ML20096H013

Nonproprietary Fracture Mechanics Evaluation of Inservice Insp Indication Indian Point Unit 2 Reactor Vessel

ML20096H013
Person / Time
Site:	Indian Point
Issue date:	08/31/1984
From:	Bamford W, Furchi E, Schmertz J WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP.
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ML100321896	List: ML20096H013
References
WCAP-10650, NUDOCS 8409110232
Download: ML20096H013 (111)

Text

_ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . _ _ _ _ _ _ _ - _ . _

WESTINGHOUSE CLASS 3 CUSTOMER DESIGNATED DISTRIBUTION WCAP 10650 l

FRACTURE MECHANICS EVALUATION OF INSERVICE INSPECTION INDICATION INDIAN POINT UNIT 2 REACTOR VESSEL AUGUST 1984 BY W. H. Bamford J. C. Schmertz l

E. L. Furchi D. C. Adamonis l

APPROVED: ik hEwm v

J. N. Chirigos, Manager Structural Materials Engineering Work performed for Consolidated Edison Co. Under IDRJ-949 Although the information contained in this report is non-proprietary, no distribution shall be made outside Westing-house or its Licensees without the customers approval.

i l WESTINGHOUSE ELECTRIC CORPORATION Nuclear Energy Systems P.O. Box 355 Pittsburgh, Pennsylvania 15230

~

8409110232 840907 PDR ADOCK 05000247 G PDR

TABLE OF CONTENTS SECTION TITLE PAGE 1 INTRODUCTION , 1 -1 2 INSPECTION RESULTS 2-1 3 FRACTURE TOUGHNESS DETERMINATION 3-1 4 FATIGUE CRACK GROWTH CONSIDERATIONS 4-1 l'

5 ALLOWABLE FLAW DEPTH CALCULATIONS 5-1 6 EXTERNAL THERMAL SHOCK FROM CAVITY FLOODING 6-1 7 SUPNARY AND CONCLUSIONS 7-1 8 REFERENCES 8-1 APP. A PRESSURE AND TEMPERATURE VARIATIONS FOR A-1 TRANSIENTS ANALYZED APP. B PROBABILISTIC ASSESSMENT B-1 l

p iii

LIST OF TABLES TABLE TITLE PAGE 3-1 Sununary of Available Weld Metal Chmistry Measure- 3-4 sents 3-2 Properties of Indian Point Unit 2 Lower Shell 3-4 1

3-3 3-4 End of Life Fluence and RTNDT at Indication 4-1 Reactor Design Transients 4-6 4-2 Beltline Surface Hoop Stresses for Level A. Level B 4-7 and Test Transients 6-1 Teperatures and Stresses for Cavity Flooding Core 6-2 (Ref.11) l l

s l

V

LIST OF FIGURES FIGURE TITLE PAGE 3-1 Identification and Location of Beltline Region 3-5 Material for the Indian Point Unit No. 2 Reactor Vessel -

3-2 Current Fast Neutron Exposure as a Function of 3-6 Depth within the Indian Point Unit 2 Pressure Vessel - 345' Azimuthal Angle - 5.33 EFPY 3-3 EOL Fast Neutron Exposure as a Function of Depth 3-7 within the Indian Point Unit 2 Pressure Vessel -

345' Azimuthal Angle - 32 EFPY 3-4 Ky , and KIc from Section XI. ASME Code 3-8

4-1 ASME Code Reference Fatigue Crack Growth Law, 4-8 Air Environment 4-2 Pressurized Water Reactor Vessel Showing Beltline 4-9 Region 4-3 Nozzle Transition, Beltline and Lower Head Regions 10 Dimensions 4-4 WECAN Finite Element Gectnetry Model 4-11 4-5 4 Loop Reactor Vessel Mechanical Boundary Conditions 4-12 4-6 4 Loop Reactor Vessel Thermal Boundary Conditions 4-13 4-7 Hoop Stress Contours - Pressure Only (3105 PSI) 4-14 4-8 Beltline Heatup and Cooldown Temperature 4-15 4-9 Beltline Inadvertent RCS Depressure Temperature 4-16 4-10 Beltline Boron Concentration Temperature 4-17 4-11 Beltline Hot Hydro Test Temperature 4-18 4-12 Cooldown Temperature Contours - Example of a Slow 4-19 --

Thennal Transient 4-13 Excessive Feedwater Flow Temperature Contours - 4-20 Example of a Fast Thennal Transient 4-14 Expanded View of Thennal Contours for Excessive 4-21 Feedwater Flow Transient (See Figure 4-13) 4-15 Hoop Stress Contours for Cooldown Transient - Ex- 4-22 ample of Slow Thennal Transient 4-16 Hoop Stress Contours for Excessive Feedwater Flow 4-23 Transient - Example of a Fast Thermal Transient -

vii

LIST OF FIGURES (cont'd)

FIGURE TITLE PAGE 5-1 Stress Intensity Factor Calculations - Surface Flaw 5-4 5-2 Combined Pressure and Thennal Hoop Stress for the 5-5 Small LOCA Transient 5-3 Temperature Distribution for Small Steamline Break 5-6 f

5-4 Combined Pressure and Thennal Hoop Stress for the 5-7 Small Steamline Break 6 Fracture Results of Cavity Flooding 6-3 s

viii

SECTION 1 INTRODUCTION A fracture analysis per ASME Section XI has been carried out to" investigate the acceptability of an indication discovered during the inservice inspection of the Indian point Unit 2 reactor vessel. The indication was found in the longitudinal weld in the lower shell region of ths vessel near the outside surface at an azimuthal angle of 345' to the cardinal axis. The analyses presented in this report are based on the information presently availabls for the indication, and have been structured to be applicable to it as presently the'racterized, or for any san 11er indication in the same location, near the outside surface of the vessel.

Ttree types of calculation are necessary for the evaluation of an indication per the requirements of Section XI, article IWB 3604:

1. Fatigue crack growth

2. Critical Flaw d4pth-normal conditions

3. Critical Flaw depth-faulted conditions Fatigue crack growth evaluation must be carried out on the indication as ,

characterized and then the indication after growth must be assessed relative to the critical flaw depths calculated. The criteria given in Section XI are clearly specified, and two alternative approaches are available, either a mergin based on flaw depth, or a margin based on applied stress inten:ity factor. The second option has been adopted here, and the required margins are:

K K

g

< d

/ 10 for normal, upset and test conditions K

Kg < b

1. 2 for emergency'and faulted conditions 1-1

The fatigue and fracture toughness properties used in the analyses were taken from the reference properties provided in Appendix A of Section XI. The irradiation damage accumulated in this region of the vessel is a function of the integrated neutron fluence, which was obtained as described in Section 3.

The indication was also evaluated relative to the primary stress limits of Section III, NB 3000, as required by Section XI to demonstrate its accept-ability.

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1-2

2 . y s : ,,. .

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SECTION 2

~INSPECTI0h RESULTS During the interval 1. period 3 inservice examinat' ion of the Indian Point Unit 2 reactor vessel, a ultrasonic reflector was detected in the lower

shell longitudinal weld at 345' vessel azimuth. The reflector was ini--

tially discovered via 45':and 60* circumferential shear wave scanning of the ad,jacent base material on the lower shell side of the intemediate-to-lower shell circumferential weld. Indcations cf recordable amplitude were apparent' in the clockwis's and counter' clockwise scan directions, approxi-mately three inches below the centerline of the circumferential weld seam.

These results were confimed during subsequent 45' and 60* examinations of the lower shell longitudinal weld ht 345' vessel azimuth.

Initially, there was some $ peculation that multiple reflectors-might be in-volved, however subsequent investigations have confirmed one reflector, located near 345' vessel azimuth, oriented axially with respect to the vessel, extending downword for a ' length of 1.96 inches from a point approximately 2.25 inches below the centerline of'the intemediate-to-lower shell weld.

Raw data predict a 2.03[Mch, through-wall' dimen,sion. Beam spread correction reduces that size to 1.2 inches. Peakf.niplitude sweep locations for the in-dications from this raflector place it at or very near the 00 surface (within .25").

