ML030910624
| ML030910624 | |
| Person / Time | |
|---|---|
| Site: | Davis Besse  |
| Issue date: | 07/31/2002 |
| From: | FirstEnergy Nuclear Operating Co |
| To: | Office of Nuclear Reactor Regulation |
| References | |
| FOIA/PA-2003-0018 | |
| Download: ML030910624 (36) | |
Text
QaI^e Ge U1,102 12.24pm named "Preliminary SSA RAI Response. PDF" PRELIMINARY RESPONSES TO SIGNIFICANCE NRC STAFF COMMENTS AND QUESTIONS ON DAVIS-BESSE SAFETY ASSESSMENT (SIA-W-DB-OIQ-301) SUBMITTED APRIL 8,2002 Question I used to determine the What is the technical basis of the failure critenon (e.g., strain exceeding 11.15%)
technical references in the literature that support failure conditions of the cladding layer? Provide specific the failure cnterion used in this evaluation.
Response to Question I curve used in the The strain value of 11.15% corresponds to the uniform elongation of the stress-strain is based largely on engineering evaluation. The use of this value as the basis for the failure criteria has through-wall strains greater than the judgment. The premise is that when any section in the cladding in load. This resisting any additional increase uniform elongation, then that section has no more capacity of strains to is redistribution of stresses and criterion is judged to be conservative because in reality, there elements strains in a particular column of adjacent elements that would prevent incipient failure when the exceed this cntenon.
(11. 15%) is Furthermore, it should be noted that the value of uniform elongation used in the evaluation in the literature, and summarized very conservative for stainless steel weld metal. Data obtained from that arc welds (SAW) is 25.7% and Table I indicates that the average uniform elongation for submerged of the populations is 27.3%. Most for shielded metal arc welds (SMAW) is 30.7%. The average for both below this with only two data points data shown in Table l indicate uniform elongation greater than 20%
value.
has been used in a Subsequent to the publication of SI Calculation W-DB-OIQ-301, the above criterion in Reference 1. The finite element analysis of disk specimens that were burst tested and documented results of wastage cavity. The geometry of the disk test specimens is very similar to the Davis-Besse RPV Reference 2 conclusion from the the evaluation are detailed in the Reference 2 calculation (attached). The test to the disk burst analysis is that the present critenon used for the evaluation is conservative compared instability results. The burst test pressures were also compared to the pressures at which numerical although slightly pressures, occurred during the elastic-plastic analysis and it was found that the instability pressure than any of the under-predicting the test failure pressures, are a much better predictor of failure proposed strain-based failure criteria.
elongation as a basis for Based on the above discussions, it is believed that the use of the 11.15% uniform the failure criterion is very conservative.
Question 2 test) account for the How does the failure crntenon (e.g., based on ultimate strain in a uniaxial tensile at the edges of the degradation effects of biaxial loading in the cladding, or triaxial loading in the cladding cavity?
Response to Question 2 biaxial or triaxial The failure criterion was based solely on the uniform elongation and did not consider compared to is conservative effects. Nevertheless, as discussed in the response to Question I, the criterion cavity geometry.
burst tests results on test specimens that are a reasonable simile of the wastage NGC-02-038
PRELIMINARY Question 3 The failure cnterion applied in SIA report W-DB-0 IQ-301 (e.g., the minimum cross-sectional strain exceeding the failure strain of 11. 15%) allows the strain levels in the cladding to exceed the critical strain value entirely through the thickness, leading to very large strains at the surface of the cladding, up to 49%
in Table 5 of the SIA report. What is the technical basis for this approach, as opposed to the average cross-sectional strain, or the maximum cross-sectional strain?
Response to Question 3 Even though the failure criterion used resulted in some elements in the cross-section exceeding the failure strain, the criterion, as compared to actual test data, was found to be conservative (see response to Question No.1)
Question 4 Did you explore a continuum damage mechanics analysis to give guidance of the failure criterion once the strains exceed the critical strain where necking/void growth starts? If not, provide the technical basis for not using a continuum damage mechanics analysis. [Poisson's ratio of 0.5 no longer applies once this critical strain level is exceeded, so the analysis is strictly not valid. (Poisson's ratio is continuously changing as the voids grow at the strains beyond the start of necking.) This results in a stress redistribution that is not accounted for in a standard elastic-plastic analysis.
