ML20079D057
| ML20079D057 | |
| Person / Time | |
|---|---|
| Site: | Comanche Peak  |
| Issue date: | 03/31/1991 |
| From: | Matt Bartlett, Burghard H SOUTHWEST RESEARCH INSTITUTE |
| To: | |
| Shared Package | |
| ML20079D042 | List: |
| References | |
| NUDOCS 9106270123 | |
| Download: ML20079D057 (145) | |
Text
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, SOUTHWESTRESEARCH IN S TIT U T E uro cutceaa noso . post orsics onawra rsiio saw auto =io.itras. usa terre.oiio isi 3 e4.siti . titex e.. ...
ANALYSIS OF CIIECK VALVE SWING ARMS I Task 1 Evaluation of Commercial Swing Arms FINAL REPORT SwRI Project No. 06 3893100 I to TU Electric Co.
Comanche Peak S.FJ.
I P.O. Box 1002 Glen Rose,TX 76043 I
March 1991 I
APPROVED:
/\ d '
Gerald R. Levermt, Director Materials & Mechanics Department I.
Prepared by: j
- 11. C. Burghard, Jr. Thomas C. Trbovich, Manager M. L. Bartlett Quality Assurance r
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SAN ANTONIO. TEXAS HOUS TON. TEXAS - DETRolf wCHtGAN
- WASHINGTON DC 9106270123 910621 l I PDR ADOCK 05000445 l P PDR I I
Table of Contents l 1.0 I NTR O D U CTI O N . . . . . . . . . . . . . . . . . . . . . . .. . . .. . . . . .. . . . . . . . . . . . . . . . .. . . . . . .. . . .. . . . . . . .. . . .. .. . . . . .. . . .. . . . . . . . . . . 1 - 1 1.1 Backgmund....................................................................................................1-1 1.2 S am ple M a te ri al s .. .. .. .. ... . .. . . . .. .. .. .. . . .. .. .. .. . . .. .... .. .. .. .. .. .. .. .... .. .. .... .. .. .. .. .... .. .. . . . . . 1 - 1 1.3 S cope of Laboratory Evaluation ............................................................... .... 1 -2 2.0 NOND ESTRUCTIVE IN S PECTION S ............ ................................... ................. 2- 1 Visual and Lic uld Penetrant inspections ...................................................... 2 1 I 2.1 2.2 Radiographic : n spections .... .. .. ...................... ...... ...... ...... .. .. .. .............. .... .. .... 2- 1 3.0 MECHANICAL PROPERTIES, COMPOSITION AND RESIDUAL STRESS .. 3-1 I
3.1 Hardnes s and Te nsile Prope nie s ................................................................... 3- 1 3.2 Che mic al Compos ition .. .... . . . . .. .. . . .. .... .. ...... .. .. .. .. .. .. .... .. .. .. . . .. .... .. .. .. .... .. .. ...... .. 3 -2 3.3 R e s id u al S tre s s . . .. . . . . . . . . . . .. . . . . .. . . . . . . . . . . . . . . .. . . .. . . .. . . .. . . . . .. . . .. .. . . . . .. . . .. .. . . . . . . .. . . . . . . . . . .. . 3 - 3 4.0 META LLOG RAPHIC EVA LUATION ................................................................. 4- 1 4.1 Mic ros truc t ural Fe at um s .. .. . . ... . . . . . . . . . . . . . .. .. .. .. .. .. .... .. .. .. .. .. .. .. .... .... .... .. .. . . .. . . . . .. 4- 1 4.2 We l d R e p airs . . .. . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. . . . . . . . . . . . . . . .. . . .. . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. .. . . . . . . .. . . . 4-4 I 5.0 FLAW CH A RA CTERIZATIO N . . . . .. .. .. .. .. . .. .. .. .. . . .. .. .... .. .. . . . ... .. .. .. .... .... . ... .. .. .. .. .. . 5 - 1 5.1 Radiographic Flaw Indications .. ................................................................... 5- 1 I 5.1.1 5.1.2 5.1.3 5 ampie No . 2 40 (l 6 -in.) .. .. .. . . .. .. .. .. .. .. .. .. . . .. . .. . . .. .. .. . . .. . .. .. .. .. . . .. .. .. .... . . 5 - 1 Sample No. 255 (10.in.) .. .. . . .. .. ...... .. .... .. . . . . .. .. .. .. .. .. .. ... . .. .. ... . .. .. .. .. .. .... 5 -2 Sampie No. 242 (8.in.) . . . . .. ...... .. .. .. .. . . .. . .. . .. ..... . .... . . .. .... .. .. .. .. .. .. . . . ... .. 5-3 5.2 OtherFlaws...................................................................................................5-4 5.3 Fl a w S u mm ary . . . . . . . . . . . . . . . . . . . . . .. . . . . . . . . . . . . .. .. . . . . . . . . . . . . . . . . . . .. . . . . . . .. .. . . . . .. . . . . . . . . . . . . . . . . . . . 5 5 6.0 FRACTU RE TOU G H NES S .. .. .. . ... .. . . .. . ... .. .. . . . . .. .. .. . . .. .. . . . . . . . . .. .. . . . .. .. . . . . .. .. .. .. .. .... .. .. . 6- 1 6.1 Te s t Proced ure s ... . .. . . . . . . .. .. . .. .. .. . . .. .. . . . .. .. . . . . .. .. .. .. . ... . . .. . . .. . .. .. .. . .. .. . . .. .. .. . .... . 6- 1 6.2 Te s t R e s uIt s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. . . . . . 6- 3 6.3 D at a S t a ti s tics .. .. . . . .. .. . .. . . . . . . . .. .. .. .. . . . . . . . .. .. . . .. .. .. .. .. .. .. . . .. .. . . . .. .. -...... 6-6 7.0 S U M M A RY A ND DlS C U S S IO N . . .. .. .. .. .. .. .. .. .. .. .. .... .. . . .. .. .. .. .. .. .. .. . . .. .. .. .. .. .. .. .. ... . .. . 7- 1 8.0 RE FE RE N C ES . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. . . . . . . . . . .. . . . . . . .. . . . . . . . . .. . . .. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 8 - 1 APPENDIX A- SWING ARM CONFIGURATIONS A-1 APPENDIX B - MICROSTRUCTUREOFSWING ARMS B-1 APPENDIX C - EQUATION FOR THE STRESS INTENSITY C-1 FACTOR (K 3) FOR A SURFACE CRACK APPENDIX D - LOWER BOUND VALUE FOR THE FRACTURE D-1 TOUGHNESS TEST DATA BY USING l STATISTICAL METHODOLOGY APPROACH I l I
LIST OF TABLES E
L Inhlt EREC e
1-1 Swing Ann Sample Identification 1-3 2-1 Summary of Weld Repairs 23 2-2 Radiographic Inspection Results 2-4 3-1 liardness Data 35 3-2 Tensile Properties Data 3-6 3-3 Chemical Composition 3-7
[
3-4 Residual Stress 3-8 5-1 Fracture Mechanics Parameters for Largest Observed Flaws 5-8 6-1 Fracture Toughness Test Results 6-7 I
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LIST OF FIGURES Figure Eage 1-1 Schematic Diagram of Check Valve 1-4 12 Swing Arm from 16-inch Check Valve 15 13 Swing Arm Samples 16 2-1 Weld Repairs in Swing Arms 25 l 2-2 Weld Repairs in Swing Arms Schematic Diagram of Metallographic Section locations 26 4-7 4-1 4-2 Microstructural Featmts of Cast 17-4PH Swing Arms 4-8 4-3 Microstructural Featums of Cast 17-4PH Swing Arms 4-9 4-4 Microstmetural Features of Cast 17-4PH Swing Arms 4-10 4-5 Micmstructural Features of Cast 17-4PH Swing Arms 4-11 4-6 Microstructural Features of Cast 17-4PH Swing Arms 4-12 4-7 Microstructure of Sample No. 260 (4-in.) 4-13 4-8 Micmvoids in Cast 17-4PH Swing Arms 4 14 4-9 Microvoids in Cast 17-4PH Swing Arms 4-15 4-10 Micmstructure of Weld Repair - Sample No. 239 (16-in.) 4-16 l 4-11 HAZ of Weld Repair in Sample No. 240 (16-in.) 4-17 4-12 Microstructures of Weld Repairs in Samples No. 243 (8-in.) and 256 (6-in.) 4-18 4-13 Micmstructure of Weld Repair - Sample No. 245 (4-in.) 4 19 5-1 Locations of Radiographic Indications in Sample No. 240 (16-in.) 5-9 5-2 Internal Fissure in Sample No. 240 5-10 5-3 Micmfissure in Sample No. 240 (16-in.) 5-11 5-4 Diagram of Internal Fissure in Sample No. 240 (16-in.) 5-12 5-5 Transverse Sections from Sample No. 240 5-13 5-6 Transverse Sections from Sample No. 240 5-14 (Continued) m aamm mur lv
- LIST OF FIGURES (Continued) 5 L
ElELIIC f.RER 5-7 Diagram of Internal Flaw in Sample No. 240 (16-in.) 5-15 5-8 Longitudinal Sections from Sample No. 255 5-16 5-9 Longittidinal Section from Sample No. 255 (10-in.) 5-17 5-10 Transverse Section from Sample No. 255 (10-in.) 5-18 5 11 Diagram of Internal Voids in Sample No. 255 (10-in.) 5-19 5 12 Transverse Sections from Sample No. 242 (8-in.) 5 20 5-13 Transverse Section from Sample No. 242 (8-in.) 5-21 5-14 Transverse Section from Sample No. 242 (8-in). 5-22 5-!5 Transverse Section from Sample No. 242 (8-in.) 5-23 5-16 Diagram of Internal Flaw in Sample No. 242 (8-in.) 5-24 5-17 Surface Connected Flaw in Sample No. 255 (10-in.) 5-25 5-18 Diagram of Surface-Connected Flaw in Sample No. 255 (10-in.) 5-26 5-19 Micmfissures in 4-in. and 3-in. Swing Arms 5-27 5-20 Surface Connected Voids in Sample No. 243 (8-in.) 5-28 5-21 Fracture Stress Vs. Crack Depth for Km = 50 kstM 5-29 5-22 Fracture Stn:ss vs. Crack Depth for Typical Crack Geometry 5-30 6-1 Proportions for Compact Tension Specimens Employed in Km and Jo Tests 6-8 6-2 Type A Fracture Toughness Specimen (1.5W) 6-9 6-3 Type B Fracture Toughness Specimen (1W) 6-10 6-4 Type C Fractun: Toughness Specimen (0.76W) 6-11 6-5 Typical Location and Orientation of Fracture Toughness Specimen in 6-12 Swing Arms 6-6 Periodic Load vs. Displacement Unloads for a Je Fracture Toughness Test 6-13 6-7 J-Integral as a Function of Crack Growth for Analysis of a Je Test 6-14 (Continued) l tvuwmmmuoc v
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LIST OF FIGURES (Continued)
Figure Eagt 6-8 Histogram of K ic Data for Cast 17-4PH Swing Arms 6-15 6-9 Fractum Surface of Ji c Specimens from 16-in. Swing Arms 6-16 6 10 Fractum Surfaces of Sub-Size Ji e Specimens fmm 16-in. Swing Arms 6-17 6-11 Fractum Surfaces of Ji c Specimens from 10-in. and 8-in. Swing Arms 6-17 6-12 Fractum Surfaces of J ic Specimens 6-18 A-1 Sample No. 239,16-inch Check Valve A-2 A-2 Sample No. 240,16-inch Check Valve A-3 A-3 Sample No. 255,10-inch Check Valve A-4 A-4 Sample No. 242,8-inch Check Valve A-5 A-5 Sample No. 243,8-inch Check Valve A-6 A-6 Sainple No. 244,6-inch Check Valve A-7 A-7 Sample No. 256,6-inch Check Valve A-8 A-8 Sample No. 257,6-inch Check Valve A-9 A-9 Sample No. 245,4-inch Check Valve A-10 A-10 Sample No. 258,4-inch Check Valve A-11 A-11 Sample No. 259,4-inch Check Valve A-12 A-12 Sample No. 260,4-inch Check Valve A-13 A-13 Sample No. 246,3-inch Check Valve A-14 A-14 Sample No. 261,3-inch Check Valve A-15 A-15 Sample No. 262,3-inch Check Valve A-16 A-16 Sample No. 263,3-inch Check Valve A-17 B-1 Typical Microstructure of Sample No. 239 (16-in.) B-2 B-2 Microstructure of Sampic No. 240 (16-in.) B-3 B-3 Typical Microstructure of Sample No. 255 (10-in.) B-4 (Continued) i mmammmmow vi
LIST OF FIGURES (Continued)
Figure Eage B-4 Typical Microstructure of Sample No. 242 (8-in.) B-5 i B-5 Microstructure of Sample No. 242 (8-in.) B-6 B-6 Microstructure of Sample No. 243 (8 in.) B-7 B-7 Typical Micmstructure of Sample No. 244 (6-in.) B8 l B-8 Microstructure of Sample No. 244 (6-in.) B-9 B-10 B-9 Microstructure of Sample No. 256 (6-in.)
