ML19254D335

Forwards Addl Info on Applicability of ASME Section XI Re Existing Feedwater Piping Sys Linear Indications,In Response to Request at 791019 Meeting

ML19254D335
Person / Time
Site:	Millstone
Issue date:	10/22/1979
From:	Counsil W NORTHEAST UTILITIES
To:	Reid R Office of Nuclear Reactor Regulation
References
TAC-11793, NUDOCS 7910250279
Download: ML19254D335 (24)

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- NOIrrHt!/L!Tr trrit.rrl!!!i ATFORO CONNEOTICUT DE ?C-

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October 22, 1979 Docket No. 50-336 Director of Nuclear Reactor Regulation Attn: Mr. R. Reid , Chief Operating Reactors Branch #4 U. S. Nuclear Regulatory Commission Washington, D. C. 2^555

References:

(1) W. G. Counsil letter to R. Reid dated September 28, 1979.

(2) Telecopy from E. L. Conner for Additional Information dated October 16, 1979.

(3) W. G. Counsil letter to R. Reid dated August 22, 1979.

Gentlemen :

Millstone Nuclear Power Station, Unit No. 2 Feedwater System Piping In usponse to Reference (2), a meeting was held in Bethesda on October 19, 1979 to discuss the NRC request for additional information regarding Reference (1).

As a result of the October 19, 1979 meeting, additional information enclosed herein is submitted in response to the NRC Staff request regarding the ASME Section II applicability of the existing feedwater piping system linear indica-tions.

Attachment 1 provides a discussion of the methods of analyses including the fatigue crack growth analysis and the determination of the critical flaw size.

Attachment 2 provides the assessment of crack growth for the worst feedwater line flaw considering both the design basis transients and the thermal loading conditions observed from the Hillstone Unit No. 2 instrumentation data.

Attachment 3 provides the critical flaw e.izes for part-through wall cracks and through-wall cracks using the established loading cor,litions.

In reference to the submitted information, we note the following:

(1) The stresses calculated and presented in Tables 2 through 5 of Attachment 2 are conservative, because they were generated from a model meant to umbrella a number of different feedwater line observations. Specifii . ally, the maximum stresses in Table 4 (Attachment 2) should be multiplN by 0.753 for specific applicability to Millstone Unit No. 2.

1208 251 7910250

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Thus, the largest value of the stress intensity factor (K) which would result from a 0.100 inch deep flaw is 34.0 ksi /in.

(2) The results of the detailed piping integrity analyses confirm that the ductile f ailure limits of ASME Section XI are met. However, the upecific LEFM criteria for flaw evaluation in ASME Section XI

-- IRE-3600 is not applicable to the feedwater piping system.

(3) Using the results of the fatigue crack growth analysis and compering them to the established critical flaw size, it is concluded that there is a large safety margin.

It is, therefore, concluded that the existing condition of the Millstone Unit No. 2 feedwater pipit.g system is in compliance with the applicable criteria of the ASME Boiler and Pressure Vessel Code as docu=ented in Reference (3).

Additional information presented at the October 19, 1979 meeting will be submitted by October 24, 1979.

We request ycur immediate attention to this matter and trust this information satisfactorily dispositions the Staf f's concerns.

Very truly yours, NORTHEAST NUCLEAR ENERGY COMPANY

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W. G. Counsil Vice President Attachments 120e m am

DOCKET NO. 50-336 ATTACIDiENT 1 MILLSTONE NUCLEAR POWER STATION, UNIT NO. 2 FEEDWATER SYSTEM PIPING e n

)}0Ei OCTOBER, 1979

METHODS OF ANALYSIS In this work, the observed indication is treated as a sharp crack, and analyzed as to its behavior in future service. Growth due to further cycling is evaluated in fatigue crack growth analyses, and then the final flaw size is compared with the critical flaw size for normal and upset and other operating conditions. In this section, the methods used in these analyses will be detailed .

