ML20079C409

Feedwater Nozzle Cladding Crack Repair Rept

ML20079C409
Person / Time
Site:	Monticello
Issue date:	12/31/1975
From:	Charnley J GENERAL ELECTRIC CO.
To:
Shared Package
ML20079A872	List: ML20079A870 ML20079C409
References
NEDC-21120, NUDOCS 9106190411
Download: ML20079C409 (30)

Text

- _

NEDC-21120 SPECIAl, REPORT Class II November 1975 75NED63 MONTICELLO FEEIMATER N0ZZLE CIADDING CRACK REPAIR REPORT NUCLEAR ENERGY SYSTEMS DIVISION GENERAL ELECTRIC COMPANY FOR NORTHERN STATES POWER COMPANY J. E. Charnley

/

Approved:

O93 4 I . R. Kobsa , Manage r Operating Plants Reactor Assmbly and Performance Evaluation Docket 50-263 BOILING W ATL H ftE AC10H SYSTEMS DE PAH TMENT e GENC H AL i LLCT HIC COMPANY SAN JOSE. CALIF ORNI A 95125

"$f2%" GENER AL h ELECTRIC 9106190411 752231 PDR S ADOCK 05000263 pop

NEDC-21120 l

TABLE OF CONTENTS LaLe.

1.0 INTRODUCTION

1-1 2.0 SUltiARY AND CONCLUSIONS 2-1

3. 0 DISCUSSION 3-1 3.1 Description of As-Found Conditions 3-1

3. 2 Repair of Indications and Description of "As-Left" Conditions 3-2 3.3 Stress Evaluation on "As-Left" Conditions 3-2

3. 4 Fracture Mechanics Evaluation 3-8 3.5 Fatigue Analysis-Crack Initiation and Crowth 3-14

3. 6 Grind-Out Weld Repair Evaluation 3-21 3.7 Corrective Actions 3-22 3.8 Corrosion Evaluation 3-23 4.0 SAFETY EVALUATION 4-1 S.0 REFERENCES 5-1 i

NEDC-21130

1.0 INTRODUCTION

The f eedwater nozzle corner radii of the Monticello Reactor were inspected for cracks during the scheduled outage which began on September 11, 1975 (based on recent experiences at the Millstone 1. Dresden 2 and 3. Quad Cities 2, and Browns Ferry 1 plants). The dye penetrant examination of the cladding on the feed-water nozzle corner radii showed many linear indications. All four nozzles had indications. The total number of indicatior was .pproximately 180. These indications were completely remcved by grinding to a maximum depth of approxi-mately 1/2-in. f rom the origina t cladding surface. The deepest penetration into base metal was 1/4-in.

This report describes the cracking, the method of removal and final surface con-ditions; stress evaluation of the effects of the grindout cavities; a probable cause of cracking; a fracture mechanics evaluation of crack penetration into base metal; and the corrective actions taken.

l 1-1/1-2

NEDC-21120 2.0 SLMt\RY AND CONC 1.US10NS Examination and analyses of the feedwater nozzles indicate that the cracking in the vessel shell to nozzle radius most probably resulted from high cycle thermal fatigue attributable to excessive bypass flow around the feedwater sparger thermal sleeves, it is believed that this bypass feedwater flow caused rapid temperature fluctuations in the af f ected area, a phenomena which will be reduced by the interference fit feedwater spargers which were installed after the cracks were ground out. The replacement spargers are similar to those in-stalled in Quad Cities 2, Millstone, and Dresden 2 and 3.

A f racture mechanics analysis based upon ASML Section Xi 1Q74 edition with addenda to and including Sumner 1975 addenda was perforned to determine tha permissible flaw depth and in no case did the detected cracks equal or exceed the end-of-life allowable flaw. Stress evaluations indicate that these grindouts do not violate the ASME code design rules.

Periodic surveillance of the nozzle corner will be performed.

The methods employed in removing the cladding cracks and the corrective act ion associated with the replacement feeawater spargers to minimize the probability of a recurrence of such defects assure a safe return to f ull power operation for Monticello.

2-1/2-2

NEDC-21120 3.0 DISCUSSION

3.1 DESCRIPTION

OF AS-FOUND CONDITIONS 3.1.I Methed of Examination The reactor pressure vessel feedwater nozzle inner blend radius was inspected in compliance with Field Disposition Instruction 323/51847.

Following removal of the teedwater spargers, the Reactor Pressure Vessel (RPV) wall and feedwater nozzles were cleaned using a high pressure hydraulic cleaning process (hydrolaser) followed by cleaning of the feedwater nozzles using " flap-per" wheels, s

Dye pene t ont inspection of the inner blend radius was conducted in accordence wi th ASME Code Sec tion XI 1974 edition with addenda to and including Winter 1974 addenda, Paragraph IWA-2222 with the acceptance criteria as stated in AS ME Code , Section Ill 1974 edition with addenda to and including Winter 1c,74 addenda, Paragraph NB-5350, but with no linear indications perwitted.

During the etack repair program and for the f inal examina tion the same examination criteria and methods were employed.