B-10 Microstructure of Sample No. 257 (6-in.) B 11 B-11 Microstructure of Sample No. 245 (4-in.) B-12 B-12 Microstmeture of Sample No. 258 (4-in.) B-13 l B-13 B-14 Typical Microstructum of Sample No. 259 (4 in.)
Microstructure of Sample No. 260 (4-in.)
B-14 B 15 B-15 Typical Microstmeture of Sample No. 246 (3-in.) B 16 B-16 Typical Microstructu:e of Sample No. 261 (3-in.) B 17 B-17 Typical Microstructure of Sample No. 262 (3-in.) B-18 B-18 Typical Microstructure of Sample No. 263 (3-in.) B-19
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1.0 INTRODUCTION
1.1 DJtchtt011ad I A program was undertaken at the Comanche Peak Steam Electric Station (CPSES) to evaluate the suitability for seivice of the swing arms in a specific group of check valves supplied by a single manufacturer and installed at the station. Con anche Peak is a nuclear-powered electric generation station consisting of two 1150 Mw units. The valve swing arms selected for testing in the program were removed from valves ranging in size fmm 3 inch to 16-inch. The particular valves involved were fmm a repmsentative population of Unit 1 and Unit 2 systems. A schematic diagram of one type of check valve, illustrating the overall configuration and function of the swing arms, is shown in Figure 1-1.
The material of the valve bodies from which the sample swing arms were obtained is cast stainless steel (Type 316) in three cases and cast carbon steel for the remainder. The swing arm material was selected by the valve manufacturer and called out on vendor drawings as 17-4PH (H1100), but there ara no certified test or inspection records for these components. Actual detailed material specifications were provided from the drawing for4-inch swing arms. The manufacturing drawings for one of the 4-inch valves called out the swing arm material as 17-4PH per AMS 5398
[l], heat treated to condition H1100 per MIL-H-6875 [2]. The curmnt ASTM speification for cast 17-4PH stainless stee; is A 747 CB7Cu-l [3].
As part of the overall CPSES program, certain nondestructive inspections, mechanical properties tests, and metallurgical evaluations of the selected swing arms were performed at Southwest Research Institute (SwRI). The procedures employed and the results obtained in those inspections and evahations are presented in this repon.
1.2 Sample n1aterials A total of 16 intact swing arms v.ere fumished to SwRI by CPSES for the purposes of this program. The particular valve sizes included in the group were as follows:
2 inch 3 inch 1 inch 4 inch 2 mch 4 inch I
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I l A listing of the individual samples with their CPSES identification numbers and the cormsponding sample numbers assigned for the purposes of the laboratory evaluations is presented in Table 1-1.
In general, the swing arms are configured as a curved bar of rectangular cross section with cylindrical bosses at each end. The axes of the two bosses are oriented perpendicular to one another with one boss drilled to accept a hinge pin and the other drilled to accept the valve disk stud. Pho:ographs of selected samples illustrating the general sizes and configurations are shown in Figures 1 2 and 1 3. Photographs of each individual sample and diagrams showing their principal dimensions are contained in Appendix A.
1.3 Scone of Laboratory Evaluation I The laboratory evaluations were organized to establish the mechanical properties and the metallurgical characteristics of all the sample swing arms. Each sample was inspected visually and at low magnification (10-50X) and by radiographic procedures.* Fracture toughness tests were performed on compact tension specimens from all 16 samples, uniaxial tensile tests were perfortned on selected samples and hardness measurements were made on all samples. Metallographic examinations of selected sections from all samples were perfonned to identify microstructural features and to characterize cenain flaws identified in the nondestructive inspections. These examinations also served to expose any flaws in the arms at the transition to the valve boss disk **
which might have gone undetected in the nondestructive inspections. In addition, the bulk chemical compositions were determined for all 16 samples.
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- Liquid penetrant inspections of all 16 swing arm samples were performed at CPSES prior to shipment to SwRI.
- This region was identified as the zone of highest in-service bending stresses in earlier I investigations at CPSES.
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SWING ARM SAMPLE IDENTIFICATION Valve SwRI R CPSES Site Sample No. Valve No.
16-in. 239 2CT-0025 240 2CT-0077 10-in. 255 2CC-0317 8-in. 242 2AF-0038 243 2AF-0167 l 244 2AF 0014 6-in. 256 2FW-0198 257 2CH-0024 245 1 AF-0098 4-in. 258 1 AF-0093 259 1 AF-0101 I
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1 AF-0106 246 1Cll-0346 3-in. 261 2CC-003 262 2CC-004 263 XDD-0048 uuncaantumwoc l-3
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2.0 NONDESTRUCTIVE INSPECTIONS All of the swing arm samples were inspected visually and by liquid Duorescent penetrant (LP) techniques, in those cases where the sisual or LP inspections indicated possible significant flaws, the areas of interest were examined at low magnincation (10-50X)in the stercomicroscope. Each individual sample was also radiographed to provide for identincation of internal Daws. In the course of the visualinspection, each individual sample was immersion etched in Fry's reagent
- to highlight any major variations in structure and possible weld repairs.
2.1 Visual and Liould Penetrant Insoections In general, no significant external flaws were noted in the visualinspections. In certain cases, (mainly the larger arms), relatively deep surface pits were noted but these were generally identified as smooth, open features in the stereomicroscope examinations and could not be classiGed as detrimental Daws. Weld repairs were evident on most of the large arms and in several cases considerable weld repair had been performed on the disk attachment bosses. The etching response of the weld deposits suggested that the filler material was an austenitic stainless steel (Type 308 or similar). A summary of the observed weld repairs is presented in Table 2-1 and photographs of representative examples are shown in Figures 2-1 and 2-2. In the case of the two 16-inch swing arms, the weld repairs in the disk attachment boss were such that. after final machining, the weld deposit extended completely through the radial thickness of the boss [ Figure 2-1(a)]. Metallographic examinations established that the depths of all other weld repairs were on the order of 0.06-inch. The liquid penetrant inspections performed at CPSES did not produce any recordable flaw indications. Certain surface features did retain penetrant to a slight degree, however, and these were marked for stereomicroscope examination 2.2 Radiocrnohic Insoettions All 16 swing arms were inspected radiographically. The inspections w ere performed with an X-ray source operated at 160-290 KV and 4-5 Ma. Either Kodak T or Kodak AA film was
- ASTM E 407, No. 79. 40 ml hcl + 5 g CuCl 2+ 30 ml H O 2 + 25 mi ethanol mmemmune 2-1
l used and the source to film distance for all exposures was 27 inches, in all cases except for Sample No. 255, the radiographs were either free of recordable indications or exhibited only isolated or scattered minor flaw indications in the arm portion of the sample. These were classified as 1.evel 1 i inclusions or voids to Level 3 shrinkage in accordance with ASTM E 446 reference radiographs. The radiograph for Sample No. 255 (10-inch ann) exhibited larger flaw indications which were classified as Level 6 shrinkage. These were located in the straight portion of the arm approximately 3 inches from the disk attachment boss. A summary of the radiographic inspection results is given in Table 2-2. I I . l l l l , 5 t l__ _
TAllLE 21
SUMMARY
OF WELD Rr PAIRS Valve Sample I Location and l Site No. Extent of Repair 239 Through-thickness repairs in disk attachment boss. 15-in. 240 10-in. 255 Shallow repairs in disk attachment boss and arm.* 242 Shallow repairs in disk attachment boss and ann. 8-in. I 243 244 Shallow repairs in disk attachment boss and arm.* 6-in. 256 No repairs. 257 245 Single repair in arm at disk attachment boss. 4-in. 258 259 No repairs. 260 I 246 No repairs. 261 Single repair in ann. 3-in.
.. 262 Shallow repairs in disk attachment boss.*
263
- Metallographic examinations established that those welds designated as " shallow" were on the order of 0.06-in. depth.
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L TABLE 2 2 RADIOGRAPillC INSPECTION RESULTS Valve Sample Size No. Radiographic Indications' Two Level 3 CA shrinkage. Scattered Level 1, 16-inch 239 CA shrinkage. 240 Two Level 3 CD shrinkage. 10-inch 255 Several Level 6 inclusion / void indications. 242 Isolated Level 2 CA shrinkage. 8-inch Scattered Level 1 inclusions and Level 2, CA 243 shrinkage. I - Scattered Level 2 CA shrinkage and Level 1 244 R, inclusions / voids. 6-inch 256 Single Level 1 CA shrinkage and scattered Level 1 inclusions / voids. 257 Isolated Level 1 inclusion / void. 245 No recordable indications. 4-inch 258 259 Scattered Level 1 inclusions / voids. 246 No recordable indications. 3-inch 260 Isolated Level 1 inclusions / voids. 261 262 No recordable indi;ations. 263
- Indications are classified in accordance with E ASTM E 446 reference radiographs. [Ref. 8, pg. 8-1) awancmemunoc 2-4
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s 3.0 MECilANICAL PROPERTIES, COMPOSITION AND RESIDUAL STRESS ~ 3.1 liardness and Tensile Properties Conventional Rockwell hardness measurements were made on longitudinal sections from each of the 16 swing arms. The sections were selected from those cut for metallographic examination (see Section 4.0) and were alllocated at thejunction of the arm with the disk attachment boss. In those cases where multiple sections were examined, five hardness measurements were l made on each of two sections, one near the edge and one near the centerline of the arm. In cases where only one se ction was taken, the hardness was measured on the side of the specimen nearest the centerline. In all cases, the section was ground to a 400 grit finish prior to measurement. The results of the hardness measurements are presented in Table 3-1. The average hardness values for the entire group of samples, c', rep f." Va. 260, were in the range of HRC 28 40 (BHN 271-372).* These values equal or exceed the minirnum value of HRC 28 (BHN 271) specified by ASTM A 747 for CB7Cu-l (H1100) cast stainless steel." There were no significant variations in hardness among duplicate sections of any one sample or among various locations on any one section. The range of approximate ultimate tensile strengths corresponding to the range of measured hardness values is 130-182 ksi. The hardness of Sample No. 260 was measured as HRC 21-HRB 96 (HHN 224). This l value is substantially below the range of values for the other 15 samples and below the specified minimum hardness for CB7Cu-1 (H1100) material. The corresponding approximate ultimate strength for this hardness level is 106 ksi. It is ofinterest to note that the metallographic examina; ion established that Sample No. 260 was improperly heat treated (see Section 4.1). I The majority of the 15 samples exhibited hardness values in the range of HRC 35-40 (BHN 322-372).
