ML052770358
| ML052770358 | |
| Person / Time | |
|---|---|
| Site: | Catawba  |
| Issue date: | 09/22/2005 |
| From: | Jamil D Duke Energy Corp, Duke Power Co |
| To: | Document Control Desk, Office of Nuclear Reactor Regulation |
| References | |
| Download: ML052770358 (27) | |
Text
D.M. JAMIL
-b DDuke Vice President SoPowere Duke Power Catawba Nuclear Station 4800 Concord Road / CN01 VP York, SC 29745-9635 803 831 4251 803 831 3221 fax September 22, 2005 U.S. Nuclear Regulatory Commission Attention: Document Control Desk Washington, DC 20555
Subject:
Duke Energy Corporation Catawba Nuclear Station, Unit 2 Docket No. 50-414 Second Ten Year Inservice Inspection Interval Steam Generator C Hot Leg Nozzle Welds
References:
Letters from Duke Energy Corporation to NRC, same subject, dated October 19, 2004 and December 2, 2004 The reference letters transmitted Revisions 0 and 1 of an analytical evaluation of a steam generator hot leg nozzle weld flaw discovered during the Catawba Unit 2 End of Cycle 13 Refueling Outage. The evaluation was contained in WCAP-15658-P, "Flaw Evaluation Handbook for Catawba Unit 2 Steam Generator Primary Nozzle Weld Regions" (Proprietary).
On May 11, 2005, the NRC transmitted a Request for Additional Information (RAI) concerning this analytical evaluation. The purpose of this letter is to respond to this RAI. The RAI response is contained in the attachment to this letter. The format of the response is to restate each RAI question, followed by the response.
There are no regulatory commitments contained in this letter or its attachment.
If there are any questions concerning this information, please contact L.J. Rudy at (803) 831-3084.
Dq Ga Catawba Nudear Station2 th www. dukepower. corn Anniversary 1985-2005
A-Document Control Desk Page 2 September 22, 2005 Very truly yours, D.M. Jamil LJR/s Attachment
Document Control Desk Page 3 September 22, 2005 xc (with attachment):
W.D. Travers U.S. Nuclear Regulatory Commission Regional Administrator, Region II Atlanta Federal Center 61 Forsyth St., SW, Suite 23T85 Atlanta, GA 30303 E.F. Guthrie Senior Resident Inspector (CNS)
U.S. Nuclear Regulatory Commission Catawba Nuclear Station S.E. Peters (addressee only)
NRC Project Manager (CNS)
U.S. Nuclear Regulatory Commission Mail Stop 0-8 G9 Washington, D.C. 20555-0001
ATTACHMENT RESPONSE TO NRC REQUEST FOR ADDITIONAL INFORMATION
REQUEST FOR ADDITIONAL INFORMATION DUKE POWER COMPANY CATAWBA NUCLEAR STATION, UNIT 2 DOCKET NO. 50-414 The Nuclear Regulatory Commission (NRC) staff has reviewed the licensee's submittals dated October 19 and December 2, 2004, regarding an evaluation of a flaw indication in the reactor coolant hot leg to steam generator nozzle connection, that was discovered on October 7, 2004, during the 13th refueling outage for Catawba Nuclear Station, Unit
- 2. The NRC staff has identified the following information that is needed to enable the continuation of its review.
- 1. In the letter dated October 19, 2004, you stated,
"[t]he indication was located near the interface between the safe-end and field weld at the bottom of the nozzle."
Please confirm whether the flaw indication is in the safe end or in the field weld.
Duke Energy Corporation Response:
The flaw is located at the boundary between the safe end (actually a stainless steel buttering of the carbon steel nozzle) and the stainless steel field weld. A sketch is attached showing the location of the flaw relative to the safe end buttering and the field weld.
1
CNS Steam Generator 2C Hot Leg Nozzle Indication SS Cladding / Buttering Film standoff of exterior = 0.25" SUrfae --
Source Location
=21.25" from back of nozzle dam ring
-24.77" from back of nozzle dam ring Reference Drawings:
- 1) CNM-2201 .01-0008
- 2) CNM-1201.01-0076 2
I
- 2. In the letter dated October 19, 2004, you stated,
'[t]his letter submits the fracture mechanics analysis to the NRC (see attachment)". The NRC staff did not find in your submittal an evaluation of the detected flaw indication (1 inch long circumferential embedded flaw, 1.01 inches from the outside diameter of the pipe) using WCAP-15658-P, Revision 1. Please provide this information. The response should include the WCAP figure number (Figure A-4.6, Figure A-4.7, Figure A-4.8, or Figure A-4.9) that you used for evaluation of the detected flaw in the steam generator primary nozzle weld region. The response should also include the depth of the detected flaw (the size of the flaw in the wall thickness direction).
Duke Energy Corporation Response:
Nozzle Configuration The nozzle connection consists of a low alloy steel casting that forms the channel head of the steam generator. This casting has been buttered with a low carbon, stainless steel weld metal. The piping to buttering field weld is made after post weld heat treatment of the steam generator channel head. The Duke weld number is 2NC-13-2. The weld is a full penetration, compound V groove weld made from the outside of the pipe. The GTAW (TIG) process was used for the first inch, followed by a "courtesy" radiograph (RT).
Subsequent welding was performed using the SMAW (stick) process to finish out the weld. After completion of welding, a final RT was performed and accepted on the weld.
In addition, liquid penetrant tests (PTs) were performed on the interior and exterior surfaces of the weld.
Flaw Geometry The flaw is located at the bottom of the pipe in the C hot leg. It is approximately at bottom dead center of the pipe. Based on the radiographic data, the flaw is one inch long and oriented in the circumferential direction. Since the examination was performed using radiographic testing (RT), a limited amount of information was available to characterize the flaw. The location of the flaw relative to the outside diameter surface was established using parallax radiographic shots. These shots support a minimal flaw depth. However, because of the uncertainty in flaw depth, a bounding case has been reviewed herein.
3
Radiography Evaluation In addition to the radiography shots made to characterize the flaw, both the original construction film and the End of Cycle 13 film (non-parallax shots) were digitized. The original construction film was digitized to determine if the indication could be seen from initial fabrication welding. Digitization of the film can greatly enhance the visible interpretation of the film in some cases. Next, the End of Cycle 13 film was digitized and reviewed to determine if the linear indication was actually separated into multiple flaws. In both of these cases, there was no conclusive evidence from the digitization process that changed the film interpretation or flaw characteristics.