..).2 litpid Pene t ra n t Test Results Th e first PT revealed indications on all f our no.zles with a t .al of 66 Indica tions. Metal was ground away in approximately 1/16-in. inctements followed by a PT. The total number of indications increased as longer indi-cations branched. The maxinun number of indications, appicximately 180, were found after the four th PT. The fourth PT corresponds to a depth of approxi-nately 3/16-in. Fif ty-three grindouts penet rated into base netal. The final PT did not reveal any indications. The maximun total depth of e, rind-out was approximately 1/2-in, and the maximum depth of base netal removed was 1/4-in.

I 3-1

NEDC-21120 All but 12 of the indications were radially oriented with respect to the nozzle centerline. The 12 were circumferential1y oriented and followed the ovet lay clad weld bead.

3.2 REPAIR OF INDICATIONS AND DESCRIPTION OF "AS-LEFI" CONDITIONS 3.2.1 Method of Refag Hand-held grinders were u;ed to remove metal in the localized area of the crack.

Grinding was done in 1/16-in. increments in base metal, and in either 1/16-in. or 1/8-in, increments in cladding. After each increment a dye penetrant test was per f ormed to aetermine whet her the crack had been removed, and the location of temaining indications. Grindeuts that were suspected of penetrating into base metal were acid-etched using Nital after crack rer. oval.

Af ter the indications were removed the bottom of the cavity was ground to a radius of twice the total depth of the cavity, and the sides of the grind cavities were ground to blend emoothly with the surrounding cladding metal surface with a minimum slope of 4 to 1 in base metal 2 to 1 in the cladding. A PT was perf ormed on the blended cavity and final determina t ion of clad base metal interf ace was made using demineralized water.

3.2.2 Description of "As-Lef t" Conditions All base metal ground areas were measured and mapped as shown in Figures 1 th rou gh 4.

3.3 STRESS EVALU ATION ON "AS-LEFT" CONDITIONS The Reactor Pressure Vessel Stress Report contains an evaluation of the feed-wa ter nozzle c.inimum design dinensions at the nozzle-to-shell junction, per-f ormed in accordance with ASME Code Sec tion 111 1965 edition including Summer 1966 edition, Pa ragraph N-4 50. It is assumed therein that 1/16-in.

depth of the unclad base metal surf aces is nonexir. tent, as corrosion allowance.

The calculation shows the requi red nozzle reinf orcement area to be 28.30 in. ,

compared to available reinforcement area of 31.20 in. Die st ress report 3-2 l

NEDC-21120 MONTICE LLO o Feedweter Nossia O Inner Blond Redaws

- ~ ~

v e f.s ym,' N

-t- M j

y CC D \

~

j/ ~%,'^%'Nx

/

I 270" - -

/ * --

f -- 9 0 N

\y \x e a

\ \

'~

,, 3 p

%2fgW //

'y j ' l12 180" 2 Tl4 "

R a d iit s D -

P = nasimus deg t h

/

s- U 1,0 inch s pacing

9. 0" =-

Fi gu re 1. 45' Nozzle Base Metal Grindouts 3-3 l

NEDC-21120 i

l l

MONT8 CELLO o l Fee 1weier Nostle 0 inner B6end Radius l

p t 'lig '

F  % s t */s g N

/

Os 'Is f

/ '

D* 'lS 2 o o m , .. -.._ _ . _ - .

90 I . 9 * '/1g u: 'Is t ,

l}

g \

/ /

De '/rg g __

180*

2 M4 "

\

B a d his p 4 l_ _ *

! 0 = mantinue depth l  %

g b ,

1. 0 inch s pacing

• 9, 0" =

t Figure 2. 135' Nozzle Base Metal Grindouts l

l 3-4 l

~

NEDC-21120 MONT ICE L LO 0

Feodwater Norsie 9 inner Blerd Rad 6us

p* h8

, .f ,,

ou f , " ' %,

p((

~

-+ x .,,

/ ~

w e~

/ ,

3 s ' Igg

- " - - -- - - I ' - - -+- --

- 90

-} -

2 */ 0 ,

h  ! ,

\ ,t u e v,g 4 '

o, f3g \

\ / e v, i

D. ,y

/

N ,.Q. ' Q C

,/

• /

-/

~

c > .r 180 2%"

as B a dj% l

1. - 0 = ma n i na,' dirtb l

be 1. 0 inch s pacing

• =

9. 0" f

I 1

Figure 3. 225' Nozzle Base Metal Grindouts l

l 3-5

NEDC-21120 l

MONTICE LLO Foodseter Nostle 0, trww 81ond bdius

'n : 4

, ,}. -

7, ~

9 e '/,,

O r ig 0 . a_ - oD 270 - ,

\

N j[o, D' gg ' D e '/3 g 180 2 9. ..

l

) Parliate h

~~

b .

\

O = maximum depth

\ ,

1. 0 inch s pacing 9,o" _

l Figure 4. 315' Nozzle Base Metal Grindoute 3-6

NE DC-21120 4

does not account for all the available reinforcement area. An additional 2.2 in. Is available. One-half of the reinforcement area n us t be on each side of the nozzle centerline.

