ML20008E420
| ML20008E420 | |
| Person / Time | |
|---|---|
| Site: | Yankee Rowe |
| Issue date: | 05/31/1962 |
| From: | Currin H, Thorp A, Tong L WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| References | |
| WCAP-1997, NUDOCS 8101070051 | |
| Download: ML20008E420 (48) | |
Text
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NEW DNB (BURN 0fS)!. CORRELATIONS (sp -
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_NEW DNB (BURNOUT) CORRELATIONS by L. S. Tong H. B. Currin A. G. Thorp II May 1962 f
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APPROVED BY:
W. E. Abbott, Manager Reactor Development e
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DI9fRIBUTION W. E. Abbott H. N. Andrews S. Bartnoff s
L. Chajson C.C.Christensen(5)
P. Cohen R. J. Creagan H. B. Currin W. F. Davis P. G. DeHuff W. J. Dollard G. H. Farbmaa D. H. Fax R. J. French l
J. M. Gallaaher, Jr. (5)
H. W. Graves, Jr.
S. J. Green P. B. Haga D. F. Hanlen R. G. Hobson B. Hoffman W. E. Johnson A.R. Jones (10)
A. S. Kitzes L. B. Kramer F. L. Langford S. M. Marshall.(15)
J. D. McAdoo G. H. Minton (5)
N. R. Nelson i
C. F. Obermesser E. U. Powell J. C. Rengel C. Roderick H. L. Russo J. W. Simpson W. E. Shoupp T. Stern R. L. Stoker G. R. Taylor A. G. Thorp II L. S. Tong i-H. J. von Hollen J. Weisman R. A. Wiesemann J.H. Wright (5)
J. M. Yadon TechnicalInfor1nationCenter(3)
Thermal & Hydraulics Group (50)
-i-
mraer As a result of a parametric study of about 3000 published DNB data points, two new DNB correlations were developed for the subcooled region and the quality region respectively pending further experimental verification.
The validity of correlations in the ran6es of various parameters were examined by plotting the ratios of ( g measured / g g predicted) and (q", measured /q"DNB predicted)versusvariousphysicalparameters in the ranges of application. The results show that there are no trends due to these parameters other than those already incorporated into the correlations.
The future test on DNB by Westinghouse was described. Joint p'anning and systematic coordination were recommended for the future DNB tests in various organisations in order to obtain more effective results.
-ii-k
i TABIE OF CONTENTS 1
)
Page No.
i Abstract 11 Nomenclature iv I.
Introduction 1
II. Analysis 1
III. New DNB Correlations ami Their Applications to a Core Design 4
IV. Discussion 6
A.. Rod Bum le DNB r*.tt 6
B.
q"-DNB Concept 7
C.
Validity of Correlations in the Ranges of Various 7
Parameters D.
Flow Instability in Parallel Channels 7
E.
F.
Comparison of New Correlations with Existing Ones 8
G.
Sources ani Accuracy of Experir:Ltal Data 8
V.
Future Work ani Recommendations 9
Acknowledgements 14 f
References 15 e
-iii-L
NOMENCIATURE A
Cross-sectional area for flow, ft c
f A
Heat transfer area, ft h
p The bubble angle between liquid and solid surfaces D,
Equivalent diameter, ft.
2 G
Incalmassvelocity,Ib/hr-ft H
Localenthalpy, Btu /lb.
H, Inletenthalpy, Btu /lb.
g H
Saturated liquid enthalpy, Btu /lb.
f H
Heatofevaporation, Btu /lb.
fg ZH (HDE-Hin), Btu /lb.
gg L
Length of heated channel, ft.
P Heated periphery, ft, h
p Pressure, psia q"DE Heat flux at DNB, Btu /hr-ft q"
Averageheatfluxoftestsection,q"=ff,b q"dz, Btu /hr-ft T
Temperature, F 8,e M coo W (Tf-T1ocal)'
T Saturation temperature, F f
3 pg Weight density of saturated liquid,1b/ft 3
p Weightdeneityofsaturatedvapor,1b/ft X
Steam quality, % by weight z
Axial length, ft.
_iv.
k
I.
Introduction ThenumberofexperimentalDNB(departurefromnucleateboiling, formerly burnout) data has increased rapidly in recent years. However, the correlations of DNB data obtained from various investigators are not in Jens ami I4ttes( ) ami Zenkevich and Subotin(3) mported agreement that DNB is strongly influenced by local sub-cooling and local heat flux.
However, Cook ) founi DNB to be independent of local heat flux distribu-tion because, for a given inlet enthalpy and flow rate, DNB occurred at a same power input for both a uniform and cosine power distribution. These
.apparently conflicting results indicate that there can be two types of DNB, namely, the q"-DNB due to excessive local heat flux and the H-DNB tiue to the high enthalpy at the vicinity of heated surface.
i Ar important relationship, which was generally neglected by previous DNB investigators, is the energy balance of the test section.