Based on the conclusions provided from the radiography review above, a best estimate characterization of the flaw has been provided. The flaw is located 1.01 inches from the outside surface of the piping in the stainless steel weld material. The flaw is oriented circumferentially with a length of 1.0 inch. The flaw is most likely the result of a slag inclusion during fabrication. It has very little contrast, which indicates a limited depth. It is located at the interface between the stainless steel buttering and the Duke stainless steel field weld.
The flaw location from the outside diameter surface of 1.01 inches was considered from three positions relative to the flaw depth (top, center, and bottom). The three positions were considered for two different aspect ratios. All six cases were found to be acceptable. Based on the method used to determine location, the center position is the most appropriate and is used in the documented flaw calculation below.
For the initial evaluation, the flaw depth will be assumed as one-half of the length and evaluated as an embedded flaw. From the flaw handbook, several parameters are necessary to determine the acceptability of the indication.
These are provided below. The appropriate figure for the purposes of evaluation from WCAP-15658-P, Revision 1 for a circumferential embedded flaw in the stainless steel material is Figure A-3.7.
a = half flaw depth (in)
= 0.25 in 1 = length of flaw (in) 4
= 1.0 in t = wall thickness (in)
= 3.25 in Note: The wall thickness is based on profiling of the weld using ultrasonic testing (UT) probe.
The value of 3.25 inches is conservative and represents the lowest reading throughout the weld region of interest.
5 = distance to flaw centerline
= 1.01 in
/ t = 1.01 /3.25
= 0.311 a t = 0.25 /3.25
= 0.077 The 5 / t and a / t parameters may be plotted on Figure A-3.7 to determine the acceptability of the flaw. This point (A) is shown on the attached sketch.
In addition to the above evaluation, the flaw depth was increased to 1 inch yielding an aspect ratio of 1:1. In this case, the parameters change as noted below:
a = half flaw depth (in)
= 0.50 in 1 = length of flaw (in)
= 1.0 in t = wall thickness (in)
= 3.25 in 6 = distance to flaw centerline
= 1.01 in
/ t = 1.01 /3.25
= 0.311 a t = 0.50 /3.25
= 0.154 Again, the 5 / t and a / t parameters have been plotted on Figure A-3.7 as point (B) to determine the acceptability of the flaw.
5
A-19 WESTINGHOUSE PROPRIETARY CLASS 2 CATAWBA 2 STEAM GENERATOR INLET NOZZLE STAINLESS STEEL PIPE WELD CIRCUMFERENTIAL EMBEDOED FLAW EVALUATION CHART Sur1ceITR*edd~d Flaw Deniaratlon Lkwe 028 0.27 _11
. 1:.... Embed&d Flaw Confl"wation
- - r I =
F r r 10.20.30 ym.
0.26
- - =-14 L
4-; -4 j _~
X7 . i
__-L : _.
025 0.24 7 023 __1~ ..
0.22 --- I -
0.21 ~-.-I 02 0.19 4-EE 0.18
- 0.17 '4 117 I: t- I ' I = -, 11'
_,ii surface A . -
l 0.16
- - ! FlwsIn#th "
(i: 0.15 o6 0.14 rr.... utb....
= 0.13 TI . .7 3_ -AV,
§ 0.12 ! ._; * ; , - . l i 0 $ ifip. 1
- 0.11 L
I !1- t-- - :i 17 I 0.1 ~i .;: ::R-: .;,:
o~o0s4-+-
_i -. :
0.07 FE 0.06 0.05 ._ Tri - - 11rr~rI 0.04 -
00t n N _. _II_}.J4 rs=.wTfr'll5 i '
0 0.05 0.1 0.15 02 0.25 03 035 0.4 0.45 0.5 Distance from Surface (St)
X Inside Surface _Surface Flaw _ Longitudinal Flaw X Outside Surface X Enbedded Flaw X Circumferential Flaw Figure A-3.7 Evaluation Chart for Steam Generator Inlet Nozzle Safe-End to Pipe Weld (StainlessSteel) C tjL*QI.I 01
. CK-ooC S tv O
?%3t. 7 WCAP- 15658.P Squmber 2004 WCAP.15658-P (Sarait) RI.m101-0I~4 Le.=cC ia-sa.o4
. G.L '.)?W10 04-6
Results In both cases evaluated above, it is clearly evident that the flaw is within the bounds of the acceptability provided by Figure A-3.7 of WCAP-15658-P, Revision 1. As a result, the piping containing this flaw is acceptable for continued service for the design life of the plant. The figure in WCAP-15658-P, Revision 1 indicates 10, 20, and 30 year acceptance lines. These lines are related to the design number of occurrences of transients used in the fatigue crack growth calculation. As such, this indication is acceptable for the life of the plant provided a prorated value (30 / 40 = 75%) of the design number of occurrences are not exceeded between now and the end of plant life.
This limit on fatigue cycle counts will be tracked under our fatigue management program.
Conclusion The flaw discovered during End of Cycle 13 is acceptable without repair for the life of the plant. Acceptance by the performance of analytical evaluation as allowed by ASME Section XI, IWB-3132.4 has been validated. Additional examinations have been performed during End of Cycle 13 to satisfy IWB-2430. Successive examinations for the SG 2C hot leg weld number 2NC-13-2 will be necessary in the subsequent three ISI periods as required by IWB-2420.
- 3. On Page 3-1 of WCAP-15658-P, Revision 1, "Flaw Evaluation Handbook for Catawba Unit 2 Steam Generator Primary Nozzle Weld Regions," November 2004", it is stated,
"[t]he stress intensity factor calculation for an embedded flaw was taken from the work by Shah and Kobayashi [6]
which is applicable to an embedded flaw in an infinite medium .... This expression has been shown to be applicable to embedded flaws in a thick-walled pressure vessel in a paper by Lee and Bamford [7]." Please demonstrate the applicability of Kobayashi's formulas for embedded flaws to your current application by addressing (1) the difference between the finite geometry of the current application and an infinite medium discussed in Kobayashi's paper, and (2) the difference between the ratio of plate thickness to crack depth, t/2a, of the current application and that discussed in Kobayashi's paper. Provide Reference 7 of WCAP-15658-P, Revision 1 (Paper 83-PVP-92 by Lee and Bamford) if you believe it would help your explanation.
7
Duke Energy Corporation Response:
Note: The response to this question was developed by Westinghouse. Refer to the enclosed Westinghouse material for the additional information to support this response.
The Lee and Bamford paper is attached for your information (Enclosure 1), and should provide a sufficient basis for the use of the Shah and Kobayashi closed form solution for the embedded flaw. This work was done for the express purpose of deciding whether a closed form solution was sufficient to model embedded flaws in finite thickness geometries, and the conclusion was that the closed form solution was indeed good for this application. This conclusion was reached by setting up a series of finite element geometries and loadings, and comparing the weight function results with the closed form solutions of Shah and Kobayashi. The detailed comparisons are provided in the paper.