C 4 ' Recognizing that ASME XI 50% DAC sizing techniwes tend to exaggerate through-wall dimensions -af small reflectors at or, near'the opposite surface; in some

~

cases as much as 10'to 1, a program,for1rther investigation of this reflector was requested bylonsolidated Edisch Cal The program which was implemented i .

included 1) application of,a 45* pitch-catch transducer arrangement to deter-f mine the' extent of "r.hads{ng" by,the' rsflebtor and 2) =a delta transducer ar-rangement to more accurately establish the depth or through-wall dimension of-the' indication. ', j

.i /

Repults'from then additional . investigations suggest the reflector through-wall dimension is significentl'y smaller than the 2.03 inches predicted by ASME XI 50% DAC sizing techniques. The delta technique measurements predict a maximum depth of 0.3 inches. Additional confimatory studies are planned v

to further demonstrate the efficacy of the delta transducer arrangement for g .t reflector. depth measurements, but for this analysis the maximum corrected

/

depth was used. .

. 2-1 f-y

i SECTION 3 FRACTURE TOUGHNESS DETERNINATION As described in the previous section, the indication is located in the vertical weld seam of the lower shell, oriented axially, and centered about 2 inches below the centerline of the intermediate to lower shell weld. The location of the indication is shown in Figure 3-1, where the reactor vessel weld locations as well as their designation numbers are also shown. The indication is in weld seam 3-042A shown in this figure, located at an angle c8 345' from the cardinal axis.

The properties of the weld seam of interest were not obtained when the vessel was constructed, since no characterization was required. It is known that this weld was made with RAC03 weld wire (heat number W5214) and Linde 1092 flux, with a Nickel 200 wire addition. A review was made of welds made with this heat of weld wire combined with the Nickel wire and 1092 flux, and it was found that several characterizations have been made. A summary of the l

chemistry results is given in Table 3-1. Based on these nine chemistry analyses, it was concluded that the average values of copper, nickel and phosphorous should be used, and these are provided in Table 3-2. Since the actual RT of this weld was not measured, a generic value of -56'F was NDT used [1]. For reference the chemistry and RT values f r the two adjacent NDT base metal plates have also been included in Table 3-2 [2].

Fluence was determined for both present life and end-of-life in the cross-section of the vessel where the indication was located, at the 345' degree trimuthal angle. Results are presented in Figure 3-2 for the current ;

exposure and in Figure 3-3 for the projected end-of-life (32 EFPY) exposure o'f the vessel.

Exposure informattom is supplied both in terms of neutron fluence (E > 1.0 MeV) and of fluence equivalent dPa. Here " fluence equivalent dPa" refers to the use of the shape of the energy dependent displacements per iron atom damage function through the vessel wall normalized to the fast neutron 3-1

(E > 1.0 MeV) flu:nco at the vessel inner radius. The use of the fluence equivalent dPa permits the application of energy dependent damage gradients in conjunction with available trend curves to assess vessel integrity.

Test neutron (E > 1.0 MeV) fluence profiles were taken from Reference [3).

Displacements per atom calculations were carried out using the calculated ,

F neutron data presented in ASTN Standard Practice E 693-79, " Characterizing l Neutron Exposures in Ferrite Steels in Terms of- Displacements Per Atom".

Although Indian Point Unit 2 has recently implemented low leakage fuel l management, the flux reduction at the 345' azimuthal location resulting from l l

the new fuel annagement scheme is not large. Therefore, for conservatism the fluence projections provided in Figures 3-2 and 3-3 were based on the non low leakage neutron flux profiles.

The effect of the fluence on the frecture toughness of the vessel is not severe, because of the location of the indication near the outside of the vessel. Therefore, the end-of-life values were used in determining the final value of RT NOT.

The fluence values are listed in Table 3-3, for the region at the tip of the assumed flow, a location of 18.2 cm from the inside surface.

The irradiation damage calculations recomended by the NRC [1] were used to determ ne the end-of-life RT i

NDT "' NOT

• end-of-life fluence is determined as the lower of the results given by equations (3-1) and (3-2).

Equation 3-1:

RT NOT = I + M + [-10 + 470 Cu + 350 Cu W 0 Equation 3-2:

0 RT NDT = I + M + 283 f . m "I" means the initial reference temperature of the unirradiated material measured as defined in the ASME Code, N8-2331. If a measured value is not 3-2

available, the following gen;ric mean values must be used: 0*F far wlds made with Linde 90 flux, and -56*F for welds made with Linde 0091, 1992 and 124 and ARCOS B-5 weld fluxes.

"M" means the margin to be added. In Equation 1. M-48'F if a n'easured value of I was used, and M-59'F if the generic mean value of I was used. In Equation 2, M-0*F if a measured value of I was used, and M=34*F if the generic '

mean value of I wks used.

, "Cu' and "N1' mean the weight percent of copper and nickel in the material.

l The source of these values must be reported. The relationship of the material on which any measurements were made to the actual material in the pressure vessel must be described.

"f' mean, the maximum neutron fluence, in units of 10l 'n/cm2 (E greater than or equal to 1MeV), at the location of interest.

The calculated value of RT at end of life for the material where the NOT indication is located is shown in Table 3-3. Using this final value of RTNOT, and the temperature, the toughness of the vessel can be determined from the reference curves of the ASME Code, which are shown in Figure 3-4.

4 3-3

TABLE 3-1 l SUWRY OF AVAILABLE WELD METAL CHEMISTRY MEASUREMENTS Cu Ni P l Weld Source (Wt.5) (Wt.1) , (Wt.5) l Indian Point Unit 3 .15 1.02 .019 i Indian Point Unit 3 .16 1.06 .01 7 Indian Point Unit 3 .15 1.11 .018 Indian Point Unit 3 .15 1.09 .018 Millstone Unit 1 .19 .98 -

H. B. Robinson Unit 2 .1 54 .99 .01 2 H. B. Robinson Unit 2 .163 .90 .011

'H. B. Robinson Unit 2 .152 1.08 .014 H. B. Robinson Unit 2 .166 1.00 .01 2 TABLE 3-2 PROPERTIES OF INDIAN POINT UNIT 2 LOWER SHELL Property Weld 3-042A Plate B2003-1 Plate B2003-2 Copper, wt.% 0.16 0.20 0.19 Nickel, wt.% 1.03 0.66 0.48 Phosphorous, wt.% 0.01 5 0.011 0.01 0 Initial RT NDT -56*F 20'F -20'F l

l TABLE 3-3 END OF LIFE FLUENCE AND RT AT INDICATION NDT l Fluence (E>l MeV) 7.5 x 10I7 n/cm2 Fluence (dPa) 1.7 x 1018 n/cm2 RT 79ep

! NDT (weld) l l

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3-4 l

FIGURE 3-1 IDENTIFICATION AND LOCATION OF BELTLINE REGION MATERIAL FOR THE INDIAN POINT UNIT NO. 2 REACTOR VESSEL CIRCUMFERENTIAL SEAMS VERTICAL SEAMS f'

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3

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v s B2003-2 t1o*

INDICATION LOCATION 3-5

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e A. at 6 a no u sat a la ao a at toona m WESLEL fmi Figure 3-2 Current Fast Neutron Exposure as a Function of Depth within the Indian Point Unit 2 Pressure Vessel - 345' Azimuthal Angle

- 5.33 EFPY 3-6

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Figure 3-3 E0L Fast Neutron Exposure as a Function of Depth within the Indian Point Unit 2 Pressure Vessel - 345 , Azimuthal Angle - 32 EFPY 3-7

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SECTION 4 FATIGUE CRACK GROWTH CONSIDERATIONS ThegrowthoftheindicationofinterestwasassessedbasedonItheassumption that it was s' flaw, located at the outside surface of the vessel. The analyses reported here utilized the reactor coolant system transients detailed in the original equipment specification of the reactor vnssel as well as i additional transients which have been developed since that time. These i transients are listed in Table 4-1. All the transients are described in detail in Appendix A. Since the assumed flaw was not exposed to primary coAlant, the fatigue crack growth was calculated using the air environment crack growth reference curves of Section XI Appendix A, as shown in Figure 4-1. The calculated crack growth was insignificant, as will be shown in this section.

