected to be  a conservative estimate of G_,  (22)  Id  The data also indicate that the ASTM E 399 thickness requirement provides a more conservative  value, as shown in  Table 4.3.  4.4.2  Comparisons Between K^^ and Statically Obtained  K^  c  Due to the relatively high fracture toughness of acicular ferrite steel and the limited thickness of the controlled rolled plates used to produce the line pipe, valid K  data is impossible  to obtain except perhaps at very low temperatures.  Indeed, no  references to the linear-elastic plane strain fracture toughness of AF pipeline steels could be found in the literature.  Diesburg^  has reported the results of a J-Integral  study of an acicular ferrite, aluminum killed, rare earth treated steel very similar in composition to the AF-2 steel examined in the present work. Using compact tension specimens and the procedure described by Landes and Begley^ ^, the quasi-static J-Integral 34  plane strain fracture toughness was  determined.  Table 4.3 K_ . VALUES FOR DIFFERENT VALIDITY CRITERIA  K  (ksi-in* )  K  2  (ksi-in* ) 2  B>1.6(K /a )2  Id(ksi-in S) P <P MAX GY  -  59.0  65.0  47.1  55.9  63.0  -  61.7  68.6  139.0  55.4  62.8  -  126.5  -  61.9  Material Code  T(°C)  AF-l-TR  -50  122.7  AF-l-TR  -60  119.1  AF-2-TR  -60  126.6  AF-l-TP  -50  AF-l-LR  -50  a ,(ksi) yd  I d  B>2.5(K /o ) Id  yd  2  I d  Id  yd  K  i  70.4  - 215 -  A comparison of Diesburg's results with the dynamic J-Integral fracture toughness obtained from the IIT approach i s shown i n Figure 4.8. An excellent correlation exists, the fracture toughness values being nearly equivalent at a l l temperatures.  Dynamic fracture toughness data i s generally expected to be slightly more conservative than statically obtained values. This was observed above -20°C, although only small differences are apparent, However, no consistent deviation i s evident below that temperature. Diesburg reported that the cleavage fracture transition temperature was -18°C.  Hahn and c o w o r k e r s h a v e shown that the rate sensiti-  vity of K^ decreases with increasing yield strength and decreasing c  temperature.  In addition, for low and medium strength structural  steels there appears to be l i t t l e strain rate sensitivity i n the plane strain fracture toughness where cleavage fracture dominates (124 123 132) '  . I t appears that the fracture toughness of this AF  steel i s not sensitive to increasing strain rate where cleavage fracture occurs. In a separate study, A k h t a r m e a s u r e d the fracture toughness (ASTM E399) of a section of a X70 pipe nominally 42-in (106.7 cm) O.D., 0.425-in (10.8 mm) wall thickness.  This steel had  a reduced-pearlite (RP) microstructure and contained 0.15 w/o C, 1.6 w/o Mn, 0.05 w/o Nb, and 0.12 w/o Cr.  The pipe had been formed  - 216 -  T  Stolic  T  ond Dynamic J-lntegral fracture toughness AF-2 - crack parallel to rolling direction  L300  o  IIT  •  sialic Kj  Kj  -200r  100 \  _l  -40  -80  l_  l_  0  TCC)  • 20  Figure 4.8  Reduced  Pearlite  Steel-Fracture  Toughness  crack parallel to rolling direction  IT {  r  * t^  1300,  d  COD  £ 300  static  r  Kg  K valid r  12002  —''  !"200r  -uoo < = >  1001  •8 i  -120  -80  Figure 4.9  -40  TCC)  0  • 40  - 217 -  by the U.O.E. process and therefore the longitudinal pipe axis was the same as the rolling direction.  However, due to the thickness limitation, Akhtar's data was estimated to be valid (plane strain) only at temperatures below -105°C!  Instrumented impact tests with the crack following the pipe axis (rolling direction) were conducted on this same material.* Both the static and dynamic fracture toughness results are shown in Figure 4.9.  For temperatures at which the static K  value was valid  (approximately -105°C and below), Figure 4.9 shows that the IIT values, Kj  and K^, yield equivalent results.  It thus appears that where fracture occurs by cleavage, K - K for both acicular ferrite and reduced pearlite steels. Ic Id T  TJ  Barsorn^*^ has shown that the fracture toughness transition temperature for impact loading tests (e - 10/s) i s shifted to higher  *  Supplied by The British Columbia Hydro and Power Authority, Materials Research Center, Vancouver, B.C.  - 218 -  temperatures than that obtained by static loading (e = 10 "Vs). He presents an equation which predicts the magnitude of this shift: T  =  shift  119 - 0.12a  (Eq. 4.20)  ys  for, 250 MPa < a < 965 MPa ys where, ^  magnitude of shift in fracture  =  toughness transition temperature between slow and impact loading conditions, °C a  =  room temperature yield strength, MPa  For this reduced pearlite steel (a sition was at approximately -80°C.  - 483 MPa) the static K^  c  The IIT  tran-  transition occurred  at approximately -20°C (Figure 4.9), a shift of 60°C.  This i s in  excellent agreement with Equation 4.20 which predicts a shift of 61°CI  The static versus dynamic fracture toughness correlations shown in Figures 4.8 and 4.9 are among the few reported for pipeline steels, and in both cases they, indicate that the IIT technique produces fracture toughness data which i s i n excellent agreement with the "accepted" statically measured values.  - 219 -  Critical Flaw Sizes  4.4.3  The Battelle ductile fracture initiation  =  equation:  ln[sec(irMa /2a )  (Eq.  3.2)  was employed to calculate the c r i t i c a l crack lengths for a sharp through-wall flaw for the RP and for the AF materials.  The Folias  correction, M, in Equation 3.2 is a function of the crack length (-(1 + 1.255 for "c".  c /Rt) ), so a graphical procedure was used to solve  For comparison purposes, c r i t i c a l crack sizes were also  calculated from the IIT fracture toughness data using Equation  4.16.  The Battelle calculation employs the standard Charpy upper shelf energy as a measure of the fracture toughness and is therefore empirical.  