I l

Removal of base metal from the nozzle corner increases the required reinforce- '

ment area by the amount of low-alloy base material removed, measured as the largest cross-section area of removed metal lying in a plane through the nozzle centerlinc. Additionally, as the low-alloy steel is no longer clad at this location, l6-in, depth more than was actually removed is assumed non-existent for cor.>sion allowance, as above. All 53 grindouts that penetrated into base metal were mapped. The nozzle corner radius location with the largest base material removal is shown in Figure ',s. As indicated. the additional required reinforcement area is approximately 0,64 in. Since the available excess reinforcement area is (31.20 - 28.30) = 2.90-in. , the additional rein-forcement area required as a result of this repair is approximately 22% of that which could be removcJ without violation of the original code calculation and 0.64/2.9 + 2.2 = 13% .f actual.

??s"NozztE ,

2.94

/

I CEPTH = J$

AREA 45

) 7*

~.

r N l

N

~

, 0,19 CLAD

(

L N R = 2 5/16 l

N N

/ GRINDOUT ARE A * .45 m.2 CORROSION ALLOW ANCE ARE A = .19 in.

TOTAL ARE A RE MOVE D = .19 + .45 6n.2 l

l Figure 5. Largest Base Material Removal l 3-7 l

NEDC-21120 3.4 FRACTURE MECHANICS EVAtt% TION All indications have been removed by grinding and the grindouts have been blended to the outside surface using a minimum of 2:1 taper. The root of the grind cavity has been blended to provide a smooth radius with no sharp breaks in con-tour. Penetration into the base material occur red at soma locations and the maximum depth of penetration into the base metal was 1/4-in.

The allovable flaw depth for indicai. ions in cladding surf aces for inservice exan.inations is 1/8-in, as prescr ibed in IWB-3517,Section XI 1974 edition with addenda to and including Summer 1975 addenda, ASME Code . For cracks which extend through the cladding into the base material, a f racture mechanics evaluation is also necessary if the crack is deeper than allowed by IWB-3512.

The applied stress intensity factor is due to a combination of the pressure hoop stress and the themal stress. The stress intensity factor due to pressure alone is given by K =

F(a r) g [5 where K = stress intensity factor due to pressure, o = hee;. . t rcs; i.. tha vessel, h

a = flaw depth measured f rom tii. nozzle corner, F(a,r) = geometrical factor given in WRC-175 (Reference 3).

PD 1.000 x 20(>

g h 2t 2 x 5.0625 20.4 ksi The stress intensity factor K due to pressure is plotted in Figure 6.

g Consider the contribution from thermal stress to the applied stress intensity factor.

In a thermal analysis performed f or (Reference 2) the Millstone feedwater nozzle, the maximum themal stress at the nozzle corner was calculated to be 44 ksi for a s tep change in feedwater temperature trom 546*F to 100*F. Since this analysis 3-8 .

i

. . NEDC-21120 l l

l i

l l

1 r - - - - - ~ ~

l g i

, i

. -4 . . , -

l l __

' \

' i I

l

. I e 4 4 4

! I l i ,

i

? I i

' ( 8 1 t l

l

[

i '

L e, . . .. I

i e i

l-i.

j , ,

{ . , - . . . _

i.

! }

l i W

t t

' e i  ! u l U

_.1' , . , _ _ _ . .-. . .

. .i

\

2

. i ..ig . _ 1., , _.

. . 9.

l ' .; ,I A 2 u e , t H 1

i m

. . - - . - - - . 7 ., , 4 u,  ; . _ . .

_ y .- g l I e i

• w c

i -1 4 .

{ i g w c.

.. i . - - J 5 a,

" i O, b. m m y y i W W l

M t

i i d m

, l &

i- . . -, . .1 , . u &

I 2

1 , m

' j '

m s

v 4

. - - . 6 . . , C, t N

{i ,

( t i

4 i

I c L

} D

, ^ \, \ u i e I j f '

{- 4 3

~

i l  :

r r er

\

w' i

N o, I

g_ g -  ! -.

i c e e d 6

'u} jsz

• Awo 3yttss3yd 013:1C X H013YJ L11'.N3.tN1 SSTHIS .)

Z/L 3-9

NEDC-21120 i

did not take the presence of the the6 mal sleeve into consideration, the calcu-lated thermal stress is conservative. The same calculated thennal stress of 44 kai is conservatively assumed for the present analy9is also.

The procedure for calculating the stress intensity f actor due to the t he rma l stresses is not clearly defined in the literature. lwo different methods will be used to characteri:e the thermal stress intensity f actor.

3.4.1 Uniform Thermal Strestjiethod t This method assumss the surface thermal stress o g to be unif ormly distributed through the entire thickness. This is very conservative and provides an upper bound on the applied stress intensity factor K . This is given by K = 1.1 o g ,U = 1.1 x 44 rU where the Loress intensity formulation for a surf ace flaw under uniform stress is used.