4.1 STRESS ANALYSIS Thermal and stress analyses were performed for the beltline region of the '

reactor vessel during 25 Level A, Level 8 and Test Condition Transients. The temperature analysis was done for a total of 221 time points. The thermal and pressure stress analyses were also done for each of these time points.

The thermal and stress analyses were performed with the WECAN finite element program N , using a two-dimensional finite element model. To assure conservative results, the insulating effect of the cladding was not included. '

4.2 DESCRIPTION

OF LEVEL A. LEVEL 8 AND TEST CONDITION TRANSIENTS The design transients used in the evaluation of the reactor vessel beltline are based on conservative estimates of the magnitude and frequency of the temperature and pressure variations result.ing from various operating conditions in the plant. These are representative of operating conditions which are considered to occur during plant operation and are sufficiently severe or frequent to be included in an analysis of cyclic stress conditions.

Further, these are regarded as a conservative representation of transients 4-1 i

"r '

e ----. - - - - , - ~

which, when used as a basis for component fatigue evaluation, provide confidence that the component is appropriate for its application over the design life of the plant.

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Transients were picked to give an upper bound to the temperatu're limits. For example, the heatup and cooldown technical specification is for 100*F/hr, whereas the actual rate is closer to 50*F/hr. As another example, loading and unloading at 15-100% power is analyzed for 13,200 cycles. It actually would occur about 500 times in a typical plant. In general the transients were digitized for input to the computer, the representations where simplified, were chosen such that the resultant stresses would be higher than for the aciualtransient,thus,givingconservativeresults.

The operational transients are broken down into the following categories:

1.) Level A Service _L.imits Level A Service Limits are applicable to any transient in the course of system startup. operation in the design power range, hot standby and system shutdown, other than Level 8 Service Limits or Test Conditions.

2.) Level 8 Service Limits f

Level 8 Service Limits are applicable to transients which deviate from those controlled by Level A Service Limits and which are anticipated to occur often enough that the design should include a capability to withstand the limits without operational impairment.

3.) Test Conditions Test conditions are those pressure overload tests including hydrostatic tests and leak tests which occur in the course of testing the system both before and after initial startup.

The total number of cycles for each transient exclusive of the preoperational test cycles was distributed over the 40-year operating life of the plant.

4-2

4.3 VESSEL GE0 METRY AND FINITE ELEMENT MODEL Figure 4-2 shows the cross-sectional view of the four-loop reactor vessel.  ;

l The vessel is axisy metric and supported at the nozzles in such a fashion as

~

! to allow radial expansion and contraction. The basic geometric data required for the thermal and stress analyses input are shown in Figure 4-3.

The model for this analysis uses two-dimensional four node quadrilateral isoparametric elements. The sy metry of the region to be analyzed permitted a 20 axisyneetric finite element idealization. The outline of the idealized structure is shown in Figure 4-4.

i f.

The model consists of a mesh of rectangular elements with g elements through the thickness of the vessel wall and one hundred thirty-seven elements in height. The selection of the finite element mesh is based on considerable practical experience with thermal and pressure transients. It is selected to efficiently provide acceptable engineering accuracy without being excessively large. In order to model the stress profiles through the thickness, the nine rows of elements are distributed such that the first four rows starting at the inner surface, are half the thickness of the remaining five Nws. This distribution was chosen because the highest stresses and greatest changes in stress generally occur at the inner surface, where the temperature and pressure changes are applied and where a finer mesh will provide a more

, accurate calculation. The model consists of 1380 nodes and 1233 elements,

! resulting in 2760 gross total degrees of freedom.

The stainless steel reactor vessel cladding was not included in the model.

This assumption is considered to add conservatism to the analyses primarily because the insulating effect of the cladding is not included. This means

! that the heat transfer is higher at the vessel inner surface which causes the

! stresses on the inside and outside surfaces to be conservatively higher during

~

f thermal transients.

j 4-3 s

4.4 BOUNDARY CONDITIONS FOR STRESS ANALYS!$

The reactor vessel is subjected to axisynnetric thermal and pressure loadings so that the lower end of the model representing the center of the lower head is restrained from moving laterally (U, = 0), and the upper end"of the model l representing the shell at the nozzle region is restrained from moving vertically (U y= 0) as shown in Figure 4-5. Prassure is applied to the internal surface and no mechanical boundary conditions are applied to the l external surface. l l

Thermal boundary conditions are shown in Figure 4-6. The outside surface 1 as[umedtobecompletelyinsulated. When the inside surface of the vessel is subjected to thermal transients, the primary mechanism of heat transfer is forced convection. The thermal properties of the metal are input as a linear function of the temperature. The heat transfer coefficient associated with forced convection is obtained by using the Dittus-80elter formula.[5] The Reynolds number and Prandt1 number are obtained based on the flow cross-section geometry, the flow rate and the temperature of the coolant.

The boundary conditions for the thermal analysis are summarized below:

! a) Initial temperature is 557*F b)- External surface is insulated c) Fluid temperatures associated with the transients are according to those shown in Appendix A.

d) Heat transfer coefficient associated with the transients is 7000 Stu/hr-ft *F.

4.5 STRESS ANALYSIS RESULTS Obtained in the stress analysis are the sum of the stresses due to temperature l effects and the stresses due to internal preu ure. The WECAN compJter

! program was used to perform the stress analysis, using the same finite element model that was used to obtain the temperatures. The temperature and pressure input data for the stress analysis were given in Appendix A. The pressure induced stress analysis was performed for each time step and is combined with the temperature induced stresses.

4-4 i

r

The results of the stress analysis provide the stress distribution history for the entire region of the finite element model. Maximum and minimum, inside and outside combined thermal and pressure hoop stresses for the beltline are given in Table 4-2 for the twenty-five transients and the steady-state condition. Note that in some cases, the steady-state value is used in the calculation. This occurs when the steady state condition is either the1 maximum or minimum condition in the transient, for example when it occurs from steady state conditions.

Temperature and hoop stress contour plots are shown in Figures 4-7 through 4-16. These figures include hoop stresses for pressure only at p = 3105 psig for the cold hydrostatic pressure test, the maximum combined hoop stress for Heatup and Cooldown as an example of a slow thermal transient, and the maximum combined stress for Excessive Feedwater Flow as an example of fast thermal transient.

4.6 FATIGUE CRACK GROWTH RESULTS The fatigue crack growth analysis was carried out using the guidelines pro-vided in Appendix A, Section Xi of the ASME Code. A semi-elliptic surface flaw was assumed to exist at the outside surface of the reactor vessel where the actual indication was observed. The assumed flaw was oriented axially, and was assumed to be exposed to an air environment. Crack growth was calculateo using the stress intensity factor expression of Section XI Appendix A, as well as the reference fatigue crack growth law for air environments contained therein. The transients listed in Table 4-1 were used in the analysis, and the stresses used as input are listed in Table 4-2. Results of the faitgue crack growth analysis showed that a flaw initially 1.45 inches deep would grow to 1.454 inches in ten yars, to 1.455 inches in twenty years, and to 1.457 inches in 30 years, the re-maining design lifetime. Therefore, the crack growth is insignificant.

1115E:10/081684 4-5

Es---

l OCCURRENCE OPERATING CONDITIOrv

-TRANSIENT-(CYCL.ES) SERVICE LIMITS Heatup and Cooldown 200 Level A Loading and Unioading 13,200 Level A Reduced Tenp. Return to Power 2,000 Level A Step Load Decrease and Increase 2,000 Level A Large Step Load Decrease 200 Level A-Initial Steady State Fluct. 150,000 Level A

- Random Steady State Fluct. 3,000,000 Level A Feedwater Cycling 2,000 Level A 80 Level A Loop Out of Service 80 Level B Loss of Load 40 Level B Loss of Power 80 Level B

' Partial Loss of Flow Reactor Trip 230 Level B with No cooldown 160 Level B with Cooldown No Safety Inj.