This Charpy energy is related to the fracture toughness  , (96) by the equation :  K  where, : C A  v c  (12C E/A ) v c upper shelf energy (ft-lb)  =  ET  area of Charpy ligament (= 0.124-in ) = B(w-a)  E  (Eq.  elastic modulus  4.21)  - 220 -  Incidently, this Equation i s equivalent to that used to calculate Kj (see Equations 4.9 and 4.10) if_ the initiation energy EI i s set equal to ^ C^.  However, this equivalence is not well supported by  the energy data presented in the previous Chapter (Tables 3.3 and 3.5), EI being much less than h C .  Equation 4.21 is therefore  empirical.  Since the Charpy upper shelf temperature is employed in the Battelle equation, i t i s strictly valid only at upper shelf temperatures. The equation has merit, however, in that the Folias correction accounts for bulging which occurs around defects in pressurized cylinders (pipelines).  This bulging can cause increases in the stress at the  crack tip and therefore can result in smaller c r i t i c a l crack sizes (97) than required for a similar flaw in a flat plate . Furthermore, the Battelle relationship has been shown to accurately predict the c r i t i c a l flaw sizes in f u l l scale burst tests for certain grades of (94) pipeline steels  Equation 4.16  ( J J f( ) )» K  =  a  7ra  2  011  t n e  other hand, utilizes  a true material property (K-^) instead of an empirical estimate of the fracture toughness (Equation 4.21).  Its application i s limited,  however, as the equation assumes a through-wall crack in an infinitely (144) wide plate  and is therefore nonconservative for pipeline geometries.  - 221 -  Table 4.4 l i s t s the c r i t i c a l crack lengths for the two AF steels and the RP s t e e l as determined  from the empirical B a t t e l l e  equation and from the three IIT fracture toughness parameters, K , COD' * I d *  K  a m  K  lations:  T w  ° f *-'a  ure  s  t  r  e  s  s  levels were employed i n the calcu-  the specified minimum y i e l d strength (SMYS), 70 k s i (483  MPa),  and the t y p i c a l design pressure (hoop stress) for a X70 gas p i p e l i n e , 56 k s i (= 0.8 SMYS). For the AF-1  s t e e l , a reduced stress l e v e l of  28 k s i (193 MPa) was also employed f o r determining the c r i t i c a l crack length for a crack p a r a l l e l to the r o l l i n g d i r e c t i o n .  At +20°C, i . e . , i n the region of e l a s t i c - p l a s t i c fracture for  the pipe axis orientations, the IIT K^,^  c r i t i c a l crack lengths  were i n good agreement with those predicted from the B a t t e l l e equation. The K  crack sizes are, however, much larger.  This i s probably due to  J  the non-conservative nature of Kj at temperatures  where crack i n i t i a t i o n  occurs p r i o r to maximum load.  The B a t t e l l e c r i t i c a l crack lengths are quite d i f f e r e n t from those predicted by the IIT data for the AF-1 ness r o l l i n g d i r e c t i o n .  s t e e l i n i t s lowest tough-  The B a t t e l l e relationship predicts c r i t i c a l  crack sizes that would e a s i l y be located as leaks p r i o r to unstable crack growth.  Indeed, the IIT fracture toughness calculations predict  r e l a t i v e l y large c r i t i c a l crack lengths for operating stresses of 28 k s i (193  MPa).  Table 4.4 CRITICAL CRACK SIZES  Material & Orientation  Failure Stress (psi)  C r i t i c a l Crack Size(in) Battelle  C  V  K  J  C0D  K  K  Upper Shelf  T  (ft-lb)  (°C)  88  +20  121  +20  20  +20  94  +20  62  -20  -  96  -20  Id  AF-2 pipe axis AF-1 pipe axis AF-1 rolling direction  70000 56000 70000 56000 70000 56000 28000  4.0 6.7 3.5 6.3 3.2 5.1 8.8  9.0 14.0 7.9 12.4 0.7 1.1 4.2  4.4 6.9 4.6 7.2  RP pipe axis  70000 56000  3.4 5.5  7.7 12.0  4.1 6.4  AF-2 pipe axis AF-1 pipe axis AF-1 rolling direction  70000 56000 70000 56000 70000 56000 28000  4.0 6.7 3.5 6.3 3.2 5.1 8.8  3.7 5.8 9.7 15.2 0.7 1.0 4.2  3.4 5.3 5.6 8.7 0.8 1.3 5.0  _  0.6 0.9 3.8  20  -20  RP pipe axis  70000 56000  3.4 5.5  2.1 3.2  1.5 2.3  0.6 0.9  41  -20  —  -  —  -  -0.6 0.9 3.8 _  — —  - 223 -  However, i t should be pointed out that the c r i t i c a l crack sizes predicted by the Battelle equation for the AF-1 steel in the rolling direction are considerably larger than those predicted from the IIT parameters.  The Battelle equation i s therefore not conser-  vative.  The c r i t i c a l crack lengths predicted from the IIT data for the AF-1 steel in the rolling direction are less than one inch for stresses above 56 ksi!  If one introduces the real possibility of  strain age embrittlement, tensile residual stresses, frost heave and subsequent buckling, or dynamic loading from machinery, the IIT fracture toughness data suggest that the AF-1 i s highly susceptible to unstable crack propagation in the rolling direction.  The Battelle  equation does not suggest this, however, and this discrepancy warrants further investigation.  Furthermore, the Battelle equation i s insensitive to material properties; similar crack sizes are predicted for Charpy upper shelf (95)  values from approximately 30 to 80 ft-lb (41-108 J)  I The fracture  toughness values for materials of such widely different Charpy energies should not be equivalent and therefore the c r i t i c a l crack sizes should be different.  It i s unlikely that the AF-1 steel in the rolling direc-  tion would have the same c r i t i c a l crack length as i t would have along the pipe axis since the Charpy upper shelf values for the two orientations  - 224 -  are 20 ft-lb and 121 f t - l b , respectively (27 and 167 J ) , and the  i^  I,  fracture toughnesses are 68 k s i - i n and 189 k s i - i n 2  2  (K  U\JL)  ).  Nevertheless, the Battelle relationship predicts very similar crack sizes of 3.2-in and 3.5-in for a 70 ksi hoop stress.  The IIT data,  however, predicts c r i t i c a l crack sizes of 0.6-in in the rolling direction and 4.6-in along the pipe axis.  