** The hardness values and strength requirements specified for A 747 CB7Cu-l (H1100) material are cited in this section for reference only. The swing arms were not manufactured to specifically conform to that specification. Neither AMS 5398 cr MIL-H-6875 specify hardness or strength requirements for the H1100 condition.
newnmom>wx 3-1
s liardness measurements were au .nade on two laboratory heat treated specimens p U which had been cut from one of the 16 in. swing arms. One specimen (No. 239-4a) had been given - a complete heat treat cycle including homogenization, solution annealing and aging to produce the ill100 condition. The second specimen was simply reaged at 1100*F. In both cases, a slight increase in hardness was noted (HRC 35 36 vs. IIRC 32, see Table 31). This change in properties demonstrated that the thermal treatment that produced the initial condition was somehow different l from the laboratory full cycle treatment and that the initial condition was not fully aged. These observations at: consistent with the fact that both of the 16-in. anns exhibited a unique microstnicture (see Section 4.1). liowever,it should be noted that the initial hardness is consistent with the H1100 condition and that the microstructure indicates a heat-treated condition (not as-cast). Unia@d tensile tests were perfomied c.n specimens from four of the swing ann l samples. The particular samples selected for testing were those where sufficient material for tensi'c specimens remained after removal of the fracture toughness specimen blanks. These included one 16-in.,two 8-in., and one 6-in, ann The tests employed sub-size,round specimens per ASTM A 370 with 0.25-in dia. x 1.0-in. Ig. gage sections. The results are presented in Table 3 2. In two cases (Sample Nos. 239 and 243), the measured tensile properties conformed to the requhrments of ASTM A 747 for cast CB7Cu-l stainless steelin the 111100 condition. The yield strength for Sample No. 242 was below the minimum value specified for CB7Cu l (H1100) material (104.5 ksi vs 110.0 ksi) while the ultimate strength was relatively high (174.6 ksi vs 135.0 ksi min.). The large dif ferenee between the ultimate and yield strength is unusual for cast 17-4 (111100) material. In the case of Sampie No. 244, the yield strength met the minimum specified by A 747 CB7Cu l (111.7 ksi vs 110.0 kri) but the ultimate strength was below the minimum value requirrd by that specification (129.1 ksi vs.135.0 ksi). 3.2 Chemical Composition Chemical analyses were performed on specimens from each of the 16 samples to determine the bulk chemical composition. Particular care was exercised in the removal of the individual specimens to avoid inclusion of any repair weld deposits. In one case, an additional m ummumuu 3-2
s specimen containing only weld metal was analyzed. The r*sults of the analyses are listed in Table 3 3. The composition limits specined by A 747 Cll7Cu-l and AhtS 5398 and those for Type 308 austenitic stainless steel welding filler are included in the table for comparison. The composhionallimits of A 747 CBCu l and AhtS 5398 do not correspond exactly, The principal difference is that the maximum chromium content for AhiS $398 is lower by 1% (16.70% vs. 17.70%). The composition limits do overlap, however, so that simultaneous confonnance to both l specifications is possible. The compositions of 100f the sixteen samples confonned to the requirrments of A 747 CllCu l and AhiS 5398 except for a few minor variations.* Five of the remaining six samples (Nos. 239,240,257,262,263) had chromium contents slightly below the minimum values required by the two specifications cited.*
- l In the case of Sample No. 260, the nickel content was substantia'ly abave the specified maximum (6.12% vs 4.70% max.) and the copper was below the ndnimum (1.96% vs 2.50% min.).
The composition of the weld deposit sample is completely consistent with the sperided composition for Type 308 stainless steel bare filler wire. 3.3 ltesidual Stress The level of residual stress in the disk attachment boss was investi gated by saw cutting the boss and measuring the opening or closure of the cut. In this procedure, small punch marks were made on the top of the boss on each side of a position 90' fmm the long axis of the ann. The initial spacing of the punch marks was measuted and a saw cut was made on the boss centerline between the marks. The spacing of the marks was measured again after the saw cut. The typical locations of the punch marks and saw cut are shown in Figure 2 2. t
- hiinor deviations were noted in the nickel or chromium contents for Samples No. 255 and 256 but the values detennined were only slightly outside of the range on the specifications considering the product analyses allowances. These deviations are not considered as significant.
** Comparison with the requirements of A 747 CB7Cu l is made for reference only. The swing arms were not manufactured in strict accordance with that specification.
ow-e:.e wc 3-3
The spacing measurements obtained for each sample are listed in Table 3 1 Over the L entire group of samples, the maximum opening observed was 0.003 inch and the maximum closure noted was 0.004-inch. A first approximation of the residual stress in the boss ring is given the following formula.* [ Et
- AC
~
""' 2 n(1 -v ) D,+ AC/a where E = modulus of clasticity i v = Poisson's ratio D, = initial diameter AC = Opening or closure t = thickness of ring The residual stress values calculated in this manner are also listed in Table 3 4. In the case of 10 I of the 16 samples, the indicated level of residual stress (tension or compression) was only 1.0 ksi or less. Of the remaining six samples, the maximum tensile residual stress indicated was 3.6 ksi and occuned for one of the 4 inch arms (Sample No. 245). The maximum compressive stress l indicated was 3.8 ksi and occurred in one of the 6-inch amn.
In view of the fact that these approximated values are small relative to anticipated in service stress,** no funher evaluation of the residual stress was made. l
- Based on elastic stress formulas for a thin ring considering the opening or closure as a change in circumference.
+* See Section 5.0.
ownuummumum 3-4
~ TABLE 31 L ll ARDNESS DAL Ilardness IIRC Valve Sample Center Edge Siu No. I 2 3 4 5 Avg. 1 2 3 4 5 Avg. I 239 30 30 33 32 34 32 37 35 37 37 36 36 16-in. 240 37 37 38 37 37 37 36 37 239 4a* 34 33 35 36 34 35 - -- -- - -- -- 239-4b*
- 36 36 36 36 37 36 - -- - -- - -
10-in. 255 40 38 39 41 40 40 39 41 39 39 42 40 l 8 in. 242 39 38 41 39 4C 39 30 39 40 39 37 39 243 39 41 41 40 41 40 -- -- -- -- - -- 244 30 30 31 31 31 31 27 29 30 28 28 28 6-in. 256 34 33 34 35 35 34 32 32 30 32 31 31 257 41 40 39 42 39 40 38 38 39 39 39 39 245 34 35 34 34 36 35 36 37 34 38 37 36 4-in. 258 33 33 34 33 34 33 28 29 32 33 31 31 259 35 35 34 34 35 35 33 34 33 34 34 34 I , 260 21 22 21 21 20 21 97t 95t 97t 96t 96t 96t 246 30 32 32 32 36 32 31 32 33 36 30 32 3-in. 261 36 34 36 36 36 36 35 34 35 37 35 35 262 36 35 35 35 37 36 31 34 31 33 31 32 263 31 35 36 35 36 35 31 34 37 36 35 35
- Reheat treated specimen (2150*F 2 hrs., FC to 1900*F 30 min., AC; 1100'F 4 hrs., AC)
** Reaged specimen (1100*F 4 hrs., AC) t Rockwell B scale (HRB) i auramw3m uu 3-5
g 'm susu - e t_J--- I I- t i TABLE 3-2 TENSILE PROPERTIES DATA Gage 0.2% Yield Ultimate Ew% Reduction of Vahe Sample Strength-ksi Strength- ksi % Arts - % size No. Section 151.8 17.0 43.5 129.5 15-in. 239 174.6 8.0 12.4 0.25-in. x 1.0-in. 104.5 8-in. 242 174.5 11.0 28.5 148.2 243 129.1 14.0 30.6 111.7 6-in. 244 __ 110.0 135.0 9 min. [ ASTM 747 CB7Cu-1 (H1100)* mm. mm.
- ne tensile properties specified for A 747 CBCu-1 mm-nal are listed for reference only. He swing arms were not man in strict conformarce to that specification.
DM68URGIARD3893-RDoC
I TAllt.E 3 3 CllEMICAl, COMI'OSITION Composition . w t.c4 Conforms" Valve Sample to llsted Sire No. C Mn 51 Ni' Cr* Mo Cu Ch P S Specs. 16.in. 239 0.06 0.14 0.18 4.07 14.54 0.10 3.05 0.11 0.016 0.014 N 240 0.06 0.20 0 30 3.93 14.85 0.09 3.02 0.17 0.015 0.013 N 10.in. 255 0.05 0.25 0.48 4.97 15.27 0.12 3.19 0.21 0.021 0.018 N 8.in. , 242 0.N 0.48 0.58 4.15 15.81 0.07 2.99 0.25 0.013 0.010 Y 243 0.05 0.32 0.61 4.12 15.31 0.10 3.N 0.21 0.021 0.016 Y 244 0.06 0.45 0.54 4.32 16.85 0.28 2.79 0.17 0.020 0.018 Y 6in. 256 0.N 0.63 0.M 3.80 15.30 0.17 3.02 0.22 0.017 0.0')4 Y 257 0.05 0.22 0.23 3.83 15.08 0.07 2.87 0.18 0.013 0.013 N 245 0.N 0.50 0.M 4.06 16.20 0.10 3.09 0.23 0.020 0.013 Y 4 in. 258 0.06 0.45 0.52 3.76 15.84 0.10 2.86 0.26 0.016 0.011 Y 259 0.05 0.45 0.52 3.70 15.94 ().1 I 2.86 0.27 0.016 0.011 Y 260 0.06 0.37 0.24 6.12 16.47 0.84 1.96 0.16 0.016 0.012 N 246 0.N 0.57 0.54 4.22 16 06 0.17 3.22 0.24 0.024 0.012 Y 3.in. 261 0.05 0.38 0.50 3.83 15.50 0.10 2.95 0.22 0.014 0.011 Y 262 0.05 0.20 0.22 3.89 14.88 0.11 2.93 0.18 0.013 0.012 N 263 0.06 0.21 0.19 3.80 14.84 0.11 2.94 0.17 0.013 0.012 N ASTM A 747 0.07 0.70 1.00 3.60- 15.50- - 2.50- 0.15- 0.035 0.03 CB7 Cu.1 mas mas max 4.70 17.70 3.?0 0.35 rnax max AMS 5398t 0.06 0.70 0.50- 3.60- 15.50- - 2.50- 0.01 0;Ji 0.03 max max 1.00 4.70 16.70 3.20 0.35 man rnax 240-W t t 0.06 1.02 0.38 9.47 20.33 0.05 0.13 0.01 0.021 0.005 -- Type 308 SS 0.08 1.00- 0.30 9.00- 19.50- 0.75 0.75 0.03 I 0.03 -- Filler Wire max 2.50 0.65 11 00 22.00 mas max man man Variation of 0.18% and 0.25% above the maximum and below the minimum specified values for Cr and Ni, respectively, are a' w ed for product analpes in accordance with ASTM A 781.