- 4. On Page 3-4 of WCAP-15658-P, Revision 1, it is stated, "NRC procedures exist for addressing the impact of thermal aging on fracture toughness for full-service life. The approved procedures were applied to the nozzle safe end to pipe weld, as well as to the cast piping itself." Provide the specific document (e.g., NUREG number) and parameters used (e.g., ferrite content) in your determination of fracture toughness for full-service life using NRC procedures. Explain how you use these NRC procedures to determine the first set of proprietary Juc and Tot given on Page 3-4. It is further stated on this page, "[e]ven with thermal aging, equivalent to full service for SAW welds, the tearing modulus remains high (>100) and the unaged toughness, J1c, is not significantly reduced." Provide information supporting this statement.
Duke Energy Corporation Response:
Note: The response to this question was developed by Westinghouse. Refer to the enclosed Westinghouse material for the additional information to support this response.
The cover letter transmitting the first SER issued by NRC on this subject is attached (Enclosure 2); many others have been issued over the years which provide a similar 8
endorsement of the Westinghouse approach. This approach was developed over an extended period of time with Westinghouse internal funding, and included interactions with the Staff on many occasions to clarify the methodology, so it was important to keep the methodology out of the public literature. Westinghouse has used this same approach to describe the assessment of thermal aging which was done for a number of different flaw evaluations, over a period of many years.
9
ENCLOSURE 1 LEE AND BAMFORD PAPER
I THE AMERICAN SOCIETY OF MECHANICAL ENGINEERS 83-PVP-92 345E.478LNwYwkY.1017 i.
IxsI Te soce ale ot be reep4 ibe for vtee es or "*monosomwed inPsM or in dosc son aIt tin et mmsaci pIca0lM Ds i w of ari ODe uie or stlana or p ntd hi It s intsd ol If 1w pepw is pubi ahed hIn U AtI J ouL elees for "er*[* P llcen Mon pWtesetn Ful ffedit shud be gien to ASME, tVW Tetnical Diion. end th 44s~ Ppe we ovae ft_ A&E h nine Warmonh srftw eeng.
Prited hi USA STRESS INTENSITY FACTOR SOLUTIONS FOR A LONGITUDINAL BURIED ELLIPTICAL FLAW IN A CYLINDER UNDER ARBITRMY LOADS Y. S. Lee, Mem. ASME W. H. Bamford, Mem. ASME Westinghouse Electric Corporation Nuclear Energy Systems P.O. Box 355 Pittsburgh. Pennsylvania 15230 ABSTRACT included the effect of the outside4 surface of the cylinder. For example. Underwood' ' assumed that Elastic stress intensity factor solutions were the effect of a given crack shape on KI (stress obtained for a burled elliptical flaw of major to intensity factor) of a pressurized cylinder isthe semi-minor axis ratio 6. located in the longitudinal same as that due to the shape effect on the K1 ofa plane of a cylinder of radius to thickness ratio 10. plate under uniform tension. Kobayashi et alt6 7 I The flaws were located at distances of 3/16. 1/4, 1/2 have estimated KI for a longitudinal semi-and 3/4 times the wall thickness from the inside of elliptical surface flaw in a cylinder under both the cylinder wall. The results were compared with pressure and thermal shock loadings. In this study.
those for an infinite elastic medium available In the the authors determined the solutions for flaws in literature. The results presented in this report cylinder from the solution for similar flaws in flat show that the infinite medium solutions can be used plates subjected to identical stress profiles. The for evaluation of cylindrical vessel geometry flat plate results were then modified with curvature considered in this report for flaws located within 25 correction factors obtained from a two-dimensional to 75 percent of the wall thickness. For flaws analysis. Both front and back surface effects were located outside this region the stress intensity considered, but the curvature correction factors were factors appear to increase. A comparison between the most accurate for only the deepest point along the ASME Section XI membrane correction factors and those semi-elliptical flaw.
of this work shows that the former are conservative for the eccentricity limit specified in the code. More recently direct three-dimensional solutions to the semi-elliptical surface flaw problem in a INTRODUCTION cylinder have been obtained by several investi-gators Among these are the re jlts reported by The burled elliptical flaw is one of the most common Ayers.19) 1U1ckburn and Hellen ,II1 and Atluri.
flaw types found in many structure. Accurate stress et al.Jl Ayers used a condensed quarter-point intensity factor expressions under all applicable element to determine the stress intensity factor loading conditions are necessary to establish the distribution of two semi-elliptical flaws In a actual safety margin. tongitudinal buried flow in a thermally shocked cylinder with ratio of outer to thick-walled cylinder is of particular interest inner radius equal to 1.90. Blackburn and Hellen because of its application in the structural used a conventional three-dimensional finite element integrity evaluation of reactor vessel beltlines. code and a virtual crack extension procedure to determine the stress intensity factor expressions for Many Investigators have studied the surface flaw in inner and outer surface flaws In cylinders under cylinder. Due to the perceived difficulties of three pressure with the ratio of outer to Inner radius dimensional analysis, the early Investigators assumed equal to 1.461. Alturi used a hybrid-displacement the naw to be continuous ai used two dimensional crack tip element to determine the stress Intensity analysis. Bowie and Freese') studied a continuous factor distribution around the flaw border of semi-longitundinal flaw In a cylinder under pressure. elliptical flaw in pressurized cylinders with the This work was generalized to both continuous ratio of outer to inner radius equal to 1.6 and 2.0.
longitudinal and continuous circumferential flaws in For the same geometries, these direct three-a thick cyl 1er under arbitrary l.4lings by Buchalet dimensional approaches all give similar results and Bamford' ' and Labbens, et al. '. Initial (approximately 10 percent variance).