(HDNB - Hgg) G A
=P I
I" d*
c h
and rearransing, (HDNB - Hgg) = 4 h (1) e The above equation reveals an important fact. ForaconstantL/D,andG, (HDNB - Hin) is directly proportional to 7. Hence, the H-DNB and q"-Dd cahnot be differentiated from the data of a uniform heat flux test. In d'
order to identify H-LNB or q"-DNB, the data point should be tested on two different power distributions. This is because H-DNB depends on average F
7 but not the shape of power distributions.
l II. Analysis The q"-DNB phenomenon can be described by Helmholtz instabilityb) with an additional mixing effect from the coolant stream passing over the heating
. l-I
surface. Since high pressure reduces the bubble size, the q"-DNB flux can be. increased at high pressures before Helmholtz instability is reached.
The local sub-cooling affects the condensation rate of the bubbles, hence this is a strong parameter of DNB.
The fact that H-DNB occurs at high local enthalpy can be interpreted as a change of flow pattern due to the existence of a larSe local void at the vicinity of heating surface. The mechanism of this DNB should be of ag hydrodynamic nature. Therefore, the Navier-Stokes' equation, Energy equa-tion, and Continuity equation of two phase flow with proper boundary condi-tions should describe H-DNB. The dimensional analyses by normalizing these equations i boundary conditions were performed independently by Griffith(U and Zenkevich However, the former did not develop a correla-tion froct his analysis and the latter developed a correlation which cannot correlate experimental data very well. Their non-dimensional groups are listed as follows:
Griffith:
q" "O
L i e"g i cf f' i p"f P
g' D,'
c
,f 2T E oV 8
pD k
8 2
p p
I g g,Vg y D p
fg g g
,f g
f g
~
g e
Zenkevich:
1/2 f
P 1/2 3"
(pfGge)
=f 6
f #
^
l G
C f
o'H
' u ('p
-)
j (3)
H, f
fg p p-pg The non-dimensional greiaps obtained by Griffith can be recombined with themselves or with energy talance equation in the following forms:
bx ief k(T -T,)
Cp H -H O
f x
pf f in (4) x
~ u pD k
fg g,V Of.
"f H
1
. s.
i
d c,"
bE xh,bE.
e x
(5)
I PY 4hh P
H 4L fg g i f
fg e
e D
(a/V)
(GVD,}o!(GVD,)"
g c/V AccordingtoChan6(9}, surface tension a is insignificant in forced convec-tion boiling, because the bubb1.e has high stability where its size is small in a forced convection flow. If th'e DNB correlatica is limited to vater only, viscosity and surface tension a of the saturated vapor acd liquid are unique functions of the pressure and hence the roups containing p or a
.can be replaced by functions of p pf.
For the surfaces of same caterial and roughness, the angle $ between liquid and solid surfaces are constant.
Henceequations(1)throuEh(6)canbecombinedas:
f kNB L
f' in G e
~~ " f ]s go
~
p g
s go 7 U
[ff.
-f e
fg o
co-Since a high value of G can reduce the thickness of water layer along the wall, this causes early DNB. The heat of evaporation of water, Egg, is higher at lower pressures. This means an equal amount of water can absorb more heat during evaporation and thus increase H at lower pressures.
DE Amm11 values of D, p l e M ues of (H
/Hbulk).
s 2e near vall apparent H f a smaller cha - 1, usually measured from the bulk stream, DE is higher. The high inlet enthalpy, Eh, y uld give a larger void fraction at exit with the same heat input and thus a high Hin causes le bNB*
i f I
o.
m III. New DNB Correlations ani their Applications to a Core Design As a result of a parametric study of about 3000 Published DNB data points, two new DNB correlations were developed for the subcooled region and the quality region respectively. These correlations ani their ranges of parameters are listed as follows:
1.
DNB Correlation in Quality Region tHDE = 0.529 (Hf-Hin) + (0.625 + 2.36e- # D ) H
-1.5G/10 e
e fg L/D
- 0.41 H e e - 1.12 H fg fg p/pf+054PHfg l
The above equation correlates the existing data within 25% and with a 95% confidence as shown in Figure.s 1 and 2.
The correlation was also s
plotted with data at 2000 psia in Figure 3 This plot shows that the data spread is. improved to 20%.
31nce tBDE = HDE - H,, M s i
g correlation can be expressed explicitly either by tE rH DNB DE' However, it is more convenient to calculate the percentage of accuracy in gg than in H The ranges of parameters of the correlated DE.
experimental data are as follows:
- metries: circular tube, rectangular channel, annular channel, and rod bunile.
Axial heat flux distribution: uniform ani non-uniform has velocity = 0.2 to 4.0 x 10 lb/hr-ft Pressure = &;0 to 2750 Psia L/D,=21to656 Inlet subcooling (H -Hin) = 0 % M m/D.
f 2
Local heat flux = 0.1 to 1.8 x 10 Btu /hr-ft Exit quality = 0 to 0.90 by weight
6 2.
DNB Correlation in Subcooled Region 6
q"DNB=(0.23x10 +0.094G)(3.0+0.01S,,)(0.435+1.23e-0.0093I/Dg3 (17-1.4e'*)
"h'#'
-3/L
-1/3 a = 0.532 (Hf-Hin)/I (A /A )
fg g f The above equation correlates the existing uniform heat flux DNB data within 20%_ accuracy and with a 95% confidence as shown in Figures 4 and 5.