" Y = yes; N = no. Comparison to listed speu$ cations is made for reference only. 'Ihe swmg arms were not manufactured in strio conformance to these specifications.
t AMS 5398D also specifies maximum contents for three other elements as follows. 0.05 mas Al. 0.02 max Sn,0.05 max N. Analyses were not paformed for these elements. tt Weld metal from repair toi.e. I I I
- - - n
i 'l TAllLE 3 4 RESIDUAL STRESS l Valve Site Sample No, Initial C.in. Final AC In. Approximate Stress . ksi g 16 in. 239 0.460 0.460 0 0 240 0.524 0.523 0.001 0.6 10 in. 255 0.467 0.463 0.004 3.2 8.in. 242 0.501 0.500 0.001 1.0 I 243 244 0.505 0.451 0.507 0.447
+0.002 0.004
+2.0 3.8 6-in. 256 0.444 0.444 0 0 257 0.459 0.459 0 0 245 0.389 0.392 +0.003 +3.6 4in. 258 0.440 0.440 0 0 259 0.503 0.502 0.001 1.2 260 0.532 0.532 0 0 246 0.460 0.459 0.001 1.6 3in. 261 0.590 0.591 +0.001 + 1.6 262 0.472 0.472 0 0 263 0.522 0.524 +0.002 +3.1 I
I I 'I
, _ _ _ . _ 38
I 4.0 META 1,1 OGRAPillC EVAL UATION Metallogmphic sections were taken from each of the 16 swing arr . samples and examined to 1 characterize the microstructural features in the region of die junction of the arm with the valve disk l boss in each case, the sections were taken completely through the thickness and in planes parallel to the long axis of the arm. One sample arm from each valve size group was sectioned at threc ; locations; on the centerline and near each side of the arm (typically 0.1 inch from the outside). A schematic diagram illustrating this method of sectioning is shown in Figure 4-1. In cases where noticeable surface Haws were present, the location of one of the three sections was shifted to pass through the flaws. The remaining samples in each size group were sectioned at only one location. The single section was taken either at the centerline or at the location of visible surface flaws. The l sections were ground and polished by conventional metallographic techniques, and examined at 501000x in the optical metallograph. These examinations were made both in the as polished condition and after etching in Vilella's reagent.* In addition to the systematic metallographic sectioning of all 16 samples, selected samples were sectioned in a manner to locate and characterize the microstructural features coricsponding to selected radiograpiic flaw indications. Also, in cases where the routine sections revealed any linear-type flaw, the sections were sequentially ground and examined to establish the flaw size. The results of the metallographic characterization of specific flaws are presented in Section 5.0, 4.1 Microstructural Features 17-4Pli stainless steel is a CrNiCu alloy steel in which tne strength and hardness is developed by a combination of an austenite-to-martensite transformation followed by precipitation hardening. Typical heat treatmentsinvolve an initial solution heat treatment at 1900*Fto austenitize the steel and completely dissolve the copper. The part is quenched from the solutionizing temperature to develop a martensitic structure and then reheated to a temperature in the range of 9001150*F to temper the martensite and precipitate a copper rich phase within the martensite. The Onal hardness (and strength) depends on the precipitation treatment temperature, with the higher II -
- ASTM E 407 No. 80. SmlliC1,1 g picric acid,100 ml ethanol.
omtuw=mooc 41
I hardness corresponding to the lower end of the temperature range. The typical rnicrostructure for this alloy is a martensitic matrix with dispersed ferrite islands or stringers. The ferrite content can range from 1% to 8% depending on the composition of specific heats of material and the distribution of ferrite varies with fabrication processes. The copper rich precipitate is submicroscopic in size for the lower aging temperatures and is only barely resolved by conventional metallography for the higher aging temperature. In general, all but one of the swing arm samples were characterized by a martensitic type matrix with dispersed ferrite. liowever, a wide variety of detailed mierestructural features were observed among the sixteen samples. Representative microstructural features selected from the entire group of samples are illustrated in Figures 4-2 through 4 7. The particular microuructural features observed for each of the sixteen samples are presented in Appendix B. The shape and distribution of ferrite saried considerably among th; several samples and,in some cases, within a single section of or.e sample. Figure 4 2 illustrates both stringer type ferrite and irregular ferrite islands observed in a sin;;te section. This difference is attributed simply I to the orientation of particular grains relative to the plane of polish. The island features in Figure 4-2(b) probably represent cross sectional views of stringers such as those in Figure 4 2(a). Another similar situation with ferrite stringers and discrete ferrite islands is shown in Figure 4 3. In this l case, the stringers are thinner (rigure 4 3(a)) and the islands appear both as small, ellipsoidal shapes and larger irregular shapes. Again, the differences are consistent with different grain orientations. Other variations i~n ferrite morphology are shown in Figure 4-4. Distinct intergranular network ferrite was often obseved as in Figure 4-4(a). In other cases, the ferrite was distributed in long, thin, connected stringers to develop a cell-type network within the matrix, see Figure 4-4(b). These two types of fertite distribution are also likely to be associated with grain morphology and orientation. Intergranular ferrite in an clongated grain structure, such as columnar grains in a casting, would present the network feature of Figure 4-4(a)in a plane oriented transverse to the elongated greins and the cell-type structure in a plane parallel to the grains. The nearly complete network evident in Figum 4-4(c)is simply another manifestation ofintergranular ferrite. I
--- 4a
In certain cases, the ferrite content of the microstructure was found to be very low. L Examples illustrating that situation are shown in Figure 4-5. The ferrite content in those photomicrographs is on the order of 1% or less. Three of the four 3 inch swing arms (Nos. 261, 262,263) were characterized by a very low ferrite content. In general, all but one sample exhibited a distinctly martensitic matrix. In most cases, the matrix consisted of the usual tempered martensite with very nne-scale features and some l evidence of a very fine precipitate. Examples of this type of matrix structure ate shown in Figures 4 2 and 4-3. A similar tempered martensite structure is also evident in Figure 4-4(c) but,in this case, there is a variation in etching response with light etching in the immediate vicinity of the grain boundaries. This latter feature suggests a slight compositional variation across individual grains and probably indicates incomplete homogenization. The degree of variation in the case of Figure 4-4 l is slight, however, and probably daes not significantly innuence the mechanical propenies. Of all the 16 swing amis,10 sampler, including No. 255 (10 in.), No. 243 (8-in.), Nos. 244 and 256 (both 6-in.), Nos. 245,258 and 259 (all 4 in.), and Nos. 246,261,262 and 263 (all 3-in.) exhibited the I typical tempered martensite structure. In the case of three of the 3 inch samples, the structure was somewhat coarser with more distinct martensite needles (Figure 4-5). Three samples [Nos. 239 (16-in.), No. 242 (8-in.) and No. 257 (6-in.)] cxhibited a matrix with distinctly different microstructural features from those desaibed above. Examples of l this type of microstructure are shown in Figures 4-6(a) and 4 6(b). These structures are also martensitic but are characterized by coarse martensite needles some of which are extremely long in panicular grains. These features probably renect some differences in thermal history from the group described earlier, but it should be noted that there were no consistent, notable differences in composition or hardness of these three samples as compared to the entire group. The hardness of two of the samples (Nos. 242 and 257) did tend toward the high side of the hardness range, however. One 16-in sample (No. 240) exhibited mixed matrix ieatures with typical tempered martensite- and coarse, lath-type martensite at various locations in the same section. A photomicrograph from that sample is shown in Figure 4-6(c). ourmmum mxc 4-3
Although detailed differences in microstructum we e observed among 15 of the 16 r L samples, all of the features may be classified as one or another variation of a martensitic structure. Most difference s in ferrite morphology can be attributed to grain orientation effects and the variations in matrix structure. Although indicative of differences in thennal history (including heat treatment), these microstructural variations do not necessuity indicate significant differences in properties. This factor is born out by the generally uniform hardness values m:asured for the 15 samples. All of the microstructural features of this group are normal for heat trehted 17 4Pil castings. Sample No. 260 (4-in.) exhibited distinctly dendritic features as illustrated in Figure 4-7 (also see Figure 11 14, Appendix B). The original dendritic grains of the as-cast condition I were not evideat but the dendritic pattern was retained in the etching response. This factor indicates that the sample was probably solution annealed after casting, but it is clearly evident that it was not adequately homogenized. Such a lack of homogenization would definitely affect response to subsequent heat treatment. Itis ofinterest to note that the hardness of this sample was significantly lower that of the other 15 (llRB 96 vs. liRC 34 40)and thatit was the one sample with a composition I distinctly different from the typical 17 4 pil con: position. Also, when tested with a small magnet, Sample No. 26J was noticeably less magnetic than the rest of the samples. A general observation among all 16 swing arm samples was that shrinkage microvoids were observed at random locations in the metallographic sections. Typical examples are shown in l Figures 4-8 and 4-9. These were always small (0.001005 in, across) and generally isolated with a maximum concentration such as shown in Figure 4-9(a).* Micmvoids of this type are a common feature for steel castings and are not considered as detrimental flaws. 4.2 Weld Repairs Certain of the routine longitud >aal sections (A-A, B-B and C-C in Figure 4-1) intersected weld repairs made in the vicinity of the valve disk boss. In addition, one metallographic Except for those cases where distinct mdiographic indications were obtained. Those cases are treated separately in Section 5.1. oumwmmm ua 4-4
section was taken from the valve disk boss of Sample No. 240 for the specific purpose of examining b the ilAZ of the through thickness weld repair. Photomicrographs from the several sections a containing weld repairs are shown in Figures 4-10 through 4-13. In every case, the weld metal exhibited a totally different etching response from that of the base metal. The only effect the etchant had on the welds was to delineate a columnar structure. l Otherwise, the weld metal was unetched. These features indicate that the weld repairs were made with an austenitic stainless steel filler metal; not an alloy to match the 17 4Pil base metal. Tids observation is consistent with the one case where chemical analysis positively identified the weld deposit as Type 308 austenitic stainless steel (see Section 3.2 and Table 3-3). In the case of Sample No. 239 (16 in.), the microstructure of the llAZ consisted of equiaxed grains outlined by intergranular ferrite and light etching zones (Figure 410). This l microstructure is totally different from that of the base metal remote from the weld. (Compare Figmes 4-10 and B-1.) This factor, together with the strong columnar etching response of the weld metal, indicates that the casting was not heat treated (or was improperly heat treated) after the weld repair was made. The microstructure of the IIAZ in the other 16-in. swing arm (Sample No. 240) was also distinctly different frcm that of the base metal, in this case, the IIAZ was characterized by equiaxed, martensitic grains either outlined by light etching boundaries or without distinct boundaries (see Figure 4-11), These features are distinctly different fmm the base metal which exhibited coarse martensitic features with distinct intergranular ferrite pools (compare Figure 4-11 and B-2). Again, this difference in structure between the base metal and the llAZ indicates that the casting was not properly heat treated after the weld repair. la the case of the other three weld repairs examined, the microstructure of the llAZ was generally similar to that of the base metal. The llAZ of Sample No. 243 (8-in.) exhibited equiaxed grains outlined by intergranular ferrite essentially identical to the base metal structure [see Figure 4-12(a)). Distinct ferrite stringers, typical of the base metal structures were present o w--m ume 4-5
immediately adjacent to the fusion lines in Samples No. 256 (6-in.) and No. 245 (4 in.), see L Figures 4-12(b) and 4-13. These microstructural features at the weld repairs indicate that these three samples (No. 243,246 and 256) were properly heat treated after the weld repairs. It should be noted that, even though the weld was made with an austenitic filler metal, no specific defects such as voids, lack-of-fusion or underbead cracking were observed in those weld repairs which were sectioned. Also, Figure 4-13 illustrates that the weld repair in that case was l only 0.06 in. deep. Similar, relatively sh ilow penetration was also observed in other cases of repairs at locations outside of the disk attachment bosses. I I l I L l I I i 1 l ____ .e
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I 5.0 FLAW CilAllACTEltlZATION 5.1 11adlographkfhtn.htdkniluns Six of the sample swing anus exhibited radiographic indications which were classified as Level 2 or higher shrinkage and ve.ds, see Section 2.2 and Table 2 2. Three indications rated as Level 2 and Level 3 (Samples No. 240 and No. 242) and the one case with the large, Level 6 indication (Sample No.255),were selected for metallographic examination. in each case, specimens containing the Gaw indication were cut from the swing arms and sequential metallographic sections were prepared and examined to characterize the Daw. The observations made in . lese multiple section examinations are described in the following subsections. 5.1.1 Sam & No. 240 H6-ind Sample No. 240 exhibited two radiographic flaw indications classified as Level 3 shrinkage. The locations of these two indications are shown in Figure 5-1. Section A,- A,was cut transverse to the long axis of the arm at the approximate position of the Gaw near the end of the straight portion. Examination of both sides of this initial cut did not reveal any microstructural Daw. Ilowever, a small subsurface shrinkage crack (hot crack) was present after removal of an additional 0.10 inch beyond Section A, A,. A photomicrograph of this feature is shown in Figurt 5-2. The crack was segmented in this plane and the cumulative length of the several segments was 0.14 inch. Other than the shrinkage crack, there were no microstructural abnonnalities at this location. The character of the Gaw changed dramatically after removal of an additional 0.006 inch, in this new section, the Daw appeared as a very tight, segmented, intergnmular microcrack with a cumulative length of 0.07 inch, see Figure 5 3. The feature was not detectable at magnifications below 200X. Apparently, this metallographie section intersected the extreme outer fringes of the Daw. Further grinding in increments of approximately 0.05 inch did not reveal any other microstructural Daws. Based on the measmtment of the material removed between each section, the maximum dimension of this Oaw in a direction parallel to the long axis of the swing i ow-mm woc 5-1 I
ann is 0.16-inch. The features observed in the sequential sections show that, as the worst case, this flaw may be considerrd as a small, internal, planar crack located near the center of the cross sectiori of the amt A diagram of the orientation and maximum size of the Gaw is shown in Figure 5-4. The second radiographic indication in Sample No. 240 was located within ~ the curved portion of the ann near the hinge boss, see Figure 5 1. Section II -B, was cut to intersect the Daw indication. A cluster of voids was clearly evident on each side of this initial cut and sequential sections were examined in both directions from Section B, II.. Photomacrographs from these sections are shown in Figuirs 5-5 and 5-6. In each of the three sections closest to the center of the indication, the void cluster consisted of one or more relatively large voids (larbest dimension on the order of 0.05-inch) with several much smaller voids in the immediate vicinity [see Figures 5 5(b) and (c) and 5 6(a)]. Of the two outer sections, one exhibited a cluster of very small voids (Figure 5 6(b)] and the other contained a small fissure [Figur 5 5(a)i. These fer.tures inmcate that the outennost sections were located at or near the extremitie of the flaw and serve to establish the maximum dimension of the Dawed zone. A diagram showing the extent of the flawed mne is shown in Figure 5 7. Since the metallographic sections identified the Daw as a clustcr of voids rather than a single large void or a discirte crack-type feature, the Oaw zone in Figurr 5-7 is dimensioned to represent a volume which completely contains the void cluster. Note that ths outlinec areas in Figurr 5 7 represent the projections of the volume on the top and side of the arm. The cross sections shown in Figures 5 5 and 5-6 must be considered to view the actual shape of the Hawed zone. In this case, the flaw may be considered as a thin, irregular volume lying on a warped surface and containing a concentration of small voids. Considering the volume as a single, open void, the maximum cross sectional void area would be only 0.01-in.2 or less, 5.1.2 Samph1c,255 (10-in.) The radiographic inspection of Sarnple No. 255 revealed Level 6 voids or inclusions within the straight portion of the arm approximately 3 inches from the centerline of the valve disk boss. Multiple longitudinal and transverse sections were taken through the zone of radiographic indications to size the Daws. ow-wunm uw 5-2
The longitudinal sections revealed a large, tiongated void oriented essentially nomial to the long axis of the arm. Photomacrographs from selectedlongitudinal sections are shown j in Figure 5 8. In these r N.s. the void is roughly circul m cross section with a maximum dimencion on the order of 0.16-mch. Near the centerline of the arm, the character of the flaw changeu irom a single large void to clusters of very small voids as shown in Figure 5-9. Further
} grinding eliminated the flaw. ._ )
Transverse sections in the flaw zone revealed a second, large,iragularly
] shaped void on th o a '.te side of the arm. Photomacrographs showing the cross section of that void in transvece sections are shown in Figure 5-10.