attempts to approximate solutions for the till three dimensionalgiroblem were made by Underwood'4J and As noted above, the geometries treated in the full Kobayashi.' 0' Neither Underwood or Kobayashi three-dimensional manner are only for limited loading
i conditions and for geometries that are widely Problem Statement different from connercial pressure vessels (ratio for outer radius to inner radius is 1.1). To verify the Shown in Figure 2-1 is a cylindrical pressure vessel structural integrity of these pressure vessels, a containing a buried flaw in the longitudinal considerable range of loading condit)fs were cross-section of the vessel wall. The buried flaw investigated by McGowan and Raymund.'"' They shape is assumed to be elliptical, as shown in Figure assumed the aspect ratio of the flaw to be 6.0 as 2-1, with major and minor axis of 2c and 2a, reconmended by the ASlE Boiler Pressure Vessel respectively. In this Investigation, the aspect Code,Section III, Appendix G.'"w Using boundary ratio of the flaw, defined by the ratio of the major integral equatiop mthod, Heliot. Labbens, and axis to the semi-minor axis f2cIa), is assumed to be Pellissier-Tanontl ' solved the identical problem 6. The flaw is assumed to be located in the and obtained results which agreed well with the longitudinal cross-section of the wall of a cylinder macrpJe]ment results. More recently, Newman and with dimensions representing Pressurized Water Raju'I'0 obtained stress intensity factor solutions Reactor (PWR) vessel beltline. The inside and for a wide range of semi-elliptical surface cracks on thickness are and T. respectively. The ratio of the inside of pressurized cylinders. R%/T is chosen to be 10. The dimensional values chosen for RI and T are 65 in. and 6.5 in.,
Investigations of the fracture mechanics charac- respectively. The stress intensity factor solutions teristics of the buried elliptical flaw have not been are obtained for the buried flaw located in four undertaken as extensive)y as those for surface positions in the longitudinal cross-section shown in flaw. Kobayashi. et alW and Ishida('t' obtained Figure 2-1. The flaw location is defined by the correction factors for a hurled flaw In an Infinitely distance, 4T, measured between the inside surface long strip subjected to membrane stress loading. A of the cylinder and the flaw major axis. The values colpptson of the reults obtained by Kobayashi, et of C chosen for this investigation are 3/16/, 1/4, a and Ishida' shows that the results 1/2 and 3/4. The ligament thickness measured, differ by a wide margin depending on the location of respectively, from the inside and outside surfaces of the crack and the ratio of the flaw minor axis to the the vessel are represented by t. and of.
strip thickness. The largest difference were observed for both larger values of eccentricity The stress intensity factor solutions for the buried (measured in terms of distance between the flaw axis flaw are sought for four different crack opening and the strip center line) and ratio of crack depth pressure loadings, e Cx} defined by:
(minor axMs 1to strip thickness. Shah and Kobayashi obtained in analytical expression for a(X) = AO (2.1) the stress intensity factor for a buried elliptical e x) - Al x (2.2) flaw in an infinite medium subjected to an arbitrary (x )
- A2 x3 (2.3) internal pressure using three-dimensional elasticity a(x) -A 3 x3 (2.4) theory.
9 where AO. Al, A2 are arbitrarily coefficients Naturally, Shah-Kohayahsi(l ) solution did not and x is the coordinate defined in Figure 2-1.
contain curvature effects. The objective of the present analysis is to obtain the stress intensity MACROELEMENT TECHNIQUE factors (or magnification factors) for buried elliptical flaws at various locations in the A detailed formulation of the macroelement technique longitudinal cross-section of a finite thickness developed by Hall, Raymund and Palusamy is contained vessel. The aspect ratio defined by the ratio of the in M20I. The macroelement technique consists of flaw najor axis to seni-minor axis employed in this dividing the flawed three-dimensional structure into analysis is 6.0 as recommended by the ASME Boitr and two or more substructures and modeling the region Pressure Vessel Code,Section III, Appendix G."1) containing the flaw by one or more macroelement The ratio of outer to inner radius of the cylinder is substructures. The solution process begins by 1.1. The stress intensity factor solution is obtaining a condensed stiffness matrix for each of obtained by the Macroelement Technique developd by the substructures followed by global displacement Hall, Raymund and Palusamy. This method have solution. The mode I crack-tip stress intensity lieen shown to give good results for various factors (K1) are then determined from the three-dimensional crqct4 problems as reported by llisplacemeenr solution using Parks stiffness Palusamy and Heliot.' 't The stress intensity derivative 2 method.
factor distribution around the elliptical crack front m
is calculated using the stiffnqll1 derivative The macroelement is built out of 45 microelements of procedure as proposed by Parks." which 37 are blended bricks and 8 are wedge elements, Figure 3.1. The details of these microelements are The physical dimensions of the buried flaws contained in (20], and it suffices to know that the considered in this study is identical to an ellipse wedges have 45 degrees of freedom (d.o.f.), and the with major and minor axes of 3/4 and 1/4 vessel wall d.o.f. of blended bricks can be varied subject to thickness, respectively. Stress intensity factor minor restrictions specified in [20J. The undeformed soliitfans are presented for the flaws located at macroelement shown in Figure 3.2 contains a built-in three-sixteenth, one-quarter, one-half anti quarter-elliptical crack. The region surrounding the three-quarter positions of the vessel wall and crack tip is modeled by a channel of 28 blended subjected to constant, linear, quadratic and cubic hricks enabling the analyst to wary the density of
.'fstrihutions of crack surface loadings. These nodes to achieve a desired combination of accuracy results are then compared to the corresQqding and cost. The total of d.o.f. corresponding to the results obtained by Shah and KobayashiI1'" for choice of minimum density is 1296.
infinite medium. Comparisons are also presented hetieen the present results and those due to the The principal characteristics of the macroelement are AppeneMxA of the Section XI of the ASI1E Boiler the following: it is compatible with the 20-node isoparametric element; it has the option to vary 2
I crack tip region nodal density: it is parametrically displacement solution was obtained. Thirdly, given defined so as to allow curved faces: it permits a the global displacement solution, local displacement wide variety of crack surface loadings (any combi- solution was obtained for each of the microelements.
nation of terms of a bivariable cubic polynominal), Finally Parks' [22] stiffness derivative technique and it significantly reduces the man-time needed to was applied to determine the stress intensity factor formulate the finite element model. at each of the crack tip nodes. This process was carried out for each of the load components FINITE ELEMENT MODEL individually and the stress intensity factor were determined which were then combined based on The finite element model of the reactor vessel superposition principle to obtain the stress beltline containing the buried flaw consisted of a intensity factor solution for a combined loading.
45' sector of the beltline of length 69 inches and of thickness 6.5 inches. Taking advantage of the Experience shows that a third degree polynominal sytnetry condition, only one half of the flaw was obtained by the summation of loadings defined by modeled. The crack tip region was modeled by two Equation (2.1-4) is adequate to represent the stress macroelements and each one was treated as one distribition in the cross-section of a PWR heltline substructure. The remaining region called mother under all operatin conditions of loadings including structure was modeled using 20-node isoparametric postulated accident condition loading. Therefore the brick and 15-node isoparametric wedge elements. A crack opening hoop stress a can he represented as typical finite element model (for t = 1/2) is shown follows:
in Figure 4-1.
a = AO + Alx + A2x2 + A3x3 (5.1 I The details of finite element models 1 through 4 corresponding to g equal to 3/16/. 1/4, 1/2 and The resulting stress intensity factor KI is then 3/4, respectively. are presented in Table 4.1 A expressed using the principle of superposition as typical model used for the buried flaw located at follows:
( - 1/2 is shown in Figure 4.1. This model consists of 7 layers of elements each in the KI( ) - KID(*) + KI(l() + K12(.) + K13(') (5.2) circumferential and axial directions. The first layer defined by 9
- ec contains the macro- where KID, KIl, KIz and Ki3 are the stress elements. The dimensions of the macroelement in the intensity factor val ued due to each of the loadings.
axial and radial directions are denoted by Lc and It is convenient to express KI as a function of ic. The values of Wc, Lc and oc were elliptical angle e in the following familiar form:
chosen to be 3 inches, 1.791 and 2.031 inches, respectively, for all four locations. For the case of X equal to 0.5, the value of Wc was chosen to Kill) M(*,ac)[AdHO(#) + .' A H (, (5. 3) be 0.609.