Since the data for this correlation was obtained from uniform heat flux test sections, it is not clear whether this is due to high heat flux or high enthalpy'. Hence, the q", predicted 1y equation (9) should be considered as either a local DNB flux or a i
average heat flux up to the point of DNB Tha ranges of parameters of 2
the correlated experimental data are as follows:
Geometries: circular tube, rectananlar channa1, an=lar channa1, and rod bundle.
Axial heat flux distribution: uniform 2
- ss velocity = 0.2 to 8.0 x 10 lb/hr-ft i
Pressure = 800 to 2750 psia
{
% = al to 365 Inletsubcooling,(Hf-Hir.)=0to700sta/lb.
2 Local host flux = 0.4 to 4.0 x 10 Btu /hr-ft s&cmH== at DNB, a
= 0 to 228 F se 4
3.
Application of DNB Correlations to a Core Design In determining the thermal safety margin of a core design, a DNB ratio should be calculated along the hot channel by using either quality-DNB correlation or subcooled-DNB correlation. The quality-DNB correla-tion is used to calculate the H.DNB ratio as g predicted /bt h1 in a quality region. The subtooled-DNB correlation is used to calculate
- /
L
the heat flux DNB ratio in t.he subcooled region in the following two forms:
Incal q"-DNB ratio = q"DNB predicted!S" local Average q"-DNB ratio = q"DNB predicted < average up to DNB point
}
e It vill be noted that the q"DNB predicted! average up to DNB point is also the H.DNB ratio and hence one may infer that the subcooled DNB mechanism is of similar hydraulic nature to the quality type.
'5qs. (10) and (11) v111, of course, be equal froci uniform heat flux tests without giving any indication as to whether the q" concept or H concept is control 13 g.
Before a specially designed DNB test $s conducted to identify the nature of q"-DNB and to further verify the accuracy of the correlations, all these new DNB correlations are recocmended to be used as an additional check of thermal safety margin of a core desi5n.
IV. Discussion A.
Rod Bundle DNB Data The fuel assembly geometries of many per.rer reactors are rod bundles.
It is of general interest how these new DNB correlations apply to a rod bundle geometry. The available DNS data obtained from re'. bundles are listed in Table I.
The measured enthalpy rise of a hot channel was evaluated based on a unit cell formed by the neighboring fuel rods.
The heat input to the channel comes from these rods. No clixing between channels was assu=ed.
The cocparison of the predicted gg and the measured g g were plotted in Figure 6 which verified the applicability of the new correlations to the rod bundlem.
-6 t
L
(
,. 1 B,
q"-DHB Concept The q"*DNB concept was developed from pool boiling DNB. However, the DNB fluxes of pool boiling in pressurized water calculated by using 102tateladze equation ( 0) are usually higher than a million 2
Btu /hr-ft as shown in Figure 7.
There are many subcooled DNB fluxes measured in a uniform flux test section less than 800,000 Btu /hr-ft.
It is very difficult to explain hcv a subewled DNB under flow can be lower than a DNB in pool boiling. This leads one to consider H-DNB
) conducted rome hot patch DNB even in the sutcooled region. BAP tests and the results can be predicted by the new correlations very well as shown in Figure 8 where the DNB ratio in subccoled region was calculated by using Eqs. (9) and (11). Eq. (11) ir. effect is ba sed on the dI-DNB concept. Hence, the existence of a local q"-DNB in a reactor channel is in dot;bt.
C.
Validity of Correlations in the Fanges of Various Parameters In order to show the validity of the new correlations, both ratios
' ( b MB measured b NB predicted) and (q"DNB measured! DNB predicted}
are plotted versus various physical paraneters in the ranges of applicat.on as shoien in Figures 9 through 18. Tnese figures show that there are no trends due to these patameters oth r 'than those already incorporated into the correlaticns.
The subcooled DNB correlation Eq. (9) vas used in predicting DNB in quality region by setting S
= 0.
Tne result gives a much larger spread ee in the quality region than in the subcooled region as shown in Figure 19 This indicates that the saturation temperature should be the upper limit of application o* the subcooled DNB correlation.
D.
Flow Instability in Farallel Nnnels BAPL tested the DNB in a 27 in. lcng rect,aD6ular channel with parallol channel effect as reported in WAFD-188, table 27 The data
. L
-e were compared with the prediction of H-DNB correlation as shown in Figure 20. The result shows that at a low pressure (800 to 1200 psia)
M is 1 wer than predicted probably due to flow instability, and DB that at a high pressure (1600 to 2000 psia) $
is n t affected by DE parallel channels. Since flow instability is strongly influenced by channel geometry, the result plotted in Figure 20 can only be considered as a special example but not as a general conclusion.
N E.