A dias. ram of the flawed zone showing the locations of the metallographic sections and the extent and shape of the voids derived fmm the sequential sections is shown in Figure 5 1l The overall shape and locations of the two voids show that certain transverse sections v,ould intersect both voids. Thus, this flaw may be considemd as an internal, rounded shrinkage 2 cavity with a maximum cross-sectional area of 0.23 in , 5.1.3 Samp!e No. 242 (8 in.) Sample No. 242, one of the two 8-in. swing arms, exhibited a single, Level 2 radiogrephic indication similar :o those recorded for Samples 239 and 240. This indication was located in the straight portion of the arm approximately 2 inches from the centerline of the valve disk boss. Multiple transverse sections through the zone of the indications were examined and representative photomacrographs are shown in Figures 3-12 through 5-15. In each cross section, the flaw appeared as a cl.ister of small voids with the most severe condition shown in Figure 5-15.
"I It shou!d be noted that in Figures 5-14 and 5-15 some of the voids are alignedin the verticaldirection, an orientation that would contribute to a stro: g radiographic indication in a through-thickness radiograph.
The location, extent and orientation of the flawed zone in Sample No. 242 are illustrated in Figure 5-16. The dimersioned zone in tlat diagram represents the volume ; J containing the void clusters in the same manner as for Samp'e No. 240 in Figure 5 7. If the flaw I
] mutamemasoc 5-3 1'
zone is considered to be a single void with dimensions as shown in Figure 5 16 and cross sections ] compatible with the void clusters seen in transverse sections, maximum cross sectional area would 2 be determined by the acction shown in Figure 5-15. That area was determined to be 0.02 in . 5.1 Other Flaws Othet sinall flaws were encountered in the course of the routine metallographic examinationr, These were observed in the longitudinal sections taken through the boss-to-ami l transition at the valve disk end of certain swing arms (see Figure 4-1) and had not been detected in the nondestructive inspections. A distinct, surface connected shrinkage crack (hot crack) was present at the valve disk I boss of Sample No. 255 (10-in.), see Figure 5-17. The crack was located in the radius at the boss and was oriented transverse to the long axis of the arm. In this case, a hght etching zone was present around the crack indicati~.ig a localized variation in composition. The light etching zone extended beyond the crack tip along on intergranular path. Sequential sectioning established that the crack extended almost half-way across the arm with a maximum length on the top surface of 0.93-inch. The maximum depth measured in the several sections was 0.07-inch (Figure 5-17(a)]. As a worst case, this flaw may be considered as a tight, surface a nnected crack oriented normal to the axis of the swing arm. This case is diagramed in Figure 5-18. The dimensions of the flaw are sucli that it may be conservatively represented as a semi-elliptical crack. 0.93-inch long and 0.09-inch deep. A small, internal fissure was present near the top surface of the boss in one of tht. 4-inch swing arms, see Figure 5 19(a). This microfissure was less than 0.02-in across in tnat section and sequential grinding established that its maximum length in a direction transverse to the long axis of the arm was 0.10-inch. This flaw was considerably smaller than the fissure-type flaws observed in other s.uaples. Figure 5-19(b) shows a small, unique flaw which was found in one of the 3-in. swing arms. In this case, the naw consisted of a highly irregular fissure which formed a nearly closed nuraunauamem w 5-4
I path. It was located at an outer corner of the valve disk boss and the maximum dimension across tne path was 0.025 inch. Considering the small size of this feature and its non-criticallocation, no further metallographic sectioning was performed. As noted in Section 2.1, relatively deep surface pits were noted in certain of the swing arm samples during the visual and LP inspections. In the course of the routine metallographic examinations, cenain of the longitudinal sections were located to pass through such pits so that the depths and profiles could be established. A photomacrograph of a typical group of surface pits is shown in Figure 5-20 together with a photomicrograph from the section through the pits. In this particular case, the maximum depth was only 0.01-inch and the pit profile was relatively blunt. Also, there was no evidence of cracking from the pits. These features were generally representative of tho'c in other sections of other arms. In view of the small size and blunt character, these flaws are not considered to be detrimental. 5.3 Daw Summary The flaws identified by the metallographic examinations were of two distinct types, namely; 1) small fissures or cracks, and 2) shrinkage cavities or clusters of small shritikage voids. The s;gnificance of the two fissure-type flaws [ Samples No. 240 (16-inch) and 255 (10-inch)) may l be directly evaluated by linear clastic fracture mechanics (LEFM) since they both represent sharp cracks of known dimensions. In this regard, a surface-connected crack represents a more severe condition than an internal crack of equivalent dimensions.* As a result, evaluation of both fissure-type flaws as surface-connected cracks is a conservative app oach. LEFM techniques are not directly applicable to the open, rounded features of the cavity in Sample No. 255 o'; the void clusters in Samples No. 240 and 242. However, as an initial, conservative approach, LEFM may be applied to these flaws by considering each flaw as a tight, I I I
- The stress intensity factor for a surface-connected semi-elliptical erack exceeds that of an internal elliptical crack of the same dimensions by a factor of 1.12.
l aunewwmmwoc 5-5 1 I
l elliptical crack of a size equivalent to the length and breadth of the void or void cluster. An additional degree of conservatism is introduced by treating the flaw as a semi-elliptical surface-connected crack. Following these concepts, calculations were made to determine the stress intensity factor (K t) associated with each of the flaws described in Figures 5-4,5-7,5-11 and 5-16. Each flaw was treated as a semi-elliptical, surface-connected crack oriented transverse to the swing arm and the calculations were made for the case of a pure bending load with the flaw on die tension side. The expression used for the determination of the stress intensity factor for such a case is. I K, = Ho, n " F ",", b Q(tc j I where: c3 = bending stress 2e = crack length t = section thickness Q = crack geometry factor b = section width and F and H = boundary correction functions a = crack depth The basis for development of this expression is described by Newman and Raju [1] and the individual equations for Q, and the functions F and H are presented in Appendix C. This expression takes the geometry of the swing arm into consideration as well as # : crack geometry. The results of these calculations, using a value of a = 60 ksi for the maximum bending stress, are presented in Table 5-1.* The stress intensity factors determined by these calculations ranged from 18.6 to 31.1 ksi6. The fracture mechanics evaluation of the flaws in the swing arms is presented graphically in Figures 5-21 and 5-22. Figure 5-21 illustrates the critical stress for fracture (c ) for l U The value of 60 ksi for the tensile stress at the outer fiber due to bending was selected as I representative of the surface impact stresses based on earlier investigations at CPSES, tmw:nc<umannoc 5-6 1
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d i I cach crack geometry as a function of creek size Figure 5-22 demonstrates the interdependency of
- the critical stress (o ), crack size 01) and fractun: toughness (Kc) for a conunon flhw shape in Table 5 1.
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TAllLE 51 FRACTURE MECll ANICS PAR AMETERS FOR LARGEST OllSERVED FLAWS Valve Sample W t ; 2c a I Figure l Kti Site No, in. in, ir., in. Fio, alt i a/2c ksl6 0.16 0.07* 5-4 0.05 0.44 18.6 1 16-in. 240 2.15 1.30 0.70 0.10*
- 5-7 0.08 0.14 31.1
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0.93 0.09* 5-18 0.09 0.10 30.5 10 in. 255 1.78 1.(X) 0.94 0.08t 5-11 0.08 0.08 29 5 l _ -__ 8-in. 242
' 50 1.00 0.67 0.12 *
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m om Q. om n..mw x, aQa. v$$. m. u.a; m)eq +m y...n%.;a,+nv..a<. a1 m s. a , a ,,u.,p... .,px ,./,e. c . . 8 N 46662-64 200X FIGURE S 3. MICROFISSURE IN SAMPLE NO. 240 (16-in.). Transverse section 0.006 inch beyond section of Figure 5 ;. Three matching photomicrographs. The upper two are displaced to the left to fit the single page format. Arrows mark matchpoints in each photomicrograph. Etchant: Vilella's reagent owmmmamervoc 5 11
m w> > x >
= A, ym .
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View A in Figure 5-1 DIAGRAM OF INTERNAL FISSURE IN SAMPLE NO. 240 (16-in.). See Figure 5-1 for location. Twice size. FIGURE 5-4. Dimensions in inches. owrawn-wooc
w m - 1 rn r i ; 1
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~7963 (a) Section BcB, - 0.34-in. As polished. 16X _ , r u* . ,x . ..g y pp_y gg n .
