ML20084L444
| ML20084L444 | |
| Person / Time | |
|---|---|
| Site: | Hatch  |
| Issue date: | 03/31/1984 |
| From: | Marisa Herrera, Obermiller K, Ranganath S GENERAL ELECTRIC CO. |
| To: | |
| Shared Package | |
| ML20084L427 | List: |
| References | |
| CRF-137-10, DRF-137-10, MAR-84-05, MAR-84-5, MLH01.DA, MLH1.DA, MLHO1.DA, TAC-54150, NUDOCS 8405140464 | |
| Download: ML20084L444 (22) | |
Text
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a DEF #137-0010 -
M&R 84-05 i4 MJ01.DA i
3 l
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ARE.YRIS W m m m Amma m aarmen l
M EA145 2 l
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March 1984
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Prepared by:
f L. Eerrors. Engineer i
b ehanics Analysis W
I
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4 Q ormiller. Engineer l
E chanics Analysis
[
Approved by:
- 8. Ranganath. Manager behanics Analysis I
i I
i 8405140464 840507
{DRADOCK05000pg GENERAL h ELECTRIC
)
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.d TART.R OF OstTIDml 1
1.
INIRODUCTION 2
i 2.
TWRRMAL STRESS ANALYSIS l
t 3.
LIKELIBOOD OF T10t00GH-WAIL CRAMIM 7
4.
MITICAL MAM SIZE Pot LOSS OF CDE, ANT ACCIDENT 7
10 5.
CW4Q,USION 10 6.
REFEREN MS I
11 7.
FIGURES i
a 5
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.e 1.
INTRODUCTION
[
Examination of the Hatch 2" vent header revealed a through-wall flaw which was nearly 360' of the header circumference in length.
Figure 1 shows the vent header, and Figure 2 shows a sketch of the flaw.
Metallurgical examination at General Electric Company of samples removed from the pipe containing the crack f aces revealed that f ailure was in a brittle manner, and that crack initiation i
occurred near the butt weld at top center of the header.
The most likely j
cause for brittle fracture is that the temperature of the vent header f
material, SA516 Gr. 70 carbon steel, dropped below the Nil Ductility Tempe r a ture (NUT). Below NUT, a material is susceptible to brittle fracture given that an initiated flaw and a driving force for crack propagation are present.
Examination of the torus area in the vicinity of the failure showed
~
that a nitrogen line was directly above the initiation location.
Further investigation of the nitrogen injection system showed that cold nitrogen was deposited on the vent header on several occasions.
The temperature of the ni.trogen was capable of dropping the material temperature well below the NDT, and producing thermal stress that could cause crack extension.
I I
Although the presence of the cracked vent header does not impede normal operation of the plant, depressurization during a Loss of Coolant Accident l
(LOCA) could be hampered by leakage of steam through the cracked area.
I i
This analysis is for the purpose of addressing three key areas:
- 1.. Determination of thermal stresses produced by cold nitrogen injection and if the observed cracking can be explained.
2.
Determination that existing cracks, if any, grow through-wall and thus l
become detectable by leak monitoring.
I I
3.
Determination of maximum through-wall crack length that could be tolerated without pipe rupture during a LOCA event.
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's 7
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Responses to the above crats are required to esplaim the itketste of the went
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s header and to evaluate tk sensequence of 'na existing orack during a LOCA
,s--
I ovest.
2.
'IltliBMAL MMESS AN* LYSIS '
l
%mm JEsthod i
Due to the complex three dimensional' geometry of the structure, and applied thermal loading, the finite element method,was used to determine the therust l
He ANSYS eospoter program was used to perform the analysis. He f
stre sse s.
(STIP 63) 'ef the ANSYS element library was three dimensional shell element selected to model the vest header. this 'obsent is espable of both. membrane Figure 3' showy the finite element model developed and bending type stresses.
for the smalysis.
Due to symmetry about the ?bstt weld, at which initiation is thought to have occurred only half et the pipe, sy:ssetris about the butt veld f
eenterline, is required, he stif fener plate which lies.on the bottom side of the pipe is simulated by increasing the thiekusas of the shell eleuests from
.25 to.88 inch. He diameter of the pipe is $4 inches. He elements which oonprise the stiffener piste are enclosed by the bold line in Figure 3.
i
' Appropriate bus 4sry conditions at each end of the model were seleeted to I
, disulate symmetry at one sad and a long pipe at the other.
i j
l fatorialprophrtie,e for SA516 Gr. 70 CS were used for the smalysis, i
l Thereal LosditLC j
s s.
s The selection of tempergture boundary conditicas is a very erseial step in the l
De temperature analysis to noeurately determine ress1 ting stresses.
l distributies duo' to the injectJos of the altrogen is not aneurately known.