3.4.2 Equivalent Hoop Stress Method in this method the ef fect of the thermal stress is considered by calculating an eouivalent hoop stress and using the stress intensity formulation for a nozzle under pressure loading. The equivalent hoop stress is obtained by dividing the nozzle surface thermal st ress by the stress index chich is conservatively h

assumed to be 2.5 (Ref erence 3).

o eq

c th/2.5

2.5

= 17.6 ksi The thermal stress intensity factor is then K

g

= F(a,r) c .E By calculating the equivalent hoop stress we are assuming that the thermal stress distribut!cr. is similar to that c iused by the stress concentration due to pres- l l

l I

3-10

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

NEDC-21120 i sure loading. This is a more reasonable assumption than considering the thermal ,

stress constant. A detailed strees analysis is expected to indicate that thermal l stress decreases more rapidly than pressure stress, but models are not available to determine the distribution.

i The total applied stress intensity factor K is obtained by adding the contribu-tions due to the pressure stresses and the thermal stresses.

i Figure 7 shows the total applied stress intensity f actor calculated by the two different techniques. The thermal analysis performed in Reference 2 as well as the actual temperature measurements have shown that the nozzle corner is at a temperature high enough to retain a toughness level of 200 kai-in. .

The critical flaw size a is then at the intersection of the 200 ksi-in.1/ line with the applied stress intensity curve.

According to the first method where the thermal stress is assumed to be unif orm, the critical flaw size is 2.05 in. It must be remembered that this calculation is based o extremely conservative assumptions that thermal stress is distrib-uted smift rmly with flaw depth. Even the calculation of the thermal stress ,

value is conservative since the beneficial ef fect of the thermal sleeve is not included.

Figure 7 also shown that for the second method based on the equivalent hoop stress technique the critical flaw does not occur (i .e., the crack would penetrate the wall and leak without growing unstably) . ,

The fact that a nozzle corner flaw would " leak before break" is supported by the results of the HSST program. For example, in the test on vessel 5 at 190'F with a nozzle corner flaw of 1.2 in. , f ailure occurred at 2.75 times the ASME design i pressure and the mode of f ailure was leakage. The "ler.x before break" feature provides additional assurance of safety, 1 3-11

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

9 NEDC-31130 -

I I

v* ,

'] l e- 1

$ o 3 d

\

g. . . ,

..~

4 o 5!- I

. e 3 o

$ ~

e=

)

• i

. A t +

i I o"  !

[a 8

+ h

~ f f$ ,

!E 5 EE

,. $ . . f 2a

• e g

E - .

id - =

l l$ 0  % i ,

E 4 E k l h l4 \' l 1 i

e x I, l

o E

iE 5 8 N IM o w I

lE  ! D i $

% i 9 L' E g N I i e u

\ e M l

I

.g . . . . . j

' } o, "

i h

s l s . I 1  %

• l v , N s

j s i e w

I N

's \ l a

I 'o,

\ . o. c i s i

54 m. t l '

,N, l

• Ds

t .

h

, t.,T .,

! _ _ .. ___ _ . _ _ _ . _ __ a J . **

• w ~ _

8 n

R 8

89 R n

gg 'ui-1sq 'n Altsg31g] ss3yis g333agy lV101 i

I l

3-12 l

NEDC-21120 It can therefore be concluded that the critical flaw size is not less than 2 in. and is probably closer to 8 in. The exact value can be determined only by detailed stress analysis.

According to Section XI ASME Code the allowable end-of-life flaw size can be obtained by dividing the critical flaw size by a factor of 10 or alternately by calculating the critical size for the case where the toughness is taken to bc )__ tines the actual available fracture toughness.

,- 10 Because of the non-linear dependence of the stress intensity on the local stress these two approaches are not equivalent for the nozzle corner crack problem. Ibpending on the method used the allowable end-of-life flaw size ranges f rom 0.17 in.

1 to at least 0.8 in. (g the wall thickness) as shown on Figure 7. The equivalent hoop stress method is the more realistic .nethod and therefore

, based on a conservative interpretation of Section XI the allowable end-of-life flaw size is 0.8. Thus all observed cracks were less than the end-of-life allowable.

Since the critical flaw size is greater than the wall thickness, a flaw cannot grow to the critical size. Thus the end-of-life allovable flaw size cannot, strictly speaking, be determined and the criteria are not applicable, i

l 3-13 1

L _-

NEDC-21120 3.5 FA11GUE ANALYSIS-CRACK INITIATION AND GROWTH Temperature fluctuations with ranges up to 125'r and f requency of up to 1 Hz, have been observed in the vicinity of the blend radius of the feedwater nozzle in tests run under normal reactor operation at Millstone with feedwater spar-gets similar to those seplaced at Monticello. This design used a slipfit be-tween the thermal sleeve and the nozzle which permitted bypass leakage past the thermal sleeve. It is believed that the thermal cycling observed at Millstone resulted in part from the bypass leakage flow into the nozzle. Cracks were dis-covered in the Millstone feedwater nozzle cladJing near the point s where sig-nificant therasl cycling was observed. Analysis also showed that t hese cracks could have been initiated by thermal f at igue resulting f rom the observed tempera-ture cycling.

Although temperature measurtments have not been made at Monticello, it is be-lieved that Monticello has also experienced thermal cycling comparable to that observed at Millstone because of the similar slipfit spatger design used at Monticello. He a t transfer calculations based on the original Monticello geometry have been performed. These calculations considered a range (0.008 to 0.030 inch) of radial gap between the leakage land and the thermal sleeve. At rated condi-tions the temperature of the leakage water in the vicinity of the blend radius was calculated to be 130'F to 150*F colder than reactor water. Thus fluid tem-perature cycling of this magnitude is predicted, which corresponds to blend radius temperature cycling of approximately 90'F to 130'F. This result sup po r t s the conclusion that leakage flow is the cause of the observed thermal cycling.