Response to Question 4 The analysis performed was judged to be conservative as validated by the disk burst test results discussed in the response to Question 1, and as such, it was judged unnecessary to consider the application of continuum damage mechanics analysis to this evaluation.
Question 5 How would the strain values change if the stress free temperature was assumed to be the stress relief temperature instead of 70F, and the analysis accounted for the differential thermal expansion of the cladding and head steel at the operating temperature of 605F?
Response to Question 5 As can be seen in the SI report W-DB-OIQ-301, and as further clarified by the above responses to Questions 1 through 4, the strains at the failure pressures from both the analyses and experiments are very large (on the order of 11% or greater). The strains corresponding to thermal expansion effects, at either temperature. are expected to be much smaller (on the order of 0.1%). Therefore, the effects of changing from a stress free temperature of 70F to 605F will not have any significant impact on the results of the analysis Question A Does the size of the degradation cavity and the transition from the cladding thickness to the head thickness that was used in the SIA report reflect current knowledge regarding the cavity geometry, in particular the undercut area described in Figure 13 on page 103 of the Davis-Besse Root Cause Analysis Report (CR2002-0891), dated April 15, 2002? What is the transition geometry assumed in the analyses?
NGC-02-038
PRELIMINARY Response to Question A The size of the degradation cavity and the transition from the cladding thickness to the head thickness used in the calculation reflected what was the best available at the time of the calculation. More work is currently in progress on the removed damaged cavity to determine the exact size and geometry of the cavity and the transition regions.
Question B Is there sufficient mesh refinement through the cladding thickness to adequately capture the bending and shear strains at the edge of the cavity? Describe any sensitivity studies used to demonstrate the adequacy of the mesh refinement.
Response to Question B In the analysis of the wastage cavity, six elements were used through the thickness of the cladding. A convergence study, using both an axisyrnmetric model and a three dimensional model was performed in Reference 2 to evaluate the impact of the number of through-wall elements in the thickness of the test specimens. The results indicate that there is no significant difference in the burst pressure predictions when the number of through-wall elements is increased from six to 12. Therefore, it is concluded that the analyses of the wastage categories with six elements through the thickness represents a converged solution.
Furthermore, when fewer elements than six were used in the convergence study, it resulted in conservative estimates of the burst pressures.
Question C Was the cladding deposited by weld wire? Do the thinner cladding thickness measurements from UT coincide with the locations of weld bead toes? In what direction do the cladding weld beads run relative to the long axis of the degradation cavity?
Response to Question C The cladding was deposited by weld wire. It is difficult to determine if the thinner cladding thickness measurements from the UT coincided with the location of the weld bead toes since the UT measurements were taken on one-inch grids and as such, there was not adequate resolution to make such a determination.
It is also difficult to determine the direction of the cladding weld beads from the available information.
Additional investigation of the removed damaged cavity is currently in progress that might provide more information NGC-02-03R
PRELIMINARY References
- 1) P. C. Riccardella, "Elastic-Plastic Analysis of Constrained Disk Burst Tests," ASME Paper No.
72-PVP-12, Proceedings of Pressure Vessel and Piping Conference, New Orleans, LA, September 17-21, 1972.
- 2) Structural Integrity Calculation W-DB-OIQ-304, Rev. 0, "Evaluation of Failure Criterion Used in Elastic-Plastic Analysis of Davis-Besse RPV Head Wastage."