Table 4.4 also l i s t s the c r i t i c a l crack sizes at -20°C. The pipeline industry assumes that the Battelle ductile fracture initiation equation (Equation 3.2) i s valid at a l l temperatures above the 85% shear transition temperature  (as defined by the DWTT)  since that transition specifies the regime of ductile fracture. However, Equation 3.2 employs the Charpy upper shelf energy (through Equation 4.21), a value which may not apply at the minimum design temperature of -18°C, even though ductile failures may be expected at that temperature  (85% shear obtained in the DWTT). The c r i t i c a l  flaw sizes from the Battelle equation at -20°C are therefore equal to those at +20°C, although, as Table 4.4 shows, there were significant decreases in the Charpy energy of the steels between those temperatures. However, using fracture toughness data obtained at -20°C, different, smaller flaw sizes are obtained.  The use of fracture toughness  parameters determined by IIT to determine c r i t i c a l crack sizes i s therefore more objective, less dependent upon empirical  assumptions,  and more responsive to fracture toughness temperature transitions.  - 225 -  The use of fracture toughness data in the pipeline industry i s virtually nonexistent, however.  4.4.4  Empirical Correlations Between  and Other  Material Properties  4.4.4.1  versus Charpy Energy  The difficulty in obtaining fracture toughness data and the wide popularity of the simple Charpy impact test have prompted many workers to attempt correlations between K.^ and the total absorbed energy obtained from a Charpy test, C  v  (=ET).  Such cor-  relations are necessarily empirical since comparisons are being made between tests which have significant differences.  The Charpy test  is conducted under impact loading, whereas K^ data i s obtained c  under slow strain rate conditions; the Charpy specimen has a relatively blunt notch, the fracture toughness specimens require fatigue precracks; and, the energy absorbed i n a Charpy test i s a measure of the entire fracture event, whereas K i s related to the initiation Ic of a crack. For these reasons, empirical relationships between C T  and K_ can at best have limited application. Ic  Nevertheless, many  such correlations can be found i n the l i t e r a t u r e a n d their use in the absence of fracture toughness data i s often s u g g e s t e d .  - 226 -  Sailors and Corten  have correlated the dynamic  fracture toughness with the corresponding Charpy energy obtained at the same temperature  (transition and lower shelf range).  This  correlation i s noteworthy because i t was derived from data from eleven low alloy structural steels and two pressure vessel steels, and since similar strain rates were used in obtaining the correlating data.  Their relationship i s :  K  where-,  d  =  15.873(C )°' v  375  (Eq. 4.22)  expressed in ksi-in C in ft-lb v  Those workers obtained a surprisingly good linear regression correlation coefficient of +0.94 i n their study which employed data from the thirteen different steels and both precracked and standard Charpy specimens!  The  data obtained in this study (which also comes from  the transition and lower shelf temperatures) has been plotted against the corresponding standard Charpy total energy in Figure 4.10. Equation 4.22 does not f i t this data.  An empirical correlation  was fitted to the acicular ferrite pipeline steel data by linear regression and i s of the form:  - 227 -  1—r r  r  T  -  T  Dynamic fracture toughness vs charpy energy Acicular ferrite steels K = 33.67(C,j ld  K = 15.873 (C,j ld  tf  IOO80-  I  60"  %  40  o oo-o° OJ>o o ^—  20h  10  J  4  6  l_L  8 10  20 C, (ft-lb)  J t_L 60 80100  40  Figure 4.10  1  r  Dynamic fracture toughness vs  yield strength  Acicular ferrite steels  2.18 - 2.14 o 2.10  cr, =l07 a  (Ky/c^J  2.06 2.02 1.98  J  .56  I  -48  I  I  I  -.40 -.32 -.24 \oq{Ku/cr ) yi  Figure 4.11  L  -.16  -.08  0  - 228 -  K_, Id The  =  33.67(C ) ' v 0  1 6 2  (Eq. 4.23)  c o r r e l a t i o n c o e f f i c i e n t was +0.74.  4.4.4.2  K vs Y i e l d Strength Id T J  Fracture toughness values are also often correlated with the y i e l d strength.  The y i e l d stress i s believed to be the single most  important mechanical property governing the fracture toughness of a material ^ " ^ . 2  The r a t i o (% /a ) i s often used as a fracture conc  y  t r o l c r i t e r i o n since i t i s a d i r e c t measure of both the p l a s t i c zone size  ahead of a crack t i p and the c r i t i c a l flaw s i z e for unstable  crack extension  For cleavage fractures, many of these correlations have been reduced to the f o r n / " ^ ^ :  o*/a  r  where,  y  =  a(K_ /a ) Ic y  3  =  cleavage fracture stress  a  =  y i e l d stress  y a, 3  =  empirical material constants  'f  (Eq. 4.24)  The cleavage fracture stress i s independent of temperature and s t r a i n rate^  1 2  ^ , so Equation 4.24 can be rearranged to give:  - 229 -  CT  where,  C  y  =  =  C ( K  Ic  / a  y  )  ( E q  - ' 4  2 5 )  /a  For the wide range of materials examined by Hahn, et a l ^ " ^ Equation 4.25 was shown to reduce to:  a  =  a */2.35(K^ la ) *  y  The  t  ic  3 3 3  (Eq. 4.26)  y  data obtained i n this study was similarly correlated  with the dynamic yield strength and the resulting graphical representation of Equation 4.25 i s shown in Figure 4.11. The equation representative of this data i s :  - 9n? . V  "  1 0 7 ( K  Id  / a  yd  )  "  ( E q  ' ' 4  2 7 )  The correlation coefficient was only - 0.62.  Note from Equation 4.25 that the constant 107 i n Equation 4.27 is equal to  /a. Using a value of 2.35 for a  , the cleavage strength  of the acicular ferrite pipeline steels i s estimated to be on the order of 251 k s i (1734 MPa). An independent estimate of the cleavage fracture * (156) stress, c?