(1) FATIGUE CRACK GRO'n~H ANALYSIS The f atigue crack crowth analysis was conducted in the same manner as suggested by ASME 3ection XI, Appendix A. The operating transients which af f ect the feedwater line are all considered, and scheduled out over a 40-year period. The initial flaw depth assumed was that of the original indication, but slightly greater depths were also considered, to give further information.

Crack tip stress intensity factors (K ) Iwere calculated using an expression for a continuous flaw oriented circumferentially at the inside surface of the pipe. The stresses were linearized through the pipe wall thickness, and used to calculate KI and AK I . The fatigue crack growth for any single transient was calculated from a crack growth rate law deter-mined to be applicable for the materials of the pipe, exposed to a water enviro nment.

(2) DETERMINATION OF CRITICAL FLAW SIZE The feedwater piping and welds are fabricated from carbon steel and operate at elevated temperature. A great deal of study of the failure aspects of piping and tubing have been undertaken in recent years and considerable experimental data are now available. A large number of failure theories have been developed, both analytically based and empirically based, with

- varying degrees of success. This section will briefly review various types of theories, and provide the basis for use of the plastic insta method in predicting the critical flaw size for the area of interest.

I208 25;

Fracture mechanics was first developed for the case of low energy fracture which involved small deformations. This is termed brittle fracture, and the theory is called linear elastic frc ture mechanics (LETM). The theory is generally applicable only to brittle materials for example high strength alloys and those which operate at low temperature.

A further requirement for strict applicability of LEFM is the presence of a heavy section, where the stress state is plane strain. Failure in these materials and geometries is abrupt, and the crack propagates at speeds approaching the speed of sound. The theory predicts that failure will occur when the applied stress intensity factor exceeds the material's fracture toughness, or resistance to failure.

There is a large range of materials and geometries where the conditions necessary for linear elastic or brittle fracture do not exist. This happens in lower strength carbon steels, stainless steels, and Inconel, particularly those with high ductility, and also in structures with thin sections with low constraint on the opening of a crack. A good example of this geometry is piping and tubing. In this case, once a crack is loaded to the point where it begins to propagate, failure does not occur at once. Instead, as the crack propagates, the plastic zone ahead of the crack grows with the crack, and a steady increase in the magnitude of the load becomes necessary to overcome the increasing resistance of the material to fracture. Consequently, a toughness oriented single para-meter fracture criterion becomes totally inadequate to deal with the problem of ductile failure.

A number of concepts have been developed for the prediction of ductile failure, and these are reviewed in detail in a number of recent works, for example references 1, 2, 3, and 4. Two of the more popular parameters for ductile fracture are the J-integral [5] and the crack opening displacement (COD)[6] concert. These parameters have been shown to be successful at predicting the onset of ductile crack propagation, but are only now being extended to the prediction of final failure. Extensions to the point of unstable propagation and final failure have thus fai been centered on R-curve technology [7] and development of the Tearing Modulus concept by Paris [8] is an extension of this trend.

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Based on the level of Charpy energy at 0*F from tests of the actual material as well as experience with results from similar materials, the transition f rom orittle to ductile behavior should occur at room temperature or below.

The operating regime for the feedwater lines as well as their material and geometry places the fracture mode squarely in the large strain -

general yield regime. As such, the crack will generally not become unstable until beyond the point where the entire remaining ligament becomes plastic.

If this occurs, the failure will b'e well predicted by the plastic limit load of the structure, corrected to account for the material strain hardening behavior.

There is considerable body of experimental data which shows that the governing mode of failure for ductile cracked pipes and tubes is that of plastic instability. Several series of experiments on piping geometries were completed by both General Electric and Bat;1ile Memorial Institute as early as 1968, and these results, as well as other more recent results are well-predicted by the plastic instability method, as discussed in Appendix A. Therefore, the approach taken in this analysis was to evaluate the propensity for failure by the plastic instability mode.