Annulus DNB Data Analysis The annulus DNB data are mostly obtained in quality region vbere the H.DNB correlation applies. However, the H-DNB correlatAon was developed for a channel vitt all vetted surfaces heated. In calculating the measured B in an annular channel with one side heated, only a half thickness of the annular water layer is considered to be heated and the other half to stay cool and unmixed. All annulus data were evaluated p s way and compared with the prediction of Eq. (8) as sh" a in Figure 21. It can be seen that for a high heat flux in a long test section, the assumption of " unmixed" gives an over corrected value of the measured M. This effect of unheated surface in an annular channel vill be investiEated in greater detail in the future.
F.
Comparison of New Correlations with Existing Ones Other existing DNB correlations develepad by Jens and Lottes(2)
Zenkevich and Subo..n(3), zenkevich @, ra,l G.E.
) were plotted with experimental data in Figures 22, 23, 24, and 25. By comparing these figures with Figures 1 and 4, one can see that the new DNB correlations Eq. (8) and (9) rit date in the present rances of parameters better than the other existing correlations.
G.
_ Sources and Accuracy of Experimental Data The sources of experimental data used in present study are listed in Table II.
e The accuracy of these experimental data was first examined by an energy balance.
P., f q"dZ = A G if!
DE Ifallwettedsurfacesareheated,D,=4A/P*
c h f q"dZ = G b NB D,/4 For a uniform heating, above equation becomes Y = G gg D,/4L Any data point which was out of energy balance by greater than 10 percent was excluded in the present study to avoid being mislead by inaccurate measurement.
Future Work and Recosunendations V.
A.
Future Work in Westinghouse Westinghouse Atomic Power Division has built several 5-rod 10 foot DNB test sections which are scheduled to be tested in the high pressure water loop in Argence National Laboratory. The objectives of this test are:
1.
To verify the accuracy of the new DNB correlations in a full-scale rod bundle at the operation conditions of a large NR.
2.
To identify the existence c, H-DNB and q"-DNB in various conditions.-
. 1, o
B.
Recommendations 1.
Theoretical enalysis of boiling heat transfer and flow pattern studies of a two phase flow shoul. be encouraged in exploring various mechanisms of DNB.
i 2.
It has been indicated that earb DNB due to flow oscillation is one cause of poor correlation. It is recommended that a transient pressure measurement system should be installed for all DNB tests.
3.
In review of the existing data, it was founi that there has been considera,ble duplication in test conditions in various-organizations, and that there are cert c 1 cocbinations of various parameters in which no testing has been performed.
Itsohappensthatitiswiththesecombinationsthatmany commercial power reactors are now being designed. In order to obtain more effective results, it is recommended that the future DNB tests in various organizations thould be jointly 2
planned and systematically coordinated..
3_
- e L
TABIJ!l I - ROD BUNDLE DNB DATA
Reference:
WAPD-TH-478 P = 2000, L = 9.25", D, = 0.202", 3 x 3-rod bundle E,
3x E
g exit exit bMB kNB
- 551 0.403 793
.26 8e5
.792 33 0.994 753
.17 766
.803
Reference:
DP-615 P=1000,L=37",D,=0.255",7-redbucale(virevrapped)
H, 3 x 10' E
E gg r 10~
g dt ett gg 503 0.535 815.4 42 952
.391 473 0.587 802.4 40 941 448 498 0 939 750.4 32 858
.556 494 1.59 665.8
.19 739
.617
Reference:
DP 625 P=1000,L=37",D,=0.255",7-rodhmale(ceramictubesusedasspacers)
H, I x 10 E
T E
kB x 10~
~
g edt dt DE 507 0 510 782.4
.37 902
.342 511 1.41 672.4
.20 741
.562 511 1.00 717.4
.27 to6
.513 511 1.00 737.4
.30 836
.553 Reference EW-72255 P = 1200, L = 18", D,
= 0.208", 19-red bundle (vire vrapped)
H x 10 kotCh.x10 E
H YNB
- in exit exit DNB 570
.498 384 750
.29 1009 560 525
.480 370 708
.22 971 549 477 491 379 679
.17 970
.6el 401
.500 386 620
.08 933
.683 525 492
.R9 722
.24 1006
.606 305
.498
.384 551 906 767 566
.991 764 674
.16 829
.668 520 1.010 779 63.'
.09 790
.698 499 973 750 62 5
.09 808 770 kk5 1.000 771 585
.02 785
.872 557 1.980 1 526 619
.08 708 766 519 1 970 1 519 594
.03 702 922 539 2 930 2.26 596
.04 677 1.037
- D, is measured at hot channel-
- c. _
r o
a TABLE II - SOURCES OF EXPERIMEMPAL DATA POWER CHANNEL CHANNEL L/D' PRESSURE G/lo QJALITY NUMBER REMARK 2
SOURCES DISTRIBUTION SHAPE LENGTH, IN.
PSIA lb/hr-ft AT EXIT, % oF POINES Circular and 3 0 to 21 to 500 to o.2 to
-0.45 to N71 WAPD-188 Rectangular 27.4 365 2750 8.0
+0 97 0.49 fo Non-Uniform-Recthngular 27 245 2000 2.2 42
~
Rod' o.24 to
-0.25 to Only 2 pts.