10 Level B with Cooldown and Safety.Inj.

20 Level B-Inadvert. Depressurization Inadvert. Startup Inactive 10 Level B~

p Loop Inadvert, Safety Injection 60 -Level B Actuation Control Rod Drop 80 Level B Excessive Feedwater Flow 30 Level B 26.400 Level A Boron Concentration 80 Level B

l. Refueling 20 Test L Turbine Roll 230 Test Hot Hydrostatic 10 Test Cold Hydrostatic Test i'

TABLE 4 REACTOR DESIGN TRANSIENTS 4-6

~

TABLE 4-2 BELTLINE SURFACE HOOP STRESSES FOR LEVEL A. LEVEL B AND TEST TRANSIENTS Maximum Correspond- Minimum Correspond.-

TRANSIENT- Outside ing Inside Gotside ing Inside Stress Stress Stress Stress ksi ksi ksif ksi Heatup and Cooldown 16.22 26.10 8.70 -6.98 Loading and Unloading 22.35 23.36 21.99 26.89 Reduced Temp. Return to Power 23.01 23.49 19.97 29.88 Step Load Decrease and Increase 23.09 23.17 21.93 26.28 Large Step Lead Decrease 23.50 21.91 22.06 24.98 Initial Steady State Fluct 22.55 26.17 22.17 23.42 Random Steady State Fluct 22.40 25.01 22.31 24.54 Feedwater Cycling 23.75 21.98 20.43 31.93 Loop Out of Service 21.61 23.5 21.22 27.47 Loss of Load 25.99 10.99 22.35 24.77

. Loss of Power 26.52 19.26 25.27 27.15 Partial Loss of Flow 23.16 24.92 21.39 18.79 Reactor Trip with No Cooldown 22.34 24.77 20.86 18.87 with Cooldown No Safety Inj. 20.69 21.17 17.53 27.86 with Cooldown and Safety Inj. 20.5 20.77 16.50 42.28 Inadvert. Depressurization -0.824 1 .631 -7.045 46.20 Inadvert. Startup Inactive 22.53 21.24 21.28 29.17 Loop Inadvert. S.I. Actuation 23.48 25.59 22.05 19.43.

Control. Rod Drop 22.34 24.88 20.07 21 .51 Excessive Feedwater Flow 15.37 21.8 11.35 62.88 -

Boron Concentration 22.65 25.10 22.35 24.77 I Refueling 0.909 -32.11 .926 32.66 Turbine. Roll 14.38 22.04 7.437 42.30 Hot Hydrostatic 21.54 27.07 6.80 -2.35 Cold Hydrostatic Test 29.97 33.21 0.0 0.0 4-7

8000 -. ...'

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SUB. SURFACE FLAWS .

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(Air Environment) -

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s_ a 3 4 saTeoh so 30 40 so so70aosoco STRESS INTENSITY FACTOR RAsect , am, (usa g) ,

!. - UPPER SOUND FATIGUE CRACK GROWTH DATA FOR REACTOR VESSEL STEELS Figure 4-1 ASME Code Reference Fatigue Crac,k Growth Lew Air Environment 4-8 .

l.

Control Rod Drive i M

Techanisms hi., $ @

Closure Head I.

% N

'"1 Thickness Transition l}

notatt Inlet g

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Vessel Wall Thickness Transition

1. JfLt1 EQllLE f "'l "

- - - < ~. RADIUS I w Vt5stL wall Mt1 tlEE ' " -

Tuttsuiss 88s1918 TeaaliTI0n I I Section

, Analyzed BEACTOR Cett 1- '

p _. . _ _ _ _ _ _ _ _._

, 1i ,

a l

- toute sta0 Vessel Wall assits '(

Lower Head I I ,"

Thickness Transition i _. ,________

Figure 4-2 Pressurized Water Reactor Vessel Showing Beltline Region 4-9

• 10.7f ~

,, . v,

.. " 85.6*N n0ZZLE SNELL

l. TRANSITION REGION l

e SELTLINE

.~

l

• = 8.825"

: *6.656"A '

i. (SA$E ETAL) 80TT04 HEAD j i

! ANIOTION

.28 R

(845E ETAL) 5.375"

f l

Figure 4-3 Nozzle Transition Belt 11ne aM Lower Head Regions-Dimensions 4-10 l

ggrf - -- . .

s.y- ,

w 0

55 h __

< t 1

1 "o

T 4

l $

.. .b 3 , ,

g t,,3

- .

Figure 4-5 4 Loop Reactor Vessel Mechanical Boundary Conditions 4-12

~

N///, i APPLIED TEWERATURE " [

/

V

- /

f l/

i- 7 WT TRANSFER COEFFICIDIT  %

X100 BTU /NR *F FT2 litSULATED SUtFACE

/

^

/

- /

/

/

i i

~

i N

N

\

l Figure 4-6 4-Loop Reactor Vessel Thennal Boundary Conditions 4-13

l, .

i le S E

Contour Stress -

flumber (ps!)

1 21000 1 4 2 24000 '

3 27000

[ 4 30000 5-5 33000

~

X = Max Stress = 33215 PSI

/

/ 3 /

i f

2-2  !

f 3 3q I ,

~

/ .

i 9

Figure 4-7 Hoop Stress Contours - Pressure Only (3105 psi) 4-14

( -

inc.o ... ., , ., ., ., , .,. .,

l Eh8 - -

me - -

b me - -

r d

as '- -

k.

4 lab 33 -

g, . . .. ... ... ... .. ... ....

60 61 &2 63 &4 &S &4 &7 &g &g Le DISTRCE 0(M 4

Figure 4-8 Beltline Heatup and Cooldown Temperature i

e 4-15 c~e--- a .,. - , - - - - - - . - , - - - - - - - - - __

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1 i

amo -

s r, -l -,

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DISME OVD 1

i Figure 4-9 Beltline Inadvertant RCS Depressure Temperature 4-16

i l

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&O at h2 63 64 &5 66 &7 L3 &9 he msmCE (XM Figure 4-10 Beltline Boron Concentration Temperature 4-17

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Figure 4-11 Beltline Hot Hydro Test Temperature e

4-18 '

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1. TEW CO.NTOUR r i

1 532 -

2 534 I

( -

3 4

535 i

537 5 SM-s 16 6 540

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Figure 4-12 Cooldown Temperature Contours - Example of a Slow Thennal Transient 4-19 O

r

~

• X Contour Temp k

No. *F

/

435 1

0- L , 2 442 l

,  ! 3 450 4 457 5 465

- ,' 17 6 7

472 480 16 8 487

- - - - - ' 9 495 7 15 10 502 7~' 11 510 i K 14 12 517

./ E ,

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Figure 4-13 Excessive Feedwater Flow T perature Contours - Example of a Fast Thermal Transient

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4

' 4 ,

\ , < g.20 ,, . ,1 I, 4 i

i N .

s

(

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Contour Temp no. *F 1 435 '

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10 . 502 2 11 510 0

.. 3 12 13

'l 517 l

525 4 L 14 532 15 540 5 [ ,

16  !

M7 17 555 6  !

7  !

8 9 '

l -

4 1

12 13 /,

14 15 f 16 17 Figure 4-14 Expanded View of Thermal Contours for Excessive Feedwater Flow Transient (See Figure 4-13) 4-21

s 4

Contour

• l Stress '

No. PSI l 1 .M X = 26.105

- 3 i

.r; 1

.t :.

, -t ..

i. s 1 1 X---

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.5

. 1-

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+

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$ t Figure 4-15 Hoop Stress- Contours for Coo 1down Transient - Example of Slow The# Transient

. < c e 0 i

4-22 s

i s

_a-4. -

SECTION 5 ALLOWABLE FLAW DEPTH CALCULATIONS 5.1 NORMAL. UPSET AND TEST CONDITIONS (LEVEL A AND B) .

For these conditions the stress analyses for the fatigue crack growth considerations in the previous section were used. The criterion which must be met for these conditions is K

g <

1 I

( 10 The most severe of all normal, upset and test conditions is the cold hydrotest condition, a constant temperature pressure test, generally carried out only at the beginning of plant life. All the other normal, upset and test transients are less severe, because no pressure exceeds the 3105 psi used for the hydrotest, and thermal stresses will always result in compressive or near zero stresses near the outer surface of the vessel.