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%.?; l*j @[^k'h._%. ~ drA M z g f ji g'4-< f:7thf[ k'b c.N NO.740. Sec Figure 5-1 forlocations. Arrows
- .fyM(
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g~ , {g.N
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l Arrows mark larger voids. I
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b I - 47962 16X (b) Section Do-B o + 0.36-in. As polished. FIGURE 5 6. TRANSVERSE SECTIONS FROM SAMPLE NO. 240. See Figures 5-1 and 5-7 for locations. ow ncam m a m rooc 5-14
-w - -,1 ga em 1 -- i r
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L 4 .- I i View B in Figure 5-1 DIAGRAM OF INTERNAL FLAW IN SAMPLE NO. 240 (16-in.). Out!ined areas represent FIGURE 5-7. velume containmg concentration of srna 11 voids. Set hgun:s 5-5 and 5-6. Scale is approximate. j
-wnwomuooe 1
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47831 16X (b) FIGURE 5 9. LONGITUDINAL SECTION FROM SAMPLE NO. 255 (10-in). See Figure 5-11. Section Ai -Ai + 0.88-in. As polished. emuu=weerax 5-17
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l 2X 47814 2X 47900 2X 47813 l (b) Section B,-B, + 0.19-in. (c) Section B,-B, + 0.29-in. (a) Section B,-B Note: Grinding to B,-B, + 0.40-in. removed flaw. FIGURE 5-10. TRANSVERSE SECTION FROM SAMPLE NO. 255 (10-in.) See Figum 5-11. Etchant: Vilella's reagent o nmuwassemoc
= 1.0
= 4.0 f-=- i 0.10 0 i h L
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g I i ta- : O+ FIGURE 5-11. DIAGRAM OF INTERN AL VOIDS IN SAMPLE NO. 255 (10-in.). See Figures 5-8,5-9 and 5-10. n - encu nu m enoc 5-19
I
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FIGURE 5-12. TRANSVERSE SECTIONS FROM SAMPLE NO. 242 (8-in.). See Figure 5-16 for locations. As polished. uwscuoianne nxr 5-20
,,,q y,-, .<q.g. y.my .
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y . .. . 1 , ...s 16X 47913-915 Section A,-A, + 0.64-in. FIGURE 5-13. TRANSVERSE SECTION FROM SAMPLE NO. 242 (8-in.). See Figure 5-16 forlocation. As polished. Imae8L1tGIARf).1991FDOC
p 5
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I in. , E I I FIGURE 516. DIAGRAM OF INTERNAL FLAW IN S AMPLE NO. 242 (8 in.) Outlined areas represent wlume containing concentration of small voids. See Figures 5-12 through 5-15. Scale is approximate. omtat:acuum 91r. ooc 5-24 I
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1 FIGURE 518. DIAGRAM OF SURFACE-CONNECTED FLAW IN SAMPLE NO. 255 (IG-in.). Transverse section of swing arm. Dotted line connects points of measured depths in metallographic sections. Shaded area represents flaw as a semi-elliptical shape. I natuwam usoc 5-26
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a > B s U..q ( ,Og Sample No. 263 (3-in.). (b) ! 47381 100X Section C-C, Figure 4-1. l (a) Sample No. 245 (4-in.). ! Section B-B, Figure 4-1. ! Right-hand edge is bore of valve disk boss. Repair wcld is at top of boss. 3 FIGURE 5-19. MICROFISSUREF iN 4-in. AND 3-in. SWING ARMS. Etchant: Vilella's reagent i.
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a l 6.0 FRACTURE TOUGilNESS Fracture toughness testing provides a measure of the resistance to fracture for a given material. Fracture toughness tests can be broken into two categories, i.e., plane-strain fracture toughness tests (Kr), determinations of fractutt toughness by means of the parameter J ci in cases where stable crack extension occurs prior to fracture. The property K ci characterizes the resistance of a material to fracture in the presence of a sharp crack in a neutral environment under conditions of a state of stress approaching tri axial plane strain near the crack front, and a small crack-tip plastic region compared with the crack size and specimen dimensions [5]. Kci represents a lower limiting value cf fracture toughness and is determined directly in tests where plane strain conditions irevail at fracture, if, however, the specimen exhibits stable crack growth during the test, a plane-strain analysis is not appropriate. Instead, a J icanalysis is required, where J i c is the critical value of the J-integral near the onset of stable crack extension [6]. Subject to certain limitations, the parameter Jg: may be converted to an equivalent value for plane-strain fracture toughness, K c. i 6,1 Test Procedures Both K ci and Jotesting procedures are similar and employ compact ter sion specimens of the type shown in Figure 6-1. In both types of test, the specimens are loaded in the manner shown and the crack opening displacement (COD) is measured as a function of load. Prior to testing, the specimens are fatigue precracked to initiate a sharp crack at the root of the deep notch. The specimens are machined to provide specific B/W and a/w ratios. In this investigation, compact tension specimens of the designs shown in Figures 6-2,6-3, and 6-4 were used to determine fracture toughness values. Three different specimen sizes were required because of the variation in swing arm dimensions among the 16 specimens. The distance (W) from the centerline of the pins to the back edge is 1.5 in.,1.0 in., and 0.76 in. for Type A, Type B, and Type C, respectively. Type A and B specimens are of the sizes commonly used for fracture toughness testing. While Type C is of the same geometry as Type A and B specimens, the size is slightly less than accepted standards. The adequacy of these smaller specimens was investigated during the course of the program. twammmmwr 6-1
I 1 The location selected for removal of the specimen coupon from each swing arm was based primarily on the individual sample size and defect locations. Specimen locations were chosen to avoid radiographic indications and areas of weld repair. The orientation of the crack path for each specimen was consistently chosen such that crack propagation would be across the width of the arm. Figure 6-5 shows the typical location and orientation of a fracture toughness specimen in a swing arm. Prior to testing, each specimen was polished on both sides to facilitate crack length measurements and then fatigue precracked. The fatigue crack was allowed to grow until the ratio of crack size to specimen width (a/W) was between 0.45 and 0.55. While the cyclic load necessary to initiate a crack may be reasonably high, the load was decreased as the precracking progresses, such that the final precrack load is well below the anticipated test condition. All tests were computer controlled and were analyzed by stendani analysis software packages. A displacement gage was clipped onto the fmnt face of the specimen to monitor crack opening displacement. This displacement was then converted to a change in crack size by the compliance method.* The major difference between the K icand J ci test procedures is that in the Ke test the specimen fails with no significant slow crack growth prior to catastrophic fracture, while the Ji ctest exhibits significant stable crack growth Each test was conducted under the assumption that it would behave like a Ji c test. If instead the specimen failed without any crack extension, a Kci analysis was performed. The only data required for a Kranalysis are the crack size and the load at which the specimen failed. In a Ji c type fracture toughness test, periodic unloads are employed to generate periods of stable crack growth. The compliance of each unload is then used to calculate changes in crack length throughout the test. By monitoring the load and changes in compliance, a plot of J versus crack extension (Aa) can be generated. An example of the raw data and the final analysis to obtain aJevalue are graphically shown in Figures 6-6 and 6-7, respectively. Compliance is defined as the slope of the load versus crack opening displacement curve and is obtained by a clip gage mounted on the load line of the specimen [7]. omsamumm um 6-2
Figure 6-6 presents the data obtained during a Jac test as load versus crack opening l displacement, During the periodic unloads, the load is dropped 40%, to ensure that a good slope of the compliance can be obtained. In Figure 6-7 the calculated value of J is plotted against crack growth for each point. Both a linear regression and a power law fit to the data are shown. Line A is the blunting line, which is a function of yield strength and crack growth. Lines B and D air exclusion hnes, outside of which data are discanied. Line C is a 0.02 inch offset fmm the blunting line. The value of J ic is defined as the ir.tersection of the power law fit to the data with Line C. A Je value can then be converted to an equivalent Kr by the fomiula iK c = ilcFJ(1 i -v 2), where E= Youngs Modulus and v= Poisson's ratio. For more specific details on fracture toughness testing, refer to the applicable test standards used for this program (ASTM E 399 for K ci and ASTM E F13 for Jac, see Refs. 5 and 6). 6,2 Est Results Specimen coupons were sec:ioned from all 16 swing arms and machined into fracture toughness specimens. The test results for these swing arms are shown in Table 6-1. Two Type A specimens were machined from each of the 15-inch valve swing arms. For each of these two arms, one of the specimens was tested at roem temperature and one was tested at 40*F Since the operating temperatures for the sw uig arms range as low as 40 F, these tests were necessary to investigate the temperature dependence of fracture toughness down to that lower temperature. As can be seer. from Table 6-1, no appreciable difference was found between the room temperature tests (239-2 and 24C"2) and the 40*F tests (239-3 and 240-3). For each pair of tests, the K ci values were within the scatter band for the complete series of tests and the 40*F values were, in fact, slightly higher than the contsponding room temperature values. These results indicate that the fraciure toughness of the cast 17-4PII material does not exhibit a strong temperature dependence for the range of temperatures of interest. The remainder of the fracture toughness tests were conducted at room temperature, with confidence that the results are applicable to actual swing arm operating conditions. m aam mena 6-3
Since some of the swing amis were too small to provide sta,Jani size specimens, a non-conventional specimen geometry (Type C) was necessary for some of the tests. Although all of the relative dimensions corm spond to ASTM specifications, there was some concem as to whether the small specimens would provide reliable fracture toughness values. In order to evaluate the reliability of these subsite specimens, Type C specimens were machined from broken halves of two Type A specimens (239-2 and 240-2). Table 6-1 shows that no appreciable differences exist between the results obtained with standan! and subsize compact tension specimens for this material. Table 6-1 provides a population of data for the 16 swing arms investigated. The K ic values range from 53.5 to 115.3 ksi6 with a median value of 74.2 ksi6 and a mean value of I 78.4 ksiE. There are no apparent correlations between Krvalues and either valve size or specimen geometry. Thus, the data are treated as a single population. A histogram of the data is plotted in Figure 6-8 and shows a general grouping in K ci values between 60 an 180 ksi6. For this matrix of tests, scatter in the data does not come from specimen size or valve size, but can instead be attributed to the inherent variation in material propenies for cast steel. Specimen No. 260-2 failed during the precracking process. However, since the precrack load and final crack le gth befom failure were known, a gocx! estimate of K ei wasobtained for this swing arm. It is of interest to note tnat Sample No. 260 was the one sample which did not confonn to the compositional and hardness requirernents. - Certain requirements are specified by the ASTM specifications to have a truly valid Kci orJ i c test. A f:wof these requirements were not met by this material. Crack lengthmeasurements were made at nine positions across the fractured specimen surfaces to obtain an average value for the initialcrack size. According to specification ( ASTM E 399 and E 813),these nine measun:ments should not vary from the mean value by more than 10'7c. None of the specimens met this criterion. Another validity check is in mgard to the number of valid data points during the Ji c tests. For strict conformance to ASTM E 813, four data points during the test must have changes in crack length in the range from 0.006 to 0.06 inches. Six of the specimens (Nos. 239-2,239-3,239-2A,244-2, c m w a m m noc 6-4 1
245-2, and 259-2) did not meet this requirement. However, the pri. nary purpose of these two L requiirments is to eliminate scatter in the results. p The most imponant validity check is a restriction on the minimum thickness required. L For a K i c test to be in strict conformance with ASTM E 399, the thickness must e xceed 2.5 (Kdo ,)2 y where o,,is the yield strength. For a Jac test (E 813), the thickness must exceed 25 (Jdo,,). All Ji c tests met the thickness requimment, but 6 out of the 8 Kr tests did not meet this requirement. The last column in Table 6-1 tabulates the ratio of the specimen thickness to the thickness required to meet the validity requirement for the appropriate ASTM standard (E 399 or E 813). Thickness ratios greate-r than 1 indicate conformance to the specified thickness limitation and values less than 1 indicate non-conformance. Large values for this ratio are common for J ci tests, since the thickness requirements for the J ic tests are much les s stringent than the K,c thickness requirements. In Table 6-1 note that, for all cases where a J ci analysis was requin d, the thickness ratios are much greater than i 1. The six tests for which the thickness ratios are less than 1.00 do not meet the strict E 399 thickness requirement for Ki c tests (240-2A, 240-3, 255-2, 242-2, 257-2, and 260-2). However, five of these ratios are greater than 0.60 and are considettd reasonably close to the strict l validity requirement. Situations such as these, where the specimen thicknesses are hot far out of bounds, usually provide adequate constraint at the crack tip so that reliable fractutt toughness data is obtained. The one case where the thickness ratio was only 0.51 (260-2) was the one swing ann which exhibited an abnormal microstructure and where the test specimen failed during precracking. Even there,the K icvalue determined falls within the scatter band of the data set. Adequate constraint for reliable test results was demonstrated by the fact that none of the fractured test specimens exhibited significant through-thickrf ss plastic deformation, see Figures 6-9 through 6-12. The fact that adequate constraint did prevailis also evident in the lower values of K ci seenforthosespecimens with thickness ratios less than 1.0. If inadequate specimen thickness led to inaccurate K ci results, it would be expected that elevated valaes of K i c would be obtained due to the loss of constraint. t>umwamm nuc 6-5
k Thus, the lower K e i values associated with the low thickness ratios show that the specimen thicknesses did not adversely affect the results. ASTM validity checks and requirements are designed to be guidelines in determining t 4 whether or not tests are yielding accurate results. Even if these validity mquirements are not met, r tests can still produce reliable results. Engineering judgement must be used to determine if those tests failing specific validity requirements yield results any different fmm these tests which meet the ASTM requirements. 6.3 Data Statistim in order to evalualcuthe statistical characteristics of the test results, the set of 16 Ke i N. values was compared to four standa1U distributions to determine a goodness of fit. The distributions evaluated were normal,lognormal, Weibull, and EVD (extreme value distribution). Of the four, the EVD provided the best fit to the data. The details of the statistical evaluation are given in Appendix D. l Utilizing the EVD, the tolerance limits on the lower bound of the data set were determined. The results are that, for 95% of the similar swing arms stillin service, there is a 95% confidence interval that the lower limit of fracture toughness is as follows. (See Appendix D.) EVD distribution: Ki c 2 44 ksi6 This lower limi; value exceeds the maximum K values i which are expected to be encountered in service (see Section 5.0). Lower bounds of fracture toughness for other levels of population and other confidence intervals are given in Appendix D, Table D-2. I I owaammuooc 6-6
I TABLE 61 FRACTURE TOUGilNESS TEST RESULTS I - Valve Specimen Specimen Thickness I Size No. Geometry Jic Value (in lblin') K ic Value (ksi6) Ratiot 239 2 Type A 255 88.2 14.42 239-2A Type C 225 82.9* 8.26 16 in. 239-3 Type A 280 92.4 " 13.16 239 Avg. -- -- 87.8t -- 240-2 Type A 61.3 1.22 I 240-2A Type C -- 58.7* 0.67
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** Test conducted at 40*F.