As reported previously each of the macroelement is a2 4,ra 3 treated as a separate substructure and contained 45 microelements. The combined macroelement degrees of 2 2 H2 ( ) +-y- A3 H3 (o)J A
freedom varied from 3045 to 3384. The number of elements in the mother structure (the third where the function, (4. a, c) represents the structure) ranged from 173 to 194. One of the stress intensity factor factor for a buried features of the Macroelement Technique is that the elliptical flaw of major axis, 2c and minor axis, 2a number of crack tips nodes can be chosen at will in an infinite medium subjected to a uniform crack subjected to minor restrictions specified in opening stress of unit magnitude and is defined by Reference 20. In this investigation each of the (29]
macroelements had 25 crack tip nodes.
The first verification of this method was made with reference to a plate containing a semielliptical surface flaw of aspect ratio. 5. fractional W~e ,a.c) U a (cos2 + 4 sin #)/4 (5. 4a) through-wall depth, 0.6, and subjected to remote uniform tension loading. It has been shown that the V/-7 KI values obtained by the macroelemet technique 4
agreed with fih)e of RaJu and Newman- Z i and Smith Q .o (cos+ + ;a sin2s )l /: do (5. 4h)
J within 3 and 8 percent, - 0 C and Sorensen respectively. Recently, the macroelement technique has been successfully applied to the Battelle bench The quantities Ho({) , N(e), H2(.) and H3(4) are mark problem (26) on surface cracked plate (27] magnification factors of position along the crack semielliptical surface cracks in a cylinder (l31 and front and are obtained by comparing Equations (5-2) a single quarter-circular corner cracked hole in a and (5-3) as follows:
plate '283.
DETERMINATION OF STRESS INTENSITY FACTOR SOLUTION NO($ ) d~~$M+ ,C} (S. 5a)
The stress intensity factor determination using K01(4) nacroelement is carried out in essentially four steps. First, the condensed stiffness matrices were obtained for each of the three substructures. l - Al M ___ __ __ (5.5b)
Secondly, these three substructures were connected, w Al M (e a~r.)
the prescribed loads were applied and a global 3
Shah and Kobayashi (19). Specifically, for this K224) (5.Sc) comparison macroelement results due to g equal to HS (* ) * ; 1/2 was used. The values at q equal to 0' and 180'
-r A2 M (e ,a~c) are lower than those due to the infinite medium values. The deviations range from 8 to 16 percent lower. Another observation is that the macroelement results are symnetric with respect to the major axis
- X (1) of the (flaw) ellipse within about 2 percent. Since 13 no mesh convergence studies were made as it is C5. 5d) required in any finite element analysis for H !,) establishing the accuracy of the numerical results.
.1 3 the question remains whether the lower values 42*
A 11(oac) obtained for the cylinder geometry is either due to real effects arising from either curvature and free 3 3} b)oundary conditions or inadequate finite element modeling. In general, mesh convergence studies using macroelement are expensive and therefore it is not A diagranatic description of the load components and performed. However, a number of solutions by this the associated magnification factors are shown in technique have been compared with results obtained by Figure 5-1. other investigators (20,13,21,28]. These comparisons show that the macroelement results can be lower by STRESS INTENSITY FACTOR RESULTS about 5 to 10 percent.
Iagnification factors were obtained for 6 to 1 aspect In order to quantify the discrepancy the macroelement i ratio (3 to 1 major to minor axis ratio) buried technique was applied to the buried flaw tn an elliptical flaws located at .
- 3/16, 1/4. 1/2 and infinite medium. All the four loading cases were 3/4. The results obtained for the magnification analyzed. The resulting magnification factors are factors Hn, Hi, H2 and H3 are presented in compared with the solution due to Shah and Kobayashi Figures 6-1 through 6-4. The values of magnification r19] in Figures 6-6 through 6-9. The values of Ho.
factors vary as a function of elliptical angle. S. HI, H2 and H3 at # equal to 0 90 and 180 The positions defined hy o equal to 0 and 180' are degrees due to the macroelement method are compared the farthest and nearest locations on the crack minor with those due to Shah-Kobayashi method in Table axis iith respect to the inside surface of the vessel 6-3. The deviations range from -16 percent for H3 wall. For all positions the factor Ho is the least to 13 percent for HI. It should be noted that varying with respect to o and the value is about these deviations occur at
- equal to 0 or 180 unity. The fact that the value of Hf is about unit degrees whereas the deviations corresponding to 90 suggests that the influence due to curvature and degrees is 2 percent for Ho)and zero percent for Foundary surface nay not be very significant for H1, H2 and H3. In general. whereas the these cases. deviation at 90 degree location is consistent with previous macroelement results (13, 20, 28), the A comparison of the values of respective deviations at 0 and 180 degrees are somewhat larger magnification factors due to various positions shows than expected. It Is believed that the larger than that the respective values for positions defined by expected deviation is due to inaccurate mapping of
- 1/4. 1/2 and 3/4 are the same within 2 microelements in these locations.
percent. For the positions C equal to 3/16, the values of Ho are about 7. 0 and 12 percent higher In Figure 6.10. the membrane correction factors ate equal to O0, 90' and 180' than the respective obtained for a cylinder of radius to thickness ratio values due to other positions. The value of H2 due 10 are compared with the ASME Section Xi [23] values to c equal to 3/16 are found to he about 71 0 and '1 derived from a plate solution. The comparison shows percent at . equal to 0'. 9' and 180' than the that the ASME Section XI values are conservative for respective values due to other positions. A similar the range specified in the code 12e/T - 0.6, Figure but a smaller Increase in the values of Hi and H3 6.10) and the extrapolation of this value beyond the due to X equal to 3/16 are found and these percentage Increases are summarized in Table 6-1. specified limit may not be conservative.