NGC-02-038
PRELIMINARY Table 1: Tensile Test Data for 304 Stainless Steel at 550°F Reference YS ksi UTS ksi Elong % RA % Matl Type NUREGICR-6235 20.8 62 38.4 70.8 Base NUREG:CR-4538 22.2 67.3 39 70.8 Base NUREGfCR4538 22.8 68 8 40.5 70.8 Base NUREG/CR-4687 20 1 65.2 53.8 71.3 Base EPRI NP-4768 23.1 61.3 47 74 Base EPRI NP-4768 24.8 62.6 45 70 Base EPRI NP4768 33.2 72.7 42 67 Base ASME 72PVP12 34 84 54 75 Base Ave.Base 45.0 71.2 EPRI NP-4668 44.8 62.9 22 46 SAW EPRI NP4768 36 61.8 25 67 SAW EPRI NP4768 40.8 70.3 25 69 SAW NUREG/CR-6098 37.4 68 26.4 SAW NUREGiCR-6389 49.1 68.1 30 46 SAW NUREG'CR-6389 45 67.1 33 42.4 SAW NUREGCR-6389 54.3 74 15.5 63 SAW NUREG!CR-6389 51.8 71.8 13.7 54 SAW NlUlREG/CR4878 471 67.6 31.5 44.2 SAW NUREGICR4878 28.3 67.5 34.5 47 SAW-Ann Ave.SAW 25.7 53.2 EPRI NP4668 45.7 65.1 26 58 SMAW EPRI NP4768 46 8 61.4 37 48 SMAW EPRI NP-4768 49 4 64.7 35 46 SMAW NUREGICR-4878 40.8 70.3 24.8 68 6 SMAW Ave.SMAW 30.7 55.2 NUREG/CR-4538 44.3 65 4 33 74.3 Weld NUREGICR-4538 42.2 64.3 30 72.9 Weld Ave.SAW&SMAW 27.3 53.8 NGC-02-038
PRELIMINARY C STRUCTURAL CALCULATION FILE No: W-DB-OIQ-304 INTEGRITY Associates, Inc. PACKAGE PROJECT No: W-DB-O1IQ PROJECT NAME: Operability and Root Cause Evaluation of the Damage of the Reactor Pressure Vessel Head at Davis-Besse CLIENT: First Energy Corporation CALCULATION TITLE: Evaluation of Failure Criterion Used in Elastic-Plastic Analysis of Davis-Besse RPV Head Wastage PROBLEM STATEMENT OR OBJECTIVE OF THE CALCULATION:
Develop a finite element model to simulate actual test data to evaluate the effectiveness of the failure criteria used in the elastic-plastic stress analysis of Davis-Besse RPV head wastage cavity.
Project Mgr. Preparer(s) &
Document Affected Revision Description Approval Checker(s)
Revision Pages Signature & Signatures &
Date Date 0 1 - 28 Original Issue Al - A2 BI -B9 Project CD-Rom PAGE I of 28
PRELIMINARY 1.0 Introduction During recent in-service inspections of the reactor pressure vessel (RPV) head and penetrations at Davis-Besse, significant wastage was observed in the vicinity of control rod drive mechanism (CRDM) No. 3. A calculation package was prepared for First Energy [I] to determine the limiting pressure load of the damaged RPV head.
Based on the review of this calculation package, the NRC raised a number of questions (See Appendix A), the majority of which were concerned with the failure criteria used in the evaluations.
The purpose of this calculation is to develop a better understanding of the failure criteria as used and its relative "conservativeness" in regards to the failure pressure.
2.0 Technical Approach The failure criterion used in Reference 1 was set such that the maximum strain could not exceed the ultimate tensile strain. Hence for the stainless steel cladding where the maximum strain is expected to occur, the maximum equivalent total strain is limited to the maximum strain of 11.15%
(corresponding to the ultimate strain for the stainless steel cladding in Reference 2) through the thickness of the component.
In order to evaluate the reasonableness of this failure criterion, the results of the failure pressures predicted with this criterion were compared against test results of very similar geometries. Disk burst test, similar to the Davis-Besse head wastage geometry were performed under the auspices of the PVRC Subcommittee and documented in and ASME publication [3] (see Appendix B for the actual publication).
Described in Reference 3 were a series of burst tests using machined disks of various materials. The test disk dimensions and the illustration of the test setup are shown in Figure 1. The materials tested included 304 Stainless Steel, A-533 Grade B Low Alloy Steel and A85 Grade C Carbon Steel. For the purposes of this calculation, only the 304 Stainless Steel testing will be reviewed.
As can be seen in Figure 1, three basic disk geometries wvere tested. In order to evaluate the effectiveness of the failure criteria developed for Reference 1, the same failure criteria will be used to determine the disk burst pressures. As a result, a series of finite element models were developed using the test disk dimension provided in Reference 3. The models were created and evaluated using the ANSYS finite element software [4]. The actual evaluations and subsequent failure criteria comparison are included in the following sections.
