ML20147D966
| ML20147D966 | |
| Person / Time | |
|---|---|
| Site: | Clinch River |
| Issue date: | 11/30/1978 |
| From: | Elder G, Prevenslik T, Todd M ENERGY, DEPT. OF, CLINCH RIVER BREEDER REACTOR PLANT, WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML20147D956 | List: |
| References | |
| WARD-D-0178, WARD-D-178, NUDOCS 7812200251 | |
| Download: ML20147D966 (212) | |
Text
{{#Wiki_filter:Io WARD-D-0178 WARD-D Category 11 REVISION 3
'V Clinch River Ih
- I Breeder Reactor Plant CLOSURE HEAD CAPABILITY FOR I STRUCTURAL MARGIN BEYOND l DESIGN BASE LOADIN'G I
I NOVEMBER 1978 I I Prepared for the United States Department of I Energy under contracts EY-76-C-15-2395 and EW 76 C 15 0003. Any Further Distribution by any Holder of this Document or of the Data Therein to Third Parties Representing Foreign Interest, Foreign Govern-I ments, Foreign Companies and Foreign Subsidi-arles or Foreign Divisions of U.S. Companies Should be Coordinated with the Director, Division of Reactor Research and Technology, United I States Department of Energy. I = == @ Westinghouse Electric Corporation J ADVANCED RE ACTORS DIVISION c 3' WJ B0X 158 MA0lSON, PF NNSY LV ANI A 15663
, 781220025(.(
INFORMATION CONCERNING USE OF THIS DOCUMENT PRELIMINARY DOCUMENT This document contains information of a preliminary nature prepared in the course of work for the U.S. Energy Research and Development Administration. This information is subject to correction or modification upon the collec-tion and evaluation of additional data. l NOTICE ' I This document was prepared as an account of work sponsored by the United States Government. Neither the U.S. Energy Research and Development Administration, nor any of their employees, nor'any of their contractors, subcontractors, or their employees, makes any' warranty, express or implied, or assumes any legal liability or responsibility for the accuracy, com-I pleteness or usefulness or any information, apparatus, product or process disclosed, or represents that its use would not infringe privately owned rights. I I I ' WESTINGHOUSE ELECTRIC CORPORATION ADVANCED REACTORS DIVISION B0X 158 MADISON, PENNSYLVANIA 15663 I I ^
I WARD-D-0178 I Revision 3 I Closure Head Capability For Structural Margin Beyond Design Base Loading I November 1978 .I l I lI Prepared by: M.A. Todd T.V. Prevenslik I G.G. Elder I I Approved by: // , fj, ,
- 7) . [ . ' Bi tn'e r , f fan ~ager
/ Structural Analysis I '
I I Westinghouse Electric Corporation I Advanced Reactors Division P.O. Box 158 Madison, Pennsylvania 15663 I E' l 1
I. ! Revision Record Sheet
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I Revision No. Revision Date - _ _ Pages Affected Remarks Original Issue June 1977 Total Document 1 December 1977 Cover, Title, Sunnary, Change TLSM to I i, iii, vii, viii, ix, 1, 2, 11, 12, 14, 15, 16, 19, 27, 28, 34, 37, SMBDB and revised pages to reflect the results of 39, 43, 44, 45, 46, 47, a Fabrication I 50, 52, 53, 54, 55, 60, 64, 75, 96, 98, 100, 103, 105, 107, 109, 110, 118, Non-conformance action to change geometry of I 120, 121, 125, 126, 127, 151, 152, 153, 154, 163, 165, 167, 168, 169, 171, reactor vessel flange in region of the margin 184-192 shear ring groove. 2 August 1978 Cover, Title, ii, iii, Revised text to I vi, vii, viii, 11, 16, 18, 19, 55, 56, 100, 102, 107-110, 117-130, include SRI test results and to eliminate 135 ambiguities about analytical model geometry. I 3 November 1978 Cover, Title, 1, 2, 132, 133, 134 Revise Introduction and Conclusions to reflect I completion of SRI tests. I I I I I IO I
SUMMARY
The results of the most recent and detailed analysis of response of the Clinch River Breeder Reactor Plant Closure Head System Structural Margin Beyond the Design Basis Loading are presented and assessed for Closure Head structural Adequacy. Results obtained from scale model experiments dre sut1Yrlarized and correlated with analytical predictions. Analytical and experimental results both indicate that the Closure Head is capable of withstanding 661 MJ Structural Margin Beyond the Design Base loads without structural failures. I ' I I I l 8 i ' I I J O w
TABLE OF CONTENTS Page
1.0 INTRODUCTION
. . . . . . . . . . . . . . . . . . . . . . . . . . I 1.1 Ba c k g ro u n d . . . . . . . . . . . . . . . . . . . . . . . . . . I 1.2 Purpose . . . . . . . . . . . . . . . . . . . . . . . . . . . I 1.3 Scope , . . . . . . .
. . . . . . . . . , . . . . . . . . . 1 2.0 CRBRP CLOSURE HEAD AND RETENTION FEATURES . . . . . . . . . . . 3 3.0 STRUCTURAL MARGIN BEYOND THE DESIGN BASE LOADING -
- II 4.0 ANALYTICAL PREDICTIONS OF CLOSURE HEAD RESPONSE . . . . . . . . 14 4.1 System Analysis . . . . . . . . . . . . . . . . . . . . . . . 16 4.1.1 Model . . . . . . . . . . . . . . . . . . . . . . . . . . 16 4.1.1.1 Geometry. . . . . . . . . . . . . . . . . .... .. 16 4.1.1.2 Fo rmu l a t i c n . . . . . . . . . . . . . . . . . . . . . 16 4.1.1.2.1 Rotating Plugs. . . . . . . . . . . . . . . . . . 16 4.1.1.2.2 Shear Rings . . . . . . . . . . . . . . . . . . . 20 4.1.1.2.3 Risers ..................... 27 4.1.1.2.4 Shield Plates . . . . . . . . . . . . . . . . . . 27 4.1.1.2.5 Spacer Ears . . . . . . . . . . . . . ..... 28 4.1.1.2.6 Support Skirts. . . . . . . . . . . . . . . . . . 34 4.1.2 Ma terial Properties . . . . . . . . . . . . . . . . . . . 37 4.1.2.1 Stress-Strain Curves 8 4.1.2.2 Damping. . . .
............. 37
.................. 39 4.1.3 Loading Conditions ................. .. 43 i 4.1.4 Pea k Response Resul ts . . . . . . . . . . . . . . . . . . 43 4.1.4.1 Plug Bending Strains ................ 45 4.1.4.2 Shear Ring Deformations . . . . . . . . . . . . . . . 46 4.2 Component Analysis ..................... 46 4.2.1 Models, Material Properties, and Load Conditions . . . . 47 4.2.2 Peak Strains . . . . . . . . . . . . . . . . . . . . . . 47 5 4.3 Structural Criterion and Eiraluations . . . . . . . . . . . . . 47 4.3.1 Criterion. . . . . . . . . . . . . . . . . . . . . . . . 48 4.3.2 Evalua tions . . . . . ................. 51 4.3.2.1 LRP. . . . . . . . . . . . . . . . . . . . . . . . . 51 4.3.2.2 IRP Notch Root . .................. 52 i
I _ TABLE OF CONTENTS (Con't) Page 5.0 EXPEP,IMENTAL PREDICTIONS OF CLOSURE HEAD RESPONSE. . . . . . 55 5.1 Margin Ring Strip Model Tests. . . . . . . . . . . . 56 l 5.1.1 Margin Ring Drop Tests. . . . . . . . . . . . . 57 5.1.1.1 Test Setup and Procedure. . . . . . . . . . . 57 5.1.1.2 Results of Tests . . . . 60 g 5.1.1.2.1 Test No. 1 . . . . . . . . . . . . . . 60 W 5.1.1.2.2 Test No. 2 . . . . . . . . . . . . . . 64 5.1.1.2.3 Test No. 3 . . . . . . . . . . . . . . 64 5.1.1.2.4 Test No. 4 . . . . . . . . . . . . . . 69 5.i.i.3 Caiculation of Drop Weights and Heights. 5.1.2 Margin Ring Static Tests . . . . . . . 75 80 h 5.1.2.1 Test Setup and Procedure. . . . . . . . . . . 80 5.1.2.2 Results of Tests . . . . . . . . . . . . . 81 5.1.2.2.1 Test No. 1 . . . . . . . . . . . . . . 81 5.1.2.2.2 Test No. 2 . . . . . . . . . . . . . . 82 5.1.2.2.3 Test No. 3 . . . . . . . . . . . . . . 84 5.1.3 Discussion of Results . . . . . . . . . . . . . 91 5.2 Margin Ring Dynamic TesL . . . . . . . . . . . . . 96 5.2.1 Test Specimons . . . . . . . . . . . . . . . 96 5.2.2 Conduct of Test . . . . . . . . . . . . . . . 98 5.2.3 Selection of Dynamic Test Conditions . . . . . . . . 98 5.2.4 Results and Discussion. . . . . . . . . . . . . 100 5.2.4.1 Test No. 1 . . . . . . . . . . . . . .
. 104 5.2.4.2 Test No. 2 . . . . . . . . . . . . . .
104 5.2.4.3 Test No. 3 . . . . . . . . . . . . . . 104 5.2.4.4 Test No. 4 . . . . . . . . . . . . . . 104 5.2.4.5 Static Load-Deflection Test. . . . . . . . . 106 . 5.2.4.6 Discussion of Model Response Time. . . . . . . 106 5.2.5 Conclusions . . . . . . . . . . . . . . . 106 5.3 Reactor Scale Model Tests * - - - - - - . - . . - 110 5.3.1 Test SM-1: Static Closure Head Response . . . . . . 112 5.3.2 Tests SM-2 and SM-3: Dynamic Response of a Simplified Reactor System . . . . . . . . . . . . . . . 117 5.3.3 Tests SM-4 and SM-5: Dynamic Response of Complex Models. 123 5.3.4 Conclusions . . . . . . . . . . . . . . . . I '"' ii E
TABLE OF CONTENTS (Con't) Page 1 6.0 ASSESSMENT OF CLOSURF HEAD CAPABILITY. . . . . . . . . . 131 g 6.1 Capability of Margin Rings. . . . . . . . . . . . . 131 r 6.2 Capability of Rotating Plugs and Vessel Flange . . . . . . 132 6.3 System Capability. . . . . . . . . . . . . . , . 132 6.4 Independent Assessment of LRP Margin Ring. . . . . . . . 134
7.0 REFERENCES
. . . . . . . . . . . 135 i Appendix A Prediction of Structural Response to Transient Excitation by Scale Model Testing. . . . . . . . .
