ML20138H543

Forwards Response to 970403 RAI Re Pressurizer Thermal Transient on SONGS Unit 2

ML20138H543
Person / Time
Site:	San Onofre
Issue date:	05/01/1997
From:	Rainsberry J SOUTHERN CALIFORNIA EDISON CO.
To:	NRC OFFICE OF INFORMATION RESOURCES MANAGEMENT (IRM)
Shared Package
ML20138H550	List: ML20138H543 ML20138H557 ML20138H566
References
TAC-M98232, NUDOCS 9705070218
Download: ML20138H543 (17)

Text

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s- momess EDISON Manager, Plant I.icensing An IDISON lYllR%4710%4t

• Ctanpuny May 1, 1997 U. S. Nuclear Regulatory Commission Attention: Document Control Desk Washington, D.C. 20555 Gentlemen:

Subject:

Docket No. 50-361 Pressurizer Thermal Transient San Onofre Unit 2 (TAC WO. M98232)

Reference:

April 3, 1997 letter from Mel B. Fields (NRC) to Harold B. Ray i (Edison),

Subject:

Request for Information Regarding the Pressurizer Thermal Transient on San Onofre Nuclear Generating Station Unit 2 (TAC NO. M98232)

Enclosed is Southern California Edison's (Edison's) response to the referenced NRC request for information regarding the March 4,1997 pressurizer thermal trainsient on San Onofre Unit 2. The evaluation performed by Edison demonstrates that the pressurizer remained within the acceptance criteria of Section III, Appendix G of the ASME Code, 1989 Edition, no Addenda.

Please let me know if you have any questions or would like additional information.

d Very truly yours,

\) ~

9705070218 970501 ,

PDR ADOCK 05000361 P PDR s Enclosure 670001  !

cc: E. W. Merschoff, Regional Administrator, NRC Region IV K. E. Perkins, Jr., Director, Walnut Creek Field Office, NRC Region IV J. A. Sloan, NRC Senior Resident Inspector, San Onofre Units 2 & 3 M. B. Fields, NRC Project Manager, San Onofre Units 2 and 3 i ,

5a re Nuc! car Generating Stution San Clemente, CA 92674-0128 714. % 8-7420

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, ENCLOSURE PRESSURIZER STRUCTURAL INTEGRITY ANALYSIS l SAN ONOFRE UNIT 2 Question 1: Using the reported reference temperature (RTNDT) of 60 F for the surge

, nonle forging, a 1/4 thickness (1/4T) deep flaw and the thermal transient i that represents the cooldown of the surge nonte, provide the following information to demonstrate that the surge nonle will meet the Appendix G t

criteria for the entire transient: .

1 (a) the values of the primary membrane stresses due to pressure Response: The primary memorane stress due to pressure at the critical nonle 1 l section 1-1 selected for the evaluation, opt , , is l opres = PR/T

= 1.47 ksi i

where 1

P = pressure

= 0.35 ksi R = radius

= 5.565" T = thickness

= 1.327" The nonle section is shown in Figures 1 and 4 as Section 1-1. An additional sect on at the base of the nonle, Section 2-2, was also evaluated, and it was concluded that Section 1-1 is more limiting.

(b) the values of the thermal stress due to the thermal gradient, Response: The finite element method was used to perform the thermal transient ans!ysis of the nonle due to a unit load thermal shock of 1000*F on the inside surface of the nonle. A scale factor of 0.254 was applied to the calculated stresses to reflect the actual accelerated temperature drop during the ,

cooldown transient (254*F). Figure 2a shows a plot of peak hoop stress acting on the inside surface of the nonle at Section 1-1 as a function of time, and Figure 2b represents the first 40 seconds of the same plot. The figure shows that the-maximum peak stress of 88.5 ksi on the inside surface of the nonle j occurs at less than two seconds from the onset of the transient.

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(c ) the values of the membrane stress intensity factor Response: The following value of Kpi was calculated K, = 1.75 ksi{id For more details, see response to Question 2 below.

(d) the values of the thermal stress intensity factor Response: The following value of Krr was calculated Kg = 62.36 kvi{id For more details, see response to Question 2 below.

(e) the temperatures of the surge nozzle at the 1/4T location ,

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Response: The temperature time history at the T/4 location, on Section 1-1, is  !

given in Table 1, and plotted in Figure 3. This time history is conservatively I based on a thermal shock from 430*F to 110'F on the inside surface of the l nozzle. The actual rapid drop in the surge line temperature is only 254*F (from  !'

429'F to 176'F). This drop occurs at a linear rate of approximately 25'F/ min.