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ll lI i lI ,I I ! f 4 LI a> i y 3- ! S 8 lI j 3 2- !g- e i 5 !g $ 55 60 65 70 75 80 85 90 95 100 105 110 115 ll K ic Value, kslVin l FIGURE 6-8. HISTOGRAM OF K ci DATA FOR CAST 17-4PII SWING ARMS
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SUMMARY
AND DISCUSSION The observations made and the data obtained in this investigation are sununarized as follows:
- 1) Visual and liquid penetrant inspections did not detect any significant external flaws.
- 2) Radiographic inspections revealed Level 6 indications of voids or inclu:, ions at one location in the swing ann from a 10-in. valve (Sample No. 255). Isolated Level 3 indications were noted for the two arms from 16-in. valves (Nos. 239 and 240) and both of the anns from 8-in. valves (Nos. 242 and 243). The I remaining twelve swing arms were either free of flaw indications or exhibited only isolated faint Level 1 and Level 2 indications.
- 3) Weld repairs were evident in 10 of the 16 swing arm samples. In the two 16-inch arms, weld repairs in the disk attachment boss extended through the full thickness. All others were on the order of 0.06-in. deep. Etching response of the weld deposits indicated that the repairs had been made with an austenitic stainless steel filler metal. Chemical analysis of one weld deposit identified the filler metal as Type 30S austenitic stainless steel.
- 4) The average hardness values for 15 of the 16 swing anns were in the range of HRC 28-40 (BHN 271-372). One sample (No. 260) exhibited an average hardness of HRC 21 (BHN 224). For comparison, ASTM A 747 specifies a minimum hardness of HRC 28 (BHN 271) for CB7Cu-l (H1100) stainless steel (17-4PH).*
- 5) The properties determined in two out of four uniaxial tensile tests [No. 239 (16-in.) and No. 243 (8-in.)) met the requirements for A 747 CB7Cu-1 material (17-4PH) in the H1100 condition. In the other two tests, the yield strength for l one 8-in. sample (No. 242) and the ultimate strength for one 6-in. sample B (No. 244) were below the minimum of t. st reference specification.
- 6) The bulk chemical composition for 9 of the 16 arms confonned to the requirements for A 747 CBCu-l materialin every respect. Six other samples conformed to that specification except for slightly low-chromium contents (14.54-14.88?c vs.15.50% min.) or slightly high nickel contents.** The composition of one 4-in. sample (No. 260) was clearly out of range for 17-4PH stainless steel.
- 7) In general, the microstructural features of all but one of the swing arms were normal for cast 17-4PH stainless steel heat treated to the H1100 condition. One sample (No. 260) exhibited distinctly dendritic microstructural features indicating an inadequate solution-annealing treatment.
The requirements for ASTM A 747 CB7Cu-l (H1100) material are cited for reference only. The swing arms were not manufactured to specifically conform to that specification.
** The requirements for ASTM A 747 CB7Cu-l (H1100) material are cited for reference only.
The swing arms were not manufactured to specifically conform to that specification. AMS 5398D does not specify hardness values for the HI100 condition. . I n uu m a w m anac 7-1
- 8) Multiple sectioning characterized the flaws associated with level 2 and Level 3 radiographic indications as small shrinkage Ossures (major dimension = 0.2-in.)
or clusters of small shrinkage voids (maximum dimensions of individual voids ( on the order of 0.05-in.). The flaw associated with the Level 6 radiographic indication in Sample No.255 was found to consist of two relative large r shrinkage cavities, the larger measuring 0.94-in. Ig. A.15-in. dia.
- 9) Metallographic sectioning also identified a small, crack-type flaw measuring 0,93 in. Ig. x 0.09-in. deep near the disk attachment boss of the swing arm from the 10-inch valve (Sample No. 255).
- 10) All of the flaws observed were associated with fabrication processes as opposed to being service-induced, and there, was no evidence ofin-service growth of any flaw.
I1) The larger of the observed flaws (Level 2 and larger) were mainly confined to the 16 inch,10-inch, and 8-inch swing arms.
- 12) Room temperature (=72*F), fracture toughness tests on specimens from all 16 sample swing arms resulted in values of the critical stress intensity factor (Kc) in the range of 53.7 to 115.3 ksiE with a median value of 74.2 ksiE and a mean value of 78.4 ksid The K ci value for the one swing arm with a typical hardness, composition and microstructure (Sample No.260) fell within the scatter band for the data from the other samples.
- 13) Fracture mechanics analysis was employed to evaluate the individual flaws.
Based on the worst-case scenario in each instance, the calculated stress intensity I values (Ki ) were in the range of 18.6 to 31.1 ksi6 l 14) Statistical analysis of the fracture toughness test results indicated that for a 95% confidence interval, the lower bound of fracture toughness (K ic) for 95% of the l 56 s ving arms rerr Jning in service is 44 ksi6.
- 15) Duplicate fracture toughness tests on the 16-in. swing arms conducted at 40'F resultedin K ci values well within the scatter band of all of the room temperature K ic data.
- 16) Duplicate fracture toughness tests on the 16-in, arms, employing both the standard size test specimens and the sub-size specimens, established that the sub-size specimens produced reliable Ko values.
In general, all but one of the swing arm samples included in this investigatian were free of gross defects and all but one exhibited compositions, microstructures, and hardness values comparable to the requirements of ASTM A 747, the current specification for cast 17-4PH (H1100) material [A 747 CB7Cu-1 (H1100)]. Only one sample, a 4-inch swing arm was found to deviate significantly from the typical t.omposition and hardness of cast 17-4PH (H1100) stainless steel. In this case, the composition clearly did not meet the requirements for 17-4PH and the hardness was owmamummawc 7-2
significantly below that of the other 15 samples. This panicular sample also exhibited an abnormal microstructure, indicating improper heat treatment. Six samples Lviated slightly fmm the compositional requirements of the reference specifications but the hardness for each of these was above the specified minimum. All remaining samples conformed to the compositional and hardness requirements for CB7Cu l (H1100) material in every respect. Weld repairs had been made to ten of the sixteen sample swing arms and,in each cue, the l repairs had been made with an austenitic filler metal rather than one with a composition matching that of 17-4Pil stainless steel. However, most of the weld repairs were shallow and cosmetic in nature and no welding-induced Daws were noted, even for two cases with through-thickness repairs. I in the perfonnance of the fracture toughness tests, complete conformance to all of the validity checks of the governing ASTM specifications was not accomplished. However, the validity requirements in question are strict test specifications and nonconformance does not necessarily indicate invalid data. In those cases which did not strictly conform, the character of the fracture, h and the general nature of crack extension data was typical for valid fracture toughness tests. Overall, the progress of the tests and character of the fractured specimens indicate that the data obtained is reliable and adequate for use in design analyses. Radiographic inspection and subsequent metallographic sectioning identified a few distinct Daws in the larger swing arms. In general, these were relatively small and, except for one case, were embedded within the cross section of the arm. One Gaw was identified as shrinkage cavity with dimensions on the order of 0.1-in. x 0.9-in. All other cases consisted of small fissures or clusters of shrinkage voids. One case was a surface-connected fissure with dimensions on the order of 0.9-in. long x 0.09-in, deep. The remaining Daws examined were all embedded and had major dimensions on the order of 0.7-in. x 0.1-in. The significance of the observed Daws was evaluated by linear clastic fractum mechanics (1.EFM). The evaluation of each flaw was based on a worst-case scenario. For the LEFM analysis, all flaws were assumed to be tight, sharp, surface-connected cracks with dimensions to envelope each complete flaw zone. In addition, the Daws were considered to be oriented transverse to the t maaammmt 7-3
I long axis of the swing arm and located at the point of maximum bending stress. With these very conservative assumptions, and for a bending stress of 60 ksi, the range of ' tress iut msity factors (KJ sssociated with the Daws is 18 to 31 ksi6. These values are well below the range of Ke values detemiined by the fracture toughness tests. The general trends of the dependency of critical fracture stress on crack size, crack shape and inherent fracture toughness are illustrated in Figures 5-21 and 5 22. Figure 5 21 demonstrates that, l for the complete range of crack geometries observed in the group of samples and for a toughness level of Kg. = 50 ksiE, a transversely-oriented flaw would require a through-thickness dimer.sion on the order of 0.4t before failure could occur at an operadng stress of 60 ksi. A second view of the significance of Daw size is presented in Figure 5 22. Here, it can be seen that, for a flaw of g typical geometry with a depth of 0.25t, the inherent toughness of the arm material would have to be 40 ksi6 or less before fracture at an operating stress of 60 ksi would be likely. In the tests and examinations performeu in this program, the largest flaw depth noted was approximately 0.lt and the lower value of K ci measured was 53 ksi6. Again,it sho@i x noted I that all conditions assumed in the LEFM analyses were very conservative. 5 I I I I I E I
- - - - u
8.0 REFERENCES
- l. Aerospace Material Specification 5398D, Steel Castings, Sand and Centrifugal, Corrosion Resistant,16Cr4.1Ni0.22 (Cb+Ta) 2.8 Cu.
- 2. hn sy Specification MIL-II-6875, Pmcess for Heat Treatment of Steel.
- 3. ASTM A 747, Standard Specification t'or Steel Castings, Stainless, liecipitation liardening.
Newman, J. C., Jr., and Raju,1. S., "c.n Empirical Stress-Intensity Factor Equation l 4. for the Surface Crack," Engineering Fracture Mechanics, Vol.15, No,1-2, pp, 18f-192 (1981).