136 Appendix B Mechanical Properties of Materials used in Test Models. . 149 Appendix C Significance of Excluding Momentum Transfer in Analytical Predictions of CRBRP Closure Head Capability Under SMBDB Loading. . . . . . . . . . 157 Appendix D Effect of Fabrication Non-Conformance on Response of Reactor Vessel Flange to Structural Margin Beyond Design Base Loading . . . . . . . . . . . 190 I I 1 i 1 1 I l l l3 r iii M
l LIST OF FIGURES ' Figure No. Title Page 2.0-1 Reactor Schematic. . . . . . . . . . ...... .. 4 gl 2.0-2 CRBRP Closure Head Plan. . . . . . . . . . . . . . . 5 m' 2.0-3 CRBRP Closure Head Elevation . ........... 6 2.0-4 CRBRP Riser Layout . . ....... .... .... 8 ) 2.0-5 Margin Ring Sections . . ...... .. ... . .. 10 3.0-1 HCDA Force-Time History . .... ... ..... .. 13 4.1-1 3-D Half Symmetry Model of the Closure Head Assembly 17 4.1-2 SRP, IRP, and LRP 3-D Finite Element Model . . . . . 18 4.1-3 SRP Shear Ring Locally Adjacent SRP and LRP Region . 21 4.1-4 SRP Shear Ring 2-D Axisynunetric Finite Element Model 22 4.1-5 Standard and Modified ANSYS (STIF 40) Dynamic Combi na ti on Elements . . . . . . . . . . . . . . . . 23 4.1-6 SRP Shear Ring Static Force-Deflection Curve and Bilinear Approximation. . . . . . . . . . . . . . . 24 h 3 4.1-7 IRP Shear Ring Static Force-Deflection Curve and Bilinear Approximation. . . . . . . . ....... 25 4.1-8 LRP Shear Ring Static Force-Deflection Curve and Bilinear Approximation. . . . . . . . . . . . . . . 26 4.1-9 SRP Shield Plate Model. . . . . . . . . . . . . . . 29 4.1-10 IRP Shield Plate Model. . . . . . . . . . . . . . . 30 4.1-11 LRP Shield Plates. . . . . . . . . . . . . . . . .. 31 4.1-12 SRP Upper Spacer Bar Static Force-Deflection Curve l' Bilinear Approximation . . . . . . . . . . . . . . . 33 4.1-13 IRP Upper Spacer Bar Static Force-Deflection Curve Bilinear Approximation . . . . . . . . . . . . . . . gl W 35 4.1-14 LRP Upper Spacer Bar Static Force-Deflection Curve Bilinear Approximation . . . . . . . . . . . . . . . 36 l m 4.1-15 Rotating Plug Stress-Strain Curves . . . . . . . . . 40 , 4.1-16 Shear Ring Stress-Strain Curve Average Properties at 400 F. . . . . . . . . . . . . . . . . . . . . . 41 4.1-17 Upper Spacer Bar Stress-Strain Curve Average $ Properties at 400 F. . . . . . . . . . . . . . . . . 42 3 4.3-1 Triaxiality Functions vs. Experimental Local Ductile Rupture Failure Data. . . . . . . . . . . . 49 g iv E'
I LIST OF FIGURES (Con't) Figure No. Title Page 5.1 -1 Margin Ring Drop Test Specimen Components. . . . . . . 58 5.1-2 Drop Tes t Sc hema ti c . . . . . . . . . . . . . . . . . . 59 5.1-3 Test. No. 1 Specimen After Test. . I 5.1-4 Test No. 1 Right Hand Margin After Test. . . . . . . . 61 62 5.1-5 Test No. 1 Left Hand Margin After Test . . . . . . . . 63 1 5.1-6 Test No. 1 Deformed Shape of Right Hand End. . . . . . 65 5.1-7 Test No.1 Deformed Shape of Left-Hand End . . . . . . 66 5.1-8 Test No. 2 Deformed Shape of Right Hand End. . . . . . 67 5.1-9 Test No. 2 Deformed Shape of Left-Hand End . . . . . . 68 5.1-10 Test No. 3 Specimen After Test . . . . . . . . . . . . 70 5.1-11 Test No. 3 Right-Hand Margin After Test. . . . . . . . 71 5.1-12 Test No. 3 Left Hand Margin After Test . . . . . . . . 72 1 5.1-13 Test No. 3 Deformed Shape of Right-Hand End. . . . . . 73 5.1-14 Test No. 3 Deformed Shape of Left Hand End . . . . . . 74 I 5.1-15 Test No. 4 Drop Weight Initial Impact Sequence . . . . 74a 5.1-16 Test No. 4 Deformed Shape of Right-Hand End. . . . . . 76 5.1-17 Test No. 4 Deformed Shape of Lef t-Hand End . . . . . . 77 5.1-18 Test No.1 Force Deflection Curves Low Carbon I 5.1-19 Steel Specimens. . . . . . . . . . . . . . . . . . . . Test No. 2 Force Deflection Curves SA-508 Specimens. . . . . . . . . . . . . . . . . . . . . . . 83 85 5.1-20 Test No. 2 No Load . . . . . . . . . . . . . . . . . . 86 5.1-21 Test No. 2 PEL . . . . . . . . . . . . . . . . . . . . 87 5.1-22 Test No. 2 PI. . . . . . . . . . . . . . . . . . . . . 88 5.1-23 Test No. 2 Test Set-Up-Front . . . . . . . . . . . . . 89 5.1-24 Test No. 2 Test Set-Up-Back. . . . . . . . . . . . . . 90 I 5.1-25 Test No. 3 Test Set-Up-Back. . . . . . . . . . . . . . 92 5.1-26 Test No. 3 Force-Deflection Curves SA-508 Specimens. . . . . . . . . . . . . . . . . . . . . . . 93 5.1-27 Test No. 3 Shear Ring Rotation Second Loading Cycle. . 94 4 I V .I
LIST OF FIGURES (Con't) I Fiqure No. Title Page 5.2-1 Margin Ring Dynamic Test Exploded View of Specimen . . 97 5.2-2 Margin Ring Dynamic Test, Test Set-Up . . . . . . . . 99 5.2-3 ] Force-Deflection Characteristics . . . . . . . . . . . i07 5.2 4 Force Applied to Head Specimen . . . . . . . . . . . 108 5.3-1 Static Test SM-1: Test Model Schematic . . . . . . . 113 1 5.3-2 SM-1 Static Test SM-1: Instrumentation . . . . . . . 114 l 5.3-3 Static Test SM-1: Deformed Geometry . . . . . . . .. 115 5.3-4 Static Test SM-1: Pressure-Volume Relationship. . . . 116 5.3-5 Schematic of Test SM-3 Showing Instrumentation , Layout . . . ... ................ 118 5.3-6 Schematic of Test SM-2 Showing Instrumentation . Layout . ...................... 11'3 ' 5.1-7 Comparison of loadir.g Pressures: SM-2 and SM-3 . . . 121 5.3-8 Comparison of Accelerations: SM-2 and SM-3 . . . . . 122 5 W 5.3-9 SM-4 With Instrumentation. .
............ 124 5.3-10 SM-5 With Instrumentation. . . . .......... 125 5.3-11 Comparison of Loading Pressures: SM-4 and SM-5 . . . 126 5.3-12 Comparison of Accelerations: SM-4 and SM-5 . . . . . 127 5.3-13 SM-5 Vessel Head Accelerations and Displacement. .. 129 9-1 Comparison of Stress-Stra.in Curve.s of Experimental and CRBRP Closure Head Materials. . . . a 152 g B-2 Comparison of Stress-Strain Curves for Experimental and CRBRP Shear Ring Materials . . . .
153 C.2-1 LRP Shear Ring Region 2-Dimensional Axisymmetric Model Including Momentum Transfer . . . . . . . . . 163 C.2-2 Continuous and Plane Stress Static Force Defl ection Curves . . . . . . . . . . . . . . . . . . 164 C.2-3 LRP Shear Ring Region 1 Dimensional Model g Excluding Momentum Transfer . . . . . . . . . . . . . 166 g C.2-4 LRP Shear Ring Region Non-linear Spring Representation . . . . . . . . . . . . . . . . . . . 167 C.6-1 Significance of Momentum Transfer Force-Deflection Curves LRP Impact Velocity
- 73 in/sec. ...... 179 C.6-2 Significance of Momentum Transfer Inner LRP Periphery Displacement-Time Histor Velocity s 73 in/sec. . . . . . . y LRP Impact
,......... 180 C.6-3 Significance of Momentum Transfer Notch Root Equivalent Strain vs. Inner LRP Periphery Displacement LRP Impact Velocity s 73 in/sec. . . . . 182 vi
LIST OF FIGURES (Con't) Figure No. Title Page, C.6-4 Significance of Momentum Transfer Equivalent Stress Distributions (Contours 4000 psi) LRP Impact Velocity s 73 in/sec. . . . . . . . . . . . . 184 C.6-5 Significance of Momentum Transfer Force-Deflection Curves LRP Impact Velocity s 73 in/sec. . 185 C.6-6 Inner LRP Periphery Significance of Momentum i
- Transfer Displacement-Time History LRP Impact Velocity s 300 in/sec. . . . . . . . . . . . . . . . 187 j C.6-7 Significance of Momentum Transfer Notch Root B Equivalent Strain vs. LRP Periphery Displacement LRP Impact Velocity s 300 in/sec. . . . . . . . . . . 188
[p C.6-8 Significance of Homentum Transfer Equivalent Stress Distributions (Contours s 4000 psi) LRP Impact Velocity
- 300 in/sec. . . . . . . . . . . 189 I D.1 Reactor Vessel Flange Groove for LRP Margin Ring As-Built Dimensions. . . . . . . . . . . . . . . 192 D.2 Reactor Vessel Vlange Symbols and Definitions . . . . 196 i
I l vii
l I LIST OF TABLES _TabieNo. Title Page ! 4.1-1 Material Temperatures and Elastic Constant for the I' Plugs and Shield Plates. . . . . . . . . . . . . . . . . . . 38 4.1-2 SMBDB Loading Pressure Time History 8
... ......... 44 5.2-1 Dynamic Test Conditions. . . . . . . . . . . . . . . . . . 101 ,
5.2-2 CRBRP Head Conditions Modeled. . . . . . . . . . . . . . . 102 gl 5.2-3 5i Test Conditions. . . . . . . . . . . . . . . . . . . . . . 103 ' 5.2-4 Residual Plastic Deformation . . . . . . . . . . . . . . . 105 ; A.1 Important Variables in Structural Dynamics . . . . . . . . 142 A.2 Dimensionless Products for Structural Dynamics . . . . . . ' 143 A.3 Scale Factors for Structural Dynamics. . . . . . . . . . . 144 B.1 CRBRP Closure Head Material Properties . . . . . . . . . . 154 B.2 Test Specimen Material Properties. . . . . . . . . . . . . 155 B.3 Comparison of Test and CRBRP Properties. . . . . . . . . . 156 C.2-1 Bilinear Material Properties . . . . . . . . . . . . . . . 161 C.2-2 LRP Shear Ring Region Spar Non-Linear Properties . . . . . 165 , I I I I E I viii E
LIST OF ACRONYMS ASME -- American Society of Mechanical Engineers ASTM -- American Society for Testing and Materials CRBR -- Clinch River Breeder Reactor CRBRP -- Clinch River Breeder Reactor Plant CRDL -- Control Rod Driveline CRDM -- Control Rod Drive Mechanism ERDA -- V. S. Energy Research and Development Agency EVTM -- Ex-Vessel Transfer Machine HCDA -- Hypothetical Core Disruptive Accident IRP -- Intermediate Rotating Plug IVTM -- In-Vessel Transfer Machine LRP -- Large Rotating Plug NRC -- U.S,. Nuclear Regulatory Commission SMBDB -- Structural Margin Beyond the Design Base SRP -- Small Rotating Plug VIS -- Upper Internals Structure B B E I I 1 B
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1.0 INTRODUCTION
l
1.1 Background
The Clinch River Breeder Reactor Plant (CRBRP) uses liquid sodium as its coolant. Because this substance must be maintained at an elevated tempera-I ture and in an inert atmosphere, an under-the-head refueling capability was incorporated into the design. The reactor Closure Head consists of three nested rotating plugs, eccentrically arranged, by means of which the refueling machinery may be positioned over any assembly in the reactor core without the necessity of removing the head. The U.S. Nuclear Regulatory Commission (NRC) has expressed concern about I the adequacy of the Closure Head Design to withstand the underhead loads imposed by a Hypothetical Core Disruptive Accident (HCDA). This concern has been expressed in numerous requests for additional information on the CRBRP Preliminary Safety Analysis Report (PSAR). 1.2 Purpose The twofold purpose of this report is (1) to present the most recent analytical and experimental information on the adequacy of the CRBRP l Closure Head to withstand the loads associated with a particular 661 MJ HCDA, the Structural Margin Beyond the Design Base (SMBDB) condition; and (2) to provide a comprehensive technical reference document to substantiate the CRBRP Project answers to Round 2 PSAR Questions related to SMBDB evalua-tion of the CRBR Closure Head. I 1.3 Scope This report contains a section describing the CRBRP Closure Head design in its present configuration. The Structural Margin Beyond the Design Base scenario is briefly described and the predicted underhead force-time history is presented. The results of a detailed analysis of the Closure l Head in three dimensions is presented, along with supporting analysis to verify the simplifying assumptions embodied in the system model. A con-siderable amount of experimental work has been performed, including 1 1
-l static and dynamic tests of scaled strip specimens of the plug margin ring area and of scaled axisymmetric models of the same. In addition, a series of reactor scale model tests have been performed. The results of these tests are incorporated into this report. The section on evaluation of structural adequacy contains an assessment by an independent consultant, retained for this purpose.