SUMMARY
AND CONCLUSIONS Since there was virtually no difference in the results obtained for positions e equal to 1/4, 1/2 Elastic stress intensity factor solutions were and 3/4, question was raised whether there will he obtained for a buried elliptical fnaw of aspect ratio any difference between this set of results and those 6 (major to minor axis ratio 3), located in the for a similar flaw in an infinite medium. Shah and longitudinal plane of a cylinder of radius to Kobayashi f119 have reported a general stress thickness ratio 10. The flaws were located at intensity factor solution for a buried elliptical distances of 3/16. 1/4, 1/2 and 3/4 times the wall flaw in an infinite medium and subjected to a thickness from the inside of cylinder wall. The two-dimensional third order crack surface loadings. macroelement technique of the three-dimensional finite element analysis was used in obtaining the From this closed form solution, the values for the numerical results. The loading on the crack surface the magnification factors Ho, Kl N2 and H3 consisted of a crack opening pressure field were derived for a 6 to 1 aspect ratio flaw buried In represented by a one-dimensional third order an infinite medium, and are presented in Figure 6.5. polynomial. The accuracy of the Macroelement results Comparisons are shown in Table 6-2 for X equal to as well As the effect of curvature and free boundary 0', 90' and 180'. The last column of this table shows the ratio of the macroelement calculated values were investigated by comparing the results of the for the cylinder geometry over the closed form macroelement three-dimensional finite element 14 results obtained for an infinite elastic solid by analysis with the closed form results obtained for an II.
4
, II
identical flaw located in an infinite elastic medium High Stress Gradients", Significance of Defects and subjected to identical loading conditions. The in Welded Structures (ed. T. Kanazawa and A. S.
results of this evaluation investigation lead to the Kobayashi). pp. 127-43, University of Tokyv following conclusions: Press, 1974.
- 1. A comparison of magnification factors for the 5. Kobayashi, A. S., Polvanich, H., Emery. A. F.,
Cylinder and infinite elastic medium shows that, and Love, W. J., 'Stress Intensity Factor of a for flaws located within 25 to 75 percent of the Surface Crack in a Pressurized Cylinder",
wall thickness, the stress intensity factors Computation al Fracture Mechanics, June, 1975, agree within + 3 percent. Therefore in this region the Infinite elastic medium solution can 7. Kobayashi, A. S., Emery. A. F. Polvanich, N.. and be used for engineering fracture mechanics Love, W4.J., "Surface Flaw in an Pressurized and evaluation of cylindrical vessels of comparable Thermally Shocked Hollow Cylinder', Int. J. Press geometry. Ves. and Piping. Vol. 5, op. 103-172, 1977.
- 2. For flaws located outside of the region defined 8. Kobayahsi, A. S., A. F., Polvanich, N., and Love, by 25 to 75 percent wall thickness, the stress W. J., "Inner and Outer Surface Cracks in intensity factor appears to increase at 0 and 180 Internally Pressurized Cylinders", J. Press. Yes.
degree locations as the flaw gets closer to the Tech., ASME pp. 83-89, Feb. 1977.
surface. For the one case analyzed where the flaw was considered to be located at a distance 9. Ayers. 0. J., "Three-Dimensional Elastic Analysis of 18.7 percent of wall thickness from the of Semi-Elliptical Surface Flaws Subjected to inside, the magnification factors were higher by Thermal Shock", Computational Fracture Mechanics, percentages ranging from 2 to 12 percent. No ASME, led. E. F. Rybicki and S. E. Benzley), pp. S increase was observed at the 90 degree location. 133-143, 1975.
- 3. A comparison between macroelement and closed from 10. Blackburn, 14. S. and Hellen, T. K., 'Calculation results from a buried elliptical flaw in an of Stress Intensity Factors for Elliptical an-'
infinite elastic solid medium confirms the Semi-Elliptical Cracks in Blocks and Cylinders",
previously published conclusions that the Central Electricity Generating Board Report No.
macroelement method can give results that are RD/B/N3103. July 1974.
accurate within 5 to 10 percent at 90 degree locaton. However, the inaccuracies at 0 and 180 11. Atluri, S. N., Kathiresan, K., Kobayashi, A. S.,
degree location have been observed to be as large and Hakagaki, M., "Inner Surface Cracks in an as 15 percent. Internally Pressurized Cylinder Analyzed by a Three-Dimensional Displacement-Hybrid Finite
- 4. A comparison between the ASHE Section XI membrane Element Method", Proceedings of the Third correction factors and the macroelement International Conference on Pressure Vessel correction factors shows that the former values Technology (Tokyo, Japan, April 19-22, 1977),
are conservative for the eccentricity limit ASME, pD. 527-53.
specified in the code and that the extrapolation of the current code values may not be 12. Atluri, S. A., and Kathiresan, K., "Outer and conservative. Inner Surface Flaws in Thick-Walled Pressure Vessels", paper G 5/4 Transactions of the Fourth REFERENCES International Conference on Structural Mechanics in Reactor Technology. San Francisco. Cal. 1977.
- 1. Bowie, 0. L. and Freese, C. E., "Elastic Analysis for a Radial Crack in a Circular Ring", 13. McGowan, J. J. and Raymund, M., "Stress Intensity Engineering Fracture Mechanics, Vol. 4, No. 2, Factor Solutions for Internal Longitudinal pp. 315-322, June 1972. Semi-elliptical Surface Flaws in a Cylinder Under Arbitrary Loading", STP-677, ASTR4, Philadelphia,
- 2. Buchalet, C. 8. and Samford, W. H.. "Stress 1979, pp. 365-380.
Intensity Facotr Solutions for Continuous Surface Flaws in Reactor Pressure Vessels", Mechanics of 14. ASME Boiler and Pressure Vessel Code, Section Crack Growth, ASTM-STP-490, pp. 385-402, American III, App. G "Protection Against Nonductile Society for Testing and Materials. Philadelphia, Failure", 1980.
1976.
- 15. Heliot, J., Labbens, R. C. and Pellissier-Tanon,
- 3. Labbens, R., Pellissier-Tanon, A., and Heliot, A., "Semi-elliptical Cracks in a Cylinder J., "Practical Method for Calculating Subjected to Stress Gradients', STP-677, ASTMh Stress-Intensity Factors Through Weight 1978, pp. 341-364.
Functions", Mechanics of Crack Growth, ASTM-STP-590, pp. 368-384, American Society for 16. Newman, Jr., J. C., and RaJu, I. S.,
Testing and Materials, Philadelphia. 1976. "Stress-Intensity Factors for Internal Surface Cracks in Cylindrical Pressure Vessels", Trans.