£ , for these steels can be made from IIT data :  - 230 -  o  * f  (ksi)  =  72.5 P  when, P ^ / P e y  From this equation, assuming ^^^^QY  =  GV  (lbs)  (Eq. 4.28)  0.8  =  at approximately -100°C  and that the dynamic yield strength for the acicular ferrite steels is approximately 130 ksi (897 MPa) at that temperature, the cleavage fracture stress i s estimated to be 283 k s i (1952 MPa).  The conclusion to be drawn from these correlations between and other material properties are that they are empirical and unreliable (poor correlation coefficients).  Although such correlations  may be useful for crudely measuring the relative fracture toughness of materials, they are not necessary with an instrumented Charpy machine since the fracture toughness, yield strength, and absorbed energy values may be obtained simultaneously.  - 231 -  5. CONCLUSION  5.1  Conclusions  An instrumented impact test machine with a drop tower design was statically and dynamically calibrated to accurately measure the energy absorbed by a Charpy specimen.  The importance of adhering to proposed ASTM IIT validity criteria was assessed. When the fracture time, t^, i s less than the electronic system response time, T , the measured results can K  be seriously inaccurate: measured loads and fracture toughness parameters are attenuated and fracture times increased. The validity criterion, t^ > 3T ( T = period of inherent specimen o s c i l lations) appears to be quite conservative.  An unnecessarily high impact velocity, v , i s the single most detrimental test parameter.  The effect i s to decrease the  fracture time such that t^ < T^ and/or t^ < 3T, thus invalidating the test results.  High amplitude specimen oscillations are also  generated which hinder data analysis.  It appears that the proposed plane strain fracture toughness validity requirement, P y < Ppy' MA  n  o  t  ^  e r e s t  r i c t i v e enough to  - 232 -  ensure plane strain conditions and the measurement of the most conservative fracture toughness parameters. 2 requirement, B > 1.6  (^j^/cf^)  The specimen thickness  » does appear to be adequate to  ensure plane strain conditions, however. The effect of deviating from the standard Charpy specimen notch dimensions was evaluated. Measurable increases in the absorbed energy were obtained with increasing notch angle and decreasing root radius.  In general, increasing the specimen thickness from 10 mm to  13.7 mm caused a decrease in a l l components of the upper shelf energies per unit area, although the lower shelf energy and the transition temperature were not significantly affected.  Crack growth studies confirmed that for general yield failures, crack initiation occurs prior to maximum load; the crack extends to f u l l specimen thickness at maximum load.  Estimates of the initiation energy  based upon the assumption that a crack initiates at maximum load are therefore nonconservative.  The initiation and propagation components of the total absorbed energy showed transitions with decreasing temperature.  This suggests  that EI may have particular significance in terms of a transition temperature approach to the fracture initiation event.  - 233 -  A comparison study of the dynamic properties of two acicular ferrite steels clearly demonstrated the usefulness of IIT.  The tests  revealed that present pipeline toughness specifications may be inadequate for ensuring fracture control.  Very low initiation energies  were obtained i n tests parallel to the rolling direction - a test direction not included in the present toughness test requirements in one of the AF materials. For precracked specimens, the initiation energy remained at a lower shelf condition even at room temperature. The tests indicate that more stringent pipeline toughness specifications are necessary in a l l directions i n the pipe.  It i s suggested that the  acceptance criterion be based upon the magnitude of the initiation energy obtained from a precracked Charpy specimen.  This would ensure  a conservative estimate of the initiation energy and be applicable to the most severe in-service defects.  Instrumented impact testing also showed that strain aging the semi-killed acicular ferrite pipeline steel decreased only the propagation energy; the initiation energy was unaffected.  This indicates that the  potential for crack initiation i n this steel i s not increased by strain aging.  The total fracture energy as obtained from a standard Charpy test was shown to often mask the fracture toughness value of a material. In some instances, materials of equivalent fracture toughness had dissimilar Charpy energies.  In others, similar Charpy values were obtained  - 234 -  with materials of widely d i f f e r e n t fracture toughness. the  IIT indicated  i n i t i a t i o n energy obtained from testing precracked Charpy speci-  mens accurately denoted the r e l a t i v e magnitude of the fracture toughness; the precracked i n i t i a t i o n energy t r a n s i t i o n temperature also coincided with that of the fracture toughness.  Although more work i s  required to establish the significance of the EI parameter, these tests indicate that i t could be a basic parameter for assessing true fracture initiation.  The equation used by the p i p e l i n e industry to predict c r i t i c a l defect sizes i s based upon a material's Charpy upper shelf energy which i s not representative of the fracture toughness, and hence, the c r i t i c a l defect s i z e .  In general, the c r i t i c a l crack sizes  predicted from fracture toughness data obtained from IIT were more conservative than those obtained from that empirical equation and were responsive to toughness t r a n s i t i o n s .  Good correlations between the fracture toughness values from IIT and those s t a t i c a l l y obtained by conventional techniques were observed.  