WAPD-TH-478 Uniform Bundle 9 25 45.8 2000 0.99
+0.43 18 reached DNB TID-12092 Uniform Circular 16.2 19 to 1000 0 75 to 0.19 to 916 (CISER-27) 153 32 o.97
,f0to DP-555 Uniform Annular 42 175 1000
-o.
t 27 DP-615 Uniform 37 114 1000
[,*h
- 9$*
0 "I
y,"pped 5
1, b,
DP 625 Uniform
[$ t
[20 t 37 n4 1000 4
Fe w e DP-645 Uniform Circular 24 27 to 85 to 2.2 to 0.07 to n
75 1000 8.5 o.16 DP-665 Unifom Circular 24 27 to 500 to 3 0 to
-0.24.to' a
75 1000 7.4
-o.08 ARL-6063 Uniform.
. Circular 18 59 2000
- 0.69 to
-o.32 to g
1 95
+0.23
_=-
e G I l
TABLE II - SOURCES OF EXPERIE MPAL DATA (CONFIRTED)_
SOURCES POWER CHANNEL CHAKEL L/D*
PRESSURE G/10 SIALITY NUMBER REMARK DISTRIBUTION' SHAPE LENGTH, IN.
. PSIA lb/hr-ft AT EXIT, % OF POINTS 122 to 500 to
-O to 11
-01E 22079 Uniform Rectangplar.
'CU-MPR.I-8-61 Uniform Annular' 42 113.5 1000 0 70 to
-0.02 to 34 (DP685) 1 90
+0.22
- "d 0*8 to o
t GEAP-3558 "j"i {f,
Anrmiar 108 322 1000 8
0.88 to 0.11 t Uniform
!=-1 =e 108 322 1000 24 1.71 0.28
%,, w Polonik Non-Uniform Annular 108
'322 1000 16 2
1500 t 0.42 to 0.26 t TID-7529 y,1for, 3,,m 3,,-
252 656 8
(Pt.2) 2000 1 70 0.4 t
0 92 to O
t CEA-1853 Uniform
~ Circular.
19 7 80 30 9
gg 0
to to W-72255
.Unifors' Rod ~ mmdle -
18 60 1200 2
GRAP-3771-2 Uniform Anr=1 =r 29 and 58 to 1100 and 2.0 to
.12 to 60 36 1W 1400 6.2
+.12
.h '
e
l.CKNOWLEDGEMENT The authors wish to thank Miss Polly Sedlak for her help in checking the energy balmnee of the data and plotting numerous curves during the development of these correlations.
W e
w Nee en
REFERENCES 1.
meleonics, Vol. 20, No. 2, "DNB Correlations Disagree", February 1962.
2.
Jens, W. H., and Lottas, P.
A., " Analysis of Heat Transfer Burnout Pressure Drop and Density Data for High Pressu e Water", ANL-4627, 1951.
3.
Zenkevich, B., and Subbotin, V., " Critical Heat Flux in Forced Circulation of Water; Subcooled to Boiling", Atoc:1c Energy No. 8, USSR, 1957.
4.
Cook, W. H., " Fuel Cycle Program - First Gparterly Report" GEAP-3558, September 1960, p. 35.
5.
Zuber, N., "On the Stability of F,iling Heat Transfer",
Trans. ASME 80,, 1958.
6.
DeBortoli, R.
A., Green, S.
J., LeTourneau, B. W., Troy, M., and Weiss, A., " Forced Convection Heat Transfer Burnout Studies for Water in Rectangular cha - 1s and Round Tubes at Pressure above 500 Psia",
WAPD-188, October 1958.
7.
Griffith, P., "A Dimensional Analysis of the Departure from kcleate Boiling Heat Flux in Forced Convection", WAPD-TM-210, December 1959 8.
Zenkevich, B., " Similitude Relations for Critical Heat Ioading in Forced Liquid Flow", Atom. Energy 4, No. 1, pp. 7k-77, January 1958.
9.
Chang, Y. P., " Final Report, Section 1 - An Analysis of the Critical Conditions and Burnout in Boiling Heat Transfer", TID-14004 July 1961.
- 10. Kutateladze, S.
S., " Heat Transfer in Condensation and Boiling",
ABC-tr-3770, 1952.
11.
" Big Rock Point Plant Hazard Summary Report", (3/12/62 Revision),
. m
a6 y
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s.== x*L l '= t .I. p. i _._ i.. ;. 3 l 7... . _ s i e. .i i ,,.L j i 5 ..{ p. ...9 i i. i,i l 1 .s i i i. f 12 3 4 4 .i i. 3.. .j i i 4 ~ l~' j' 3-Predicted q"DNB, 10 Btu /hr,ft ._4 t O l L
WCAP-285) R W ( YAIZEE PRESSURIZER CIADDII;G EVALUATICI; ~ e ' .m Im. c.n(p~s) 42,ws. u/n-3a5.sg fai~~ l'A cs g i US O E 9 2 4'Oy,j 196Sa } s., ( t a // IIovember 2,1965 l WESTINGHOUSE D P.CTRIC CORPORATION Atomic Pouer Division P. O. Box 355 J Pittsburgh 30, Pennsylvania 1 D>.,,s.