Figure 8, which is based on ASME Section Ill f atigue curve extrapolated to 10 cycles and on temperature fluctuations of 1 Hz, shows that 125'F cycling at rated conditions would begin initiating cracks between 300 and 45,000 hours0 days <br />0 hours <br />0 weeks <br />0 months <br />. It is expected that some c racks would initiate in less than one year of rated operation. Thermal cycling during startup and other conditions can be signifi-cantly higher than 125'F thua decreasing the time required to init iate cracks.

Based on the frequency and amplitude of the thermal cycling it is concluded that the high cycle thermal stresses would drive the cracks to a depth of approximately 0.1 inch. Further growth must be attributed to another cause.

3-14

, 1iI l:l  !'  ! l . l!lill!i lii! Ii 1!i)!l! ll!lllIl

, k x0?= O 0

~

!, ~i,-l: {-  :! , .I. > _

A- _

T- _ _.

_ A - _ - '

0

_ 1 Z D-H e> o:

R -

- _ I .

- I - -

A- - -

- 5 S 0

_NA-

- - 1 N E O

_ M- -

_ - I T

- A N- .

- - - m> m O- I

~

- T

_. a>

. - - m,i'0

- I F

- D -

i E- - ed -

1 N n

- S- _

I o i

A .

- t a

B-K C

i t

- i 3 A n 0 R I .

.c . I 1

C k

_ c a

- _ O C r

T

\_

_ 2 .

_. S e

u g

- - - R

- ~ i 0 U t EE a

. ,I

_ . - a

- MV O F 1

_. - SR . H

- AU

- 8 N

- C e r

ON ._ u g

- G 0 i

. 9 _> I ,'}  ;. .!

F OI. 1

_ ' _. - - ES SE

- AD -

B p'

• y F j . ; i t' jl

,ii iti I

0' 0 0 0 0 O 5 0 5 0 5 2 2 1 1 ng Ov woz<e . gws wv<ke2m g< swr w'G

NEDC-31130 Metallurgical examination of the material near the cracks at Millstone also supports the conclusion that the cracks resulted f rom f atigue cycling.

The second mechanism is large thermal and pressure stress cycling due to reactor startup. Crack growth in the blend radius cladding has been calculated using:

(based on Reference 6) ag - (ag ) 0.5 - (0.5) (1.4 x 10-9) N (2.5 Go) where:

ag a final crack size ag = initial crack size N = number of stress cycles o = stress range using the equivalent hoop stress method

= + c (pressure plus thermal cycle) h or

= (thermal cycle)

Crack growth in the base metal at the Llend radius has been calculated using:

(ASME Section XI, Appendix A) a f

= (ag )" - (0.863) (0.3795 x 10~9) N [F M .% )3.726 Crack growth is plotted on Figure 9 using the values in Table 1. Each startup cycle cc .ists of one cycle of pressare plus thermal, and five thermal cycles.

The dat6 points do not all lie on a scooth curve as the geometrical factor was changed in a stepwise manner. Figure 9 shows that a 0.1-inch crack will grow to a 0.44-inch crack in 78 cycles and to a 0.50-inch crack in 88 cycles.

l Thus the existence of cracks up to a total depth of 0.5 inch is attributed to I

initiation by high cycle f atigue and growth by pressure plus thermal cyclir.g.

i 3-16

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

' ' I NEDC-31120 l

\

i i

)

l

)

1 -

i 1

l

- - 1:5.. . . . . . . . - . . . . _ - ,

I

+-

. I asa_. . . . . . . . . . . . . . . . . .

I i (

U6 - .l M _- _ _- 1

_._4 l  ! t f I

. i N l 1 10 0 - - -

-i I l

(ks+er or i Cyclesi I l  !

t t i i I e .

i i

__. 04 . . . . . . . ..

g . . . . . - . . -

t 50 45 i

i  :

i i

_y_. Ciud

. . . . . _. ~_ . .. . _

M.9r ?ktal .

l

( . 1 ,2 . 3 . 1 . h . ti a (inch) l Figure 9. Crack Crowth 3-17 I

i

. _ _ _ . . _ _ - . _ . -___.___.m__ ..-_

NEDC-21120 Table 1 CRACK CROWTH CLADDING "i "f N th h j lin.) (in.) (cycles) Qsi) (ksi) 0.05-in. 0.057 10 96 20.4 0.078 30 96 20.4 0.112 50 96 20.4 0.174 70 96 20.4 0.183 72 96 20.4  !

BASE METAL

• i "f N th h (in.) (in.) (cycles) (ksi) (ksi) F(a , r) 0.19 0.0 73 44 20.4 NA 0.238 83 44 20.4 2.4 0.282 93 44 20.4 2.3 0.324 103 44 20.4 2.2 0.416 113 44 20.4 2.2 0.441 123 44 20.4 2.1 0.501 .. 44 20.4 2.05 0.555 143 44 20.4 2.00 3-18 y - - . . . - . - . - -

---

.y

. fr.DC-21120 The as-left condition has been evaluated for f atigue lif e. Two conditions were consider- : the bottom of a grindout, and the surface of the cladding.