PRELIMINARY 3.0 Finite Element Models A series of finite element models were constructed to determine burst pressure for the various disk configurations. Initial studies were performed using an axisymmetric model but subsequent evaluations included three-dimensional modeling similar to that used in Reference 1.
The elastic material properties for all evaluations were for 304 stainless at room temperature as defined by Reference 5. These values used were as follows:
l Modulus of Elasticity, E, e6 psi: l 28.3 l Poisson's Ration, v: 1 0.3 The plastic material properties for stainless per Reference 3 were:
0.25 Y.S. SuitCult Reduction A 1' (psi) (psi) (in/in) In Area (psi) 34,000 84,000 0.54 0.74 193,060 0.494
[1] Stress Strain Curve Assumed to be of form v = A (E) n Therefore the stress-strain curve used in all of the evaluation is shown in Table 1. Any additional model specific conditions will be described in the following sections.
Table 1 Stress Strain Curve for 304 Stainless Steel 131 Strain (in.in) Stress (psi) 000O 0 0.025 31208 63 0 050 43952.49 0 075 53699.79 0 100 61900 24 0 125 69113 97 0 150 75627.79 0 175 81611.83 0 200 87176 84 0.225 92399 68 0.250 97336 26 0 275 102028 8 0 300 106510 0 325 110805.8 0 350 114937.5 0.375 118922.4 0 400 122775 0 425 126507.5 0 450 130130 6 0 475 133653 1 0 500 137083 0 525 140427 1 0 550 143691.6 0575 146881.9 0.600 150002 7 0 625 153058 4 0 650 156052 8 0 675 158989.5 0 700 161871.6 0 725 164702 2 0 750 167483.7 0.775 170218 7 0 800 172909.5 0.825 175558 0 850 178166 2 0 875 1807358 0 900 183268 6 0 925 185766 0 950 188229.5 0 975 190660 4 I 000 193060 1 025 1954294 1.050 197769.7 1 075 200082 I 100 202367.3 1 125 2046264 1.150 206860 2 1.175 209069 7 1 200 211255 4 1.225 213418 2
PRELIMINARY 3.1 Axisymmetric Finite Element Model Structural Solid I The axisymmetric models were developed in ANSYS using the 2-D 8-Node evaluated as was the element, PLANE82. All three geometries described in Reference 3 were A total of 5 effects of the finite element mesh density on the onset of numeric instability.
each evaluation was evaluations for each disk geometry were made, the only difference between the thickness of the mesh density, which can be simplified to the number of elements through where 4, 6, 8, 10 the thinned portion of the disk. As such, the mesh densities that were evaluated of mesh density for and 12 elements through the thickness. Figure 2 shows the progression geometry-A.
simple vertical restraint The mechanical boundary conditions for these evaluations consisted of assumed to into the entire region of throughout the approximate clamp region. This region was an example of the applied the disk, which remained at the full 1 inch thickness. See Figure 3 for model.
boundary conditions on the 4 element through thickness, geometry-A 3.2 Three-Dimensional Finite Element Model 8-Node Structural The three-dimensional models were developed in ANSYS using the 3-D 3 were evaluated as was Solid element, SOLID45. All three geometries described in Reference instability.
the effects of the finite element mesh density on the onset of numeric geometries were also Only a 30° section of the total disk was modeled since the loading and only difference between symmetrical. Two evaluations for each disk geometry were made; the the number of elements each evaluation was the mesh density, which again can be simplified to the mesh densities for the 3-through the thickness of the thinned portion of the disk. As such, the thickness. It should dimensional models that were evaluated were 4 and 6 elements through used 6 elements be noted that the stainless clad for the actual Davis-Besse cavity evaluation through the thickness. Figure 4 shows the two mesh densities for geometry-A.
vertical restraints as The mechanical boundary conditions for these evaluations used the same were applied to the axisymmetric evaluations. In addition, axisymmetric boundary conditions This results direction.
the free ends of the disk, the preventing translations in the circumferential Figure 5 direction See in the centerline of nodes being limited to translation in only the vertical thickness, for an example of the applied boundary conditions on the 4 element through geometry-A model.
4.0 Loading pressure was applied to the All of the evaluations were loaded in the same manner. An incremental an example of the applied cavity surfaces until numeric instability was reached. See Figure 6 for pressure.