5.2  Suggestions for Future Work  More data i s necessary to confirm the effectiveness of the proposed IIT v a l i d i t y c r i t e r i a .  Fracture toughness parameters should  - 235 -  be obtained by conventional  test procedure and should be compared  with those generated by IIT employing each of the p o t e n t i a l plane  1 j  strain validity c r i t e r i a .  The s i g n i f i c a n c e of the i n i t i a t i o n energy as measured by IIT needs to be better defined.  A complete fractographic analysis  of the specimens tested i n t h i s work should provide a better understanding of the conditions necessary to control both the fracture i n i t i a t i o n and fracture propagation event.  A l l fractured specimens  have been coded and desiccated to permit a future study.  Specific areas of i n t e r e s t to the p i p e l i n e industry have been revealed by t h i s study.  Additional testing i s required to  establish the minimum i n i t i a t i o n energy needed to protect against crack i n i t i a t i o n .  More correlations between IIT and Drop Weight  Tear Testing are required.  An i n v e s t i g a t i o n of the usefulness of  a precracked f u l l wall specimen i s necessary.  The value of IIT i n assessing fracture toughness behaviour has also been shown.  This approach could be e a s i l y applied to a study  of the toughness properties i n the weld bead and HAZ of the s p i r a l seam welds and g i r t h welds i n the p i p e l i n e s t e e l s , p a r t i c u l a r l y f o r the AF-1 s t e e l i n i t s low toughness orientations.  - 236 -  A l l future data from the Department of Metallurgy's instrumented impact machine should be filed in the computer so that correlations may be more efficiently generated and so that sophisticated statistical analyses can be made^"^.  The incorporation of a dual-beam oscilloscope would be quite useful for providing both load-time and energy-time data and would relieve the necessity of measuring the energy with a planimeter.  - 237 -  REFERENCES  1.  D. Leavitt: Audubon, 1977, vol. 79, No.2, p.145.  2.  Instrumented Impact Testing, STP 563, ASTM, Philadelphia, 1974.  3.  M.G. Dawes,Ed.: Int. Conf. on Dynamic Fracture Toughness, Welding Inst., London, 1977.  4. R.A. Wullaert, D.R. Ireland, and A.S. 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Akhtar: B r i t i s h Columbia Hydro and Power Authority, Materials Research Centre, Vancouver, B.C., unpublished research, 1977.  154.  F.M. Burdekin, et a l : Weld. In the World, 1975, v o l . 13, p.29.  155.  R.H. Sailors and H.T. Corten: i n Reference 97, p.164.  156.  S. Ensha and A.S. Tetelman: Report UCLA-ENG-7435, School of Engineering and Applied Science, UCLA, Los Angeles, CA, 1974.  157.  R.A. Wullaert, et a l : i n Reference 3, p.31.  APPENDIX A DYNAMIC CALIBRATION RESULTS coot f10070 TI029O u302u2 UiO 7 8 0 v roevs V 7 0963  . ?  TtKP c  YSS  -uo. -40.  79,0 79.0 79,0 79,0 7 9,0 79.0  -10.  -uo. -uo. -uo.  6  c Ti a o 7o  r io29o UJ02U2 liio  7bo  V 7 0293 V /096i c a n t  -uo. -UO. -'10. -UO. -uo. -UO.  2.° 2.4 2.0 3.3 3.6 3,1  TtMP  J  I.vi f 1 4 r I O N UF.F" i.t c Won IM.E-3 IU  NORMALIStO tT  CvJOC  -UO , -uo. -uo. -•4,1, •"»• -uo.  1 068, eu 7 , 23-47. 1 691., 3c u 0 , 3351 . T£G"Ai  FRACIURE  I1007U T 1 ll^vjo  CRI r iVfO'l  U30 786 V/i-'293 V '096 1 Extv.>jriUN  0.116 O.Cbu 0,316 0.228 0 , U Ob 0,u9S TfcKMINATEi)  UU.5 35.3 97 ,8 7 0 .U 135,0 139,0  0.260 0,266 0.32b 0.336 0,527 0 ,U /o  RTrt  ItR  » * «  —» ——  3,5b 3,o5 2.32 2,98  0,06 0,06 0.08 0,07 0,09 0.09  2.68  2.33  TOilUriStSS Cui)»  P«I  lui.a 157.1 M2U.7 M25.2 > I 1 7 .9 »116.2  1-43.8 >1 7 U . 4 157.1 >19IJ,9 230..1 205.8 19 .0 179,d 2 7 0 .8 2 7 J,1 27U.8 25'J.9 0  PL A N t STRAIN  If•  0.260 0.266 0,169 0.217 0.195 0 , 170  CCD J-  tT  fcl FT-t.nS  IN  0.979 0,9 79 1,310 1.36 1 1 ,36fl 1, uua  182 .6 162.6 270 .6 . 229 .6 317 .9 323.a  0.872 l.Ouo 1.785 1.363  3,075 2.739  IN  0.979 u .979 u . uSo 3 . 3 3 U  7.213 fl , 09U  3,616 3,02b 6.382 6, 035 11.570 11.969  A  A/w  IN  5,5 U.U 12,1 8,7 16,8 17.3  8.5 9.5 37.U 38,2 5U.5 54,9  0,079 0.079 0,079  0.20 0.20 0,20 0.20 0.20 0.20  0,079  0,079 0.079  VO  TR  OSULLAT 1UNS  FT-LH  Ft/S  *S  MS  201,6 201.8 201.8 201,8 252,2 252,2  11,3 11,3 11.3 11,3 12,7 12,7  tU  .0/29 ,0729 .0729 ,0729 ,0729 .0729  SIM (K.5I-SQR1 (lN)/S)«t+5 5,54 __ 5,9u o,3U 5.36 5.13 5.3b FLO* STRESS  CKIIt-HIA P*I IN  1U.0 13.9 U9.5 Ub,9 71,3 72,3  tf  INT  * 5 I - S H 1 ( 1 '<) 3,tfo U.23 7. H 6 b. 1 t 14.07 13,6i  TM  ti Ft-LH/IN.IN  112.8 111.7 396.8 37 7,9 57U.2 582.u  .01  c I'J-I.H/IN*I.T 1 0 vi 7 0 1 1 ii i 9 o Ub0242 U J 0 7 8 e> V /0<i93  TliV  7n27.. 7027 , 8331. 8331. o b i ? . 12197, obUJ, 1 0 3 9 B , • o 2 5 u , lu3i>'l. 6 1ou , 1 u575 ,  16.03 29 23 .13 55,5b 51 .6 4 p.,  p*i  "S  >7t>27. >8331. 57lo, 5o33. 5290,  >2?9.8 >251.1 172,3 169,7 159.(4 152.7  Tfc"H  PM  pt; V  H.SI  uSI >229.6 >251.1 185,8 184.9 173.9 169,2  .0332 ,0332 . 0332 ,0332 .0332 .0332  <  - 247 -  APPENDIX B DERIVATION OF THE CORRECTED ENERGY (EQUATION 2.9)  (57)  Notation: Subscript o indicates at moment of impact Subscript f indicates at moment of f i n a l rupture x  displacement  t  time  a  acceleration  v  velocity  E E  C  a  corrected  energy  energy calculated assuming constant velocity, V  E  Q  available impact energy  m  mass of tup assembly  F  applied  From f i r s t p r i n c i p l e s ,  E  Q  c o  (ma)dx x o  force  - 248 -  (minus sign due to fact that a < 0 during  impact)  (ma)vdt t o  >f  /if  mv(dv/dt)dt t" o  =  -  1 mvdv v  - h m(v-r  Also,  =  - v  o  o  )  t. E  a  =  - ^ mav dt Q  t o *f \ m(dv/dt)dt t" o  =  - v  Tf  I o  mdv  V  o =  - mv  =  2 —1 2 mv o  o  and, Eo  (v,. - v ) f o  N  (Eq. 2.8)  - 249 -  Now, mavdt  +  mav  o  dt -  \ mav dt \ o  o t. (mav  - mav )dt o  -  ma(v  - v )dt o'  - 1 \  ma(v  - v )dt o  +  \ mav dt 1 o  o t. mav dt o  o t. So,  E a  and, E  - E  c  Multiplying by  4E  a  =  - m  \ (v - v )adt \ o'  4 E ( = 2mv ) gives, Q  o  (E - E ) c a  Q  =  2 2 - 2m v o  \ (v - v )adt \ o o t,  =  - 2m V 2  2 Q  \ (v - V ) ( d v / d t ) d t Q  250 -  2 2 1 2m v \ (v - v )dv o \ o v o 2m v  o  \ (v - v )d(v \ o v  o (v = constant, so dv = 0) o o So, 4E (E - E ) o c a'  2 2 = - 2m v [ o  4E (E - E ) o c a  2 2 = - mv (v - v ) o f o  x  ( v  c  and, therefore.  