I I. C.ummary of Analyses l The very minor amount of cracking found in the Yankee prescurizer vill not t affect operation nor pcesent any kind of a cafety hazard. l l Actually, only 6 cracks were found by dye penetrant, all in the spot velding cladding of the upper head. Although they may have been acecntuated or even cauced by operation, there is no reason to accume that they were not there when l l I the unit was placed in service. Corrosion of any carbon steel exposed vill not contribute noticeably to the general crud Joad, and hydrogen pickup froa corrosion and the H pr s ure in the 2 precourl:er is lower than has been shown in WCAP-2855 to be of no concern to the j A302 B steel from hydrogen embrittlement effects. i A fatigue evaluation has been made to determine the maximum crack depth that can be tolerated at the beginning of life in the pressurizer vessel vall. This crack depth is in excess of.6" in the steam phase, and over. 35" in the water phace portion. An ntegrity analysis was performed showing the chance of fast brittle, frac-ture o.' the precsurizer cannot occur under any transient or steady state condi-tions. l l All defects have been accurately charted with respect to location and type l Further exploratory investigaticn of questionable areas will be made before e the pressuriner is returned to service. I 1 F I 1 l d 1-1 ( i
II. Gen ral D:scription of Cladding Crneks When the Yankee pressurizer was opened for inspection, several linear deposits of rust were noted, and it was decided to carefully examine the upper portion for cracking by dye penetrant inspection. The entire head area, and the shell area down to 6" below the head-shell g* rth veld wns inspected in this me.nner. Nine y dye penetrant indications were found, all in the head area in the B & W spot velded cladding. No indications were found in any veld deposited cladding, or in the spot velded cladding below the head-shell joint. Details of the penetrant examination and description and location of all indications found tre described in Appendix A. There were 9 indications found in all, of which one was a pit, and two were classified as inclusions or foreign material embedded in the cladding during manufacture. Of the remaining 6, 4 were cracks in the ligament area between spot velds, and 2 were apparently partially in the veld nugget. It is not known how deep the cracks in the ligament (unbonded) areas are, but it could be presumed that some at least are completely through the cladding, ex-posing the SA302 B steel to the pressurizer environment. In addition, there were penetrant indications at isolated spots on the edges of the veld deposited cladding over the head-shell girth veld. As the surface st these edges was rough, with some veld undercut or overlap, spurious '.ndica-tions would be expected. Although it is felt that no cracks exist in this lo-cation, further exploration of these edges vill be performed locally f or con-firmation. Several locasions vill be ground smooth and re-inspected by dye penetrant. The cracks in the stainless cladding where it is not bonded to the base metal cannot, of course, cause propagation into the base metal. The two cracks par-tially in the nuggets vill be explored by local grinding to confirm that they do not penetrate into the carbon steel, and to alleviate any stress concentration effect of the cracks. II-l
There are several possible causes for the cracking. Spot velded cladding is l usually both highly stressed and severely sensitized as a result of manufacturing j operations. Intergranular cracking of such caterial by pickling operations or j corrosive environments would be expected. There is no reason to expect that me-l I chanical stressing or thermal cycling vill cause failure of sensitized naterial, l therefore there would be no reason to expect any cladding cracks to propagate into the A302 B alloy steel. Further discussion of these pointr is presented l ' in Appendix B. i In addition to the cracking, several other condi+. ions were noted. In or.e area, in the head, a group of spot velds had apparently been r.2ptured, and a " blistered" or " wrinkled" condition exists. As evidenced by local grinding, l this condition was probably present when the component was put into service. l l One crack was found in this row of 8 " blisters". Several veld repaired or l patched areas were also noted in the head area. One of the cracked areas was found adjacent to a weld repair. I \\ l i O II-2