The fracture toughness Kg, was determined from the reference toughness curves in Appendix A of Section XI, which were shown in Section 3. To use these curves, RT must be calculated, and in Section 3 it was shown that NOT the RT a en a was M F. D e leak test temperature for M NOT vessel is required to be at least 310*F for up to 5 effective full power years of operation [6), and thereafter will increase. Hydrotest temperature would be higher than this value, but at any higher temperature the fracture toughness would still be on the upper shelf, which has been set at 200 ksi/

in for this steel. -

Therefore using the criterion, we find E = 63.2 ksi/ in >K

/ 10 I and the maximum size of an allowable flaw will be that for which the maximum calculated stress intensity factor is 63.2 ksi/ in.

5-1*

The stress intensity factor for an outside surface flaw in the Indian Point Unit 2 reactor vessel was calculated using the expression published recently by Newman and Raju [7). This expression is applicable for a range of aspect ratios, and for this analysis the length to depth ratio was set at 2, even though the actual dimensions result in a ratio.of 1.4. This assumption provides some small conservatism in the results obtained.

1 The hoop stress distribution for an internal pressure was calculated directly l in the finite element analyses detailed in Section 4.

Results of the stress intensity factor calculations using the 3105 psi hydrotest condition are shown in Figure 5-1. Using the relationship and the l

previously calculated allowable stress intensity factor, we find that a flaw l with a depth up to 31 percent of the wall, or 2.67 inches is acceptable.

1 5.2 EMERGENCY AND FAULTED CONDITIONS l

l The emergency and faulted transient categories considered in this evaluation are:

Large Steam 11ne Break (LSL8)

Small Steamline Break (SSL8)

Large LOCA Small LOCA Steam Generator Tube Rupture (SGTR) l Since plant specific transients were not available, the generic transients i developed in References 8, 9 and 10 for the Westinghouse Owners Group (W0G) were used. All of these transients involve severe cooling of the inside vessel I

surface which results in compressive loadings near the outer surface. For j example, the low temperature and low pressure characterizing the small LOCA l

transient result in compressive stresses near the outer surface as shown in i Figure 5-2. For transients with low temperature and high pressure, there is a potential for tensile stresses to exist near the outer surface.

l 5-2

For all of the emergency and faulted transients considered, the lowest temperature at the outer half of the vessel well for which the stresses are tensile is approximately 240*F. This result is obtained from the evaluation of the small steamline break transient. Figures 5-3 and 5-4 show the temperature and stress distribution for this transient. The ta'aperature, 240'F, results in the fracture toughness g K , ucurring on W W sW which has been assumed to equal 200 ksi/ in.

using the Section XI criterion for faulted and emergency conditions (level C and D), we have

>K I

/2 or E = 141.4 ksi/ in >K g

/2 Stress intensity factor calculations were carried out for the worst case faulted condition, small steamline break transient, which includes a repressurization to 2350 psi, using the same methods as previously described for normal, upset a.ad test conditions.

For a postulated outside surface flaw.the results are shown in Figure 5-1.

The K g values for this transient never reach 141.1 ksi/ in, therefore the emergency and faulted conditions considered do not affect the integrity of Indian Point 2 vessel.

To verify the acceptability of the flaw indication relative to vessel in-tegrity following emergency and faulted conditions, the results of avail-able probabilistic analysis were reviewed. The review shows that on a probabilistic basis the affect of the indication on the risk of significant flaw extension or vessel failure is negligible. The details of this analysis are discussed in Appendix B.

l 5-3

1

'5.3 PRIMARY STRESS LIMITS In' addition to satisfying the fracture criteria, it is required that the pri-mary stress limits of.Section III, paragraph NB 3000 be satisfied. A local

.l area reduction of the pressure retaining membrane must be used, equal to the '

area of the detected indication. Specifically, two criteria must be met: ,

Pg +-Pb< l.5 Sm f r design conditions Pg+Pb < : 3.0 S, for nonnal, upset and test conditions.

To evaluate these criteria it was assumed that the indication extended along the entire length of the vessel, reducing the net section by 1.45 inches. Even under this extremely conservative assumption the indication can easily be shown to be acceptable.

From Table I-1.1 of NB 3000, S for A533B Class 1 steel at 600"F was found to be m

26.7 ksi. The applied surface stresses for the governing transients were taken directly from Table 4.2, and increased linearly in proportion to the reduction in area. For design conditions a pressure of 2500 psi was used. From Table 4-2, the maximum surface stress for a pressure of 2500 psi can be obtained from the results for " Hot Hydro Test", and can be shown to be a conservative value for P, + Pb . : After reducing the net section appropriately, we find that Pg+Pb = 32.14 ksi < l.5 S,= 40 ksi The second criterion, for all nonnal, upset and test conditions, is less limit-ing. The governing condition is- the " Cold Hydrostatic Test" found in Table 4.2.

After reducing the net section appropriately, Pg+Pb = 39.92 ksi < 3.0 S, = 80.1 ksi Therefore, the primary stress limits of NB 3000 are met.

l

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SECTION 6-EXTERNAL THERMAL SHOCK FROM CAVITY FLOODING

' Since the indication'in the Indian Point Unit 2 reactor vessel is located near '

the outside: surface, the effect of an external thennal shock due to inadvertent flooding of the reactor _ vessel cavity is of interest, because such an event would produce positive thermal stresses in the region of the indication. Even though such an event is not a design transient, it has occurred, at Indian Point Unit 2, and will be evaluated for completeness.

A detailed fracture analysis of such an incident was recently completed (11],

~

and those results willEbe reported here, to show that this type event is not of concern to structural integrity.

In the earlier analysis a flaw was assumed to exist at the junction of the lower shell and bottom head region, oriented axially. This is a somewhat more severe stress location than that of the actual indication. The cavity flooding water boiled as it came in contact with the vessel, which was assumed to be operating at steady-state pressure of 2250 psi, at 550*F. The stress and temperature results for_ this case are shown in Table 6-1, 1The fracture toughness for_this case is equal to the upper shelf toughness, which is 200 ksi/ in.- Using the emergency and faulted condition criteria,

'the allowable flaw ' depth is I!b = 141.4 ksi/ in >K g

/2 The stresses were linearized through the vessel wall thickness, and the stress intensity factor expression of Section XI of Appendix A was used for an external surface flaw. The stress intensity factor is presented as a function of flaw depth in Figure 6-1. It can be seen here that the allowable flaw depth is 1.77 inches, and therefore the external thermal shock from cavity

- flooding is not a threat to the integrity of the Indian Point Unit 2 vessel, even with the observed indication.

6-1 v-m.- --- -- - ------*- -a---

TABLE 6-1 TEMPERATURES AND STRESSES FOR CAVITY FLOODING CORE (REF 11)

Node Temperature: 200*F 284*F 370*F 458'F [550*F (Outer Wall) (Inner Wall)

Hoop Stresses: Outer Wall Inner Wall 84.4 ksi -41.3 ksi Linearized Stress Components e,= + 21.6 psi a b = + 62.8 ksi 6-2

e

. . 9 . - + - . . es. . . . . . . . . .y. .

6 * *

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h

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m_.__._.a.__...._

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+-

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h a x

e _-

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a___- . .. -_. _

b m.

s _

=

o.e M -

y -

E

. m N*

P '

.h

_.h__

Crack Depth / Wall Thickness Figure 6-1 Fracture Results of Cavity Flooding s

6-3

o l

SECTION 7 SUMNARY AND CONCLUSIONS An evaluation of the recently observed indication in the beltline region of the Indian Point Unit 2 reactor vessel has been completed. The indication is

-located near the outside surface of the vessel, oriented axially, and for the purpose of this evaluation is assumed to be a flaw. The through-esell dimension is 1.2 inches, axial length 2.0 inches and the edge of the indication is 0.25 inches from the vessel surface. The proximity of the outside surface requires the indication to be characterized as an outside suffaceflaw. The characterized dimensions of the surface flaw are therefore a depth of 1.45 inches and length equal to 2.0 inches. The characterized

[ depth is 16.8% of the well thickness.