(3) SAFETY ASSESSMENT Once the growth of the assumed crack-like def ect has been calculated, the resulting flaw is compared with the critical flaw size to determine the margins of safety for further operation. T' *.s assessment metted is similar to that used in Section XI of the ASME code, but the details of the calculations are different, especially the critical flaw size calculation for ductile failure. Note that there are presently no rules or guidelines in the ASME code for such calculations in secondary systems. The assessment method used is, therefore, based on good engineering practice.

1208 256

APPENDIX A DETERMIMTIOS OF THE MOMEST CAPACITY OF FRESSURIZED FIFINS WITH CIRCUMFERENTIALLY ORIESTED THROUGH WALL FLAWS A straignt section of pipe with a circumf erential1y oriented through wall flaw; as shown in Figure A-1, is considered. It is assumed that plane sections re=ain plane during deformation, and that the flaw is not tov Iarge in comparison with the pipe circumference. For flaw lengths which approach one-half of the circum-ference, the present method is not accurate.

The pipe is loaded by internal pressure, P, an axial force, F, and a bending moment, M. Because of the bending moment, the axial stress will be compressive somewhere in the cross section. The point of de=arcation between tensile and compressive stresses is the neutral ayls, as shown in Figure A-1. To determine the location of the neutral axis, the axial force on the pipe from the internal pressure and other loads, N, is equated to the integrated stresses over the cross-sectional area of the pipe, No as follows:

N = PnR2+F (A-1)

Where:

P = Internal Pressure R = Mean Radius of the Pipe F = Other axial force (if any) n -8 N =2 of Rt de

+ 2[ - of Rt d6 (A-2)

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-8 2 Where:

t = Pipe Thickness of = Flow Stress = 0.4 (cy , + ou) a = Crack angle as shown in Figure A-1.

B = Angle to located neutral axis, Figure A-1.

1208 57 9

Equating the quantities in (A-1) and (A-2) leads to the definition of the neutral axis which is:

Of to- (A-3) 0 " 2 cf t - PR Figure A-1 also illustrates that the angles a and 8 are related at the limit moment; therefore, equating areas above and below the neutral axis results in the following:

a = 28 (A-4)

The fully-plastic limit moment capacity, M ,b is obtained by taking moments about t'

the aeutral axis as follows:

(90 - a) (2n - 6)

Mb=2 [-S ( t cf sin 0 de - [R (n + S) m t of sin 0 de (A-5)

Where: of = 0.4 (cys + Ou), that is, of is the flow stress.

After integration and substitution of the limits, the moment capacity for a pipe without internal pressure is found to be:

Mb = 2cf Ro 3

t (2 cos 6 - sin a) (A-6)

For simple pressure loading with no bending, the limiting force is equal to:

No - 2 (n - a) R m E f (A-7)

For any arbitrary pressure, P, the force produced is:

N = WR 1 2 P (A-8) 1208 75:

Then, the ratios of axial force to li=it axial force and moment to limit moment are defined as follows:

n=bm=b ho Mb (A-9)

Since the internal pressure and bending moment interact, the combined effect will cause a reduction in the moment capacity Mb to M. From Hodge's interaction theory (1) . The corrected limit moment is determined from:

M= (1 - n 2) Mb (A-10)

Substituting for all the parameters from equations (A-6) through (A-9), we obtain:

2 2 2 2 4 9 4 (v - a)2 Rm t of -n Ri P' ML= 2 [Ro 2 (2 cos 8 - sin a)] (A-ll) 2 (n - a)2 Rm E Of (1) Hodge, P. G., Plastic Analysis ot' Structures. McGraw-Hill Book Company, 1959, pp. 130 - 190.

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ECFET No. 50-336 ATTACleiEST 2 MILLSTONE NUCLEAR POWER STATION, UNIT NO. 2 FEEDWATER SYSTEM PIPING 1208  :". 6 1 OCTOBER, 1979

ASSESS"ENT OF GRLTH OF FEED:ATER LI' E FLAWS MILLSTONE II U.H. Bamford The purpose of t'his work is to estimate the future growth of a flaw located in the counterbore region near the feedseter nczzle safe-end-to-pipe weld. The flaw cf interest has been confirred by UT to be approximately 0.10 inches deep, and oriented circurferentially.