- 5. ASTM Standard E 399 83. " Standard Test Method for Plane-Strain Fracture Toughness of Metallic Materials," Annual Book of ASTM Standards, Vol. 3.01.
- 6. ASTM Standard E 813-88. " Standard Test Method for J,c, A Measure of Fracture Toughness," Annual llook of ASTM Standards, Vol. 3.01,
- 7. liance Information for l Saxena, Common CrackA., and Hudak, Growth S. J., International Specimens," " Review and Extension of Comf of Fracture, Vol.15 Journa No. 5, pp. 453-468 (1978).
- 8. ASTM Standard E 446-81, " Reference Radiographs for Steel Castings up to 2-in.
Thick," Annual Book of ASTM Standards, Vol.1.02. I I l l l l l o m ecua u m m mooc 8-1
L s I I APPENDIX A SWING ARM CONFIGURATIONS l I 16-inch, Sample Nos. 239 and 240: Figures A-I and A-2 10-inch, Sample No. 255: Figun: A-3 8-inch, Sample Nos. 242 and 243: Figures A-4 and A-5 I 6-inch, Samp!c Nos. 244,256, and 257: Figun:s A 6, A-7, and A-8 4-inch, Sample Nos. 245,258,259, and 260: Figures A-9, A-10, A-11, and A-12 3-inch, Sample Nos. 246,261,262, and 263: Figums A-13 A-14, A-15, and A-16 A'I owmmmmmwr
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I APPENDIX C EQUATION FOR TIIE STRESS INTENSITY I FACTOR (K i) FOR A SURFACE CRACK I
I I
I D%tSERGitM380CDOC C-I
The development of the equation for the stress intensity factor for a surface crack, abstracted from Newman and Raju [1] is presented below. The equation is developed for the general case of c
a semi-clliptical surface crack in a member with a rectangular cross section, subjected to combined bending and uniform tensile stresses.
The notations for the deviation are as follows:
S, = remote uniform tensile stress 2c = crack length S. = remote outer fiber bending stress Q = parametric angle of ellipse t = section thickness Q = crack geometry factor 2b = section width and F and H = boundary correction functions For the special case of pure bending and K% at $ = j, the generalized equation reduces to that shown in Section 53 of this report.
STRESS-INTENSITY FACTOR EQU ATION FOR Tile SURFACE CRACK An empirical equation for the stresvintensity factors for a surf ace crack in a nnite plate
[ subjected to tension and bending loads has been 6tted to the 6 nite. clement results from Raju g
and New manl.Ml for a!c salues from 0.2 to 1.0. lo account for the limiting behasior as alc approaches zero, the results of Gross and Srawle>l81 for a single-edge crack hase also been used Two types of loads w ere applied to the surf ace-eracked plate: remote uniform tension and remote bending. The remote unForm tension stress is S, in Fig. 2ia); the remote outer-6ber bending stress S in Fig. 2(b) is ca'eulated from the applied bending moment Af. The stress intensity factor equation for combined tension and bending loads is
& = ( S, + HS. ) (1) nfF('f.f.f.6)
I for O c a'r i 10. O t a'r < 10. c!h c 03 and 0 S 6 ; r A useful approximation for Q.
9 -
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- Fy : Surfur-coded riate suNeaed o tension or vndy toads i
DMBURGRAm3mCDOC C-2
+
deseloped by Rawe and used in Ref [9L n l m L (2)
Q a 1 + 1464(h (" t 1). ,
The functions F and H are defined so that the boundarpeorrection f actor for tension a equal to F and the boundary-correction factor for bending n equal to the product of H and F The function F was obtained from a sptematie curse fitting rrocedure by using Jouble series
~
polynomials in terms of alc. a t, and angular f unctions of 6. The choice of functions was beed on enginee:ir g judgment The function F was taken to be i .
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on is)
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I e=1+ 0.1 + 035 { ' iI - sin e i !. m The function f, an angular function from the embedded elliptical crack solutionll0) n
, s f, = (j):cos 6 + sin 6 @
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l The fun tion H. descioped herein also bs curse ntting and engineermg juJement. h.n the form
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C-3
s I
L 1
I APPENDIX D LOWER BOUND VAL UE FOR Tile FRACTURE TOUGilNESS TEST DATA BY USING STATISTICAL METilODOLOGY APPROACII l
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by Tony Yi Torng I
DWTsatlRGIARD3891 D DI h-1
- l. INTRODUCTION r
L The questions arises as to what would be the lowest fracture toughness expected from any swing arm that has not yet been replaced. In order to search for an appropriate lower bound value for the sixteen experimental fracture toughness tests, the statistical methodology approach is the most reasonable tool. The purpose of this analysis is to provide some statistical assessments and to select the most reasonable characterization of the data. Inherent statistical uncertainty can be characterized, for purposes of analysis, by an analytical model describing the distribution, based on a random sample of data. But the choice of a distribution is very subjective and not at all reliable.
It is necessary to determine which of the several competing statistical models best fit a random I sample of data by using a goodness of fit test methodology. After the distribution model and corn sponding parameters (mean and standard deviation) are determined, a one-sided a% lower bound with fl of confidence can be identified. This lower bound can also be called the " tolerance limit." The reason for using this tolerance limit is due to its ability to provide a certain confidence in handling the uncertainty. Note that there still exists some Gnite possibility of Gnding values below the chosen limit, no matter how small the limit is. In other words, the choice of the limit depends on engineering judgement, since there are lower limits to which K,c values can occur.
- 2. MODEL PARAMETER CALCULATION A total of five distributions hue been selected for this study. They are Normal, Lognormal, Weibull, and Extreme Value Distribution (EVD). The parameters for each distribution are calculated I using the maximum likelihood method. These parameters and the corresponding cumulative distribution function (CDF) are summanzed as follows:
- a. Nonnal p(mean) = 79.1 I c(standard deviation) = 18.76 i CDF=c
( 0 s i
D%1SBURGilARUD-3893 DOC D-2
N
- b. legnormal L p y(mean) = 4.345 o g(Standard Deviation) = 0.22985 F
L CDF=$
f log (X) y'
- ( 04 s
- c. Weibull a = 4.16519
= 86 A2621 CDF = 1 -exp (iT) i
- d. EVD a = 0.06176 p = 70.62699 CDF = exp[-exp(--u(X - p)))
where X represents the fracture toughness random variable.
I
- 3. MODEL SELECTION l To determine the most appropriate model, a g(xxiness of fit test using Cramer-Von Mises statistics was applied. The test uses a statistical value W to select the model and is defined as 4
l W,2=
, N E D,2 i=1 where e
g D, = G, X, ; 6 - F,
< ~)
2 a G, X,; O is the cumulative distribution function (CDF) of the defined distribution model, O is the
( s maximum likeliho(xl estimate of the model, N is the total number of data points, and F is the 3 empirical CDF defined as outs at3GHARIJD-38nIXX' D-3
I F'.='~N , i = 1,N l All the CDF values including the empirical values are listed in Table D-1. Based on the results of the statistics; EVD produces the best fit.
Table D.1 Empirical Data CDF Normal Lognormal Weibull EVD 1 53.5 3.125 % 8.64 % 5.60 % > 12.68% 5.610 %
2 57.1 9.375 % 12.08 % 9.57 % 1630% 9.%0%
3 63.8 15.625 % 20.79 % 20.51 % 24.60 % 21.770 %
4 65.4 21.875 % 2332% 21.71 % 26.88 % 25.130 %
5 65.6 28.125 % 23.64 % 24.12 % 27.18 % 25.560 %
6 67.4 34375% 26.91 % 27.96 % 29.88 % 29.500 %
7 70.1 40.625 % 31.64 % 33.% % 34.16 % 35.590 %
8 74.2 46.875 % 39.77 % 43.39 % 41.12 % 44.840 %
9 74.2 53.125 % 39.77 % 4339% 41.12 % 44.840 %
10 78.2 59.375 % 48,10 % 52.47 % 48.28 % 53.450 %
81.1 65.625 % 5432% 58.73 % 53.57 % 59.230 %
I 11 12 87.8 71.875 % 67.93 % 71.42 % 65.60 % 70.730 %
13 100.4 78.125 % 87.23 % 87.48 % 84.50 % 85.290 %
14 101.2 84.375 % 88.10 % 88.17 % 85.47 % 85.950 %
15 109.7 90.625 % 94.80 % 93.76 % 93.27 % 91.435 %
16 115.3 96.875 % 97.32 % %.00% %39% 93.860 %
W, = 0.07241 0.05303 0.07254 0.04502
- 4. TOLERANCE 1,lMIT It is obvious that when only sixteen samples are available, the sample mean and sample
.I standard deviation are themselves random variables. Thus, there is uncertainty in the parameters themselves, which is renected in the corresponding lower bound of the data. The tolerance limit is the value above which we may predict with a% confidence that Y70 of the population willlie, that is, I
D-4 I
D4tTJiURGil ARI7D 3893 D(X'
1 Prob..,(X t< X < Xu) = #
whereX tandXu repn:sent the lower and upper bounds of the predicted variable,X. For analyr.ing fracture toughness da:a, we are concemed only about the lower bound; therefore, we exp:N te I construct the following:
Prob.o (X > Xt ) = n In Reference s D-l and D-2, tolerance limit tables for Normal distributions are given. However, I the EVD distribution was selected to best reptsent the data. To simplify the estimation of the tolerance limit, a simple assumption has been made for both types of distributions, that is:
The uncertainty occurring in the Normal distribution will be applied to the other distribution types, by assuming the cumulative distribution value at the lower bound of the Normal distribution is the same for other distribution types. The first step is to find the K value from the tolerance limit table, based en the required population and confidence, and find the corresponding CDF value,
' Xm - y cdf = G UN a @(-K)
( s l
where C is the cumulative distribution function for the standard Normal, y and ou are the mean and standard deviation values of the Normal distribution, and Xm is the lower bound value for the Normal distribution. Based on this assumption, the lower bound value,Xt for other distributions can be solved by the following function:
F(X1) = c(-K) where F(Xt) repn:sents the CDF function for other distributions.
This assumption was proven by using the Lognormal distribution to run a Monte Carlo I simulation program to compare with the Normal distribution results. Based on the comparison, it shows the assumption was accurate.
According to the assumption, several levels of population and confidence are solved using the Normal (for comparison) and the EVD distributions, as shown in Table D-2.
I I
I
, _ _ _ _ D.s
Table D 2
[ Lower Bound Normal EVD
~
90% confidence 90% population 44.54 50.72 1 K = 1.842 .
90% confidence 95% population 35.97 46.15 K :. 2.299 90% confidence 99% population 19.59 38.69 K = 3.172 95% confidence 90% population 40.98 49.03 K = 2.032 95% confidence 95% population 31.76 44.09 K = 2.523 1
95% confidence 99% population 14.13 36.5 1 K = 3.463 99% confidence l 90% population 32.99 44.67 K = 2.458 99% confidence 95% population 22.29 39.83 K = 3.028 i 99% confidence 31.96 99% population 1.734 K = 4.124 1
- 5.
SUMMARY
AND DISCUSSION Based on the results shown in Table D-2, the following observations are evident:
. The Normal model predicts smaller lower bounds of K,c, thus, leading to overly conservative results.
owsyn;nonenwn.noc D-6
. The EVD provides the be< goodness of fit to the data. It is seen that for the 99%
l confidence and 99% population case, the EVD predicts a lower bound of fracture toughness of about 32.0 ksi6. From the statistical distributions evaluated, we can thus predict with 99% confidence that 99% of the remaining swing arms presently in service have fractum toughness values gnater than i 32.0 ksi6. Other values of the lower bound for different confide nee levels and different population levels are giver in Table D-2.
- 6. REFERENCES D- 1. Hine., W.W., And Montgomery, D.C., Probabilitv and Statistics in Engineering and Management Science, John Wiley,1980.
D-2. Wirsching, P.H., Reliability Methods in Mechanical and Structural Design, class note, University of Arizona,1985. I 1 I I I l ontsJ;URGHARDD 38nDoc D-7 1}}