4.' Underwodd. J. H., 'Stress Intensity Factors for ASME, J. of Pressure Vessel Tech., 102, 1980, pp.
Internally Pressurized Thick-Walled Cylinders", 342-346.
Stress Analysis and Growth of Cracks, 17. Kobayashi, A. S., Ziv, M. and Hall, L. R.,
ASTM-STP-513, pp. 59-70. American Society for "Approximate Stress Intensity Factor for an Testing and Materials, Philadelphia, 1972. Embedded Elliptical Crack Near Two Parallel Free
- 5. Kobayahsi, A. S., "A Simple Procedure for Surfaces" International Journal of Fracture Mechanics Vol. 1, 1965, p. 81-95.
Estimating Stress Itensity Factors in Regions of 5
A
- 18. Ishida, M., "Stress Intensity Factors for the Kobayashi (19] with those obtained in this report.
Tension of an Eccentrically Cracked Stripe, Journal of Applied Mechanics. Sept. 1966, p. The normal stress ez at I - 0 which is 674-675. perpendicular to the crack plane x and y is expressed as:
- 19. Shah, R. C. and Kobayashi, A. S., "Stress Intensity Factor for an Elliptical Crack Under Arbitrary Loading', Engineering Fracture Mechanics, Vol. 3, 1971, P. 71-96.
- 20. Hall, C. A., Raymund, M. and Palusamy, S. S., A
-o CZ Izzo 2G = 1 nMI nuo AN(Cj)ym n(A-I ) I Macroelement Approach to Computing Stress Intensity Factor for Three-Dimensional Structures", Int. J. Frac.. Vol. 15. No. 3, 1979. if x O thena = a.
1A aa 0A y pp. 231-245.
- 21. Palusamy, S. S. and Heliot, J.. Two iihere iis stress function-and Stress-Intensity Factor Calculation Methods and Solutions for Various Three-Dimensional Crack x one of the ellipsoidal coordinate system Problems, presented in the 5th Intl. Conf. on p(x,y) a arbitrary loading in x-y plane Frac., CANNES, France, 1981. G - shear modulus of a medium x " maJor axis
- 22. Park. D. M., "A Stiffness Derivative Finite y
- minor axis Element Technique for Determination of Crack Tip Stress Intensity Factors", International Journal The stress function 4 is expressed in terms of the of Fracture, Vol. 10, 1974, pp. 487-502. higher derivatives of the potentia) fYnction with rIspecl to x and y which is known. I Therefore
- 23. ASME, Boiler and Pressure Vessel Code, Section a '/az Ix a 0 Is obtained through the XI, Appendix A, 1977 Edition. algebraic operations.
I equation (A-1) the unknown constants Cjj in
- 24. Raju, I. S. and Newman, J. C., Jr., "Stress ahz Ix O are obtained by equating the Intensity Factors for a W1ide Range of cQefficzents of ym which are Aom and thos eof ym in Semi-Elliptical Surface Cracks in Finite a hz2 1X0. A more detailed derivation and Thickness Plates, Engineering Fracture Mechanics, explanation is given in reference [19],
Vol. 11, 1979, pp. 817-829.
The coefficient, A00 in equation (A-1) is equal to
- 25. Smith, F W., and Sorensen, D. R.. "The that in a32/jzl IxO and it is given as:
Semi-Elliptical Surface Crack - A Solution by the Alternating Method," International J. of Fracture, Vol. 12, 1976, pp. 47-67. Also see 316 C02 (A-2)
Colorado State Univ. Tech. Report 4, NASA Grant = Aoo 0 11 COO + 14 C20 +
NGL-06-002-063, 1973.
The coefficient, AlO is equation (A-l ) is:
- 26. Hulbert, L. E., "Benchtark Problems for 3-0 Fracture Analysis". Int. Jl. of Fracture, 13, 1977. pp. 87-91. I CI2 (A-3)
'K AI0 e 622 CIO * £27 C3 0 ' "2
- 0
- 27. McGowan, J. J. and Raymund, J. J., "Stress Intensity Factor Solutions for Surface Flaws in Repeating the above procedure, the following inite Thickness Plates under Arbitrary Loading simultaneous equations are established:
WCAP °31R, Westinghouse Electric Corporation.
Pittsburgh, PA April 1978. I/2G A01 a £33 CO1 +438 C21 3 C0 3 310 2R. Palusamy, S. S. and Raymund, nl., "Stress l/2G A20 °
- 44 C20 + 46 C02 0 (A-4)
Intensity Factor of Corner Crack at the Edge of a Hole in a Plate", STP 743, ASTM, Philadelphia, l/2G All °0 55 Cl1 PA 1981.
1/2G A02 £ 664 C20 + 66 C0 2
- 79. Sadowsky. M. A., "Stress Concentration Around a I
Triaxial Ellipsoidal Cavity", Trans. ASME J. of A2 1/2G a 0 a*88 C21 + 6810 C03 Aop. Mech.. 1949.
Al 2 /2G 0 a 897 C30 + £99 C12 APPENDIX I A03/2G aloe C21 + £1010 C03 STRESS INTENSITY FACTOR SOLUTION FOR ONE DIMENSIONAL STRESS A30/2G 0 £77 C30 + 679 C1 2 DISTRIBUTION OBTAINED FROM SHAH-KOBAYASHI RESULTS where .ai are known and are defined by equations This section discusses the derivation of the reduced (A-7) mn6 (A-8 to follow.