PRELIMINARY 5.0 Mesh Density Results For each evaluation, the pressure was allowed to rise incrementally until numeric instability occurred. The points of instability, as compared to the actual disk burst tests, are shown in Table 2.
Table 2 Mesh Density Effects of Numeric Instability Model Pressure (psi) Predicted/
MThrough-Wall Numeric Actual Test Test Result Model Type Elements Instabili Burst (%)
Axisymmetric 4 12725 84.8 Axisymmetric 6 13942 92.9 Axisymmetric 8 14004 93.4 Axisymmetric 10 14022 15000 93.5 Axisymmetric 12 14005 93.4 3-Dimensional 4 13979 93.2 3-Dimensional 6 13997 93.3 Axisymmetric 4 5929 87.2 Axisymmetric 6 6630 97.5 Axisymmetric 8 6695 98.5 Axisymmetric 10 6695 6800 98.5 Axisymmetric 12 6694 98.4 3-Dimensional 4 6688 98.4 3-Dimensional 6 6671 98.1 Axisymmetric 4 6317 82.0 Axisymetric 6 6962 90.4 Axisymmetric 8 6997 90.9 Axisymmetric 10 6998 7700 90.9 Axisymmetric 12 6997 90.9 3-Dimensional 4 6976 90.6 3-Dimensional 6 6974 90.6 The results are also shown graphically in Figure 7.
PRELIMINARY 6.0 Total Strain Results Based on Section 5, only the highest through-wall element count cases wvill be further evaluated. As a result, Figures 8 though 10 show the total Von Mises Strain just prior to onset of instability for the 12 through-wall element axisymmetric model and Figures 11 through 13 show total Von Mises Strain for the 6 through-wall element 3-D model.
7.0 Strain Criteria Comparison The original failure strain criterion described in Section 2.0 indicated that when the through-wall total strain exceeded the uniform elongation percentage, the structure would be considered to have failed. As a check of this criterion, the total Von Mises nodal strains as they varied with pressure were extracted from the middle of the modeled disk at the top, middle and bottom of the wall thickness. The resulting strains were then plotted versus the pressure and compared to the actual burst pressure measured in Reference 3 and the failure pressure as defined by the Failure Criterion in Section 2.0.
From the definition of material properties used in the disk burst test, the uniform elongation for 304 stainless steel was 54% (see Section 3.0). Therefore, the failure of the disk will occur when the through-wall total strain exceeds 54% throughout the thickness.
An examination of the 3 geometries for both the axisymmetric and 3-D modeling can be seen in Figures 14 though 19. The results are further summarized in Table 3.
Table 3 Failure Criteria Comparison Model Model Failure Pressure psi)
Type Geometry Burst Test 131 Instability Failure Criteria Axisymmetric A 15000 14005 -11000 Axisymmetric B 6800 6694 -5500 Axisymmetric C 7700 6997 -5750 3-Dimensional A 15000 13997 -11000 3-Dimensional B 6800 6671 -5500 3-Dimensional C 7700 6974 -5750
PRELIMINARY 8.0 Conclusions Based on the summary in Table 3 of Section 7.0, the use of the uniform elongation limit as the basis of failure criteria in an elastic-plastic finite element analysis results in conservative failure pressures as compared to actual test results. For the three geometries, the uniform elongation criteria predicted a failure pressure that was in the range of 73% to 84% of the actual failure pressure.
A better prediction of actual failure pressure is the pressure at which numeric instability was reached in the ANSYS program. Assuming a numeric instability criterion, failure pressure would range from 90% to 98% of actual failure pressure.
V Revision Preparer/Date Checker/Date 0
RLB 5/31/02
_ File No. W-DB- OlQ-304 Page 8 of 28
PRELIMINARY 9.0 References
- 1) Structural Integrity Calculation W-DB-OIQ-301, Rev. 1, "Elastic-Plastic Finite Element Stress Analysis of Davis-Besse RPV Head Wastage Cavity."
- 2) Email of from B.R. Grambau (Framatome ANP) to N. Cofie (SI), "308 Stress -Strain Curve," March 15, 2002, SI File W-DB-O0Q-202.