At the bot tom of a base metal grindout:

k a

h+ th alt 1 where:

alt = alternating stress k = 5.0 max c 20.4 ksi h

o = 44.0 ksi alt = 73 ksi N,yyg

= 1450 cycles The required number of remaining cycles is taken to be 90% of t he total number of Monticello design cycles, which is equal to (0.9) (1500) = 1350 cycles Thus the usage f actor is UF = = 0.93 If50 Therefore new cracks are not expected to initiate at the bottom of a base retal grindout if the as-lef t usage f actor is less than 0.07 and thermal cycling of the base metal at rated conditions is less than 45'F, which corresponds to the endurance limit.

On the surface of the cladding:

ko +O gg ,

O " ~~~ u  !

alt 2 l

3-19 ,

9 NEDC-21120 k = 3.1 (stress index) o = 20.4 kai h

th

= 96.0 kai

= 79.6 kai o,1g N,yg

= 2800 cycles UF = = 0.48 f00 Therefore new cracks are not expected to initiate in the cladding if t he as-lef t usage factor is less than 0.52 and thermal cycling of the cladding at rated conditions is less than 50'F which corresponds to the endurance limit.

'Since the pressure stress at the blend radius is predominantly in the hoop direction, the cracks are expected to be oriented radially with respect to the nozzle and this is confirmed by the crack observation at Monticello.

The circumferential cracking inside the nozzle bore is attributed to thermal cycling coupled with lack of fusion in the original cladding.

The analysis presented here illustrates the importance of restricting the local surf ace temperature fluctuations in order to prevent further crack initiation.

With the use of the interference fit feedwater sparger design, it is believed that the fluid temperature fluctuations will be reduced to acecptable Icvels.

As discussed in Subsection 3.7, operating data f rom Millstone 1, which uses an interference fit sparger design similar to that just installed at Monticello, showed that the fluid temperature fluctuation range was within 50*F, thus con-firming the effectiveness of the new design. It is therefore reasonable to conclude that thermal cycling problems are less likely to occur at Monticello with the new design spargers.

Even with minimal thermal cycling it is possible that additional cracks will initiate as the remaining material has experienced an unknown amount of fatigue damage.

l 3-20 l

Nr.DC-21120

3. 6 GRIND-0UT WELD REPAIK INALUATION Based on an evaluation f the completed nozzle grindouts and availabic weld repair methods, it was concluded that veld repair of the areas excavated to remove cracks is not justified.

For this application, two types of weld repair after final post-weld heat treat-ment are possible under Section XI rules. Procedure No. 5 in 1WB-4430 allows recladding those areas where the low-alloy base metal has been exposed. There-fore, this repair would be justified only if required for corrosion resistance.

This is shown not to be the case in the examination of base metal corrosion characteristics given in Subsection .h 8 . It is additionally shown in Sub-section 3.3 that 1/16-in. depth of the exposed base metal is considered as cor-rosion allowance.

Procedure No. 4 in IWB-4420 could also be used prior to the above recladding to first restore the removed portions of the base metal. This would be neces-sary if the structural adequacy of an excavated area were insufficient. Ilow-ever, it is shown in Subsection 3.3 that the excavated areas satisfy the ap-plicable Section 111 Code design limits, and that the amount of material re-moved was relatively minor.

Consideration was also given to the effects of performing weld repairs dis-cussed above. Use of the temper-bead methods of Section X1 is generally felt to involve substential dif ficulty, and uncertainty.

As this type of process is infrequently applied, its qualification for this ap-plication would be partly developmental. The significant amount of structural 3-21 l

l NEDC-311du .- i restraint of the repair veld areas from the adjacent vessel wall would require l careful study to avoid cracking. Substantial thermal and stress analysis effort and testing would also be needed to develop the techniques for the required pre-heat, interpass, and the 450 to 550'F thermal treatments to avoid harmful gradi-ents and distortion. Because of the restraint and beat sink effects caused by the vessel wall, the requirements for heating, cooling, and location of insula-tion are likely to be complex.

In summary, it is believed that weld repair of the grind-outs need not be per-formed, and that consideration of the design adequacy does not warrant such a repair in this case, i

3.7 CORRECTIVE ACTIONS i

The methods utilized to locate and remove all detected cracks assure that local discontinuities, capable of propagation, do not exist at this time.

As discussed in Subsections 3.4 ano 3.5 of this report, it is import-ant that the local surface temperature fluctuations, such as were likely to have been experienced during operation with the original spargers be reduced. The replacement feedwater spargers have been installed with an interference fit based upon measurements of the diameter of the thermal siceves and the nozzle bores.

This procedure should assure continuous contact around the circumference of the thermal sleeve, except during intermittent periods of very low feedwater tem-perature experienced during startup.

As a result of these design improvements, the leakage through the gap is expected to be small during full power operation. The anticipated leakage at

! low power IcVels is not expected to cause significant annular fluid temperature fluctuations in the thicker regions of the nozzles, because the small amount of i leakage flow will be mixed with hot reactor water in the annulus. The effective-l l ness of the above design improvements has been confirmed for the similar sparger f

to nozzle fit-up used in the Millstone 1 sparger replacement. Operating data showed the fluid temperature fluctuations in the nozzle annulus were reduced to a maximum of 50*F for any operating condition. The effectiveness of the inter-ference fit design will continue to be evaluated as part of the ongoing sparger design evolution.