The initial RTNDT for the longitudinal weld in the beltline region is estimated by the NRC thermal shock report [2] to be -56*F, and the copper content is 0.16 wt. percent. Using radiation damage estimation methods the change in RT was conservatively assessed to be 79'F. With this NOT information, along with the temperatures of the applicable transients, (Teh > 240*F) the vessel steel toughness will be on the upper shelf. in the region of the indication. Therefore, Kg ,= 200 ksi/in Kg , = 200 ksi/in The allowable crack size was obtained from IWB 3600 of ASME Section XI.

Specifically, for normal, upset and test conditions we have:

K Kg< = 63.2 ksi/i'n (normal, upset, test) l For emergency and faulted conditions:

l K g<[K = 141.4 ksi/in (emergency, faulted) 7-1 l

l

The stress 17.t:nsity facter, gE , is the driving fcrce fer the crack aftsr it

, is subjected to a fatigue crack growth analysis. The fatigue crack growth analysis of Section 4 indicates a growth of about less than 15 in depth for the remaining. lifetime of the vessel. This growth is essentially insignificant,andwasbasedonaconservative,up-to-dateset.ifdesign transients.

For normal and upset conditions, the worst case condition (hydrostatic test) I shows an estimated allowable crack depth of 315 of the wall thickness. The flaw as characterized is only 16.g5, so this criterion is set.

Fofemergencyandfaultedconditions,thegoverningconditionisasas11 steam 11ne break transient, as described in Section 5. The results of the fracture evalua't ion for this transient show that an external surface flaw l would nat be affected by this transient, since the applied K gnever reaches the allowable, so the criterion for normal, upset and test conditions will be

governing.

, .Therefore, it is concluded that the indication is acceptable without repair. i l

i l

l i

l 7-2 i

l o

SECTION 8 I 1

l REFERENCES l

1

1. NRCPolicyIssue,EnclosureA,"NRCStaffEvaluationofP[essurizej Thermal Shock," SECY-82-465 Nov. 23,1982.

2.

• Calculation of Operating and NTOL Vessel RT alues, i

NOT er

WDG-82-290, 8ecember 1982.

l

3. " Analyses of Neutron Flux Levels and Surveillance Capsule Lead Factors for the Indian Point Unit 2 Reactor," Westinghouse NTO letter SAO-RSA-655, July 1979.

4

4. WECAN, Westinghouse Electric Computer Analysis User's Manual, Westinghouse R&D, Pittsburgh, Pa., Sept. 17, 1979.

5. Dittus F. W., and Boelter, L. M. K., " Heat Transfer in Automobile Radiators of the Tubular Type," Calif. Univ. Publication in Eng. 2 No.13, pp. 443-461,1930.

6. Norris, E. S., " Reactor Vessel Material Surveillance Program for Indian Point Unit 2 - Analysis of Capsule T,* Southwest Research Institute Project 02-4531 Final Report, JrJne 30, 1977.

1. J. C. Newman and I. 5. Raju

• Stress Intensity Factors for Internal 4

Surface Cracks in cylindrical Pressure Vessels.* Trans. ASME, Vol.102,

huv. 1660.

t

8. Cheung, A. C., et. al., *A Generic Assessment of Significant Flaw Extension, Including Stagnant Loop Conditions, from Pressurized Thermal Shock of Reactor Vessels on Westinghouse Nuclear Power Plant," WCAP 10319, 8ecember,1983.

] 9.

• Summary of Sas11 Steam 11ne treak Analysis," perforised for the

! Westinghouse Owners Group, January 1984.

j 8-1 i

10. Meyer T. A., "Sumary Report on Reactor Vessel Integrity for Westinghouse Operating Plants". WCAP 10019. December 1981.

11. D. T. Entenmann, et. al., " Indian Point #2 Reacter vessel;3tructural Evaluation for Accumulation of Water in Containment Incident,'

Westinghouse Electric Corp., WCAP 9822, November 1980.

12. Jouris, E. N. and Witt, F. J., "An Application of Probabilistic Fracture Mechanics to Reactor Pressure Vessels Including Multiple Initiation and Arrest Events,' Transaction of the Seventh International Conference on structural Mechanics in Reactor Technology, August 1983.

l l

l I

1 1

j 8-2

t

' APPENDIX'A-

' TRANSIENTS DESCRIPTIONS ,

s LEVEL A AND B' TRANSIENTS - FIGURES A-1 TO A-46 LEVEL C AND:D TRANSIENTS - FIGURES A-47 TO A-51 1

I' e

f A-1

t

- _ a,c.e 7

E i

w E

N w

E i m

-1 0 1 2 3 4 5 TIME (H0URS)'

Figure A-1 Plant Heatup - Reactor Coolant System Pressure Versus Time A-2 I

z.

-' a,c.e

.g.

3 i

H.

-1 0 1 2 3 4 5 TIE (HOURS)

Figure A-2 Plant Heatup - Reactor Coolant System Temperature Versus Time A-3

a,c.e 2

E

. t.

-w E

w Y

U ac

-l 0 1 2 3 4 5 TIME (HOURS)

Figure A-3 Plant Cooldown - Reactor Coolant System Pressure Versus Time A-4

4 -

-L L

, ^ -

a,c.e I 1

t

.I'

- O 5

i w em i

e

! -M d

i I.

(-

0- 1 2 3 4 5 h TIME (MO.URS) 1 d

Figure A-4 Plant Cooldown - Reactor Coolant System Temperature-Versus Time i-t' t

A-5 t

, - , - . . ...y,,_. , , , , - , . . . . , . . . . , , , , , , , , , .,,, ._+._m...,w.,....,...,.ym- _

_.,-.,_.-<_.-.-..._,._,_-m.w, ,-.w.-.,.,.

l a,c.e l I

1 i

E E

p 5

E w

5 W

a.

f 0 500 1000 '1500 2000 2506 TIE (SECONDS)

Figure A-5 Unit Loading - Reactor Coolant Pressure Versus Time A-5

a,c.e b

a p

E

=

w-E

=

a o 500 1000 1500 2000 2560 TIME (SECOMOS)

Figure A-6 Unit Loading - Cold Leg Temperature Versus Time A-7

- a,c.e

.E.

E E

5 -

5* ,

Y R

O E.

o 500 1000 1500 200u 25TN TIE (SECONDS)

Figare A-7 Unit Unloading - Reactor Coolant Pressure Versus Time A-8

a,c.e C

• L E-S 5

w E

Cas b

W 0 500 1000 1500 2000 3do TIE (SECONDS)

Figure A-8 Unit Unioading - Cold Leg Temperature Versus Time A-9

a,c.e E

5 G

1 5

Y

R O

E 500 1000 .1500 2000 2500 0

TIME (SEC00105)

Figure A-9 Reduced Temperature Return to Power - Reactor Coolant Pressure Versus Time A-10

'.;w  : '

3p ..-. U ,y s *o s \ ,1 N

.. 3 fi C 7 3. . J.g '.

-, s - n ,\ r,

0; p- y -;.

i

%; p s;,.s 4: .,0 ;q s ,

• - g' N ,

~

is p

-A g $s k j !.

., .y ,

.i ( N 1 [q s Y,

\,. i i.

a,c.e 1 _ _ ( /. .M, y< -

. /'

u -

c .6 .

l It' i t 'N ,,s; I A, i , '

s

%- Q.

-- i s, '

0 ,

'- l ,g 1h x' f._ "

(

.1 , ,

/

• s . n, <t ya s e ..

- .s

-s..

3

,/ ..

5 + ,

- s > >

%, .}

\; e 1 g y ' T<j. +s ' ,p'; 'i .

j n, s f% g 6

f s. I s -

f , ,

1

' .7

u- : .

.c t

g ,

e *-

3s

/

, 'j .

4

/ A

. s J -

t g

. T((-- k. )' p. t' g Cd,: l' i, 'i c 3' ._

4 ; & -T A h, ,.; s. \500 2000 2500 Qg 1000 h1500

.s. , .

s

. . I/ o /4 TIME (SECORM) y t, n x y. fs . * *

,a 1,-

.,., (1 ,

. s- > ., \

.s .tw

.c '

4

-t g .g , ..,, ,

s. '

. ' *4 .