E  c  E c  = E  a  - E  2 /4E a o  = E (1 - E /4E ) a a o  f " 2  - 251 APPENDIX C LISTING OF FORTRAN PROGRAM "ENERGY" FOR IIT DATA REDUCTION C C C C C  THIS PROGRAM CALCULATES VALUES FROM DATA OBTAINED WITH AN INSTRUMENTED CHARPY IMPACT MACHINE. A LOAD-TIME PHOTOGRAPH OF THE IMPACT EVENT IS FIRST ANALYZED TO OBTAIN THE AREA UNDER THIS CURVE FROM WHICH THE ENERGY ABSORBED IN FRACTURING THE SPECIMEN CAN BE OBTAINED. OTHER DATA IS SUPPLIED TO MAKE OTHER STRENGTH AND TOUGHNESS CALCULATIONS. DIMENSION TEMP (99) , YSS (99) , Y SD (99) , EO (99) , VO (99) ,TR (99) ,OSCIL(99) DIMENSION PGY(99) ,PM(99) ,PSI (99) ,TGY(99) , TM (9 9) , ET (99) , EI (99) DIMENSION EP(99),A(99) ,R (99),CM(99) ,DI (99),ETN (99), EIN(99) ,RTR(99) DIMENSION PIER(99),RJ(99) ,CODM( 99) , RKPMD ( 99) ,RKPSID (99) ,R JIC(99) DIMENSION RKCODH (99) ,CCD (99) , SIR (99) ,CCTS (9 9) , RKJ (99) , SFLOW (9 9) BEAL*8 CODE(99) LOGICAL*1 BL,GT,SWA (99) ,SWB(99) ,SHC(99) ,SWD (99) DATA BL,GT/' *,'>•/ DIMENSION SI (100),SIGMA1 (99),Al (99),B 1 (99),P1 (100) DIMENSION PSCPM (99) ,PSCPSI(99) ,PSCCOD(99) DIMENSION S2 (100) SIGMA2 (99) , A2 (99) , B2 (99) ,P2 (100) LOGICAL LK BEAD(5,5)LJ 5 FORMAT(12) DO 990 1=1,99 SWA(I)=BL SWB(I)=BL SWC(I)=BL SWD(I)=BL 990 CONTINUE DO 9999 J=1,LJ C F I I S THE AREA UNDER THE LOAD-TIME PHOTOGRAPH, UP TO THE POINT OF C MAXIMUM LOAD, IN SQ-IN. C FP IS THE AREA UNDER THE LOAD-TIME PHOTOGRAPH,FROM THE POINT OF C MAXIMUM LOAD, IN SQ-IN. C PMD,PGYD,TMD,TGYD ARE THE LOAD S TIME MEASUREMENTS ON THE PHOTO TO MAXIMUM C LOAD 8 ELASTIC LIMIT, RESPECTIVELY, IN INCHES C DH=DROP HEIGHT OF TUP, IN FEET C A=CRACK LENGTH,IN C YSS=STATIC YIELD STRENGTH,PSI C E= ELASTIC MODULUS, PSI C PR=POISON'S RATIO C S=SUPPORT SPAN,IN C W=SPECIMEN WIDTH,IN C B=SPECIMEN THICKNESS,IN C RR=ROTATTONAL RATIO, FOR COD CALCULATIONS C RLCF S TCF ARE FACTORS TO CONVERT THE MEASUREMENTS FROM THE PHOTO TO C LOAD & TIME ANALOGS (LB/IN S SEC/IN) C EO=IMPACT ENERGY VO=IMPACT VELOCITY READ(5,99) CODE (J) 99 FORMAT (A8) READ(5,10) T R ( J ) , DH,PMD,PGYD, 1TMD,TGYD,FI,FP 10 . FORMAT (F20. 8) E=31200000. PR=0.30 A(O)=0.079 S= 1.574 B=0.394 r  B=0.394  TEMP(J)=20. BR=0.33 TSS (J)=77300. BLCF=2985.1 TCF=0.000597  - 252 -  C C  C C C  C C C C C C C C C C C C  C C C C  C C  C C C  C  C  CF=RLCF*TCF EO(J)=100.875*12.*DH V O ( J ) = ( ( 9 2 6 6 . 1 1*DH)**0.5) A I AND AP ARE THE VALUES UNDER THE L O A D - T I M E C U R V E , UP TO AND FROM T H E POINT OF MAXIMUM LOAD, IN L B - S E C . WRITE{6,880) F I , F P 8 8 0 FORMAT{1X, 2 F 1 0 . 3 ) AI=FI*CF AP=FP*CF AT=AI+AP UT AND WI REPRESENT THE UNCORRECTED TOTAL ABSORBED ENERGY AND THE ENERGY REQUIRED TO I N I T I A T E A CRACK (CRACK I N I T I A T I O N I S ASSUMED T O OCCUR AT T H E POINT OF MAXIMUM L O A D ) . HT= (AT) * { V O ( J ) ) WI=(AI)*(VO(J)) T H E SOURCE OF MOST OF THE FOLLOWING EQUATIONS I S : SERVER,IRELAND, AND WULLAERT,"STRENGTH AND TOUGHNESS EVALUATIONS FROM AN INSTRUMENTED IMPACT T E S T " , DYNATUP T E C H N I C A L REPORT TR 7 4 - 2 9 R , E F F E C T S T E C H N O L O G Y , I N C . ,SANTA BARBARA, CA, 1 9 7 4 . PM = MAXIMUM LOAD DURING IMPACT EVENT (ASSUMED TO BE POINT OF CRACK INITIATION) PGY=GENERAL Y I E L D LOAD P * I - " E Q U I V A L E N T ENERGY" FRACTURE LOAD TM=ELAPSED TIME TO MAXIMUM LOAD TGY=ELAPSED TIME TO GENERAL Y I E L D LOAD Y SD=DYNAMIC YIELD STRENGTH SFLOW=FLOW STRESS PM(J) = (RLCF)*(PMD) PGY ( J ) = (RLCF) *(PGYD) TM ( J ) = (TCF) * (TMD) TGY (J) = (TCF) * (TGYD) A/W=CRACK DEPTH TO SAMPLE WIDTH RATIO R(J)=A(J)/W C H=MACHINE COMPLIANCE ( R E F . : I R E L A N D , I N S T R U K E N T E D IMPACT T E S T I N G , ASTM STP 5 6 3 , 1 9 7 4 , P P . 3 - 2 9 ) CALCULATED FROM: TOTAL COMPLIANCE MINUS SPECIMEN COMPLIANCE CM (J) = ( (VO (J) ) * (TGY (J) ) / (PG Y (J) ) ) 1 ( ( 7 2 . * ( 1 . 8625* (R (J) *R ( J ) ) - 3 . 9 5 * (R (J) **3) + 16 . 3 7 7 7 * 1 (R (J) **4) - 3 7 . 2 2 7 7 * ( R ( J ) **5) + 7 7 . 5 5 4 * (R (J) * * 6 ) - 1 2 6 . 8 7 2 7 * (R ( J ) * * 7 ) 1*172.5325* 2 (R (J) **8) - 1 4 3 . 9 64* (R (J) **9) +66. 564* (R (J) **1 0) ) +20 . ) / (E<B) ) OSCIL=PERIOD OF SPECIMEN O S C I L L A T I O N S . TGY AND TM SHOULD BE > 3 ( O S C I L ) FOR A V A L I D T E S T IN WHICH I N E R T I A L E F F E C T S ARE A V O I D E D . O S C I L ( J ) = 1 . 6 8 * S * ( (W/S) * * 0 . 5 ) * ( ( 7 2 . * ( 1 . 8 6 2 5 * (R (J) * R (J) ) - 3. 95* 1 ( R ( J ) **3) + 1 6 . 3 7 7 7 * (R (J) ** 4) - 3 7 . 2 2 7 7 * (R (J) **5) +77. 5 5 4 * (R (J) **6) 1 - 1 2 6 . 8 7 2 7 * ( R ( J ) * * 7 ) + 1 7 2 . 5 3 2 5 * ( R ( 3 ) * * 8 ) - 1 4 3 . 964* (R ( J ) * * 9 ) + 1 6 6 . 5 6 4 * (R (J) ** 10))+ 2 0 . ) * * 0 . 5 ) /1 9 6 8 5 0 . E T , E I , A N D EP ARE T H E CORRECTED T O T A L , I N I T I A T I O N , AND PROPAGATION ENERGIES OF THE IMPACT EVENT ( R E F . : G R U M B A C H , E T A L . , R E V U E DE M E T A L L U R G I E , A P R I L , 1 9 6 9 , P . 271) ET(J) = ((WT)*(1.- (WT)/({4.)* (EO(J)))))^12. EI ( J ) = ( ( W I ) * ( 1 . - ( W I ) / ( ( 4 . ) * ( E O ( J ) ) ) ) - ( P M ( J ) * * 2 ) * ( C K ( J ) ) / { 2 . ) ) / 1 2 . EP (J) =ET (J) - E I (J) E T AND E I ARE NORMALIZED BY D I V I D I N G BY THE SPECIMEN LIGAMENT AREA E T N ( J ) = ( E T ( J ) ) / ( ( B ) *{W-A(J) )) BIN (J)= (EI (J) ) / ( B * (W-A (J) ) ) DI=SAMPLE D E F L E C T I O N AT CRACK I N I T I A T I O N DI (J) = (TM (J) ) * (VO (J) ) * ( 1 . - (WI) / ( ( 4 . ) * (EO (J) ) ) ) - (PM (J) ) * (CM (J) ) I F ( P M ( J ) . E Q . P G Y (J) ) GO TO 999 T AN= ( (PGY (J) ) / ( (TGY (J) ) * (VO (J) ) * ( 1. - ( (PGY (J) ) * (TGY (J) ) * (VO(J) ) /  - 253 -  YSD(J)=YSD(J) *0.001 YSS(J)=YSS(J)*0.001 SF10W(J) = S F L O W ( J ) * 0 . 