As a result of the location of this flaw, instrumentation was in-stalled to acnitor the terperature fluctuations in one loop. Results showed that in a certain flow rate range the water stratifies, pr du-cinc significant stresses which are potentiall ir.ncrtant for crack growth.

The types of stratification produced were typical of those observed in other plants, but not as severe. The observed stratificaticns were classified under five different types, as shown in Figure 1. The tem-perature difference from top to bottom of the pipe for profile 1 was measured at about 350*F, whereas for other plants it has been found

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-to be as high as 450 F. -

A three dimensional finite element stress analysis has been completed for each of the five temperature profiles in Figure 1, and transient studies have shown that the five profiles represent limiting conditions compared with the stress results obtained for any transient step in be-tween the profiles.

To accomplish a fatigue crack growth analysis, the system design tran-sients for normal, upset and test conditions were corbined with the

. cycles ci stress from stratification, which occurs during hot standby operaticn. As shown in Figure 2, there are approximately nine cycles of varicus degrees which for the purpose of this cnalysis, we will assccc, occur eacP time hot strndby cccurs.

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2-A tabul.ation of the cycle types used in .oe crack crowth analysis, along with acplicable stresses, is provided in Table 1. Tables 2 through 5 show the stresses et various locations around the pipe as a result cf the stratification.

Tne actual stresses from the three dimensional analysis were used for the fatigue crack growth analysis, except in two cases, where compres-sive stresses far exceeded the yield stress in ccc ression. The locat-ion is at the top of the pipe, and the condition c.ccurs only when the pipe is nearly filled with cold water (profile 1) at low floa. For this case, tensile residual stress values were assumed to exist, ecual to the yield strength. This is seen at locations 1 and 2 in Tables 2 and 4. This assumption is considered to be extremely conservative.

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Crack growth was calculated at each of thirteen locations around the pipe for periods of 1, 2, 3 and 4 years, assuming an initial flaw of 0.100 inches deep, extending entirely around the inside of the pipe.

' A fatigue crack growth law which accounts for mean stress or R ratio (omin./ o max.) as well as the presence of the water environment was used. The law is shown in Figure 3.

Results of the crack growth analysis are shown in Table 6, for each of the locations considered. These results show that the observe flaws will not grow significantly during the next years service. The final flaw size for the worst location is a factor of 5 smaller than the critical flaw size for the pipe, as shown in Figure 4.

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SUMMARY

OF TEMPERATURE PROFILE OCCURRENCES flUMBER( ) RANGE PLANT EVENT NUMBEROFOCCU_RRENCES(I)0FPROFILE OF EVENTS PROFILE 1 PRCFILE 2 2 3 4 5 1

27f-l?42 5 9 115 276- 1242 MILLST0tlE Il0T STANDBY 6 6 6 (LI.NE2). ..

L tM ca (I) fiUMBER OF OCCURRENCES (WITH ATT0P/ BOTTOM lilFORMATION.

" 300 F) IS BASE '

N os (2) NUMBER OF EVENTS IS BASED ON PRESENTLY AVAILABLE PLANT OPERATIflG HISTORY U1 RANGE = (li 0F OCCURRENCES 20%)

(3) RAtlGE FOR TOTAL tlUMBER OF OCCURRENCES OF THE PROFil.E AND IS CALCULATED AS:

4 (# EVEliTS) X S, WHERE S = EVENT SIMILARITY FACTOR AND 5.5 51.5.

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F.C T:E 2 FATIGUE CPACK GROWTH DATA FOR SA-503, CL ASS 2 AND CLASS 3 AND SA-533, GRADE 6, CLASS 1 STE'ELS

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TAELE 2 STRESS RESULTS - AXIAL DIRECTIO!;

C0!iDITI O!? 5 HOT STA!;CEY #1 Inside Surface Outsi:'e Surface Location Max. Min. Max. Min.