Figure A-la Reduced Tegerature Return to.Ptwer - Cold Leg Temperature '

q- r Versus .Tinie,' . ,3 k

-L g t g r I m

• h.

g ( *< 5 Ig k [' I .,

-I p,  %

g f

( \ i

.., y

• t- '

,'t A4s ,

s . s . si  %

• 1.; i g,. i \ 'f

.\ I f e. ,,- i y , - . . _ . . _ . . . . _ _ ., /

,, i

. . _ _ a,c.e

=

E s

E E

5 5

b I

l ,

a a a s .s r E N N N N I N TIME (SEC00605)

NOTES: 1. TRANSIENT ASST 4ED TO BEGIN AT 90 PERCENT OF FULL POWER.

2. AFTER TRANSIENT IS COMPLETED, PLANT WILL BE AT PROGRAMED OPERATING C0fCITIONS AT THE INCREASED POWERLEVEL(100PERCENTPOWER).

Figure A-11 Step Load Increase - Reactor Coolant Pressure Versus Time

\

A-12

~

a,c.e s

C u

^

E P

5.. ,

5 1\

I w

an.

a E 8 8 8 Y e

o arw i. i e g g i

. 3 TIME (SECONDS)

Figure A-12 Step Load Increase - Cold Leg Temperature Versus Time A-13

- a,c e

~

=

i E l s

l C

! 5 E

I Y  ;

e s s-8 8 8 8 l

2 E  !! N I E E TIME (SECONOS)

!- )

l t

I NOTE: 1. TRANSIENT ASSLMD TO BEGIN AT 100 PERCENT OF FULL  !

POWER.

l

2. AFTER TRANSIENT IS COMPLETED, PLANT WILL BE AT PROGRAMED OPERATING CONDITIONS AT THE REDUCED POWER LEVEL (90 PERCENT PWER). l Figure A-13 Step Load Decrease - Reactor Coolant Pressure Versus Time A-14

i a,c.e C

o U

E.

a 5

5>

I I

e 8 8 8 8 8 8~

• i i E, 8

o m e 5 E 2 TIME (SECONOS)

Figure A-14 Step Load Decrease - Cold Leg Temperature Versus Time A-15

l l

- a,c e l

l l

C E

., E 1

S E

w E

w 2

0 200 400 600 800 1000 1200 TIE (SECONDS) 1 Figure A-15 Large Step Load Decrease with Steam Dump - Reactor Coolant Pressure Versus Time

.R A-16

a,c.e C

.' E C

5 E.

3 Y.

E E

0 200 400 600 800 1000 1200 TIME (SECONDS)

Figure A-16 Large Step Load Decrease with Steam Dump - Cold Leg Temperature Versus Time A-17

a,c.e 5

c 2

E

=

w B-

-E

.E E

..- 5 S~

E E

E,

=

E E

as i

5 4

i Figure A-17 Steady State Fluctuations - Reactor Coolant Pressure and Temperature Versus Time L

, A ,. .

a,c.e l

=

2 E

r-f E,

w 5

w E

~

.0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 U TIME (H0URS)

Figure A-18 Feedwater Cycling - Reactor Coolant Pressure Versus Time A-19

4 a,c.e l

l I

1 1

W

-5 G.,

5

-5 W

8 E

as e

l

. 0.0 0.5 1.0 1.5 , 2.0 2.5 3.0 3.5 TIME (HOURS)

Figure A-19 Feedwater Cycling - Cold Leg Temperature Versus Time A-20

- a,c,a C

E' 5-G 5

E

-W R

E E

0 100 200 300 160 0 500 600 TIME (SEC0 LIDS)

Figure A-20 Loop Out of Service, Normal Loop Shutdown - Reactor Coolant Pressure Versus Time A-21

a,c.e l

O ow >

5 P

=

5 E

i!!

E E

e i

L -

0 100 200 300 400 500 600 TIME (SEC00lDS)

Figure A-21 Loop Out of Service, Nonnal Loop Shutdown - Cold Leg Temperature Versus Time A-22

_______.._____.___.__.s._____ . _ , -

- a,c.e O

2 p

5 E

n-w 4

3 w

f

~

200 300 400 500 600 0 100 TIE (SECONOS)

Figure A-22 ' Loop Out of Service, Normal Loop Startup -

Reactor Coolant Pressure Versus Time l

l l

A-23

a,c.e i l

f- l E l p l

=_ i E

.Y.

E E

w I

s 0 100 200 300 400 500 600 TIE (SECONDS)

Figure A-23 Loop Out of Service, Nonnal Loop Startup - Cold Leg Temperature Versus Time 6

A-24

a.c e 1

E:

,. 5

- _=

- E.

=

t

0. 25 50 75 100 125 150 TIE (SECONO3)

Figure ' A-24 Loss of Load - Reactor Coolant Pressure Versus Time A-25

- a,C,e m

5

.c 2

5 -

W R

2 E

e 0 10 20 30 16 0 50 60 70 80 90 15 TIE (SECONDS)

Figure A=25 Loss of Load - Reactor Coolant Temperature Versus Time A-26

s a,c.e C

.E,

.-- E r-2 5: .

w .

E -

y ..

a.  : ...

0 2000 4000 6000 8000 10000 12000 TIME (SECONOS)

Figure A-26 Loss of Power - Reactor Coolant Pressure Versus Time A-27

a,c.e C

t E

E 2

=

ll>

w 5

4 5

- 0 2000 M00 6000 8000 10000 12500 TIME (SECONDS)

Figure A-27 Loss of Power - Cold Leg Temperature Versus Time A-28

s-

.. _ - a,c.e

=

E_

E i:

I w

3 Y

Figure A-28 Partial Loss of Flow - Reactor Coolant Pressure Versus Time A-29

-. a,c.e C

o U

s c

5.

5

=

a:

k W

5 5 5 $ $

i i i i e i TIME (SECONOS) ' . __

Figure A-29 Partial Loss of Flow - Cold Leg Temperature Versus Time A-30

a,c,e -

1 h

=

E 5 2

= -

5 '

w E -

N '

W

a. -

1 l l 8

• [ 8

• g 8 8-a . . . .

6 E R R 8 E

_ R_ R 8

~

TIME (SECON05) i 2

NOTE: PRESSURE IS RETURNED TO THE INITIAL VALUE AFTER 9 MINUTES, AT A RATE CONSISTENT WITH .,

NORMAL PRESSURIZER HEATUP.

=;

a Figure A-30 Reactor Trip with No Cooldown - Reactor Coolant Pressure '

Versus Time au M

u A-31

e i

~ ~

a,c,e C

8 e

5 c

5.

E Y

a E

W 8 a M 8 8 8 8 8" o

o S o o o o o d 5 d N N N N N TIME (SECON05) .

Figure A-31 Reactor Trip with No Cooldown - Cold Leg Temperature Versus Time A-32

_ a,c.e

.E.

E 2

E:

E R

O E

i L.

0 25 50 75 100 125 15 0 175 200 TIME (SECONDS)

Figure A-32 Reactor Trip with Cooldown, No SI - Reactor Coolant Pressure Versus Time A-33

a,c.e C

t 5

e 5

4AJ E

C 5

k 75 100 125 150 175 2h 0 25 50 TIME (SECONDS)

Figure A-33 Reactor Trip with Cooldown, No. SI - Cold Leg Temperature Versus Time A-34

~

.- a,c.e O

E 5

a 2

E

=

a l0 E

0 100 200 300 400 500 600 700 800 900 1000 TIME (SECONDS)

Figure A-34 Reactor Trip with Cooldown and SI - Reactor Coolant Pressure Versus Time A-35

- ~

a,c.e C

o_

E p

5 E

E E

I E

5

~

0 100 200 300 W O 500 600 700 800 90010b TIME (SECONOS)

Figure A-35 Reactor Trip with Cooldown and SI - Cold Leg Temperature Versus Time A-36

- a,c.e

=

E E

3:

5 .

E E

}n.

8 8 8 8 8 8 8 8-

• 8 &

a - 5 .in i.