The results of analysis and experiment were found to be largely in agree-ment. Both analysis and experiment supports the conclusion that the CRBR Closure Head will withstand the 661 MJ SMBDB loads without structural failure. I I I I I I I I I . I a Og I
2.0 CRBRP CLOSURE HEAD AND RETENTION FEATURES i The primary containment of the Clinch River Breeder Reactor is provided by the Reactor Vessel and Closure Head Assemblies. The Reactor Vessel and Closure Head are illustrated in Figure 2.0-1. As shown in that figure, l the Closure Head Assembly is supported and retained by the Reactor Vessel y Flange. The CRBRP Closure Head consists of three nested plugs, called the Large Rotating Plug (LRP),' Intermediate Rotating Plug (IRP) and Small Rotating Plug (SRP). The LRP fits into and is supported by the Reactor Vessel Flange; the IRP fits into an eccentric hole in the LRP; and the SRP fits into an eccentric hole in the IRP. The overall dimensions of the three I plugs are as follows. I 1. The LRP has a nominal outside diameter of 257.38 inches and a nominal thickness of 22.00 inches. Its centerline coincides with the reactor centerline.
- 2. The IRP has a nominal outside diameter of 175.50 inches and a nominal thickness of 22.00 inches. Its centerline is offset 27.20 inches from that of the LRP.
- 3. The SRP has a nominal outside diameter of 67.94 inches and a nominal thickness of 22.00 inches. Its centerline is offset B 42.85 inches from that of the IRP.
The arrangement of the three plugs in their operating configuration is illustrated in Figure 2.0-2. Attached to and supported by the three plugs are the underhead shield I plates, reflector plates and suppressor plates. The lower shield plate, 10 inches thick, rests on an outer cylindrical support skirt which is in turn pinned to the underside of the respective rotating plug. The inter-mediate shield plates,10 inches thick, and upper shield plates, 9 inches thick, rest on spacer rings located near the outer and inner edges of the plates. The reflector and suppressor plates are suspended from columns which are bolted to the lower shield plates. This arrangement is illustrated in Figure 2.0-3. 1o , h l 3 l I ; 1
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en k Figure 2.0-3 I CRBRP Closure Head Elevation 6 08 E
Major head mounted components are as follows: (1) The In-Vessel Transfer Mechanism (IVTM) Nozzle is attached to the SRP at the location indicated in Figure 2.0-3. (2) The Ex-Vessel Transfer Mechanism (EVTM) Nozzle is attached ! to the LRP at the location indicated in Figure 2.0-3. (3) The 15 Primary and 4 Secondary Control Rod Drive Mechanisms (CRDM) and Control Rod Drivelines (CRDL) are attached to the IRP at the locations shown in Figure 2.0-3. (4) The Shield and Seismic Support is attached to the IRP and encloses the Primary and Secondary CRDM. (5) The Upper Internals Structure (UIS) and UIS Jacking 1 Mechanisms are supported by the IRP at the locations indicated in Figure 2.0-2. The UIS is suspended below the Closure Head as shown in Figure 2.0-1. The three plugs are suspended from their bearings by a system of inner and outer risers. This plug-riser arrangement is shown in Figure 2.0-4. Typically, the inner riser is a thin-walled cylindrical shell bolted to the periphery of the plug. The inner riser engages the riser bearing, which is supported by the outer riser. The outer riser is bolted to the periphery of the corresponding cutout. In this manner, the SRP is supported by the IRP, the IRP by the LRP, and the LRP by the Reactor Vessel Flange. The three plugs are independently rotatable, being driven through gears attached to the inner risers. Rotation of the three plugs allows handling machinery attached to the SRP to be positioned over any element in the core for re-moval and replacement during refueling. i Under normal conditions of operation and refueling, the plugs are held in position by the ricers at. of thet own weight. In the event that abnormally high upward loads are applied to the plugs (as during an HCDA), plug re-I tention is accomplished by the LRP, IRP and SRP Margin Rings. These are segmented rings seated in grooves machined into the peripheries of the Reactor Vessel Flange, LRP and IRP cutouts and held in place by continuous l filler rings. The locations of the margin rings and filler rings for the
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l l l l 3.0 STRUCTl>RAL MARGIN BEYOND THE DESIGN BASE LOADING If certain hypothetical conditions are assumed to exist within the reactor, then it is possible to conservatively predict overheating of the core, I with resulting partial fuel vaporization, core disassembly and pressuri-l zation of tne reactor vessel. Any event of this nature is called a Hypothetical Core Disruptive Accident (HCDA). Since pressurization of the I reactor vessel could potentially lead to a breach of the primary containment, these events though hypothetical, have been evaluated. Several such hypothetical events have been studied (Reference 8). One set of parameters was selected for detailed mechanical analysis. This set of parameters is identified as the Structural Margin Beyond the Design I Base (SMBDB) HCDA, and the corresponding structural loads are referred to as SMBDB loading. The major parameters associated with this condition are as follows. e Average core temperature following disassembly is 4800 K. e Peak core pressure is 3960 psia. e Work energy for expansion of the fuel vapor bubble to one atmosphere is 661 megajoules. The above values envelop the hypothetical events considered, but do not represent any particular accident sequence. I Expansion of a fuel vapor bubble within the core barrel will first de-form the structures surrounding the core, and then accelerate a slug of sodium coolant upward. The impact of this sodium slug against the under-side of the Closure Head Assembly imposes transient dynamic loads on the rotating plugs, plug margin rings and vessel flange. I A computer simulation was developed for the coupled hydrodynamic and struc-tural response of the reactor system to SMBDB loading. A predicted force-I time history for loading of the underside of the closure head was obtained o ( i b
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I
l \ l frem this analysis and is 111ustr6ted in Figure 3.0-1. This time history is the Structural Margin Beyond the Design Base (SMBDB) loading for the CRBRP Closure Head. I 1 l I I
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I 4.0 ANALYTICAL PREDICTIONS OF CLOSURE HEAD RESPONSE Analytical predictions of CRBRP Closure Head System dynamic response to SMBDB loading as defined by the ANL/REXCO code using various 1D and 3D finite element models have been made since mid-1974. Fundamental in the analytical methods were simplifications which permitted predictions to be readily obtained without a significant loss in accuracy. Analytical simplifi-cations are important in derivations of CRBRP closure head system response , because detailed finite element models can render a solution impractical l because of exhorbitant computing time. ' A notable simplification common to the 10 and 30 analytical system response predictions was to exclude local momentum transfer in the SRP, IRP, and LRP shear rings and adjacent plug materials during impact caused by shear ring gap closure under SMBDB loading. Accordingly, the dynamic force-deflection l behavior of the regions defined by the shear rings and adjacent plug materials as a whole was considered to not significantly differ from static l behavior. Alternately, the SRP, IRP, and LRP velocities were not considered sufficiently high to cause significant local yielding associated with local momentum transfer at impact surfaces which would not be observed in static behavior. ' In the derivation of the CRBRP Closure Head System dynamic response, the static force-deflection properties of the SRP, IRP, and LRP shear rings and adjacent plug materials ware derived with relatively detailed 2D axisymmetric models. Early static force-deflection derivations assumed linear elastic shear rings and plug materials which subsequently were up-dated to include plasticity. In 10 CRBRP closure head dynamic response derivations, the 2D axisymmetric force-deflection curves were approximated with a single 1D non-linear perfectly plastic or kinematic hardening relation. For derivations of 3D CRBRP closure head response, the full 2D axisymetric shear rings properties were represented by a number of single 10 non-linear finite elements along the SRP, IRP, and LRP shear ring peripheries. In this arrangement, the complex 2D axisymmetric finite ele-ment shear ring models were reduced to very simple 10 non-linear finite I 14 E
)
I elements, thereby providing dynamic system response solutions with minimal I e computing time and without a significant loss in dynamic response accuracy of the CRBRP closure head. The analytical predictions of CRBRP closure head system response include displacement and acceleration-time histories which were used in performing structural evaluations of head mounted components. With regard to the SRP, IRP, and LRP, the peak bending strains were sufficiently detailed in the systems response to permit direct structural evaluation of the plug materials in regions removed from the local regions of the shear rings. However, the i structural evaluation of the SRP, IRP, and LRP shear rings and adjacent plug materials required subsequent component analysis to detennine the detailed strain state in order to assure that the plugs do not separate from each other during SMBDB loading. Peak shear ring and adjacent plug displacements i obtained from the dynamic system response were imposed as static displace- , i ment loadings on the detailed 2D axisymmetric shear ring models used in de-riving the static force-deflection properties. Subsequent locntions and l magnitudes of peak strain were compared with criteria to establish the ' structural integrity of the shear rings and adjacent plug materials. In summary, ',he analytical predictions of CRBRP closure head response to i SMBDB loadin;, were based on 1D and 30 models with simplifications that ex-cluded local momentum transfer at impacting surfaces of the SRP, IRP, and LRP shear rings. An analytical justification for using static force-deflection properties of SRP, IRP, and LRP shear rings and adjacent plug materials for the range of impact velocities associated with HCDA loading is presented in Appendix C.
)
I In the following, the most recent analytical prediction of 30 CRBRP Closure Head System dynamic response to SMBDB loading in combination with a structural integrity assessment of the SRP, IRP, and LRP shear rings and adjacent plug materials is presented. a t t + 15 v.