Using a step temperature change of 320*F to calculate the crack tip temperature j is, therefore, conservative in terms of both magnitude and rate of cooling.

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Table 1 Temperature Time History at the T/4 location i

time temperature time temperature i

(seconds) ('F) (seconds) (*F) 0.001 430.00 2.943 412.44 0.004 430.00 3.660 405.68 0.010 430.00 4.537 397.82 0.045 430.00 5.558 389.48 0.078 430.00 6.723 381.02 0.121 430.00 8.081 372.38 0.177 429.99 9.698 363.46 0.255 429.97 11.637 354.28 0.445 429.81 13.971 344.86 0.536 429.67 16.851 335.06 0.772 429.08 20.547 324.59 1.135 427.48 25.543 313.01 1.417 425.73 31.457 301.77 1.837 422.53 36.000 294.40 2.337 418.14 (f) the values of the reference stress intensity factor (Kla) at the 1/4T location Response: Per Appendix G, the following requirement must be satisfied:

2 K i p + Kg < K ia j where i Kpi = primary stress intensity factor ,

Kg = secondary stress intensity factor Kai = reference (allowable) stress intensity factor  !

The material stress intensity allowable, K i g, is given by i

K, = 26.78 + 1.223eu w - 47 ,. i m p ,j gg The maximum peak stress occurs at less than two seconds from the onset of the

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transient, and the corresponding temperature at the T/4 location is greater than 400*F. For T>250*F and RTuor = 60*F, Kgi > 200 ksi (in)'d is calculated which provides a margin of safety several times greater than required by the code.

The value of RTuor = 60*F was estimated based on the acceptable estimation methods in NRC Branch Technical Position MTEB 5-2 for fracture toughness requirements.

Figure 2 shows that the stress decreases rapidly with time, which justifies the use of the upper shelf fracture toughness to evaluate the T/4 crack depth at the time of maximum stress. Later in the transient, the stress intensity decreases as the level decreases significantly.

Question 2: Describe the stress analysis and the method of converting the stresses into stress intensity factors.

Response: The finite element method, using the general purpose finite element program ANSYS, was used to perform the stress analysis part of the evaluation. An axisymmetric model, shown in Figure 4, was generated which includes the nozzle forging, the stainless steel cladding, the nozzle safe end, and the thermal sleeve. The model also included the lower head of the pressurizer vessel and part of the pressurizer surge line. The thermal transient analysis is a calculation of the temperature time history due to a step change in temperature on the inside surface of the nozzle assembly of 1,000*F. The actual 254*F drop in temperature during the March 4,1997 transient occurred over several minutes. Using a step change in temperature is, therefore, conservative.

The thermal-stress analysis, based on results of the thermal transient analysis, calculated the time history of the stress at the specified locations in the nozzle due to the step change in temperature. A scale factor of 0.254 was applied to the calculated stresses to represent the actu ! 254*F rapid drop in the temperature (from 429'F to 176*F) during the March 4 transient because the thermal transient analysis was performed for a 1,000*F step change. The stress distribution in the nozzle ' wall was calculated at several locations, and the maximum peak stress was used to perform the fracture mechanics evaluation.

The primary stress intensity factor was calculated in accordance with Appendix G as follows:

K, = a,,, = Mm

i 5

The hoop stress due to applied pressure, o g ,,, is o,g = PR/T

= 1.47 ksi where P = pressure

= 0.35 ksi R = radius

= 5.565" T = thickness

= 1.327" The parameter Q=1.033 was calcult '.ed per A-3300-1 using the following input:

a = 0.25T

, = 0.332" all = 1/6 i a/T = 0.25 The factor Mm=1.188 was calculated per A-3300-3, and the factor Mb=0.803 was calculated per A-3300-5. It follows that X,, = t.75 ksi{in The thermal stress intensity factor was calculated as follows:

The thermal stress was calculated using the finite element method. A second order curve was used to represent the stress distribution through the nozzle wall 2

o = 88.46 -201.1 x +98.71 x Using this curve, the stress at the tip of the T/4 crack (x=0.33175") was calculated to be 32.61 ksi. A straight line was drawn between the stress at the crack tip and the surface stress and extended to the T/2 location in the nozzle. The stress at the T/2 location was defined to be the membrane stress, omem, and the difference between the surface stress and the membrane stress was defined to be the bending stress,0 3,ne.