5

~ 3 TIME (SECONOS)

Figure A-36 Inadvertent RCS Depressurization - Reactor Coolant Pressure Versus Time h

A 17

l

(

a,c.e e

u 8

ii w

i 5

3 .

l-8 8 8 8 8 8 8 8-I N N N N N N N N TIME (SECONDS)

NOTE: TEMPERATURE 00ES TO COLD SHUTDOWN VALUE CONSISTENT WITH NORMAL C00LDOWN itATE.

Figure A-37 Inadvertent RCS Depressurization - Cold Leg Temperature Versus Time A-38

r. .. -

, s t

i

}

' m

+

g,,.

s .

, _\

s.

N ' -

s

.. y s s 1 .c a,c.e i% s 4 *%t 9

\

=

G s;= '

M

\

=

hpJ -

g . .

w s

1 , s

,t ,

3/23/F7 I. I. I.

\ I. I. I. $. f e' -c 3 c , . _8 C,_ S_ C TIME !3EC0405) y i '

s t

, ,s. .

~

Figure A-38 sInadve'rtent Startup of an Inactive Lcop - Reactor Coolant Pressure Versus Time 3 t . \

\

[ '

L ,

x, L. , , ,,

l\ 6 s 's

'% g  % ..

A-39 s '

y

,<t

l

~ - --

a C,e C

0 9 -

E 5

4>

w h

5 E

=

wwn

. I. a. 8 8.

8 8 8.

f.

• E b $

!  ! k TIME (SECONOS)

Figure A-39 Inadvertent Startup of an Inactive Loop - Cold Leg Temperature Versus Time A-40 t

c .,

t- ,

_ a,c.e s

' t 3rs 2

.m E

, b '

,. =

4 E

w E

m m -

g i:

a.

lI a

3 8 8 8 8 8 8-

• $ $ 5- $ $ $

o <w o ~ o $ c TIME (SECON05)

_' t_

Figure A-40 Inadvertent SI Actuation - Reactor Coolant Pressure Versus Time lI. ,

)

.?

t P

F i.

A-41

- - a,c.e i

i C

S e l 5

C" 5

E E

a:

k W

8 8 8 8 8 8

=

• ~ g s

~

g g

a e

TIME (SECONDS) l Figure A-41 Inadvertent SI Actuation -' Cold Leg Temperature Versus Time A-42

~

a ,c.e

=

E s

E E

I N

Y 8 a i & s-I $ $ $ $ $'

TIME (SECON05)

NOTE: PRESSURE RETURNS TO THE IMITIAL VALUE AT A RATE CONSISTENT WITH NORMAL PRESSURIZER HEATUP.

Figure A-42 Control Rod Drop - Reactor Coolant Pressure Versus Time A-43

4

.i

- a,c.e 1

.c

~

V 8 r-2 5

3 y

' ~

0 100. 200 300 40 500 i TIE (SECON05)

.s, t

' Figure A-43 Control Rod Drop - Cold Leg Temperature Versus Time N

.-A-44

. . . . . .-. - . - - _ . . ~ -. , . . . . . - - - .- _ . . ..

1 I

. a,c.e

=

.- g c:

E w

E E

8 8 8 8 8 8 8 8 8 8'

= i 8 i 4 i i i 4 4 o - .v m i. M un ~ as M 5_

TIME (SECONOS)

NOTE: PRESSURE RETURIls T0 INITIAL val.DE AT A RATE CONSISTENT WITH N0fDMt. TRESSURIZER HEATUP.

~ Figure- A-44 Excessive Feedwater Flow - Reactor Coolant Pressure Versus Time A-45

a,C,e

'C o

a_

s s  :

=

ha s

2 f

w 8 8 8 8 8 8 s s 8 8-a s g g i

i

~ i. .s i

- .s i g l TIME. (SECON05) -

i

l l  !

Figure A-45 Execssive Feedwater Flow - Cold Leg Temperature Versus Time 1

I t

A-46

1 1

- a,c.e

• \

l l

C g- l E  :

C 2

E I

w 5-E w

E F_

E E

S E

m-w E

=

E a.

5 w

~

0 200 400 600 800 10I0 TIME (SECONOS)

Figure'A-46 ~ Turbine Roll Test - Reactor Coolant Pressure and Temperature Versus Time A-47

_ a,c.e

. l-

-1 i%

i

'E

~!

E l 2 l

~ .

I l

. G.

O 500 1000 1600 seco TIME a,c.e

.E a

0 E

0 500 looo 1500 scoo TIME Figure' A-47. Large Steamline Break - Temperature and Pressure

+

Versus Time

.A-48

a,c.e l

-o 200o 4ooo sooo sooo toooo troio TIE a,c.e I

71K Figure A-48 Small Steamlint Break - Temperature and Pressure Versus Time A-49

, _ . .. _ ~- -.

. - a; 9

a,c,e w

-E ti 5

-y w

t-= .

~

0 100 1000 2000 3000 4000 TIME

-_ -- a,c.e i

r.

TIME Figure A-49 ' Large LOCA - Temperature and Pressure Versus Time A-50

- a ,c.e

, E 3

w i -

0 1000 2000 3000 '4000- 5000 l TIME

- - a,c.e I'

-E a

.=

E i

L l-

, '1000 2000 3000~ 4000 5000 TIME Figure A-50 ~ Small LOCA - Temperature and Pressure Versus Time A-51

v

_ a,c.e 1

l 1

N i'

5 E

'5 n

1000 2000 3000 4000' 5000 '

TIME F

a ,c .e U

R

'O

-E

~ -

1000L 2000 3000 4000 5000 TIME Figure A-51 Steam Generature Tube Rupture - Temperature ard Pressure Versus-Time A-52

- . APPENDIX 8 PROBASILISTIC ASSESSMENT

~

The results of a Westinghouse development program [1,12] can be used in order to assess the deterministic results obtained in the previous section. Using probabilistic fracture mechanics (PFM), an evaluation was perfonsed during 1983 to quantify the risk of vessel failure from transients with high pressures while at relatively low temperatures. The probabilistic model used in this evaluation is similar to the NRC model [1] except for the following:

o Constant Flaw Site In the NRC_ model, a distribution of flaw sizes was used, however, it is more appropriate in this evaluation of an indication to use a constant value estimate of tSe indication. Also in this evaluation, the flaw is assumed to be semi-elliptic with an aspect ratio of 1 to 6 on the inside surface, however, for transients in which the pressure stressas are predominant the stresses can be assumed to be constant through the thickness, therefore, an indication near the outside surface can be modeled by a flaw on the inside surface.

o K gg and Kg , Cunes The K gg and Kg , curves developed by Westinghouse were used [2]

since it is more conservative in the higher transition region.

o Fluence Attenuation The fluence distribution included the effects of dPa. This is especially important when evaluating flaws near the outer surface.

l The surface fluence used in the PFM evaluation was 1.9 x 10 '

n/cm2 which is higher than the surface flaw fluence predictec at end-of-life (EOL) for the longitudinal weld [2].

References 1 and 12 contain a much more detailed discussion of the PFM model and assumptions.

3-1

From the many transients which were evaluated in the d:velopment program, the transient which best models the ses11 steamline break (found to be the most limiting o'f the emergency and faulted conditions in sectica 5.0) is represented by a cooldoWh fren 550*F to 200*F at 100*F/hr with a pressure of 2560 psi. Using the previously discussed assumptions, the conditional probability of significant flaw extension was found to be 8.0 x 10 occurrences / reactor year given that the event occurs. This is a very 1

conservative number because (1) it is obtained using an RT which is much NOT greater than the predicted E0L RT for the longitudinal weld of co'cern n NOT

[2], (2) the frequency of the event occurring has not been taken intc consideration, and (3) the potential benefit of arrest of a propagating flaw is[notconsidered. Therefore, the results of the probabilistic analysis support the conclusions of the deterministic analyses in that the risk of vessel failure or significant flaw extension from the indication is negligible.

i 1

l l

l l

l P

I l

3-2 i

- - - - - - . _ . _ . . . . _ _ _ _ _ _ - . _ - - _ _ - _ _ _ _ _ _ _ _ _ _ _ _