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,s
s horefore, a temperature distribution will be assumed.
s l
From Referesse 1 impingement of a jet en a surfaes will eense a temperature distribution on the impacted surf ace in the shape of a Gaussian distribution l K.s(Figura 4).
Die shows that there will be a some where the temperatste of the f
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surf ace esa reach temperatures mostly egual to the jet fluid temperature, f
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A
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2 s
w.
x -..
j
.~-
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Temperatures them drop off from the peak temperstare and eventually reach To stamiste this behavior in the vent header finite
(
emblent temperature.
element model, the temperature distribution shown in Figure 3 was seed. The l
sabient temperature was ass'uned to be 100*F. while the nitrogen tamperature j
t was assumed to be -200*F.
The otrcumferentially cooled area was assumed to be from -908 to +90*
from I
top dead oester as illustrated.
From -45' to +45' the temperature was assumed t
to be -200*F.
I The assumed temperature distribution was imoorporated into the model by assisming each individual shell element a temperature.
t i
It is orpeeted that a gross bending situation vili oeour due to the top of the I
header being sold and the bottom of the header being ware. 'The stiffener plates result la a wall thickness of over three times that of the.25 inch j
i header pipe wall. This should result is bending of the cooled region inward f
sad very little displacenest and stress in the stiffened area.
i Stress Analysis Results As orpected, the thermal eyeling on the vest header esases large deflections at the top eenter location (approximately.1 inch). Figure 5 shows the l
displacement of the end model points which represent the oester of the j
eircumferential butt veld. The stiffener plates do not defloot as shows la
(
the figure.
The driving force for otroumferential crack extension as l
}
experleased in the vont header is the applied axial stress.
Crack extension t
i requires a gross bending effect, which will produce tensile stresses in the apper portion of the vest header. The axial stress as obtained from the ANSYS l
esmpster output is shown in Figure 6 for the section of elements nearest the butt veld.
Consistent with the displacement results, there is negligible
[
t stress la the stiffener plates.
The influemos of thickness and temperature l
i discontinuity can be seen la the figure by the large tensile stress of i
i approximately 19 kai in the.25 inch material. The stress deeresses to approximately 15 kai la the area of crack initiation.
The veristion l
l 1
e 1
3 l
i
r circumferentially indicates that the stress at the orack initiation location oos1d range anywhere from 15 kei to 19 kai depending on the cooled area l
loostion with respeet to erack initiation loestion.
If the cooled area eenter is not located near the e,rsek initiation loestion (i.e., stack lif tiation j
location is on the edge of the cooled area), the stress could be higher than 15 kai.
For purposes of this analysis, a value of 15 kai is used for the thermal stress.
Two additional sonroes of stress in the vest header are sold spring stress and weld residual shrinkage stress due to welding of the support l
plates on the bottom half of the vont header.
A sold spring stress of 15 kai j
was seed for the analysis. This is justified by the observed crack opening l
l and transverse displacement of the orack faces af ter fracture.
Assuming a.1 inch traarverse displacement over 20 feet gives approximately 15 kai sold l
spring stress. The determiattles of weld shrinkage residual stress due to welding of the support pistes on the bottom of the vent header requires i
sophisticated elastic plastic smalysis.
In this analysis, the veld shrinkage l
stress was somservatively omitted.
i I
f The potential for crack growth is analyzed by the use of Linear Elastic l
Fracture Mechanics (LEPN) which assumed brittle behrvior. This is consistent i
with the metallurgical finding.
l The stress intensity factor is sales 1sted by the use of the following egnation J
l from Isference 2, f
y = a, [ k + ab Y K
where e, = nombrase stress ab = bending stress l
%, = membrane sorrection f actor Mg = bending correction factor a = orack depth Q
a, shape f actor i
The membrane stress is comprised of the sold spring stress and thermal stress.
Tsid residual stress 'somprises the bendire stress component.
f i
4 i
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....,_-,_.-.__.-..-...--_..,_._.-._-,_-,,_,_.,,?
f sise' of.1 ineh in depth and 2 inches is length was assumet An initial flew This is somaistent with metallurgieel findings.
This is reasonable since'the sould assily drive initial shock stress, coupled with the residual stress, unaller flows to.1 inch in depth.