3-22 1,

, NEDC-21130 l F.ven with minimal thermal cycling it is possible that additional cracks will l initiate, e.s the remaining material has experienced an unknown amount of {

fatigue damage. Thus it will be necessary to periodically inspect the  !

not:1e blend radii. This inspection program will ensure that all flaws will  ;

be found substantially before they reach the critical size as determined by the most conservative calculations and before they reach the Section XI allowable  ;

size as determined by more realistic calculations.

3.8 CORROSION EVA1UAT10N l

3.8.1 Ceneral Corrosion in a BWR I

Ceneral corrosion raten for carbon steel and low alloy steels have been de-termined in tests pertormed by Cencral Elect ric Company. No differencen in corrosion rates were ncted between the carbon steels and low alloy steels.

The highest corrosion rates occur in low temperature, air-saturated water ,

which would be present prior te reactor startup, end during refueling outages.

At tempetatures up to 100'r, the corrosion rate of bare steel in stagnant ,

air-saturated water is 0.0015 in. per year. Very little corrosion occurs in the high temperature BWR water or steam. A thin, black oxide film forms yet rapidly at elevated temperatures and it is protective against corrosion.

The ' esured corrosion rates in 546'T BWR water or steam are less than 10 mgldr / month (a corrosion rate of 17 mg1dm / month equala 0.0001 inches per year). A very conservative rate of 0.0001 in. per year for corrosion esti- I mates on carbon steel components is assumed. '

The worst case for che total corrosion of RpV nozzles would be if all lov temperature corrosion occurred on bare, un-oxidized steel surfaces. The estimated corrosion on unclad nozzles would be as followsI f t

Assume 90% P en temperature reactor operation, then l 36 years

• 0001 in./yr = 0.0036 in. '

Assume 10% low temperature expos.,re at 100*F for startup and refueling outages, then 4 years X 0.0015 in./yr = 0,.0060_ in.

Total 40-year corrosion =

0.0096 in.

3-23

KEDC-31120 Actually, the high temperature oxide formed during reactor operation continues to provide corrosion protection when the reactor is shut down for refueling. Therefore, the total corrosion should be 1ces than 0.0096 in. for a 40-year reactor lifetime.

3.8.2 Calvanic Corrosion in BWR Environment Numerous studies ' have been made to determine whether galvanic (electro-chemical) corrosion vould be a problem in nuclear reactor systems where dissimilar metals are in contact. One of the common dissimilar metal com-binations found in Boiling Water Reactors consists of auntenitic strin?cas steel joined to carbon (including low alloy) steel.

Corrosion tests performed in high-purity reactor water have shown no detri-nental galvanic corrosion ef f ects on austenitic stainless steel-carbon steel veldments or joints. Neither the general corrosion rates, nor localized corrosion of the carbon steel have been affected. Investigators attribute th13 lack of any galvanic corrosion to the fact that high purity BVR water has very low electrical conductivity. The conductivity is too low to promote galvanic or elec tro-chemical ef fects.

The performance of stainless steel-carbon steel couples in an operating reactor has been reported by the Argonne EBVR. This boil!.r.g water reactor employed stitch welded stainless steel cladding in the reactor pressure vessel. Cracking occurred in the sheet steel cladding which exposed the bare reactor vessel steel to the water environment. Examinations performed af ter several years of operation showed no evidence of detrimental general corrosion or galvanic corrosion.

3-24

. NEDC-21120 4.0 SAFETY FVALUATION The corrective actions taken, as described in Subsection 3.7 of this document, will ensure that the original design requirements of the vessel are met during future operation of the reactor. The removal of all cracks provides assurance that further propagation will not continue in these affected areas. Replace-tren t of the feedwater spargers with improved therwal sleeve / nozzle interface will reduce the local surface temperature fluctuations to a negligible 1cvel.

With respect to ASHE Section X1, the governing code for inservice nuclear com-ponents, the st ructural integrity of the reactor pressure vessel has not been compromised, since all cracks have been completely removed. With respect to ASHE Section III, the original construction code for the reactor pressure vessel, the amount of base metal removed and consequently the additional nozzle reinforcement area required are adequately compensated for by the existing available reinforcing area which remains in excess of that required by the Code.

As such, a degradation of the original design requirements with regard to the low-alloy steel base metal has not occurred. Therefore, based on 10CFR 50.59, the cracks during this outage do not constitute an unreviewed safety question.

Installation of the new design f eedwater spargers is expected to reduce the previously experienced local temperature fluctuations to acceptable icvels thereby reducing the potential for additional cladding crack initiation. The new design feedwater spargers will also reduce the steady-state thermal stresses in the nozzle to vessel ehell region because of the reduction in flow between the nozzle and the thermal sleeve.

Therefore, based on 10CFR 50.59, operation of the reactor with the new design feedwater sparger does not constitute an unreviewed safety question.

4-1/4-2

' i

. NEDC-21130 l

5.0 REFERENCES

1. Hi'.1 stone Nuclear Power Station Unit 1. Feedwater Sparger Failure, Evaluation of Design No. 3 Test Results. Interim Report Addendum 2, September 21, 1973.

l

2. " Chloride Intrusion Incident." Special Report. Millstone Nuclear Power  !