4.1 System Analysis I The ANSYS finite element program (Reference 1 ) was used to derive the 3D dynamic system response of the CRBRP closure head to SMBDB loading. Descrip-tions of the ANSYS model, material properties, and SMBDB loading conditions are as follows. 4.1.1 Model I The 3D ANSYS model of the CRBRP closure head was based on the 180 half symmetry sector illustrated in Figure 4.1-1. 4.1.1.1 Geometry I The 3D ANSYS model of the CRBRP Closure Head was formulated in relation to the geometry specified for the Baseline Design. 4.1.1.2 Formula tion I The formulation of the 3D ANSYS model in relation to the geometry is. described for the SRP, IRP, and LRP with respective shear rings, shielding plates, spacer bars, and support skirts. g 4.1.1.2.1 Rotating Plugs The finite element mesh selected for the SRP, IRP, and LRP is illustrated In pigure 4.1.2. The SRP was modeled with 32 ANSYS (STIF45) isoparametric solid elements in two layers through the thickness. The IVTM penetration hole was in-cluded in the SRP representation. The IVTM nozzle was represented by a g W single ANSYS (STIF21) concentrated mass element suspended at the center of the penetration hole by 5 AllSYS (STIF4) beam elements of low mass and very high stiffness. The SRP Material (SA-508) was taken to be linearly elastic throughout the SMBDB loading, an expectation which proved to be consistent with analysis results. Accordingly, the ANSYS substructuring ' option was used to reduce the assemblage of 38 elements representing the ' SRP to a single linear elastic ANSYS (STIF50) superelement. l 16 I
I M M. M M M M M M M M 8 M M M M M M M M O O Y A i I I I LRP t t I IRP 8 i
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I The IRP was modeled with 83 ANSYS (STIF26) plastic flat plate elements. Since the IRP material (SA-508) was expected to yield under SMBDB loading, elastic-plastic properties were used for the flat plate elements. The region of the 19 CRDM penetrations was modeled with reduced elastic and plastic material properties to simulate the weakening effe'ct of the penetra-tions. The CRDM penetrations were not explicitly modeled. The two VIS penetrations were explicitly modeled, the UIS mass being represented by I an ANSYS (STIF 21) concentrated mass element suspended at the center of each penetration hole by six ANSYS (STIF4) beam elements of low mass and high stiffness. The LRP was modeled with 68 ANSYS (STIF26) plastic flat plate elements. Since the LRP was expected to yield under SMBDB loading, the plug material (SA-508) was modeled as elastic-plastic. The EVTM penetration hole was explicitly modeled, with the EVTM nozzle represented by a single ANSYS (STIF21) concentrated mass element suspended at the center of the penetration I hole by six ANSYS (STIF4) beam elements of low mass and high stiffness. The CRBRP Closure-Head-mounted equipment masses were modeled as ANSYS (STIF 21) concentrated mass elements at the appropriate nodes on the SRP, IRP and LRP. Rotary inertias associated with equipment of extended shape extending above the head were included as lumped moments of inertia. The rotary inertia of equipment of extended shape extending below the head was neglected. I I I I !o V 19
I 4.1.1.2.2 Shear Rings The SRP, IRP, and LRP shear rings were rot directly included in the CRBRP closure head dynamic system response model as the effects of local momentum transfer were neglected. Instead, 2D axisymmetric models of the shear rings and locally adjacent plugs were formul.ated and static force-deflection pro-perties derived. The proccdure was repeated individually for the SRP, IRP, and LRP shear rings. The locally adjacent extent of the SRP ard IRP and respective 2D axisynrnetric SRP shear ring nodel typical of the approach taken for the IRP and LRP shear ring is illustrated in Figures 4.1-3 and 4.1-4 respectively. The static force-deflection properties of the SRP, IRP, and LRP shear rings i were approximated with the classical bilinear kinematic hardening relation ! using a modified version of the ANSYS (STIF 40) dynamic combination element. l The standard ANSYS (STIF 40) element provides the classical bilinear elastic-perfectly plastic relation by combining a linear spring in series with a l slider. In addition, a viscous dashpot in parallel with the spring and slider as well as a gap in series with the combination thereof is included l in the standard ANSYS (STIF 40) formulation. The modified ANSYS (STIF 40) includes an additional linear spring in parallel with the series combination of the as formulated linear spring and slider. The standard and modified ANSYS (STIF 40) dynamic combination elements are illustrated in Figure 4.1-5. The classical bilinear kinematic hardening approximations to the statically derived force-deflection curves of the SRP, IRP, and LRP shear rings and respective locally adjacent plug materials are illustrated in Figures 4.1-6 through 4.1-8. In the 3D ANSYS CRBRP closure head model, the full 23 axisymmetric static force-deflection properties of the shear rings and respective classical bilinear kinematic hardening approximations .tre not directly applica' ale. Fractions of the full circumferential shear ring stiffnesses and yield forces in direct proportion to the distance between the finite element nodes along the plug peripheries were taken to represent local 3D properties. Each local shear ring region was modeled with the modified ANSYS (STIF 40) 20 OR I
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.40 i
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t : Static Force-Deflection Curve . __j_ _ ,t-2._2__ ___; _ 4 t
-r f
- I
; and Bilinear ADoroximation -j t } , 30 . F , . : 5 , i
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dynamic combination element. Impact damping was neglected and a nominal 1/8 in. shear ring gap was considered. The SRP, IRP, and LRP shear rings 9 were modeled along their 180 circumferential extent by 8, 14, and 15 modi-fied dynamic combination elements, respectively. 4.1.1.2.3 Risers The SRP, IRP, and LRP risers were treated as linear elastic in the 3D ANSYS model. The 2D axisynnetric axial stiffness of the inner and outer risers was determined. Mechanical bearing stiffness was considered rigid. The 6 SRP, IRP, and LRP axial stiffness were 22.47, 26.83 and 22.28 x 10 lb/in. respectively. The full 2D axisymmetric axial stiffness was simulated with a discrete number of linear elastic springs along the y riphery of the inner and outer risers. Standard ANSYS (STIF 40) dynamic combination elements with an infinite yield force and without damping were used to model the riser assemblies. A small gap was assumed in all dynamic combination elements. The SRP, IRP, and LRP riser assemblies were modeled over che 180 peripheries with 8,14, and 15 dynamic combination elements respectively. 4.1.1.2.4 Shield Plates The SRP, IRP, and LRP shield plates were modeled with ANSYS (STIF 13) flat triangular shell and ANSYS (STIF 63) flat quadrilateral shell elements. The shield plates were assumed to remain linear elastic under HCDA loading. Individual upper, middle, and lower shield plates for each rotating plug were substructured to form single linear elastic ANSYS (STIF 50) super-elements. The upper and middle shield plates were modeled with an identical I finite element mesh for each rotating plug. The respective lower shield plate, however, was modeled with a finer mesh as SMBDB loading pressures act directly on the lower shield plates. The SRP upper and middle shield plates were individually modeled with 10 quadrilateral and 1 triangular flat shell elements. Excluding the 'VTM opening, the SRP lower shield plate was modeled with 17 quadrilateral and I I I
3 triangular flat shell elements. The SRP upper, middle, and lower shield I-plate models are illustrated in Figure 4.1-9. The IRP upper and middle shield plates excluding VIS column openings were individually modeled with 40 quadrilateral and 6 triangular flat shell ele- I ments. Excluding the UIS columns and miscellaneous holes, the IRP lower shield plate was modeled with 60 quadrilateral and 13 triangular flat l shell elements. The IRP upper, middle, and lower shield plate models are illustrated in Figure 4.1-10. 1 The LRP upper and middle shield plates excluding the EVTM nozzle opening was modeled with 37 quadrilateral and 9 triangular flat shell elements. l Excluding the EVTM nozzle and miscellaneous holes, the LRP lower shield plate was modeled with 49 quadrilateral and 21 triangular flat shell elements. The LRP upper, middle and lower shield plates are illustrated in Figure 4.1-11. 4.1.1.2.5 Spacer Bars The spacer bars are separately distinct from the shield plates. Under the action of gravity, the tolerance buildup across the stack of spacer bars and shield plates prcduces a nominal 1/16 ia. clearance between the top surface of the upper spacer bars and the bottom surface of the respective rotating plug. Only the gaps between the upper spacer bars and rotating plugs were modeled. Otherwise, the spacer bars were considered connected l without separation for the duration of the SMBDB loading. In this modeling l i arrangement, the upper, middle, and lower shield plates and respective spacer bars were assumed to move as a continuous structure with impact occurring at the upper spacer bar to rotating plug interface. -l The SRP, IRP, and LRP spacer barr are of rectangular cross-section and of segmented construction along their circumferential extent. The lower and middle spacer bars were assumed linear elastic while the upper spacer bars were taken to behave in the manner of a classical bilinear kinematic harden-ing material. The full 2D axisymmetric static force-deflection behavior 28 G,I I
Upper and Middle:- i N l / \ I I
- N Lower I ~
I i l I I
/ /[:1 Figure 4.1-9 SRP Shield Plate Model l IO "
c I
1 Upper and Middle I E I
\
d'
/
l I.
- l. \ \ g I
'""N - \ g
~ ~ ,
\/ I I
R Figure 4.1-10 IRP Shield Plate Model I 30
@I I
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t
\
\
p/ - pper and IMddI* {/////$i k \
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. f/ff/t.
b piqure 4 I' ggp ShieTd pTates P 31
for the lower and middle spacer bars follows directly from linear elastic theory. However, the non-linear force-deflection behavior of the upper spacer bars assumed constant volume for plastic flow in combination with the stress-strain relations of the material for an initial area (A g) and length (l g) with the upper spacer bar material stress (c) and strain (c) relation, the force (F) and deflection relation of the upper spacer bars was derived according to the following: P = Ago( ) S = cl g where,o=a(c) The 20 axisyrrmetric linear elastic force-deflection behavior of the lower and middle spacer bars as well as the non-linear force-deflection behavior of the upper spacer bars was transformed into local 30 behavior by selecting a fraction of the properties in relation to the geometry of the shield plate finite element mesh. The standard ANSYS (STIF 40) dynamic combination element with an infinite yield and without damping and gaps was used to model local 30 behavior of the lower and middle spacer bars. The modified ANSYS (STIF 40) dynamic combination element with classical bilinear kinematic hardening and a nominal gap (1/16 in.) and without damping was used to model the local 30 behavior of the upper spacer bars. s The SRP spacer bar force-deflection properties were discretired at 6 locations around the outer peripheries of the shield plates. A total of 12 standard ANSYS (STIF 40) dynamic combination elements were used to mocel the lower and middle spacer bars. A total of 6 modified ANSYS (STIF 40) were used to model the upper spacer bars. The classical bilinear kinematic hardening approximation to the full 2D axisymmetric force-deflection of the upper SRP spacer bar is illustrated in Figure 4.1-l?. ll The IRP spacer bar force-deflection properties were discretized at 6 locations along the inner shield plate peripheries and 12 locations along the outer shield plate peripheries. Along the inner and outer peripheries of the lower and middle shield plates, a total of 12 and 24 standard 32 Og i
r- n I !