}
The membrane sad bending scereotion f actor, and the shape faster, are obtained i
from Referesee 2.
t RL = 1.74 Mb = 1.00 G = 0.84 i
The stress intensity factor for the Lesumed lands and ersek depth is
,~
56.4 keiE.
Espect testing of RA516 Gr. 70 samples removed from the Esteh vest header f
ladicates an NDI of below -50*F.
Charyy data taken from the impact testing is shown in Table 1.
This data must be sorrected due to sub-size specimens.
With correction, bmpact energy values are as low as 8 f t-lbs.
l Table 1 1
Tesserature (*F)
Innset (ft-1bs)_
r 40 12.5, 11.5, 11.0 20 11.5, 11.5, 11.0 l
0 11.5, 11.5, 12.0
-20 11.0, 11.0, 11.0
-40 11.0, 10.5, 11.5
-40 6.0, 9.0, 9.5
-80 4.0, 5.0, 8.0 From Referesse 3, a sorrelation between Charpy hardness and frasture toughne ss is obtained.
For low alloy steel the sorrelation is I
f
= 5 (CVN)
I h
1 e
r 5
l
I i
H is gives I
IC = 31.0' keik K
l This fracture toughne ss corresponds to the plane strain oondition.
Since the vent header well is only.25 inch thick, the p1'ame stress senditlom is appropriate. The plane stress fracture toughness can be found by using Irwin's sorrection factor. The plane stress, stress intensity factor, is:
\\
i k = kCEI
- 1*4 I l C
where i
SIC "
(
The sorrection f actor is valid only for 0.41 DIC 1 1.0.
Simoe PIC is greater l
This results in E,= 48 koih.
than 1 for this case, a valso of 1 is assumed.
The applied stress intensity factor ($6.4 kalk) is above the material fracture toughne ss, indicating that the part-through ersek will grow through
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i the wall thioksess and become a through-wall oraek.
[
l i
I Once a through-wall orack has ocentred.' eircumferential erack propagation f
oos1d occur if applied loads are large omossh.
The driving force for f
circumferential orack extension is the thermal stress from nitrogen injectica and oold spring stress which total to 30 kel.
From Reference 4 the stress l
intensity factor for a shell with a circumferential ersak of length 2a is j
y = Cak K
where C - 1.1 This results la a stress intensity factor of approximately 58 keih.
This is above the material tossaness of 43 keih. indicating that aireumferential erack growth will ocent once oracks have grown through-wall. The conclusions j
on brittle fracture are oomsistent with metallographic fracture evidence,
[
suggesting a brittle fracture mecasalsa.
f 6
i
,, _. - _.,. _ _,,,,,,, -. - ~ -, - - - - _, -,
r 3.
LIIELIBOOD OF 'rwangeg-VAIL GAGING Since detoottom of eracking in the vent header piplas is by visual examination or leak monitoring, it is"mesessary to know whether sold nitresh insertion l
osa cause Israe part-through oracks that may not be detooted by the visual /
{
1 1eak tests.
As shows in the earlier smalysis on the evaluation of the Esteh 2 erasking, exposures to sold nitrogen easse high thermal shook stresses. His, j
oosp1ed with the weld residual stresses, saa oasse ermok initiation. De stress intensity footors for such ersoke increases with erack depth.
l H erefore, omoe orack initiation occurs (starting from say, a weld defect), it l
will continue to propagate until the oraek penetrates the bookra11. At this f
potat, circumferential crack extension will ooost due to the gross section j
stresses such as thermal bending or sold spring stress.
Crack arrest will j
occur when the orack tip ends up in higher tougha[ss material that is not
~
i exposed to t$e cryogenic sitrogen.
i Based on the above discussion it is concinded that exposure to sold nitrogen I
will lead to through-wall eracks which can be monitored by visual examination or leak testing.
4.
MITICAL GAG BIZE POR THE VIDE WEAma DURING A LOCA EVIDE
.f he purpose of this section of the report is to determine the maximum through-wall orack that osa be sustained la the vest header without inducing i
t crack propagation. his analysis is based on licear elastic fracture l
t mechanies (LEPM) methods. He use of LEFM is conservative staos the material i
will most likely exhibit ductile behavior during any ersok propagation.
l l
he two mais loads that act on the vent header during a loss of coolant l
accident (LOCA) are the pressure dif ferential and the pool swell impact loads.
l nose loads are defined la Figure E2 4.1.1-1 and Table E2 ~.3.3-1 of Beference i
5.