- 3) P. C. Riccardella, "Elastic-Plastic Analysis of Constrained Disk Burst Tests," ASME Paper No. 72-PVP-12, Proceedings of Pressure Vessel and Piping Conference, New Orleans, LA, September 17-21, 1972.
- 4) ANSYS/Mechanical, Revision 5.7, ANSYS Inc., December 2000
PRELIMINARY
.STRAIN-GAGED CANTILEVER BEAM FOR CENTRAL DEFLEC-TION MEASUREMENT SCHEMATIC ILLUSTRATION OF TEST SETUP DISK SPECIMEN 1 Oin THICKNESS FILLET GEOMETRY (t) RADIUS A 025 in 0375 in B 0125in 0125in C 0 125 in 0 375 in 02055RO Figure 1 - PVRC Disk Test Details (Reference 3)
PRELIMINARY I T-H-'
I, III.- I I I . .,
1 -. .
. .. i 4 Element Through-Wall _
I.:.rI i ., -U .1. .
.I iIII,:.: _'1114"I'l- fi 8 Element Through-Wall 12 Element Through-Wall Figure 2 - Mesh Density Example for Axisymmetric Finite Element Model for Geometry-A
L
-- I i ., -. - - - - , II Figure 3 - Mechanical Boundary Conditions Example for Axisymmetric Finite Element Model for Geometry-A
PRELIMINARY
- .i. P, -
! I Is' I:
I I I,:
, I 4 Element Through-WVall I
,*1' I
i "I1;
- I : I I ;
i
! .: : I:
. I ::
4 I.I-t, I' I
6 Element Through-Wall
.. -- i Figure 4 - Mesh Density Example for 3-D Finite Element Model for Geometry-A
PRELIMINARY
, --,,','I' ST--:'X Figure 5 - Mechanical Boundary Conditions Example for 3-D Finite Element Model for Geometry-A
PRELIMINARY r
Applied Pressure Figure 6 - Applied Pressure Example (Axisymmetric Geometry-A Model)
Revision
~Preparer/Date ,RLB3 5/31102
<>Checker/Date IFile No. W-DB-01 Q-304 I Page 15 of 28
PRELIMINARY Mesh Refinement vs. Onset of Numeric Instability Pressure l 15000 14000 13000 12000 I_ =_
_ _ _ _t
-Axi-Geometry-B
'___i___
co C
11000 mAxi-Geometry-C
-4 3D-Geometry-A 0 _ ______-- _ -3D-Geometry-B_
10000
-43D-Geometry-C 9000 8000 7000 ,
- a ,
6000 4 5 6 7 8 9 10 11 12 Through-Wall Element Count Figure 7 - Mesh Density Effects
PRELIMINARY ANSYS 5.7 MAY 22 2002 08:48:09 PLOT NO. 1 NODAL SOLUTION STEP=1 SUB =50 TIME=.933654 EPTOEQV (AVG)
EffNu=O DMX =2252 SMN =.472E-03 SMX =1.216
.472E-03
- .135477 EM.270482
.540492
\ _1.405087
__.675497
.810503
-. 945508 Von Mises Total Strain (12x96) - (Category-A)
Figure 8 - Total Von Miscs Strain Just Prior to Numeric Instability Geometry-A - Axisymmetric
ANSYS 5.7 MAY 22 2002 08:20:47 PLOT NO. 1 NODAL SOLUTION SIEP=1 SUB =50 TIME=.44627 EPTOEQV (AVG)
EffNu-O DeMX =2.25 SMN =.183E-03 SMvX =1 A07 N_ .183E-03 am 469136
.625454
.781772 ri.938089 1.407 1x9)-
No CtgoyB .312819Stan Von Mises Total Strain -(12x96) -(Category B)
Figure 9 - Total Von Mises Strain Just Prior to Numeric Instability Geometry-B - Axisymmetric
PRELIMINARY ANSYS 5.