Kg= o,,,,,,,

• Mm + as ,,,

• M hG $ G

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The following membrane and bending stresses were calculated:

om.m = -23.24 ksi ow = 111.70 ksi lt follows that l

K,,. = 62.36 ksi5 j Finally, the stress intensity factor, Kj, for a T/4 crack is given by 6 =2%p+%7  ;

i Therefore, K, = 2(1.75) + 62.36 = 65.86 ksiS 1

The finite element analysis results show that the maximum peak stress occurs at time  !

(t) < 2 seconds. From Table 1, the temperature at the tip of a T/4 crack at t=1.8 seconds is over 400*F.  !

Question 3: Provide the basis for the conclusion that the deepest postulated flaw .

should be a 1/10T. As justification for this assumption provide 9,e l following information:

(a) the volume of the surge nozzle inspected i

Response: The required examination volume on the pressurizer surge nozzle in Unit 2 was ultrasonically examined in accordance with the NRC approved San Onofre Nuclear Generating Station (SONGS) second interval inservice Inspection Program (ISI) program. This coverage was achieved using 0,45, and 60 degree angles per SONGS procedure SO23-XXVil Rev.1. Figure 5 is a  ;

sketch of the pressurizer surge nozzle volume examined.

(b) the method of inspection of the surge nozzle Response: A volumetric ultrasonic examination was performed using SONGS procedures SO23-XXVil-20.66 Rev.1, " Ultrasonic Examination of Vessel Welds l and Adjacent Base Metal," and SO23-XXVil-20.52 Rev. O, " Ultrasonic W

7 Examination of Nozzle inner Radius Areas," on the pressurizer surge nozzle.

Beams were used at 0,45, and 60 degree angles. A copy of the procedures is provided in Attachment 1.

(c) the results from the qualification demonstration for the inspectors, the equipment, and the procedures Response: The examiners qualify their procedures and equipment by successful calibration in accordance with code requirements using a code calibration block. For the surge nozzle examination a 0,45, and 60 Jagree calibration was performed using the procedures in 3(b) above on calibration block UT-4. Before a system calibration is performed a vertical linearity is verified per procedure. A calibration is performed on a complete system prior to use of the system in the thickness range under examination, and includes instrument, search unit, search unit cable, search unit wedge, couplant, and recorder, if used. A calibration block UT-4 drawing and pressurizer w' eld designation drawing are provided in Attachment 2, equipment and material certifications are provided in Attachment 3, and Ultrasonic Testing and Penetrant examination reports are provided in Attachment 4..

(d) the probability of detection (POD) of a 1/10T flaw based on the qualification of the inspectors, equipment and procedures Response: A formal probability of detection (POD) was not prepared. Instead, the requirements of ASME Code Sections V and XI, including Appendix Vill, were used to provide a reliable inspection technique capable of detecting flaws much smaller than 1/10T.

The ultrasonic system was calibrated using a standard with a 2%T notch as required by Section V, which is invoked by Section XI.

Our procedures require all flaw-like indications to be evaluated and any indication to be reported regardless of amplitude. This exceeds the Code requirement that indications determined not to be of geometric origin must be reported above 20% DAC (screen height). Consequently, a 1/10T flaw of approximately 0.5" is well within the detection capability of this equipment.

Our examiners are qualified under the EPRI NDE Center sponsored training for Appendix Vlil, " Performance Demonstration Initiative," for carbon and stainless steel piping, including the IGSCC portions, and therefore possess appropriate, demonstrable skill levels in ultrasonic inspection methods. They were also y miified and certified according to the requirements of ASME Section XI, ASN TC-TC-1 A, and the owner approved contractor procedure LMT-QA-37.

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During the examination of the pressurizer surge nozzle documented in report numbers 296-091 UT-045, 296-091 UT-048, 296-091 UT-049, 296-091 UT-050, 4

i and 296-091PT-020, no indications were reported. All examination results were acceptable as being in compliance with the Code and applicable procedures.

! The examination reports are provided in Attachment 4.

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1 FIGURES 4

Figure 1 Pressurizer Surge Nozzle Assembly Figure 2a Stress Time History on Nozzle Section 1-1 Figure 2b Stress Time History on Nozzle Section 1-1 (First 40 Seconds) 4 Figure 3 Temperature Time History at the T/4 Location i Figure 4 Finite Element Model of the Pressurizer Surge Nozzle including the Pressurizer Lower Head

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i i Figure 5 Unit 2 Pressurizer Surge Nozzle Examination Areas e

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PROCEDURES SO23-XXVil-20.66: ULTRASONIC EXAMINATION OF VESSEL WELDS AND ADJACENT BASE METAL SO23-XXVil-20.52: ULTRASONIC EXAMINATION OF NOZZLE INNER RADIUS AREAS i

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