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, . -- .. . . . _ - + -- - -
SRP Upper Spacer Bar - F ' ' - .- :q:c l
. J . . A_,_ . .i_ ... 1 I ; I Static Force 4eNection Curve ; 'n :
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Il dynamic combination elements were used to model the lower and middle spacer bars. A total of 6 and 12 modified dynamic combination elements were used to model the inner and outer upper spacer bars. The classical bilinear kinematic hardening approximation to the full 2D axisymmetric force-deflection properties of the inner and outer IRP upper spacer bars is illustrated in Figure 4.1-13. The LRP spacer bar force-deflection properties were discretized at 12 locations along the inner shield plate peripheries and 15 locations along the outer shield plate peripheries. Along the inner and outer peripheries of the lower and middle shield plates, a total of 24 and 30 standard dynamic combination elements were used to model the lower and middle spacer bars. A total of 12 and 15 modified dynamic combination elements were used to model the inner and outer spacer bars. The classical bilinear kinematic hardening approximation to the full 2D axisynmetric force-deflection pro-perties of the inner and outer LRP spacer bars is illustrated in Figure 4.1-14. 4.1.1.2.6 Support Skirts Under the action of SMBDB loading pressures, the lower shield plates separate from the support skirts as the stack of shield plates is accelerated upward to impact the plug undersurface at the upper spacer bar. Once the peak plug g response is reached, the shielding plate stack falls back on the respective E support skirt bottom edge and tend to cause the skirt material to tear at the pin connections to the rotating plugs. The SRP, IRP, and LRP support skirts were considered linear elastic except for the skirt material locally adjacent to the pin connections which was assumed perfectly plastic. The full 20 axisymmetric axial stiffnesses of the SRP, IRP, and LRP support 6 skirts were taken as 25.96,120.8, and 545.6 x 10 lb/in while the respective 6 yielo forces were 0.975, 4.42, and 6.39 x 10 lbs. The local 3D force-deflection properties of the SRP, IRP, and LRP support skirts were discretized for the arc length between the lower shield plate finite element mesh in proportion to the full 20 axisymmetric elastic-perfectly plastic properties. The standard ANSYS (STIF 40) dynamic I 34 . i
t , ' e l 1 H[?; . .. u;..; ...Jia _aina
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=y= 242i1
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! l 9 . . .1 ._ . .l .- -l ,. 36 combination element with yield force and without damping was usec to modci 9 each discretized section along the support skirt peripheries. A nominal small gap was assumed between the lower shield plates and respective support skirt bottom edge. The SRP, IRP, and LRP support skirts were modeled with 8,14, and 15 dynamic combination elements. 4.1.2 Material Properties The CRBRP closure head system response to SMBDB loading requires an extensive description of material properties because of the relative detail included in the 3D ANSYS model. Average material properties at normal operating temperatures were used in the analytical predictions of dynamic system response. The use of minimum properties common to conservative static analysis is not justified in dynamic analysis as the peak response to , maximum properties may be more severe because of tuning to SMBDB loading content. Accordingly, average material properties providing a compromise between minimum and maximum values were used to represent CRBRP closure head materials at their normal operating temperatures. The materials for CRBRP closure head components used in the dynamic system response as identified on Reference design drawings include the rotating plugs (SA-508, CL.2), shield plates (SA-516, GR.60) and support skirts (SA-516, GR.70) spacer bars (SA-387, CL.2, GR.22), and shear rings (SA-540,CL.1,GR.8-24). The average material stress-strain curves of CRBRP closure head components including the trL tment of damping are pre-sented in the following sections. 6 4.1.2.1 Stress-Strain Curves The average stress-strain curves for CRBRP closure head materials include elastic and plastic properties and modifications thereof to reflect the ! reduced properties in perforated CRDM regions. Elastic constants for CRBRP closure head components at average operating temperature are summarized in Table 4.1-1. l l 1 o ' / : 'J 37 i l Table 4.1 1 Material Temperatures and Elastic Constants 0r the Plugs and Chield P*ates COMPONENT AVERAGE TEMP (*F) EIASTIC CONSTANTS LRP 400 E = 28.6 x 10' psi, v = .3 I LRP UPPER SilIELD 420 E = 2G.9 x lo' psi, V = .3 LRP MIDDLE S!!IELD 500 E = 26.4 x 10' psi, V= .3 LRP IOWER SifIELD 570 E = 25.9 x 10' psi, V = .3 m e 400 ' SOLID: E = 28.6 x 10' psi, V = .3 PERFORATED: E*/U = .55, v* = .3 A CSOLID: E = 26.6 x 10. 6 psi, V = .3 IRP UPPER Sif1 ELD 470 PERFORATED: E*/E = .445, v* = .26 E g l
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*/E = .445 h. 650. PERFORATED: E* r 11.2 x 10' psi E ., v* = .26 SRP 400 E = 28.6 x 108 psi, v = .3 1 SRP UPPER SHIELD 480 E = 26.5 x 10 8 psi," V = .3 SRP MIDDLE SHIELD 550 E = 26.05 x 10' psi, V = .3 Ie ' SRP IDWER SHIELD 550 E = 26.05 x 10' psi, V = .3 LRP UPPER SPACER BAR 410 E = 26.94 x 10' psi LRP MIDDLE SPACER DAR 460 E = 26.64 x 10' LRP IDWER SPACER DAR 535 E = 26.16 x 10' poipsi 4 ' IRP UPPER SPACER BAR 435 E = 26.79 x 10' IRP MIDDLM SPACER DAR 490 E = 26.46 x lo' psipsi IRP IDWER SPACER BAR E = 26.09 x 10' psi 545 s SRP UPPER SPACER BAR 440 E = 26.76 x 10' ' SRP MIDDLE SPACER BAR 515 E = 26.30 x 10' psipsi SRP IOWER SPACER BAR 550 E = 26.05 x 10' psi I SUPPORT SKIRT 500 E = 26.4 x 10' psi IRP SUPPORT SKIRT 525 E = 26.2 x 10' psi SRP SUPPORT SKIRT 550 E = 26.05 x 10' psi 3 38 9, t I, The rotating plugs (SA-508, CL.2) and shear rings (SA-540, CL.1, GR.c-24) 9 operate at a nominal temperature of 400 F. The average stress-strain curve for the solid portions of the SRP, IRP, and LRP including the perf of the CRDM region of the IRP are illustrated in Figure 4.1-15. The SRP, ions IRP, and LRP shear ring average stress-strain curve is illustrated in Figure 4.1-16. The SRP, IRP, and LRP upper spacer bars (SA-387, CL.2, GR.22) operate at nominal temperatures of 440, 435, and 410 F. Average stress-strain properties at 400 F taken to represent the spacer bar material are presented in Figure 4.1-17. 4.1.2.2 Damping The CRBRP closure head materials were assumed to dissipate SMBDB loading energy through damping mechanisms as well as by plastic deformations. A moderate viscous damping ratio (c = 0.04) was assumed in the derivation of 3D dynamic system response. The ANSYS program permits an effective viscous damping ratio (c) to be simulated for a frequency (f) of interest through a combination of mass (a) and stiffness (6) multip1ters according to the relation: l c = 4 g + nBf In order ta establish the values of the constants (a, s), a linear modal analysis with all nonlinear gaps closed was performed to determine min / max frequencies (f), f2 ). The modal analysis showed that the frequency range ~ from 20 to 300 Hz included the first 16 modes of the CRBRP closure head. - Accordingly, the fit to the min / max frequencies (f), f )2 of 20 and 300 Hz were provided a and 8 values of 9.4229 and 3.9788 x 10-5 respectively. For response frequencies between 20 and 300 Hz, the effective damping ratio is less than 4% of critical. Below 20 and above 300 Hz, the effective damping ratio is greater than 4% of critical damping. 1 I I L a ; V 39 E . .o . o : ._- l W [ 1
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m ; t.! i n. .;... : o j; i *T .s .. , u. . i. L. l, . .l . ,i P , .j p. . 42 4.1.3 Loading Conditions The SMBDB loading conditions in terms of a total CRBRP closure head force-time history associated with sodium slug impact as derived with the analytical REXCO-HEP code (Reference 2 ) was the basis for the analytical predictions of system dynamic response. The total CRBRP closure head force-time history of approximate 50 milli-second duration is illustrated in Figure 3.0-1. The total CRBRP closure head force-time history itself is not directly applicable to 30 system response predictions. The equivalent pressure-time history was constructed by dividing the total force amplitudes by the I projected area of the lower SRP, IRP, and LRP shield plates. Peak pressure amplitudes were considered uniformly distributed and in phase over all lower shield plates. The effect of the UIS columns acting on the IRP prior to and following sodium slug impact has been shown to be small and was therefore neglected in this analysis. The SMBDB loading pressure-time history is summarized in Table 4.1-2. 4.1.4 Peak Response Results The 3D dynamic system response of the CRBRP closure head over the time I duration from 0.0688 to 0.1164 seconds was derived with 11 successive solutions using the ANSYS restart option. The integration time step over the time peri.od from 0.0688 to 0.0704 seconds was taken as 100 usec. I Thereafter, the integration time step was increased to 200 usec over the 0.0706 to 0.1164 sec'o nd time duration. , / The CRBRP closure head response results included translational and rotational I displacements, velocities, and accelerations at all finite element node points representing the rotating plugs, shield plates, spacer bars, support skirts, and shear rings. In addition, system response information concerning stresses and strains throughout the CRBRP closure head components was generated. With regard to analytical predictions of CRBRP closure head capability under SMBDB loading as derived with the REXCO-HEP code, the most important dynamic response parameters in assessing the structural integrity I,[ i j 43 1 Ta bl e 4.1 -2 SMBDB Loading Pressure Time History Time Pressure (sec) (psi _ 0.0 8.0 0.0688 6.0 0.0692 672.5 0.0696 0.0700 394.2 4406.0 g 0.0718 W 46.4 0.0730 383.9 g, 0.0734 0.0736 2 91 .2 392.8 g 0.0738 368.5 0.0740 407.1 0.0744 360.7 0.0752 523.1 0.0760 435.4 g 0.0770 1450.6 3 0.0774 974.0 0.0778 1535.7 0.0788 g 1381.1 0.0800 314.3 g 0.0802 356.7 0.0818 695.7 0.0838 1646.5 0.0844 268.0 0.0862 1205.9 0.0878 3 407.1 0.0888 810.9 3 0.0896 778.1 0.0910 1015.2 0.0928 626.1 l 0.0942 1 241.9 0.0960 699.3 0.0978 1089.9 .s 3 5 0.0988 834.8 0.1000 824.5 0.1002 3 804.4 0.1010 724.0 3 0.1022 1015.2 0.1040 0.1056 0.1072 602.9 927.6 672.5 l 0.1088 881.2 0.1102 904.4 o Og I of the primary boundary are the peak strains in the SRP, IRP, and LRf as well as the respective shear rings and locally adjacent plug materials. Other CRBRP closure head components not directly forming the primary boundary are relatively unimportant in terms of CRBRP closure head capability under SMBDB loading. I In the following, the 3D dynamic system response information necessary to I assess CRBRP closure head capability by analytical methods is presented. It is important to note that the analytical predictions are based on REXC0-HEP analytical derivations of what constitutes SMBDB loading. The ERDA/ SRI Experimental Studies (See Section 5.3) will supply an additional set of load-time histories which may or may not agree in detail with REXC0-HEP predic tions . If the two sets of loads are not in close agreement, then a subsequent evaluation would be required to assess the response based on the SRI test data. 4.1.4.1 Plug Bending Strains The peak SRP, IRP, and LRP bending strain magnitudes and locations in regions removed from the shear rings are of interest in subsequent assess-ments of structural integrity. As the rotating plugs are of relatively thick sectional depth, membrane strains are minimal and tensile plastic instability is not a relevant failure mode. Only the local ductile rupture failure mode caused by positive principal strains at plug surfaces in bending that exceed an allowable fraction of the reduction in area in the uiiiaxial tensile test are of relevance in establishing the capability of the -SRP, IRP, and LRP under SMBDB loading. " I In order to determine the worst case locations in the SRP, IRP, and LRP at which maximum positive principal strains occur, the peak equivalent strains were reviewed from the 3D dynamic response results. The entire SRP was found to remain linear elastic through SMBDB loading. The IRP including the perforated CRDM region was found to experience minor plasticity in approximately 40% of the elements. The peak IRP equivalent strain was i 0.44% and occurred along the line of symmetry between the perforated CRDM region and the SRP shear ring. The LRP was found to experience minor l o) _ 45 l l I l plasticity in all elements. The peak LRP equivalent strain was 0.52% and occurred along the line of symmetry in the wider portion adjacent to IRP shear ring. Accordingly, the peak equivalent strain occurs in the LRP under SMBDB loading. The corresponding maximum positive principal strain was 0.51% with in-plane principal stresses of 67,501 and 29,205 psi. 4.1.4.2 Shear Ring Deforma tions The peak deformations in the SRP, IRP, and LRP shear rings and respective adjacent plug materials are used in subsequent 2D axisymmetric finite element component analysis of the shear rings for the purpose of assessing their structural integrity under SMBDB loading. The peak local shear ring deformations in the 3D dynamic response occurred in the LRP followed by' the SRP and IRP regions. g ur The maximum local peak deformations of the SRP, IRP, and LRP shear ring and adjacent plug materials were found to be 0.13, 0.11, and 0.27 in, respectively. The locations of all peak shear ring deformations occurred along the 1ine of symmetry in the wider section of the rotating plugs. The average deformations of the SRP, IRP, and LRP shear rings and adjacent plug materials were 0.09, 0.09, and 0.17 in, respectively. 