MAY 22 2002 08:26:25 PLOT NO. 1 NODAL SOLUTION STEP-1 SUB =43 TIME=.A66474 EPTOEQV (AVG)
EffNu=iO DMXC =2263 SMN =.206E-03
=1218 SMX--1
.206E-03
>' _ 135536
.24706195
.541524
.676854
\ E-812183
.947513 Von Mises Total Strain - (12x96) - (Category-C)
Figure 10 - Total Von Mises Strain Just Prior to Numeric Instability Geometry-C - Axisymmetric
PRELIMINARY ANSYS 5.7 MAY 22 2002 09:01:11 PLOT NO. 1 NODAL SOLUTION STEP=1 SUB =26 TIME=.933132 EPTOEQV (AVG)
EffNu=O TOP DMX =2.296 SM =.Z77E-03 SMX =1.167
.277E-03
_ 12991.1 E .389178
_ .518812
.648446
.E3778079 i.037 1.167 Von Mises Tatal Strain - (6x48) - (Category-A) .
Figure 11 - Total Von Mises Strain Just Prior to Numeric Instability Geometry-A Dimensional
PRELIMINARY ANSYS 5.7 MAY 22 2002 09:09:37 PLOT NO. I NODAL SOLUTION STEP-1 SUB =30 TIME=.444724 EPTOEQV (AVG)
Ef fNu=O TOP DMX =2.245 S =.112E-03 SWX =1.353
_ 12E-03
_ 150391 30669 M
.4A50947 iJ .751503 H901781 1.052 Von Mises Total Strain - (6x48) - (Category-B)
Figure 12 -Total Von Mises Strain Just Prior to Numeric Instability Geometry-B Dimensional Revision Preparer/Date RLB 5/31/02 Checker/Date
_File No. W-DB-OIQ-304 Page 21 of 28
PRELIMINARY ANSYS 5.7 MAY 22 2002 09:14:07 PLOT NO. 1 NODAL SOLUTION STEP=1 SUB =28 TIME-.464911 EPTOEQV (AVG)
EffNu=O TOP DMX =2157 smN =.114E-03 SX-1 =1113
- 419E-03
_ 123754
.371032
_ A94672
[E 64189311
.86559
.989229 1.113 Von Mises Total Strain - (6x48) - (Category-C) .
Figure 13 - Total Von Mises Strain Just Prior to Numeric Instability Geometry-C Dimensional
PRELIMINARY F -_ -
Pressure vs Total Von Mises Strain at Center Geometry-A (Axisymmetric) 16000 14000 I Test Burst Pressure = 15,000 psi-1
................................. ........................................................... I 12000 10000 0.
2 8000 a-6000 4000 2000 0
0 0.2 0.4 0.6 0.8 1 1.2 1.4 Total Von Mises Strain (In/in)
Figure 14 - Through-Wall Strain Results at Center of Disk Geometry-A (Axisymmetric)
Pressure vs Total Von Mises Strain at Center Geometry-B (Axisymmetric) 7000 . . ..T~est Burt Pre's's-ure-=- 6.,8.00 .psi. . . ......................................... ................ _.
I 6000 I _;&
5000
- a. 4000 0
- 3000 C.
2000 l 1000 0
0 0.2 0.4 0.6 0.(8 1 1.2 Total Von Mises Strain (inrin)
Figure 15 - Through-Wall Strain Results at Center of Disk Geometry-B (Axisymmetric)
PRELIMINARY tI Pressure vs Total Von Mises Strain at Center I Geometry-C (Axisymmetric) 8000
.... et u.r.s.t.
Pressure ..7.. I........ ......................................................
7000 I
6000
- 5000 i
e 4000 At
- 0. 3000 2000 1000 0
0 0.2 0.4 0.6 0.8 1 1.2 I1 Total Von Mises Strain (Inlin) __ _ __ __ _ _ _ _ _ ___
_ I Figure 16 - Through-Wall Strain Results at Center of Disk Geometry-C (Axisymmetric)
Revision 0 Preparer/Date RLB 5/31/02 iV Checker/Date File No. W-DB-OIQ-304 Page 25 of 28
PRELIMINARY Pressure vs Total Von Mises Strain at Center Geometry-A (3-Dimensional) 16000 Test Burst Pressure = 15,000 psi 14000 12000
_ 10000 0.
=U 8000 I r 6000 4000 I
2000 0
0 0.2 0.4 0.6 0.8 1 1.2 Total Von Mises Strain (inr/in)
Figure 17 - Through-Wall Strain Results at Center of Disk Geometry-A (3-Dimensional)
PRELIMINARY Pressure vs Total Von Mises Strain at Center Geometry-B (3-Dimensional) 7000 Test Burst Pressure = 6,800 psi I_
6000 5000 3 4000 a 3000 Q.