Other loads are present; however, they are transmitted directly to the j
[
vent header supports.
t i
i f
7
He most severe loads ooonr about.5 seeond af ter the initiation of the LOCA.
At this time the pressure differentisi in the vent header is 24 psi.
The^
exial membrane stress indsood by this pressure is:
l a, = h =
2$.
" 1*3 I*I a.
Then smalyslag the effect of the pool eve 11 impact loads, the 20 foot section of the vent header between supports is assumed to not like a uniformly loaded.
j simply supported beam. An average pressure of 11.18 psi is applied to the j
i vest header which transistes into a load of 604 lbs per inch. D e marinum bending stress associated with this load is l
i l
2 Mc 2
w1 R, i
e 8(w/4 (1,-R*g h 4
i nis gives a bending stress of 7.56 kei. The maximum stress in the vent t
header dae to these two loads is, therefore, 8.86 ksi, n e applied stress l
Intensity factor can be obtained from Referesse 6.
i y = F, (A) a h + C(a) F,(A) abb E
t where a = half crack length A=s/h y = s/xR 2 1/2 F,(A) = (1 + 0.3225 A 3 for 0 1 A f 1 i
= 0.9 + 0.25 A for 1 < A I 5 32 8
72 1 + 6.8 v
- 13.6 v
+ 20 v 1 + 7.5 13/2 - 15 75/2,33 7 7/2 l
1 as a function of orack length.
l Figure 7 shows Ey Figure 8 shows the curve used to deternise the plane strain fracture j
Although "his curve This figure was takan from Referesse 2.
t toughne ss. KIC.
+
e 8
~
l
\\
I is speelfloally for low alley steels, it also represents a conservative l
estimate of the fracture tonshases for carbon steels.
Figure E2 4.1.1-2 of Beforence 5 spoolfles the temperature ef the vent header to be 85'F at the besimming of a LOCA.
If a eosservative valse of 0'F is assumed for the Npr the correspondlag fracture tonshness is 160 kalk reference temperature, Using Irwin's oorrection f actor for plane stress behavior, the plane stress fracture toughne ss E, een be obtained.
C J
C(1+1'4I C]
where SIC "
(
)
ys i
f 1.0.
In the present esses BIC f
I" This equation is valid for 0.4 f $1C greater than one, so a value of one was used,
~~
m L
^
f
^
For RINDT = 0
- F C = [(160 kai in. )2(1 + 1.4)}1/2 = 243 kai in, j
K will be conservatively taken to be 200 kalk since at such high levels of l
KC toughness the mode of f ailure is duct!!e rather than brittle.
l
[
Cracks which are long enough to result is applied stress intensity factors t
This f
above these fracture toughne ss levels will propagate and cause f ailure.
I limit is showa la Figure 7.
The orack length correspondlag to the assumed ET of 0*F is approximately 40 inches.
If oracks exist in the vest header
[
g sad are not detected by visual inspection or other seams, they are most likely l
well below the critical crack length, herefore, they will not inhibit the vent header from performing its required function during a LOCA.
i l
1 4
e f
i l
9
u-e I
5.
CONG,USION The smalysis presented here shows that the Intch vest header erseking saa be expected, given the liquid nitrogen in8ection and the applied leadings. The resulting cracks are expected to b, through-wall so that they eas be monitored by visual observation or leak tostF.
Finally, the eritiest eraok length during a LOCA event it large enough that cracks would be detested by restine j
i visual inspection or other examinations.
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REFERENGS 1.
Seh11 sting, R., ' Boundary Layer Theory,' Pergamon Prese. New York,1955, 2.
ASIE Boiler and Pressure Yessel Code,Section II,1983 Edition.
3.
Rol f e, S.,
Barson, J., ' Fracture and Fatigue Control in Streetsres,'
l Prentice-Eall Inc., Eaglewood Clif fs, New Jerary,1977.
l 4.
Booke, D., Cartwright, D., ' Stress Intensity Faetors,' The E1111agdom Press, Umbridge, Middleses. England, 1976.
5.
' Mark I Containment Program Plant Unique Load Definition, Edwin I. Estch Nuclear Plast-Unit 2, ' September 1981, ISD0 24569 Rev. 2.
l l
i
=
6.
'The Application of Fracture-Proof Design Nethods Using Tearing
[
Instability Theory to Nuclear Piping Postulating Circumferential Through Wall Cracks, ' NUREG/CR-3464. Del Research Corporation, Mp+G, Idaho, Inc.
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