(Ksi) ( /,s i )

1 40.0 0.0 -40.0 0.0 2 40.0 0.0 -40.0 0.0 3 9.46 -23.0 8.56 8.29 4 35.12 4.63 11.23 7.02 5 68.97 - 1.43 13:20 4.66 6 65.05 - 7.28 12.61 1.71 7 46.17 - 8.06 10.83 -0.33 8 24.37 7.27 7.34 3.68

.9 24.61 7.27 8.98 2.01 10 23.67 - 2.47 6.47 -3.44 11 14.93 - 5.67 - 0.44 -7.05 12 9.30 - 5.44 - 6.03 -8.45 13 7.62 M.95 - 8.11 -8.69 ,,g; g :ig ] -l A w'.,,

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TAELE 1 TE ';SIESTS USED I" FATIG L'E C P ' C P. GF T.7P '. LYSIS MILLSTONE II Cycles Inside Surface Stress Outside Surface 5-Description (40 years) Max. Min. Max.

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Hot Standby 2 1500 )

These stresses are dependent on circumferert-Hot Standby 3 500 ) position. See Tables 2 through 5.

Hot Standby 4 2500 )

10.47 7.71 8.49  ;

Unit Lead-Unload 15000 ,

Step Increase / Decrease 2000 9.56 7.87 7.89 7 partial Loss of Flow 40 23.7 8.42 3.79 7 Loss of Load 40 23.02 8.42 3.18 7 Reactor Trip 400 22.69 8.21 2.88 7 Secondary Leak Test 200 11.23 0.0 10.04 C 1208 268

TABLE 3 STRESS RESULTS - AXIAL DIRECTIO!;

CONDITIC'; 5 HDT STA';:EY s2 Inside Surface Outside Surface Location Max. Min. Max. P.i n .

(Ksi) (::si) 1 13.49 9.78 9.62 3.24 2 12.57 9.6S 9.40 3.25 3 9.46 9.34 8.56 3.24

-4 8.76 4.63 3.21 7.02 5 8.23 -1.43 3.17 4.60 6 7.91 -7.28 3.17 +1.71 7 7.69 -8.06 3.18 -0.33 8 7.60 7.27 3.22 3.63 9 24.61 7. 71 8.97 3.37 10 2'3.67 8.02 6.47 3.53 11 14.93 8.44 -0.44 3.96 12 9.30 8.76 -6.03 4.20 13 7.61 8.86 -8.11 4.29

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(Ksi) (Ksi) 1 40.0 0.0 -40.0 0.0 2 40.0 .0.0 -40.0 0.0 3 16.19 23.71 9. 31 8.30 4 35.12 10.11 11.23 8.45 5 68.97 2.39 13.20 6.52 6 66.05 - 4.85 12.61 3.45 7 46.17 -11.04 10.83 -0.88 8 24.37 -15.29 7.34 -5.64 9 7.27 -14.93 2.01 -8.39 10 - 2.48 - 3.19 - 3.44 -4.21 11 19.41 - 5.67 7.35 -7.05 12 35.44 - 5.44 16.75 -8.45 . . , _ . . , .

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TAELE 6 F.ESULTS OF FATIGUE CFACK GC.lTH A',A'_Y::E INITIAL CF;ACK LEt;GTH = 0.100 It;CHES Crack Depth After Year Location .1 2 3 4 1 .1008 .1016 .1024 .1032 2 .1002 .1004 .1007 .1009 3 .1001 .1001 .1002 .1003 4 .1001 .1003 .1004 .1005 5 .1021 .1045 .1067 .1092 6 .1042 .1090 .1137 .1190 7 .1011 .1023 .103; .1046 8 .1001 .1002 .1003 .1003 9 .1001 .1002 .1003 .1004 10 .1001 .1002 .1002 .1003 11 .1001 .1003 .1004 .1006 12 .1014 .1030 .1045 .1062 13 .1032 .1067 .1105 .1146 U!

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DOCFET NO. 50-336 ATTACINENT 3 MILLSTONE NUCLEAR POWER STATI0li, UNIT NO. 2 FEEDWATER SYSTDi PIPING OCTOBER, 1979 1208 2/3

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