stress intensity factor expression obtained from the general solution presented by Shah-Kobayashi [19]. From equations (A-2), (A-3). and (A-4), C can be The reduced stress intensity factor expression was expressed in terms of Aij and *ij by solvigg the required to compare the results obtained by Shah and simultaneous equations. The solutions of Cij are:
6
4 8 ) ECK) - 2K C1 2 a CIO ' °
'44 j7p2r2U3 CIl ' C3 0 U (2 I ) K(K)]
C0 0 1 ALoo - 16 644 - '46 "14 h +
K R all *11T;66 64 a4r6 12
£EK)3 46AK 7 7 4 K K' TK 4 (2 - K2)K(K) - 2(K2 K4 + K4 C)
"6A 0 _ 03 a88 '310 G5'810 "a38 CO Is I (A-7) 533 '33"1010 '88 - 010e8lo) 2 066
- 8 (2(3K2-1)K(K) + (3K2 + 2-70K C 1 A0 2 44 7 7 4K4 ATK K-i~
C02 ' '66 44 ' 64 46 (2 - K )K(K) - 2(K 2 + K 4) E(K))
2
'64 8 rK C1 A0 2 , 4 6 (A-5)
C20 "M a66 '44 - '64 '46 2
A0 3 0 81 0 (A-51 533" 35'T4 E(l + K )E(K) - K' KWK)
-1 C
c12 It
-w a1010 88 1080810 4 2 . 1(2 A9 T K9 K I8K K 2+ 3K K 2+ 2K-8)E(K) 6 88 (A-5)
C0 3 1 A0 3 a88 I'cn1010 '188 '108 '810 (4K2Ks 2+ SK 2+ 3)K(K))
is given by 2-I0K2 The stress intensity factor _______2 EK]
'310 It('4K4 [2(3K - )K(K) + O3K 1(2,+ K 2)
AYTK'K K, ,1B1/2 2 2 2 O20 1/4 K = ;;T (11) (a sin a + b cos o) 038 A/T 7K K 4 r2(K + K4 )E(K) - K 2 (2K2) KWK]
420 cos s 2
_ cos2e sine 01 sine "+I 2Kk8)K(K)+(6K4-3K'+I --
n 2)E(K]
a ab tl80 sAgg~rt3K2K
'810"A99'KK 4C02 sin2a (A-6)
K 3
4C03 sid e 2 2 2
_45K K +8)K2K(K)
't101o " 6(l6K1 and minor axis of the (40_40K2-PS2 K 2+ISK2-K '4)E(K)]
%)here a and b are the major and e is the elliptical crack, respectively, In the counter elliptical angle which is positive
- 0 zero on the major 8 4 4_32+1 16 -2) E(K)+
clockwise direction and e 108 3 9 K 6(6K4 6 _
are expressed In terms of,tj the axis. .j are elliptca: a and b. The quantities 2 2 2 K 1 2K -8)KcK )W follows: (3K defined as where K'
- bat AI a Ks a b/a (A-8)
- E1K T A - atl K2AAt3 K2
- I - K'2 of the K(K)
- the complete elliptical integral 2
(K K(K) - (l1-2K )E(K)] second kind. of the
'14 = a E(K)
- the complete elliptical Integral first kind.
16 8 K2) E(K) - K K(K)I 4
A5 T5 K-'K 7
.4 Ia FIgure 3. Undeformed Macroelemnent Figure 1. Cylindrirat Pressure Vessel Containing eBuried FIsN 2
(at 4S dAtf. WEDGE (b) VARIABLE dAJ. BLENDED BRICK Figure 2. Micro Elements User' to Deign VW MaCro Element e Figure 4. Finite Element Model of Cylinder with a Longitudinal Flow, .- 1/2
00 I
v i
3 Figure 13. Comparison Of M2 Due to Mocroelament and Shah-Kobayeshi I (19) Methods for e Buried Flaw in an Infinite Elastic Solid AOLE W CLLIPFICAL Figure 10. Magnitication Factors foe a Burled Flaw in an Infinite Medium 5..MICSaVAIOI I
's - f1 l.8
- ,*b..- .*. . , .,n* ,* *0*.. "- ...
. --. ',IL,,._
0.
Figure 14. Comparison of H3 Due to Macroelement and Shah-Kobeyashi (19) Methods for a Buried Flow in an Infinite elastic Solid Figure 11. Comparlion of Ho Due to Macroelementn end Shah-Kobayashi (191 Methods for a Buried Flaw in an Infinite Elastic Solid
., ALW I
- , I X\-- S 110 a a so IS IS Figure 12. Comparison of HI Due to Macroelement end Shilh-Kobayashi FLAW ECCENTRICITY RATIO T (19) Methods for . Buried Flaw in on Infinite Elastic Solid Figure 16 Companison of Membrane Strem Correction Factors Due to ASME Code end McroleWment Method 9
I, y
I Is -j 2
SUPTCAL nCILS61 Figure 7. Magnification Factors for e Buried Elliptical Flaw Location at r- 1/4 HI (0)
Al I
H2 (
I H3 (0
-A3 *MOTOCAL AGMLI Figure L Magnification Factors for a Buried Flow Located at r - 1/2 Figure 5. Distribution of Loading Components and the Associated Magnifica-tion Factogs a
2 _I a
I a I v
t C
i Ift I
(
I al a.
LLJdrLUJ
[~f7 .5 a'..
sa.
I I I I g l I .
_
- l l e I
, 0
- to _i is oi us = iO IN lal III ' ELOTICALAWGI Figure 9. Magnification Factors for a Buried Flaw Located at r - 3/4 Figure . Magnilecation Factor, lor ABuried Flaw Located at r
- 3/16 10
f ENCLOSURE 2 COVER LETTER TRANSMITTING NRC SER
9
/ 1. UNITED STATES I NUCLEAR REGULATORY COMMISSION /.
WASHINGTON. L.C.
20655 DEC22 19G Cb Docket Nos. 50-282 and 50-306 Mr. D. M. Musolf, Manager Nuclear Support Services Northern States Power Company 414 Nicollet Mall Midland Square, 4th Floor Minneapolis, Minnesota 55401
Dear Mr. Musolf:
The Commission has completed the review of Northern States Power Company's (the licensee) request for an exemption to allow the application of the
-leak-before-break" technology as a basis for the elimination of protective devices (i.e., pipe whip restraints, jet impingement barriers, and other related changes) of the primary reactor coolant systems at the Prairie Island Nuclear Generating Plant Unit Nos. 1 and 2. These protective devices were Installed to mitigate the dynamic effects resulting from postulated large pipe ruptures. The technical information was provided by the licensee's letters dated October 24, 1984, October 21, and November 5, 1985 and supplemented by letter dated September 10, 1986 in response to staff concerns.
On April 11, 1986 a final rule was published in the Federal Reqister (51 FR 12502) amending 10 CFR Part SO, Appendix A, General Design CrTter a (GDC) 4 that became effective on Pay 12, 1986. The revision of CDC 4 allows the usE of analyses to eliminate from the design basis the dynamic effects of postulated pipe ruptures in the primary coolant system. The staff has completed the review of the licensee's submittals and concludes that the analysis of piping of primary coolant systems at the Prairie Island Nuclear Generating Plant Unit HNs. I and 2 is adequate and demonstrates compliance with the GDC 4 as amended. Therefore, an exemption to GDC 4 of Appendix A of 10 CFR Part 50 as amended that was requested by the licensee prior to the of effective date of the rule Is not necessary. On this basts, the removal pipe whip restraints, jet impingement barriers, and other associated plant hardware may be implemented at your convenience. Our safety evaluation addressing this matter is enclosed.
%i . C
.2.
This action completes our work effort under TAC Nos. 08731 and 08732.
Sincerely, Dominic C. Ditanni, Project Manager Project Directorate #1 Division of PWR Licensing-A
Enclosure:
Safety Evaluation cc's: See Next Page