2000 1000 0
0 0.2 0.4 0.6 0.8 1 1.2 Total Von Miss* Strain (inmin)
Figure 18 - Through-Wall Strain Results at Center of Disk Geometry-B (3-Dimensional)
PRELIMINARY Pressure vs Total Von Mises Strain at Center Geometry-C (3-Dimensional) 8000 -
Test Burst.Pressure. 7,70. p........s i
7000 -
6000 -
=5000 0.1 2 4000 -
Oi( 3000 2000 1 1000 I
0 0 0.2 0.44 0.6 0.8 1 1.
Total Von Milses Strain (Innin)
I Figure 19 - Through-Wall Strain Results at Center of Disk Geometry-C (3-Dimensional)
V Revision Preparer/Date Checker/Date 0
RLB 5/31/02
_ File No. W-DB-01 Q-304 Page 28 of 28
PRELIMINARY APPENDIX A NRC Staff Comments and Ouestions on Davis-Besses Safety Significance Assessment (SIA-W-DB-01Q-301) Submitted April 8. 2002
PRELIMINARY NRC STAFF COMMENTS AND QUESTIONS ON DAVIS-BESSE SAFETY SIGNIFICANCE NRC STAFF COMMENTS AND QUESTIONS ON DAVIS-BESSE SAFETY SIGNIFICANCE ASSESSMENT (SIA-W-DB-01Q-301) SUBMITTED APRIL 8,2002 FAILURE CRITERION (1) What is the technical basis of the failure criterion (e.g., strain exceeding 11.15%) used to determine the failure conditions of the cladding layer? Provide specific technical references in the literature that support the failure criterion used in this evaluation.
(2) How does the failure criterion (e.g., based on ultimate strain in a uniaxial tensile test) account for the effects of biaxial loading in the cladding, or triaxial loading in the cladding at the edges of the degradation cavity?
(3) The failure criterion applied in SIA report W-DB-01Q-301 (e.g., the minimum cross-sectional strain exceeding the failure strain of 11.15%) allows the strain levels in the cladding to exceed the critical strain value entirely through the thickness, leading to very large strains at the surface of the cladding, up to 49% in Table 5 of the SIA report. What is the technical basis for this approach, as opposed to the average cross-sectional strain, or the maximum cross-sectional strain?
(4) Did you explore a continuum damage mechanics analysis to give guidance of the failure criterion once the strains exceed the critical strain where necking/void growth starts? If not, provide the technical basis for not using a continuum damage mechanics analysis.
[Poisson's ratio of 0.5 no longer applies once this critical strain level is exceeded, so the analysis is strictly not valid. (Poisson's ratio is continuously changing as the voids grow at the strains beyond the start of necking.) This results in a stress redistribution that is not accounted for in a standard elastic-plastic analysis.]
(5) How would the strain values change if the stress free temperature was assumed to be the stress relief temperature instead of 700 F, and the analysis accounted for the differential thermal expansion of the cladding and head steel at the operating temperature of 6050F?
GEOMETRY/MESHING (A) Does the size of the degradation cavity and the transition from the cladding thickness to the head thickness that was used in the SIA report reflect current knowledge regarding the cavity geometry, in particular the undercut area described in Figure 13 on page 103 of the Davis-Besse Root Cause Analysis Report (CR2002-0891), dated April 15, 2002? What is the transition geometry assumed in the analyses?
(B) Is there sufficient mesh refinement through the cladding thickness to adequately capture the bending and shear strains at the edge of the cavity? Describe any sensitivity studies used to demonstrate the adequacy of the mesh refinement.
(C) Was the cladding deposited by weld wire? Do the thinner cladding thickness measurements from UT coincide with the locations of weld bead toes? In what direction do the cladding weld beads run relative to the long axis of the degradation cavity?
Revision 0 Preparer/Date RLB 5/31/02 Checker/Date I File No. W-DB-OIQ-304 Page A2 of A2
PRELIMINARY APPENDIX B Pressure Vessels and Piping Division Paper No. 72-PVP-1 2, "Elasto-Plastic Analysis of Constrained Disk Burst Tests,"