4.2 Component Analysis The 3D system dynamic response of the CRBRP closure head under SMBDB loading provided maximum principal strains which are directly applicable to struc-tural evaluations of the SRP, IRP, and LRP regions removed from the shear rings. However, for plug regions adjacent to the shear rings and the shear rings themseleves, detailed component analysis is required. Based on the local 3D peak deformations found in the system dynamic response, the SRP shear ring and locally adjacent SRP and IRP plug materials are worst case locations for an assessment of structural integrity of the shear rir)g junctures in the CRBRP closure head under SMBDB loading. E I 46 I I 4.2.1 Model, Material Properties, and Loading Conditions The SRP shear ring and locally adjacent SRP and IRP plug materials was analyzed with the exact same 2D axisymmetric model and material properties used to derive the static force-deflection properties for the 3D dynamic system response. The extent of the local SRP shear ring region and respective finite element model were previously described and illustrated in Figures 4.1-3 and 4.1-4 respectively. Material properties in terms of stress-strain curves at 400 F for the SRP and IRP (SA-508) and the SRP shear ring (SA-540) are identical to those used in the system response as illustrated in Figure 4.1-15 and 4.1-16 respectively. In accordance with the decision to neglect momentum transfer effects, the loading condition consisted of statically applying the peak local 3D dynamic displacement (0.13 in) to the inner boundary of the SRP region while the outer periphery of the IRP region was fixed to ground. The static loading condition must be considered conservative as the average 3D dynamic displacement (0.09 in.), which is more representative of 2D axisymmetric loading, is approximately 70% of the peak (0.13 in.) dynamic displacement. 4.2.2 Peak Strains The peak strains of interest in the assessment of structural integrity in the local SRP shear ring region are the maximum positive principal strains in relation to the local ductile rupture failure mode. The static response I shows the SRP lip to experience minor plasticity while the SRP shear ring itself was found to remain linear elastic. The notch root of the IRP was found to be the worst case strain location. The maximum. principal positive notch root strain was 11.4% with principal stresses of -83,g00, -138,000, and -191,000 psi. 4.3 Structural Criterion and Evaluations Structural evaluations of the SRP, IRP, and LRP forming the primary CRBRP closure head boundary under SMBDB loading were made in relation to the theory of local ductile rupture. The local ductile rupture failure mode occurs when peak positive strains exceed fracture ductility causing a material to I !,O crack which, in turn, can lead to gross rupture. This type of failure p) ( 4, l : I l mode is common to pressurized thick plates with notched regions. Peak bending strains in thick plates are included in this category. As such, the maximum positive principal bending strain in the LRP removed from the local shear ring region and the maximum positive principal strains in the IRP notch root are directly applicable in structural evaluations of local ductile rupture. ' 4.3.1 Criterion Il l In a general localized multiaxial stress state, McClintock (Referance 3) showed that local ductile rupture in a multiaxial stress state could be characterized P by the true equivalent plastic strain (c eq) and related to the true fracture i strain in a uniaxial test (cf ) by a triaxiality factor (TF). p SINH E (1-n)) SINH [ (1-n)TF) Whe re , n = Strain hardening exponent in the equivalent stress-strain g rel ation , eq"K('eo)" gl TF = a) + 2+ 3 I eq a), a2' 3 = Principal Stresses eq = Equivalent Stress I Experimental results from the open literature (References 4-7) for local ] ductiIe rupture in specimens of various shapes provided only maximum principal ! fracture strains (cf, max principal). Accordingly, a direct comparisen with P theory based on equivalent plastic strain (c eq) is difficult. A plot of I the experimental data in terms of the ratio of maximum principal strain in the multiaxial stress state to the true fracture strain (cf) in the uniaxial tensile test versus triaxiality factor is presented in Figure 4.3-1. o OE W W 6 T M -- 1._- 1 0 f M*E nfrMJidTi"o .'.0 %',?" Q 45 1320 .i ;j- p i: :i :f: .c .i. :L.xt: :i ml:nmumaiini:I::tiiii2itu :inu tiitil:n:1.: n i; :11 . i . : 1. .i :1: ::: . :6:.l. 4: . 1 - ..i .[ "j-H.- Figure 4.3-1 " j; ~ t I ,b Triaxiality Functions vs. Experimental Local Ductile Rupture Failure Data I F 1 'l h.l. [
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- v i i- ,, ._l . i I l A comparison of the e;perimental data in relation to the McClintock relation j l for strain hardening e <ponents (o < n < 0.5) indicate excellent agreement l ! even though maximum pr"ncipal strain was used in the correlation instead of equivalent strain as piedicted by theory, i The SMBDB criterion in ocal ductile rupture. structural evaluation is: l *
- f, max pri 1cipal 1 SF e c f, uniaxial /SJNH 3(1~n) 5 S?NH 3E(1-n)TF Where ,
"f, max principal = Trut Maximum Positive Principal Strain SF = Safe Fraction, SF = 0.7 l 'f, uniaxial = True Uniaxial Fracture Strain E f, uniaxial = in ' 100 - RA %) RA% = Percent Reduction in area in Uniaxi si tensile test TF = Triaxiility Factor l TF = c) + 2 + "3 c,e g o), 2' 3 = Principal Stresses ) o = (o) -
- 2)
- I" 1 ~ 3) +(2~ 3) n = Strain Ha dening exponent in the equ valent stress-d strain re' ation, P n o = K (c K = Strength Ct efficient E
8 s0 Og I l I 4.3.2 Evaluations 4.3.2.1 LRP The maximum positive principal bending strain in the CRBRP closure head occurs in the LRP. 'f, max principal = 0.51% i The limit on the maximum positive principal strain as defined by the local ductile rupture criterion. f i SINH d (1-n) Limit = SF e ('f, unixial)( Y p SINH d (1-n) TF % T ) l The uniaxial fracture strain of the LRP material (SA-508, CL 2) at 400 F is the true reduction of area, or, 'f, uniaxial = 68% The strain hardening exponent (n) is related to the true uniaxial uniform elongation (cu ). For SA-508, CL.2 at 400 F, the uniform elongation is 13.7%, or, n = Ln (1 + cu ) n = Ln (1 + 0.137) . n = 0.128 The triaxiality factor (TF) for the state of stress at the worst case LRP location is given by TF = o) + a2+ 3 ,eq 'h a 51 1 I TF = 67,501 + 29,205 + 0 l 58,551 l TF = 1.65 1 Using the safe fraction (SF = 0.7), the limit for the maximum positive principal strain. 4 r S l SINH d (1 .128) i Limit =0.7(68%)( 5 ) SINH d (1 .128) 1.65 / s i i I Limit = 26.9% The CRBRP closure head rotating plugs will not experience local ductile
- rupture under SMBDB loading as the maximum positive principal strain at the worst case LRP location is significantly less than the limit
- I
'f, max principal < Limit I , 0.51% < 26.9% 3 5 4.3.2.2 IRP Notch Root The maximum positive principal strain in the CRBRP closure head occurs in the notch root of the I"o plug. 'f, max principal = 11.4% I I The limit on the maximum positive principal strain as defined by the local ductile rupture criterion is given by: SINH d (1 - u) Limit = SF e ('f, uniaxial) 3 SINH d (1 - u) TF I 52 9 I I i The triaxiality factor (TF) for the state of stress in the worst case IRP notch root is given by TF = o3+ 2+# 3 I eq TF = 83,000 - 138,000 - 191,000 94,000 TF = -4.5 The negative triaxiality factor indicates that the IRP notch root is in a highly localized swte of compression which inherently reduces the tendency I for local ductile rupture to develop. Conservatively, a triaxiality factor of unity is selected in the structural evaluation. I TF = 1 1 Thus, the limit is dependent only on the safe fraction and uniaxial fracture strain. For a safe fraction (SF = 0.7) and the uniaxial fracture strain l (cf, uniaxial) for SA-508, CL.2 at 400 F. Limit = SF o ('f, uniaxial) Limit = 0.7 (68%) Limit = 47.6% The CRBRP closure head shear ring notch roots will not experience local ductile rupture under SMBDB loading as the maximum positive principal strains at the worst case IRP notch root is less than the lihiit: 'f, max principal < Limit 11.4% < 47.6% b r (h) V 53 l i ! A similar calculation for the Vessel i ~. nge notch root gave a peak strain l Of 2.63%. A nonconformance in the manufacture of the Vessel Flange required i an upward revision of this value to 5.21%. Details are given in Appendix D. l ) The values for this location are as follows. 1 i 'f, max principal
- limit = 47.6%
l 'f, max principal < Limit l1 I 5.21% < 47.6% i This location is clearly less critical than the IRP notch root. 1 I E I I1 II I I I I-54 Og I, I 5.0 EXPERIMENTAL PREDICTIONS OF CLOSURE HEAD RESPONSE Experimental studies of the CRBRP Closure Head response to SMBDB loading include three separate efforts. The first experimental work, performed i in April 1976, comorised three static tests of clane snecimans ranlicating the LRP margin rina reaion in 1/ln-scale. The results nf those tests are presented in Section 5.1. The significance of the static tests was their pronounced disagreement with previous analytical predictions, The measured l stiffness was lower than the calculated value by a factor of four but the measured proportional limit was higher by a factor of three. Investiga-tion revealed the reasons for the divergence. It was found that the original analysis had neglected the flexibility of the plug lip. In fact, g the deflection of the plug lip accounted for half the total deflection in 5 the test. Also, the original analysis had taken the onset of initial yielding to mark the proportional limit, where in fact no detectable de-viation from linearity occurred until yielding became general. Hence, the effective stiffness was further lowered by local plasticity, and the measured proportional limit, which corresponded to the formation of plastic hinges, was significantly elevated. I The fact that analysis had seriously mispredicted these two important I parameters strongly indicated the desirability of more experimental work. Three areas of uncertainty existed. I (1) It had not been established that the parameter $ obtained from a static test accurately represented the behavior of the margin I ring region under a dynamic load, as would be imposed by an HCDA. (2) It had not been established that the margin ring stiffness and proportional limits measured on plane specimens were representa-I tive of the CRBRP margin rings, since