0 0 1 RKPMD(J)=RKPMD ( J ) * 0 . 0 0 1 RKPSID(J)=RKPSID(J)*0.001 RKCODM(J)=RKCODM(J)*0.001 RKJ(J)=RKJ(J)*0.001 S I F ( J ) = S I R ( J ) * ( 1 . E-8) CODM(J) =CODM(J) * 1 0 0 0 . I F ( P M D . EQ. PGYD) SWB(J)=GT I F (PM (J) . EQ. PGY (.7) ) SWA (J) =GT IF(PM(J) .GT.PGY(J) ) SWD(J)=GT I F (PM (J) . EQ. PGY (J) ) SWC(J)=GT 9999 CONTINUE WRITE(6,100) 100 FORMAT (* 1 • , 4 X , ' C O D E ' , 4 X", ' T E M P ' , 4 X , • YS S • , 6 X, *YSD« , 6 X , ' PGY* , 5X , ' PM' , 1 6 X , « P * I , 7 X , « T G Y ' 6 X , ' T M « , 6 X , ' E T ' , 5 X , «EI« , 4 X , • E P ' , 5 X , ' A • , 6 X , « A / W • ) WRITE(6,200) 2 00 F O R M A T ( 1 5 X , ' C , 1 0 X , ' K S I ' , 1 8 X , « L B S ' , 2 0 X , • M S ' , 1 4 X , • F T - L B S • , 1 0 X , • I N ' ) WRITE(6,300) 3 00 FORMAT ( 5 X , « ',4X,' , 4 X , •« , 6X, ' 1 — ' ,6X, ' ' ,6X, ' • ,2X, ' ,4X,« •) DO 42 J = 1 , L J WRITE ( 6 , 700) CODE (J) , T E H P ( J ) , Y S S (J) , SW A (J) , Y S D (J) ,SWB (J) , P G Y ( J ) , PM ( 1J) , P S I ( J ) , T G Y ( J ) ,TM (J) , E T ( J ) , E I ( J ) , E P ( J ) , A ( J ) , R ( J ) 42 CONTINUE 7 00 FORMAT ( 3 X , A 8 , 1 X , F 5 . 0 , 3 X , F 5 . 1, 2 X , A 1 , F 5 . 1 , 5 X , A 1 , F 5 . 0 , 2 X , F 5 . 0 , 1X , F 6 . 1 0 , 6 X , F 5 . 3 , 2 X , F 5 . 3 , 5 X , F 5 . 1 , 1 X , F 5 . 1 , 1 X , F 5 . 1 , 2 X , F 5 . 3 , 3 X , F 5 . 2) WRITE(6,400) 400 F O R M A T ( ' 0 ' , 4 X , ' C O D E ' , 4 X , ' T E M P • , 7 X , « C M • , 5 X , « I N I T I A T I O N ' , 7 X , ' N O R M A L I 1SED',6X,'RTR',4X,'IER*,4X,« E O ' , 6 X , • V O ' , 5 X , ' T H ' , 5 X , 1'OSCILLATIONS') WRITE(6,500) 5 0 0 FORMAT(3 1X,«DEFLECTION•,4X,* E T ' , 1 1 X , ' E I ' ) WRITE(6,550) 550 F O R M A T ( 1 5 X , ' C , U X , ' I N / L B * E - 6 • , 4 X , ' I N * E - 3 ' , 8 X , ' F T - L B / I N * I N ' , 1 1 9 X , ' F T - L B ' , 3 X , « F T / S • , 4 X , ' M S ' , 1 OX,•MS•) VRITE(6,570) 5 7 0 FORMAT (5X , ' ',4X,' , 3X, ' « ,2X, ' •,4X,' ^ i 3X * * # 3 X * —— —— * 3 X * — — — — — • 3 X j * — — 3 X * —— ' 16X,« •) DO 52 J = 1 , L J W R I T E ( 6 , 8 0 0 ) C O D E ( J ) , T E M P ( J ) , C M ( J ) , D I ( J ) , E T N ( J ) , E I N ( J ) ,RTH (J) , R I E R ( 1J) , E O ( J ) , V O ( J ) , T R ( J ) , O S C I L (J) 52 CONTINUE 800 F O R M A T ( 3 X , A 8 , 1 X , F 5 . 0 , 5 X , F 4 . 1 , 7 X , F 6 . 2 , 6 X , F 6 . 1 , 4 X , F 5 . 1 , 3 X , F 4 . 2 , 3 X , F 14.2,3X,F5. 1,3X,F4.1,3X,F5.4,6X,F5.4) WRITE(6,580) 5 80 F O R M A T ( ' 0 ' , 4 X , • C O D E * , 4 X , ' T E M P ' , 6 X , * J * ,6X,»CODM«,8X,•FRACTURE TOUGH 1 NESS',8X,«CCD',11X,'SIR') WRITE ( 6 , 5 9 0 ) 5 90 F O R M A T ( 2 0 X , • I N T E G R A L ' , 1 0 X , • P M • , 5 X , • P * I « , 4 X , ' C O D M ' , 4 X , ' J - I N T ' ) WRITE{6,525) 5 2 5 FORMAT ( 1 5 X , » C - , 1 X/» I N - L B / I N * ! N' , 1 X , ' I N * E - 3 ' , 1 OX, ' K S I - S Q R T (IN) », 1 1 1 X , ' I N ' , 4 X , • (KSI-SQRT ( I N ) / S ) * E + 5') WRITE(6,600) 6 00 FORMAT (5X,« ' ,UX,' ' ,3X,' «,2X, • • ,2X,' 1 ' ,2X, ' ' , 2X, ' •) DO 62 J = 1 , L J WRITE ( 6 , 9 0 0 ) CODE (J) , T E M P (J) , R J (J) ,CODM (J) ,SWD{J) ,RKPMD(«J) , R K P S I D ( J t  f  1  1  1  1  1  t  #  #  9  r  #  r  t  - 254 -  1 < ( 8 . ) * (EO (J) ) ) ) ) - (PGY (J) * (CM (J) ) )) ) P S I (J) = ( T A N ) * ( ( 1 2 . * 2 . * E I ( J ) / T A N ) * * 0 . 5 ) GO TO 1000 999 PSI(J)=PM(J) 1000 YSD (J) = ( 2 . 9 9 * P G Y (J) *W) / (B* { (W-A (J) ) **2) ) C B TR=RESPONSE TIME RATIO (SHOULD BE > 1 . 1 1 FOR A V A L I D T E S T IN WHICH C T H E E F F E C T OF SIGNAL ATTENUATION IS MINIMIZED) C I E R = I N I T T A T T O N ENERGY RATIO (SHOULD BE. <.33 FOR A VALID TEST) RTR (J) =TGY (J) /TR (J) RIER ( J ) = W I / E O ( J ) SFLOW (J) = ( 2 . 99* (PGY (J) +PM (J) ) *W) / ( 2 . * B * ( (W-A {J) ) **2) ) C S T R E S S I N T E N S I T Y PARAMETERS (RKPMD,RKPSID,RKCODM, RKJ) R K P M D ( J ) = ( ( 1 . 5 ) * S * ( P M ( J ) ) * (A (J) * * 0 . 5) / ( ( B ) * ( W * * 2 ) ) ) * ( 1 . 9 3 13.07*R(J)+14.53* 1 (R (J) * R (J) ) - 2 5 . 1 1 * ( R ( J ) **3) +25. 8*(R ( J ) * * 4 ) ) PSCPM (J) = ( 2 . 5) * ( ( (RKPMD (J) ) / (YSD (J) ) ) **2) BKPSID ( J ) = ( ( 1 . 5) *S* ( P S I (J) ) * ( A ( J ) * * 0 . 5) / ( (B) * (W**2)) ) * ( 1 . 9 3 - 3 . 0 7 * 1R(J)+14.53* 1 (R (J) *R (J) ) - 2 5 . 11* (R (J) **3) +25. 8* (R ( J ) * * 4 ) ) P S C P S I (J) = ( 2 . 5) * ( ( ( R K P S I D ( J ) ) / ( Y S D ( J ) ) ) **2) C J=J-INTEGRAL RJ (J) = 2 4 . * E I (J) / (B* (W-A (J) ) ) RKJ ( J ) = ( (E*RJ (J) ) * * 0 . 5) YSG=YSS(J) IF (YSD(J).GT.YSS (J)) YSG=YSD (J) C RJIC= J-INTEGRAL VALIDITY CRITERION R J I C ( J ) = 2 5 . * R J (J) /S FLOW (J) C CODM^CRACK-TIP-OPENING— DISPLACEMENT AT MAXIMUM LOAD C V A L U E OF CODM IS VERY MUCH DEPENDENT UPON THE ROTATIONAL C R A T I O ( R R ) , THE VALUE OF WHICH I S I N D E B A T E . CODM (J) = (2. 54) * (W-A (J) ) * (DI (J) ) * (RR) I F (PM(J) . E Q . P G Y ( J ) ) RKCODM(J)= ( (CODM (J) ) * (YS G) * ( E/(1 . 1(PR**2) ) ) ) * * 0 . 5 I F (PM(J) . G T . PGY (J) ) RKCODM ( J) = ( (CODM { J) ) * (YSD (J) ) *E) * * 0 . 5 RKL=RKCODM(J) ZZ=RKPSID(J) I F (RKCODM(J) . L T . R K P S I D ( J ) ) ZZ=RKCODM(J) I F (ZZ. L T . RKJ (J) ) RKL=ZZ I F (RKJ (J) . L T . ZZ) R K L = R K J ( J ) C C C D = C R I T I C A L CRACK DEPTH (REPRESENTS C R I T I C A L S I Z E OF A C H Y P O T H E T I C A L E L L I P T I C A L SURFACE FLAW S U B J E C T E D TO THE S T A T I C C Y I E L D STRENGTH WHICH WILL P R O P A G A T E . STRESS I N T E N S I T Y I S ASSUMED C TO BE THE MINIMUM VALUE C A L C U L A T E D ) . L E N G T H / D E P T H = 6 / 1 . I F (PM (J) . EQ. PGY (J) ) CCD (J) = (RK PM D (J) ** 2) / ( (1 . 2 1) *3 . 1 4 1 6* (YSS (J) * * 2 ) ) I F (PM (J) . G T . P G Y ( J ) ) C C D ( J ) = (RKL**2) / ( ( 1 . 21) * ( 3 . 