III. C0 EROSION AND HYEETEN DI3TRIB'JIION Introduction The questions raised by the develcpment of a defect in the 304 stainless steel cladding of a 302B carbon steel pressure vessel have already been con-sidered in detail with respect to the Yankee preocure vessel an't is given in WCAP-2855, " Evaluation of Vessel Cladding Penetrations". From the cor-nsicn and hydrogen distribution standpoint, the cracks in the Yankee prescuriner cladding can be evaluated qualitatively in essentially the same manner; due censideration of the quantitative difference shows that in general the problem is less severe in the pressurizer than it is in the pressure vescel. In preparing this section, it is assumed that the render is familiar with the material presented in WCAP-2855, hence only pertinent changes in conditions and conclusions as stated in that document will be noted here. Corrosion The operating conditions to which the pressurizer will be expose d differ l from those in the primary vessel as follows: I i 1. The temperature is higher, ~ 636*F rather than 500*F. l 2. Both a liquid and vapor phase exist in the pressurizer, with the l vapor phase enriched in contained gases (H and 0 ) relative to 2 0 b I l the liquid phase. l l 3 The radiation level in the pressurizer is negligible ccmpared to that in the pressure vessel. Tne shutdown conditions in the pressurizer differ from those in the pressure vessel in that the radiation level is lcwer and the temperature may be lower, 80 F rather than 100*F. The pressurizer is vented and partially drained at shutdown, hence oxygen is present. Under operating conditions, reference to the data presented in WC/"-2855 shows that a corrosion rate of 150 mg/dm -mo for the base metal is still an acceptable and conservative value, both under the clad and at the bare of the e III-l 4 1 ~
,---n-
l I i cracks. At shutdevn cenditiens, due to the presence of oxygen, the cerrosien rate at the base cf the cracks may be as high as 1500 mg/d 2 -mo. Under t he i clad, however, the cxygen availability will be ecntrolled by diffusien frcm ( i the base cf the crack and hence will be low. Ccneervatively, for the calcu-lation of penetration rates, the high corrosien rate (1500 mg/dm"-m]) will l l be used; in calculating hydrcgen ecncentraticn in the steel h0 wever, the l I hydregen generation due to corrosicn can be calculated based on the lower corrosien rate (150 mg/d -20) since that portion cf the corrosicn which ic due to 0 does not result in the generaticn of H2 (see WCAP-2855). On this 2 basis, penetration of the base metal will occur at a late of 0.9 mil / year at operating conditions and 0 75 mil /=c during shutdcwn. i Hydrogen Distribution and Concentrations A. Parameters Used it. Jalculatinns i l Analyses of exactly the same type as were made for the pressare vessel l i are applicable to the pressurizer, with parameters as follcvs: ( Pressurizer Clad 304 SS, 0.109 inch Pressurizer Base Metal A302B Carben Steel Prc surizer Shell 3-1/8" Thickness, h2" Diameter Presearizer Heads 2" Thickness, 21-1/2" Eadium [ I Operating Temperature 636'F i Shutdcun Temperature 80'F ( l 2 Corrosion Este of Base Metal 150 mg/dm, 3 l f B. S6urces of Hydrogen l I The cnly sources of hydrogen available in the pressurizer are frce corrcaicn i i l and the hydrogen dissolved in the coolant for suppression of radiolysis. As was j i I shown in WCAP-2655, the hydrcgen frce corrosicn is given by i l %e 7.hk x 10' Y l j where / ccH (STP) generated per em corroding surface per hcur = 2 ( corrosien rate, =g/d=g Y -mc = III-2 i i i
In order to calculate the hydrcgen ccncentration and distribution in the precsurizer base metal, the average hydrc6en flux is used; as before, assuming the stainless steel claa and base metal are veld bended over 70% of their antact a rca, the qverage hydrogen generaticn with reference to the base metal i is 7 LL x 19'A' x 0 30 x 150 ~h o 1 a rea cr 3 4 x 10 ccH(STP)/cm"-hr,sincernly 2 01 ::f the base met al are t is expeced to corrcsien fluid. This is the cnly m ree c.f H during chutdown, and its mag itude is the same for bath shutdcun 2 und crerating condittens l At cr rating conditicnc, hydregen gas diccolved in the coolant must also e i b,; conridered. In the analysis of the precsure vesee3 defect, it vae shown i t FM t his scucce wac small compared to the corrosien source For the prescurizer, thic c< urce becemes ccmparable in significance with the corrosion hydrogen as will he chcwn, due net so much to an increase in this source as to a decrease in the cignificance nf the cerrosion hydregen as a result of the thinner vessel w111 ind higher diffusien crefficient (due to higher temperature) Since the preasurizer is a two phase boiling system, the hydrogen activity t ar presnure cannet be simply calculated ac it was, by use of Henry'c Law, for i i the ningle phase co71 ant in the pressure vessel. Detailed analycis of the pre scurizer sh 'ws t hat the following relation is satisfied during steady state cperatienc: p [cf - ll ) C r( m C/(m +m v n s s u'ry e C hydrogen cencentratien in vapor phase, mass H /ma s H O vapor 2 2 C hydrcgen concentratien in pressurier spray, mass H /* 8" HO g 2 2 li quid l /2 (m p [~1)/(m c( + m -m +m -m) = v ,s v s p i boilingrateinpressurizer,massH0/ time m =- 2 spray flev rate, macs H 0/ time m 2 m vent p er purge flow rate from precnurizer vapor space, mass H 0/ time 2 II" 7
i l ' Y = r' < k/P RT s p(( density of H O liquid at pressurizer conditions = 2 g density of H O vapor at pressurizer conditions = g 2 Henry's Law ccnctant, psi-liters /gm-mole i k = gas constant, psi-liters /(gm-mole K) R a i l temperature, K l T = t t j !te partial pressure of H, Pg, in psi in the vapor phase can then be 2 calculated from 6 1/2 C, f RT