the plug lips and margin ring grooves were continuous and circular and the margin rings were circular segments. I _ l _ _ _ (3) The principal mode of failure of the Closure Head, the one corresponding to the lowest aoplied load, had not been established ]j and probably could not be established by analysis, 55 Three test programs were completed. I The first, described in Section 5.1, consisted of static tests or plane specimens, mentioned above, and a series of drop-weight impart loadings of specimens of identical geometry. A direct comparison of static and dynamic response was thereby obtained, addressing the first area of uncertainty. The second test program, descr N d in Section 5.2, consisted of static and dynamic loading on fully circumferential models of the LRP and LRP margin rings. Additional data were obtained for static-dynamic comparison, and comparison between plane and axisymmetric specimen behavior was possible. These comparisons addressed qualititatively the second area of uncertainty. These first two experimental programs were limited efforts which were completed in a relatively short time period. The results obtained were primarily qualitative, serving mainly to demonstrate that no major design inadequacy existed. The exoerimental program completed by Stanford Research Institute (now SRI International), and described in Section 5.3, was of considerably broader scope and purpose. The intent of the program was to verify analytical methods and establish design parameters for possible future LMFBR projects. The results of the scale model test series are directly applicable to CRBRP, addressing the third area of uncertainty in establishing l Closure Head capability identified above. Howe /er, the utility of that work is not limited to a single design. In each of the experimental programs, material substitutions were made. The substitute materials were chosen so as to closely simulate the behavior of the prototypic head materials. These substitutions were necessitated by the difficulty in obtaining samples of prototypic materials in quantities suitable for test pruposes. A comparison of CRBR Closure Head Materials and test materials may be found in Appendix B. 5.1 Margin Ring Strip Model Tests A series of tests were performed to obtain static force-deflection relatiori-ships for the LRP-to-vessel flange interface, including the LRP lip, LRP margin ring and margin ring notch in the vessel flange. Additional tests were performed to provide qualitative indications of the difference between static and dynamic response for the LRP-to-vessel flange interface. 56 Since the LRP margin ring is 260 inches in diameter with an approximately square cross-section 3.5 inches on a side, full-scale testing was obviously impractical. A specimen geometry was chosen which, provided a 1/10-scale replica of a radial section of approximate unit thickness through the LRP lip, LRP margin ring and vessel flange. Each test specimen consisted of five separate components: one center plug, two end plugs and two shear keys. The individual components are shown in Figure 5.1-1, and the test specimen as assembled is shown in Figure 5.1-2. The center plug has a lip on each end which is a 1/10-scale replica of a plane section through the LRP lip. The end plugs and shear keys similarly replicate the vessel flange and LRP margin ring, respectively. The double-ended arrangement was originally designed to eliminate lateral loads in the static tests, and was retained for the drop tests, being readily adaptable to drop-weight loading and allowir.g direct comparison between the two test sequences. The 1/10-scale size was chosen to allow static testing on a universal test-ing machine of 60,000 lb. capacity. The size also allowed the use of easily handled drop weights (20-40 lb.) in the impact test sequence. The results of the four drop tests are presented in Section 5.1.1. Those of the three static tests are in Section 5.1.2. 5.1.1 Margin Ring Drop Tests 5.1.1.1 Test Setup and Procedure Each test specimen consisted of the five components shown in Figure 5.1-1, and assembled in Figure 5.1-2. I Shear key material was quenched and tempered AISI 4142H alloy steel in all four tests. Two different plug materials were tested, to provided direct experimental comparison of their behavior under dynamic loading. I In the first three tests, plug material was ASTM A-533, Type B, Class 1, low alloy steel. In the fourth test, the plug material was ASME SA-508, Class 2, low alloy steel forging. The SA-508 used was obtained from a trimming of the CRBRP IRP forging. The A-533 material is equivalent to the head material used in the ERDA-SRI Experimental Studies described in t Section 5.3 of this report. 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- l l- 'r--.h.' I I j l l L.. P b ! .J { j ' l , f A _ .. ,_ 9- , __ . _ . _ _ . _ _ _. f , l_J l r {_ _ _ _ . _. _ , _ 3 1 - bj N lc--I kl1 b l_..__.Ym ;3. , .. 1 , 1 1 7 , i 8 s., <. - 4 - 4 n - y \ E L-Center Plug End Plug i End Plug fiargin Ring Key 8)C Figure 5.1-2 Drop Test Schematic $9 _ _ _ _ _ - _ _ i. The test fixture with a specimen in place is shown in Figure 5.1-2. This I fixture was clamped to the table of a universal testing machine which served as a support for the components of the test apparatus. Two different drop-weight arrangements were used. For Tests 1 and 2, the I drop weight was held and released by an electromagnet attached to the cross-member of the test machine. A guide tube was added for Test 2 to maintain uniform loading at impact, since lateral drift of the weight and nonunifonn loading was observed in Test 1. Tests 3 and 4 required a different arrangement due to increased height of drop; a longer guide tube Wds added. The electromagnet could not be used with this arrangement, since the guide tube extended above the top of the test machine frame. The weight was lif ted into position by means of a cord-and-pulley arrangement and released by burning through the cord. Data acquisition was by means of before-and-after measurement in all four tests. Additional data were obtained by high-speed photography during impact in Test no. 4. Measurements of all specimens were taken by optical comparator before and after testing. The differences between before and after measurrments were tabulated as indications of plastic deformations of the specimens. In addition, dye penetrant checks for cracks were per-formed before and after each test. I 5.1.1.2 Results of Tests 5.1.1.2.1 Test No. 1 Test no.1 was performed with a drop weight of 38 pounds and a height of drop of 4.3 inches. Impact velocity was 55 inches per second. These conditions correspond to the impulse of the SMBDB initial spike on a Closure Head weighing 1,600,000 pounds, which is greater than that of the head in operating configuration. g 5 A photograph of the specimen af ter test is included as Figure 5.1-3. Figures 5.1-4 and 5.1-5 show close-ups of the shear keys, plug lips and l shear key grooves after testing. 60 I: g g -g W M W W~U T T (m 7 -._.,,,,..s N MI /I ~ 4 ~ e '5 .[b . ' . -[ ,, '8 -- .a. 7 ~ Figure 5.1-3 Test No.1 Specimen After Test I wm Lb,p1,,\ n Mgi;p MW 1r 3A yf b..;. i,f -sw m w % ' n. w,v2,~P n 47 , ;q,y,yg ;g< e 9 .n. Ci- y p e;$gbg;j www m y% p. R- l YE y Y [ ,y%yI .j ej . E7 v-V$t ( s _4}. gf a2vy : l f. lk.* F ] y . fE e L A', l yQQ .'Q l '; , un . , a up.( 't dM;t an ,f .g @why n .O e ,M fiQ d y. [7;,, : ' y 4 --W .%g/h ... MD, <nr 9[Nn. %w
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T s N. ?w ?? N _ h k . ~ . *1 m r W 1 u j ,t (! lj Visible effects were limited to a slight burnishing of the machined sur-faces in contact. Dye penetrant checks of the shear key grooves revealed no sign of any cracks before or af ter the test. Dimensional changes as measured by optical comparator are illustrated in I Figures 5.1-6 and 5.1-7. Maximum vertical plastic deformation was .0020 inches. It was observed during the test that the weight was not released uniformly by the electromagnet. Since no guide tube was used in this test, the weight struck the specimen off-center. A guide tube was added for the second test to correct this problem. 5.1.1.2.2 Test No. 2 Test no. 2 was performed with i drop weight of 24 pounds and a height of drop of 11.4 inches. Velocity at impact was 87 inches per second. These l conditions correspond to the impulse of the SMBDB initial spike acting on the head in operating configuration (total head weight of 1,004,000 l b. ) . l The specimen after test did not visibly differ from the specimen in Test 1. Again, visible effects of the test were limited to burnishing of the i machined surfaces in contact. Oye penetrant checks of the shear key grooves revealed no sign of any cracks before or after the test. Dimensional changes as measured by optical comparator are illustrated in I Figures 5.1-8 and 5.1-9. Maximum vertical plastic deformation was .0049 inches. 5.1.1.2.3 Test No. 3 Test flo. 3 was performed with a drop weight of 24 pounds and a height of drop of 90 inches. Impact velocity was 245 inches per second. These con-ditions do not correspond to any postulated accident sequence. They represent the most severe loading possible with the existing test set-up and are more severe than any accident loading postulated for the CRBRP Closure Head. 64 O I l l1 1 + + A lll 1 I n i 5 _ 0 _ 0 _ 0 _ &L f_ d d e n ) - a d m H e - o - t f t a e h r D g e i e eg 1 R p pg f a aa h h x o o S S e N t e c l d y s a el a d n mt e h n i ra T S E g oe i f r r eg 6 O D( - ,t11' 1 - 5 - e - r - u g i F n c ll l , ll' llI' ) d e e p t a a h r S e eg e p pg d a aa e h h x f m r o e d E n S S e l a el n mt dy - D d i ra n g oe 1 N o H t f a i f r O r eg D( l r e - t L _ s - e f - _ T O - 7 1 5 _ e i F r g u / [ / /, - / I-n I I l 1 I l I I I i 0 /_ 2 0 0 - I Ai \ ,I 1I I J ~_ - 1 I ' L-i.'1 - + - = g O-g l1 l 1 I l l _ 1 \l > ll + L - ll ;l l l I I - ~ - ~ _ . r I \ l J - - r1 I I L a* n i 7 4 0 _ 0 _ II I Il f s _ \ i l ! s ,1 I lI l l d e d n ) d m a e r H t _ f o - a t r D e h c e eg e _ i p ng 2 R a aa h h x _ . f S Se o o _ N l d y e p a el t s a d n mt i ra _ e h n g oe T S E i f r r eg O D( 8 _ 1 5 - e r - u - g i F e 1 1ll l i t l 1 1 Figure 5.1-9 Test No. 2 Deforrned Shape ' t of Le -Hand End . r - i 1 ! t I l I Original Shape l l --------- Defomed Shape l l / (greatly exaggerated) S I i / y ,1 _ _ _ _ , _./ s , ___,_ / ' ( .0029 in. i I l l 1 I h* I l I 1 =+ u_____________ _ _ _._ _ _ _ ___. a l l A photograph of the specimen after test is included as Figure 5.1-10. Figures 5.1-11 and 5.1-12 show close-ups of the shear keys, plug lips and l shear key grooves after testing. I l Both the plug lips and shear key grooves are visibly deformed, indicating the severity of the loading. However, the shear keys show only burnishing of the contact surfaces, and dye penetrant checks of the shear key grooves still revealed no sign of a crack. Dimensional changes as measured by optical comparator are illustrated in Figures 5.1-13 and 5.1-14. Maximum vertical plastic deformation was .0355 inches. 5.1.1.2.4 Test No. 4 Test no. 4 was performed with a drop weight of 24 pounds and a height of drop of 90 inches. Impact velocity was 245 inches per second This test differs from Test no.3 in that the plug material was SA-508 instead of A-533 and in that high-speed photographs were taken of the specimen during the interval of impact. Seven sequential photographs of the specimen, taken over an interval of 1.62 milliseconds, are shown in Figure 5.1-15. The weight is in contact l with the specimen from t=0.27 msec. to t=1.35 msec. As can be seen, the initial impact was eccentric, causing significantly greater deflection of the left-hand end of the specimen. The peak deflection photographed was k at t=1.08 msec. Deflections as scaled from the photographs were as follows: Left-hand side -- 0.12 inches Right-hand side -- 0.04 inches Center -- 0.08 inches The plug lips and shear key grooves were both visibly deformed, the pattern I of deformation being similar to that observed in Test no. 3. The shear I o\ I j' 69 I - rm - 5 .h in: . Figure 5.1-10 Test No. 3 Specimen After Test G----------------- 9 l1 , 1 llll l l il ll1,1 1 ( I l , q-~.~ 1 .g- . 4a I t.*, a, ; ( O - r O : e h.e ) s. ' .s - v D C [ ,. l-- . ,% = 5 a.- z T t s e f.,r. , r f .a . e D (('g . . f t A M f [ P.t s. i n _ g _ r a W f- j ~ : 1 1 1. 5 M 3.dn .a oH +. eN Qr I; ; . ;j ~ r t uth . ;: gsg _ iei _