Station Unit 1. December, 1972.

3. "PVRC Recommendation on T6ugt.ncas Requirements for Ferritic Materials,"  ;

WRC Bulletin 1 's .

4. " Technology of Steel Pressure Vessels for Water Cooled Nuclear Reactors " f Edited by G. D. Whitman, et al . ORNL-NSIC-21, UC-80, December 1967.

i

5. Vreeland, D.C., Caul, G.G., Pearl. W.L.." Corrosion, 17, No. 6, 269t-276t, 1961, June, i

6. NEDO-20926 Evaluation of the Structural Significance of Flaws in Nuclear Piping Welds, to be issued.

)

r L

I

{

5-1/5-2 ,

'g A . Nhc Ul?THI W 1)DN TOH P/A1 f,oa pr::;f.i ' wl,j,l/.i

. * (l l f ~)H A H f l O R M)

CO'UROLNO:)-Q3.'LU_.

8 f l L E : _. _ . . ._ _ .. ._. . _ .

f HOM: Nortnern States Pwr Cor\ DATI' OF DOC D A'I E H EC'D LTH 1WX RPT OTHER Hanneaposts, Mn ,3 g, g

__ __L.0 Maycr.- _

TO: .

OHIG CC OTHEH SENTHaC POR _ U ._ -

Mr Stello one ;s igned SE N'E L ODR PD R . _ XX CLASS ' UNCLASS PHOP INF 0 f t5U i NO CYS REC'D DOCKL i NO:

XXXXX 1 50- 2 t> 3 ,

DESCRIPTION: E NCl.O'i UHES:

Ltr trans the tottowing te incident report Supplenental Evaluation Rpt on Feedwater N ato 10-22-75........ Nozzle Claddtag Cracxs....................

f,s-- * - -

/; (4u cys encl rec'd)

PL ANT NAME: M DLIC'110

. _ . . _ _ _ _ . _ . . . TOR ACllOhlic;r( RM ATinN._ 1-5-76 _ ...._.__..chf ____ _ _

RUTLER (L) SCHWENCEH (L) .v!f M U;N (t ) REGAN C) 1:UD(L)

Wi CopW. W/ Copnet \'4Cn;w 5 W/ Cupi"s W/ COPInf CLAr;K (L) STOLZ (L) DICN E H (f.) LF Ah (L)

W/ Copics W/ Conit t, W! Copin W/ Cop;n PARR (L VASS/ ' LO (L) hN;GliTON K) n i .:t, W/ Copia W/ Copim, W/ Co,9/s W/ Capin KNIC L (L) PUhPLI. (L ) YOUNCOLOOD (E) II:t W/ Cr pia W/ Cop". , W/ Cupbs W/"Co;m INTERNAL DiST RIPUTiON D'G FD TECH Rf VIEW DENTON llc A5ST A/1 IN D

/7NHC h>H SCHRUhDER GRIMES DH All M Afs M ACCARY G AMMil.L # R. DtGGb (L) SALTZMAN yOGC, ROOF.', P 00GA H. Gl AfMN (L)

,,/GObSICK/S1 AF F KNICH f KASTNER E. GOULitOURNE (t ) MLLT2 CASE PAWLICKI DALLARD P. KRE U 12 L H (l!)

. SHAO SPANG L E R J, LEE (L) fLANR DOYD STElLO M. !E.inecP(L) MCDONALD MC OHE (L) liOUSTON E NVI RO. S. HEliD (El CilAPM AN DEYOUNG (L) NOVA 4 MUL LE H M. f.E RVICE (L) DUBE (L1r)

SKOVHOL1 (t ) ROP.S DICK L R S. SHf ppt h3 (L) E. C OUPE GOLLER (L) (Ltr) IPPOLITO KNIGHTON M. St Al E F. (E) PE T E RSON P. CO L t.lNS T E DSCO YOUNG 0LOOD H. EMI T H (L.) hARTFil LD (2)

DENISF J .CCn. LINS RFGAN S. TEL fS (L) KLECKER RI.G OPH. LAINAS PQOJECT LDR G. Wit 1.l AMS (E) LISENHUT

./ Fit L & REGION (2) DENAHOYA WIGGIN I ON ID PC VOLLMER ,5. 6 P.' E SSQ<ad V. R,INGn'M (L)

Wil SON (L)

' l' . PARRI T,r.__ (Q

. 11. tWC/J((c)

_ EXTt esNAL DISTRICI)1lON _

l

• c' / J ' ' '

/ - t OCAL

/1 PDR F ' *12.[1)(2)(10)NAllONAL

' llc (ADEHUATHY) - L ADS . _ _ _ . . . 1.. PDR SAN /L A/NY 7 1- 1 NSIC (E UCHAN AN) 1 - W. PLNNINGTON. R n EJu s G1 1 - DI.OOhM AVE N NAT L AB ASLO -

1 .. CONSUL 1AN1S 1 - G. U L R IKf,0N O R N L 1 - Newton Anth ison NT WM ARK /Rt.UME/AG U ADI AN

@ ACRS HOWWG/SEN1 /

TclA 0 55 c, a lc ! ~ ~ l b ygK

_- __ . - _ - _ . - _ . - _ - - _ -