1 416) * (YSS (J) **2) ) C CCTW=LENGTH OF CRACK EXTENDING THRU-WALL THAT WILL PROPAGATE WHEN C S U B J E C T E D TO S T A T I C Y I E L D S T R E N G T H . S T R E S S I N T E N S I T Y FACTOR USED I S C THAT CALCULATED BY J - I N T E G R A L T E C H N I Q U E . CCTW (J) = ( 0 . 7 1 4 4 * ( R K J (J) **2) ) / ( Y S S ( J ) * * 2 ) C SIft=STRESS I N T E N S I T Y RATE I F (PM (J) . EQ. PGY (J) ) SIR(J)=RKPMD(J)/TM(J) IF(PM(J).GT.PGY(J)) SIR(J)=RKL/TM(J) TH(J)=TM(J) *1000. TGY (J) =TGY (J) * 1 0 0 0 . CM{J)=CM(J) *(1.E+6) DI (J) = D I ( J ) * 1 0 0 0 . EO(J)=EO(J)/12. VO(J)=VO(J)/12. TR (J)=TR ( J ) * 1 0 O 0 . OSCIL(J)=OSCIL(J)*1000.  - 255 -  1) , S W C ( J ) ,RKCODM (J) , RKJ (J) , CCD (J) , SIR (J) 62 CONTINUE 9 00 FORMAT(3X,A8,1X,F5.0,4X,F5.0,4X,F5.2,2X,A1,F5.1,2X,F5.1,1X,A1,F5. 11,2T,F5.1,3X,F5.3,9X,F5.2) HRITE{6,505) 5 0 5 F O R M A T ( « 0 » , 4 X , ' C O D E * , 4 X , ' J - I N T E G R A L C R I T E R I O N • , 3 X , • P L A N E STRAIN 1CRITERIA', 9 X , ' C C T W , 7 X , ' F L O H STRESS•) WRITE(6,510) 510 F O R M A T ( 3 9 X , ' P M ' , 1 2 X , * P * I * ) WRITE(6,515) 515 FORMAT(22X,«IN« ^ X ^ I N ' ^ X ^ I N ' , 1 2 X , ' K S I « ) WRITE(6,520) 5 2 0 FORMAT ( 5 X , ' *,4X , • ' , 3X , ' • 1,3X,« «,8X,« «, 6X, • ') DO 72 J = 1 , L J W R I T E ( 6 , 5 5 5 ) C O D E ( J ) , R J I C ( J ) , P S C P M ( J ) , P S C P S I (J) , C C T W ( J ) , S F L O W ( J ) 72 CONTINUE 5 55 F O R M A T ( 3 X , A 8 , 1 0 X , F 6 . 3 , 1 2 X , F 6 . 3 , 6 X , F 6 . 3 , 1 0 X , F 6 . 3 , 9 X , F 5 . 1 ) STOP END ,  - 256 APPENDIX D  INSTRUMENTED IMPACT TEST RECORD  (CODE) Specimen  (TEMP) Temperature °C  Load Scale mV/div  Time Scale mS/dlv  (TR) Response Time, us  (VO) Impact V e l o c i t y ±n/s  (EO) Impact Energy in /lb  (DH) Drop Height, F t  (CF) (TCF) lb-s Time Conversion F a c t o r , s / i n Area Conversion F a c t o r , rin  (RLCF) Load Conversion F a c t o r , l b / i n  Impact Photo Measurements  OTP)  Area to Max Load, i n  Area from Max Load, i n ^  (PMD) Max. L o a d , l n  (PGYD) Gen'l Y i e l d Load.In  (S) Span, In  (B) Thickness, i n  (TMD) Time t o Max. L o a d , i n  (W) Width, i n  (TGYD) Time to Gen'l Y i e l d L o a d , i n  Crack Length Center, mm ( hi p o i n t , mm( 3/4 p o i n t , mm(  ) -( ) -C ) - (  ) = ) = ) -  (A) , Crack Length, i n  (  x(0.03937 in/mm) Shorter s u r f a c e , mm (YSS) S t a t i c Y i e l d Strength, l b / i n  2  (  (E) E l a s t i c Modulus, l b / i n  2  ) - (  )  (PR) Poisson's Ratio  Precracking Data: Max. Applied Torque, i n - l b  Kf(max.),psi-in"  Fatigue C y c l e s  )  - 257 -  APPENDIX E STRAIN-AGE  STUDY: CALCULATIONS  AF-1 Pipe Seam Welding Parameters: welding speed:  12.7 mm/s  amperage:  787.5 amps  voltage:  31.5 volts  2-pass weld  Calculation of Heat Transfer Efficiency, f^:  From:  CM. Adams, Welding Handbook, 7th Ed., Vol. 1, pp.80-98, AWS, Miami, 1976.  1  T - T p o where,  T^  =  =  A^PCtY H^ net  1 +  T - T m o  peak temperature (°C) at distance, Y(mm), from weld fusion boundary  T  q  =  i n i t i a l temperature (= 25°C)  T  FFI  =  liquidus temperature (- 1510°C)  =  net energy input  H^  et  (Eq. E.l)  E  = volts  =  I =  f^El/V amperage  f^ =  heat transfer efficiency  V  travel speed (mm/s)  =  - 258 -  i  pC  = volumetric specific heat 0.0044 J/mm -°C) 3  t  = thickness of pipe (mm)  By macroetching, heat-affected zone boundary determined to be 4.5 mm  = ^YiAZ  The peak temperature T , for the visible HAZ boundary in low alloy,steels i s approximately 730°C.  So, from Equation E . l :  1 730 - 25  _ 4.13(.0044)(13.7)(4.5) " [(y (31.5) (787.5)/12.7] f=  0.77  1  1 1510-25  for AF-1 pipe  Calculation of Peak Temperature at Various Distance from Seam Weld:  for  1 T - 25 P  Y = 0.6-in(15.2 mm), using Equation E . l :  _  T = P  4.13(.0044)(13.7)(15.2) [(.77)(31.5)(787.5)/12.7] 337°C  1 1510-25  - 259 -  Distance  From  Weld  Fusion  Y (mm)  Figure E . l Temperature gradient in Charpy specimens from near seam weld. Similarly, for the Charpy specimen shown above, for  and for  Y  =  19.2 mm (distance to notch)  T = P  285°C  Y  11.2 mm (distance to bottom of specimen)  =  T = 421°C P  Boundory  - 260 -  Calculation of Cooling Rate: for relatively thin plates:  R  =  2irkpC ( t / H  where, R = k  =  ) (T 2  N E T  C  - T )  (Eq. E.2)  3  Q  cooling rate (°C/s) thermal conductivity (= 0.051 J/m s°C for ,  steel at 300°C)* T  c  =  temperature (°C) at instant at which cooling rate applies  Approximation of the cooling rate at a point 15.3 mm from the weld fusion boundary, when  = 337°C, can therefore be calculated:  R  -  2,(.051)(.0044)[ .  R  =  3.5°C/s  (  It i s recognized that R, k  77)(31  .5)(787.5)712.7  ] 2 ( 3 3 ?  ^  = f(T) and that R i s s t r i c t l y valid only  for the weld center line. Temperature for optimum strain-age effect i s approximately 285°C^^ ^: 3  Material below Charpy specimen in Figure E . l goes from 337°C to 285°C i n : (337 - 285)°C/3.5°C/s  *  = 15 s  From: BISRA Report "Physical Constants of Some Commercial Steels at Elevated Temperatures,"BISRA, London, 1953.  - 261 -  Pipe i s welded in two-passes, so a given point near the seam weld experiences temperatures optimum for strain-aging for 2(15 s) - 30 seconds.  N.B.  Time between welding passes i s 132 s; so weld bead cools to ambient temperatures between passes.  Calcuation of Times and Temperatures to Approximate Strain-Age Conditions near Seam Weld:  Assuming that the effective aging time and temperature near the seam weld was 1 minute at 316°C(Y = 16.7 mm), the times/ temperatures required to a r t i f i c a l l y age Charpy specimens an equivalent amount may be calculated f r o m ^ ^ : 2  log t _ / t ]  where,  ^  =  2  < T  2  7500[1/T - 1/T ] 1  2  (°K)  Table E . l l i s t s equivalent aging conditions. Table E . l Time(minutes)  Temperature(°C)  60  244°  15  266°  5  285°  1  316°  0.5  330°  (Eq. 3.7)  

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