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! 17igure 5.1 15 Test No. 4 4 Drop Weight Initial Impact Sequence I' i j / . - . _ , , , ... , - , , . - , - - . , , . , - , . . . , , , - - - , - , . , ~ , , . . - . . - - - - - - - - - - - - . . - - - - - . keys show only burnishing, as before, and dye penetrant checks showec no sign of a crack. The right-hand end plug exhibited a burnish mark above the shear key groove where the center plug evidently rubbed against it ; during the impact. This probably occurred when the center plug tipped under the eccentric loading. The significance of this lateral impact is that an indeterminate amount of kinetic energy was dissipated along a load path different from that through the shear keys. I Dimensional changes as measured by optical comparator are illustrated in Figures 5.1-16 and 5.1-17. Maximum plastic deformation was measured at 0.0333 inches. 5.1.1.3 Calculation of Drop Weights and Heights The proper values for drop height and drop weight were determined from simple considerations of conservation of momentum, with a correction for dissipation of energy to the fixture and support. I The impulse of the initial spike of the SMBDB loading was calculated to be 227,000 lb-sec. Equating impulse and momentum gives I = Fdt = m v g I Or the impact velocity is given by y o I _ h m W I where W is the weight of the closure head. For W = 1,004,000 lb., v o ~ (227,000) (386) = 87.3 in/sec, l,004,000 For W = 1,600,000 lb., v = (227,000) (386) = 54.8 in/sec. o 1,600.000 I 75 t . 11 1 , 1; 1 I II l l i1 r l III ~ I ~ LI Il l I-i 5 n 4 2 0 ; 1L 1 vs _ d n d E n a H h i t g e e R p p a a f o e h S l h S d 6 p a e 1- a n m h r i 1 4 S g . o 5 i f . d r e o e O D e N m r r - u t o - g s f - i e e - 0_ F T D - y l l li!llilil !4 l 1l d ) n e E l a d c n s a H o - t t ( f e e p e p L a a f h h o S S ~ e l d - 7 p a e 1 a n n n M 1 5 4 h S d i i g r f o e e o N e m O D - r r - u t o - M a s e f e i F T D M - M ... M - 7 _ / M / / _ (Il l l l l y M ,l y 4j n M i 3 3 3 ( 0 - M I L 5[ - *lI i M - _ M . 1i1 i ; 1 M _ M M 5 - M 1' lIll - The scale factors, as defined in Appendix A, are as follows. 6= geometric scale factor = 10 A1 = ratio of elastic moduli = 1 A2= ratio of densities = 1 From the relationships deduced in Appendix A, therefore, the velocity scale I factor is unity and the impact velocities in the tests must equal those of l the conditions being simulated. The scale factor for the impacting mass includes a correction factor to allow for the fact that the line of contact is not a full circle. For precisely similar structures, the mass would scale as A263= (1) (10)3 = 1000 Since the contact area has been further reduced, an additional factor must be incorporated. This factor is nD 2tB where D is the equivalent diameter and t is the thickness of the strip specimen. Using the mean diameter of the LRP shear ring as D, the mass scale factor becomes A * " m 0) (1000) = 45,100 For a head weight of 1,004,000 lb, the weight becomes y , 1,004,000 45,100 = 22.26 1b. I For a head weight of 1,600,000 lb., the value is , 1,600,000 W 45,100 = 35.48 lb. 78 I An assessment of the aaount of energy dissipated in the fixture and supports indicated that approximately 10-14% of the total impact energy would be dissipated in this fashion due to flexibility of the fixture, the percentage r increasing with the mass. To correct for this effect while maintaining the desired impact velocities, the drop weights were increased by 12% and 16% to 25.8 lb. and 39.8 lb. , respectively. These weights include the weight of the center plug, which was 1.8 lb. Hence the drop weights were chosen to be 24 and 38 lb. The velocity Vg is the impact velocity of the plug against the shear ring. However, only the drop weight velocity could be directly controlled. It was necesscry to define the after-impact motions of drop weight and plug. The assessment of energy dissipation to the fixture also provided the infonnation that drop weight and plug moved together after .act. Since momentum was conserved, MV = (M+m) V g where M is the mass of the drop weight; V is it impact velocity; m is the mass of the center plug; and gV is the shear ring impact velocity. Also, since the drv. eight falls under the action of gravity, 2 = V 2gh Therefore, 2 h = 0 M+m 2g M For Test, h: ,1, b = (55/ 8 = 4.3 in. 2(386) 38 ] 79 1 2 L l I For Test No. 2, (87.3 [25.81 2(386f)2 h = " II'4 I"' \ 24 / For Tests 3 and 4, b = 90 in and l V = " V 2gh = g M m 2 8 V2(386)(90)' V g = 245 in./sec. 5.1.2 Margin Ring Static Tects 5.1.2.1 Test Set-Up The LRP scale model shear ring test includes a fixture bolted to the top plate of the testing machine. Bolted to the fixture are two (2) symmetric-ally arranged specimens of the vessel flange, designated as Plug 1. The LRP shear ring, represented by the shear bar, is loosely inserted in the grooves formed in the respective Plug 1. The LRP, represented by Plug 2, is bolted to two (2) plates, which in turn, rest on the bottom plate of t' testing machine. Acertures are provided in tH rixture to permit 'servation and/or photographic recording of the local deformation i Testing Machine, SN-68-255 with a maximum capacity of Available load ranges are (1) Low < 1200 lb,15 lb s .. < 20,000 lbs, i 10 lb accuracy, and (3) High b accuracy. The high range was used in the test. Starret Dial Indicators with a range of 1/2 inch and 10.00iv .at.n accuracy were used. I 80 OI Three (3) tests designated as Test No.1, 2, and 3 were performed. Test No. 1 was performed on low carbon steel specimens to de-bug set-up, equip-ment and procedures. Tests No. 2 and 3 were performed on SA-508 specimens with material traceable to the prototype CRBRP Closure Head IRP. Test 1 data, in terms of specimen loading, was taken on an intermittent-hold basis. Test 2 and Test 3 data acquisition was similar to Test 1 except as modified to improve test results. In the intermittent-hold loading of the specimens, the load was slowly increased in approximate 500 lb. increments by manually adjusting the Baldwin Test Machine and observing the load cell readout. At each of the loading increments, the load was held briefly to permit dial Indicators to be manually read and recorded. Photographs were taken at selected load increments depend-ing on the relevance of the deformation. As loading approached the plastic c instability point a rapid increase in deformation occurred. The cperator manually reduced the applied force as best possible to obtain maximum force-deflection data at the point of instability. 5.1.2.2 Results of Tests 5.1.2.2.1 Test No. 1 Test No. I was performed on low carbon steel specimens. The purpose of the test was to de-bug set-up, equipment, and procedures. Accordingly, trace-ability of the low carbon steel material and verification of equipment calibration and data recording procedures was not required or established. Test No.1 loading was increased until the plastic instability load was reached. The equipMn* originally planned for the scale specimen tests included a Satec Deflectometer which would automatically record the deflection during the test. In the course of Test No.1, it was found that the maximum test deflections were approximately 3/8 inch. As the Statec deflectometer range (0.10 inch) was not adequate to cover the test data, automatic recording of deflection was not obtained. The larger G (Va 81 q [ 1 than expected deflections also caused the range on Dial Indicator No. 2 I to be exceeded. Only Dial Indicator No.1 readings were valid and accurate over the full range of test data. Dial Indicators No. 3 and 4, which measured lateral motion of the bottom test fixture, showed motion only at I plastic instability and respective readings were not recorded. Force readings recorded from visual observation of the Baldwin Universal Testing Machine dial gauge were easily obtained. Manual adjustments of the testing machine required to achieve desired force levels, even during rapid changes at the plastic instability load, were not found to be difficult. Il B l The 1/10 scale LRP shear ring force-deflection curve for low-carbon steel, based on Dial Indicator No. 1 readings, is illustrated in Figure 5.1-18. I Features of the low carbon steel specimens force-deflection curve relevant to subsequent structural behavior are the linear elastic stiffness (K), ' proportional elastic limit (PEL), and plastic instability load (P )g and defonnation (PD). Numerical values for low carbon steel are as follows 6 K = 0.5 x 10 lb/in. PEL = ll ,000 lbs. PD = 0.400 inch. P y = 22,500 lbs. 5.1.2.2.2 Test No. 2 I' Test No. 2 was performed on SA-508 material traceable to the prototype CRBRP Closure Head IRP. Verification of equipment calibration and data recording procedures was required and established. Test No. 2 loading was increased until the plastic instability load was reached. Dial Indicators alone were used to measured deflections. Data acquhition proceeded without diff. .lty throughout the test. Dial Indi-cator No. 1 measured total deflecum between the top and bottom plates of the Baldwin Universal Testing Machine. Dial Indicator No. 2 measured the deflaction between the top of the vessel flar.ne and bottom plate of the testing machine. Diai Indicators No. 3 and 4 measured lateral motion of the top fixture and bottom fixture relative to the bottom plate of the test-ing machir,e. 82 i I 4 I gprr 7 ... M , r .yi. m q. - .rri l v p r- . o y. a; a n _ q:. i
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.f . ! 7 .1,4 i f jHb ." .... n m r v . .t "-- *r-- N W. :- - -n - "- "- - ~!- W J PP N l ,,j i M H: m[. f d : : ldF A "E I @:: d ~! F m!"- O , P1 Gh * , 3 , i The 1/10 scale LRP shear ring force-deflection curves for SA-508, based l on Dial Indicator No.1 and No. 2 readings, is illustrated in Figure 5.1-19. , Features of the SA-508 specimens force deflection cu is were based on Dial i Indicator No. 1 readings. The respective linear elastic stiffness (K), pro-portional elastic limit (PEL), and plastic instability load (P )g and defor- ! mation (PD) are as follows. 1 6 K = 0.911 x 10 lb/in. PEL = 20,000 lbs. l l PD = 0.250 in. l P g = 32,500 N . l l Corresponding values for the full-scale full-circle LRP margin are a stiff- l 6 ness of 411 x 10 lb./in. and a proportional limit load of 90 x 10 6 lb. A total of 27 photographs of the test specimen were taken at load levels ranging from 80 to 32,700 lb. Photographs of the Test No. 2 specimens at no' load (80 lbs.), PEL (20,000 lbs.), and Pg load (32,500 lbs.) are presented in Figures 5.1-20 through 22 respectively. The photographs of the front E' 5 and back of the Test No. 2 set-up after full loading are presented in Figures 5.1-23 and 24 respectively I 5.1.2.2.3 Test No. 3 Test No. 3 was performed on SA-508 material traceable to the Prototype CRBRP Closure Head IRP. Development Q. A. verification of equipment cali-bration and data recording procedures was required and established. ll Test No. 3 loading covered the range only to the PL' (20,000 lbs. ) found in Test No. 2. Two (2) cycles of loading and unloading were made to estab-lish linearity and hysteresis effects. The Test No. 3 ec,uipment utilized a total of six (6) Dial Indicators. No.1, 2, 3, and 4 were the same in terms of identification and position as used in Test No. 2. Dial Indicators No. 5 and 6 were arranged to measure the rotation of the shear ing. In order to permit shear ring rotational meas-urements, one (1) sh or t,ar was drilled and tapped to accommodate a right angled lever arm. In this arrangement. Dial Indicator No. 5 mea =ured the combined translation of the shear bar and lever arm rotation while Dial l 84 I l g . tm , ---p-7 , p -.- j -. : m -- .- . j. q- ..,. ,. .q . . ,....p ;. . . . ... .... 7 g. .--.j. . i s .. . i . . . y . . . qt ;- q .. q .. ..; ; : J .. zl . i 12_.. @ _ . jh 7 m
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