- lO P
= g s 2 i with C in parts H per millicn parts steam, and f in gis/ liter. Typical 2 3 pressurizer ecnditiens are: 2000 psi P = T 609'K (636 F), R = 1.205 i 3 O 5 3 e/ft = s 3 Fc - 38.9 f/ft l l 3,160 psi-liters /gm-mole l k = 31.6 f / = h5 J/hr m = P ISS 4/hr (1/2 gpm 6 5h0*F) m = s h,220 e/hr (60 Kw to pressurizer heaters) m. = s 2.66 ppm (30 ce(STP)/kg H 0) l C = s 2 l from which the hydregen pressure is calculated to be 0 3h6 psi. If the vent j were shut off (m = 0) the calcalated hyd agen pressure would be 80 psi; the l P c pressure calculated from Henry's Law cn the other hand is 2.6h psi. The calcu-l lations shcw the importance of maintaining a continuous vent flou from the pressurizer. I III L D
Hydrcgen analyses of the pressurizer vapor during steady vented operation have shown from 325 to 510 ce(STP)H /kg condensate which corresponds to hydrogen 2 j pressure in the vapor of from 0 91 to 1.43 psi. These results ccnfirm the adequacy of the theoretical approach, considering the sensitivity of the calcu-l lations to the assumed vent flow rate, l C. Hydrogen Concentration in the Carbon Steel Base Metal l Only steady state values of the H n entration distribution need be 2 calculated, using the above cources of H, sin these are the worst cases, 2 and are shown to be less severe than those for the pressure vessel. 1. Steady State Ccncentration From Diffusion Eates i I' The steady state' concentration of H at the inner surface of the 2 carbon steel (where it is the maximum) is calculated from ll.h % X/D C = g where C H n entration in carbon steel at inner surface; ppm j = 2 %=H flux due to corrosion, cc(STP)/cm -hr 2 X vessel thickness, em = diffusioncoefficient,em/hr D = Th values of the quantities and results of the calculations are given in Table III-1. It can be seen that the calculated H ncentration is 2 lower in the pressurizer than it was calculated to be in the pressure vessel. l 2. Steady State Concentrations From H S lubility 2 At operating conditions, it was shown that significant H pressure can 2 exist in the pressurizer, and hence it must be shown that these pressures do not result in significant dissolved H at the carbon steel surface. 2 III-5
i I I l Table III-l Steady State H Concentration Calculations Parameters and Resulte 2 For Yankee Pressurizer = o erating Cenditions 3hutdown Conditions y i x, en 7.95 7.95 %, cc(rTP)/cm -hr 3 36 x lo-(0.1062)(C )-1.63(3 36 x lo- )N 2 g D,l ) em /hr o.356 8.61 x lo-2 i e,. opm 0.0855 1.65 ppm (Pressure Vescel)(c) C 0.'42 1.99 o - *20 3/I'~T 2 ~ (a) D 5.05 e abc've 200^C em /hr 2 -7820/hT 2 i below 200 C em /hr s D 4 32 x lo xe (b) See WCAP-2855 for rationale (c) From WCAP-2855 O 1 IIT-6
1 For this purpose, use is made of Sievert's equation relating Hp solubility to H Pressure. From WCAP-2855 2 = 59.h e -6;60/rer 1/2 C p g with hydregen partial pressure, atmoopheres p = hydrogen solubility in iron, ppm C = At the maximum calculated hydrogen pressure of 80 psi, C 0 713 ppm, = g and at normal operating conditions with venting (p % 2 psi) C - 0.11 ppa. g 3 Transient Calculation of Hydrogen Concentration During Shutdown It was noted in WCAP-2855 that the hydrogen concentration in the steel at the inner surface (maximum) can be conservatively estimated in the transient conditien from et f (t )dt 1 (7) C (7f D) ! jo (t-t ) /2 where / (t ) is the H flux into the metal as a function of time. 2 To estimate C after a four month shutdown, we must make use of the g ~fcetthat'%(t')decreasesasC increases (as noted above). Assuming _/ (t ) %ss * ( ' ~ es) e (2) i = /ss s.eady state flux (= 0.047 /i) where = 2 ir.itial flux (= 3 36 x lo-ccH(STP)/cm-hr) pi 2 it can be shown that -k t ! + (%i - p ) e ~ ! ~ !e~Ydy 6TD)! 2 %ss C k y t = g as .(3) 'IIT-7
In order to evaluate k, we assume that the H flux into the 2 steel after four =cnths is 0.15 % i. From this, and Equation (2) k is found to be 7 72 x 10" hr-Fro: Equation (3), c, = o.8k ppm after four months. The original assumption that p = 0.15 pi after four months can be checked by evaluating f = 0.1062 c 63 with C = 0.84 ppm. It is found that f, which is the fraction of corrosion hydrogen entering the steel, is 0.141 when Ca 0.8k ppm, in good agreement with the assumption that fe 0.15 This is a slightly dir-ferent approach to transient calculations than was presented in WCAP-2855 but it is felt to be core accurate. e O e III-8
IV. Effects of Hydrogen An analysis of the possible effects of hydrogen on the mechanical properties of the A302 B steel was reported in WCAP-2855 for the reactor vessel case. The same criteria apply to the pressurizer. Thc maximum hydrogen levels from corrosion in the pressurizer material are lover, particularly at operating conditions, where in the reactor vessel the steady state concentration maximum was 0.32 ppa, and in the pressurizer, 0.0855 ppm. At shutdown conditions the maximum attained at steady state would be gnly 1.65 ppm, with a buildup to n.8h ppm expected for a l+ month shutdown. Because significant H pressur an xi t in the pressurizer at operating con-2 ditions, the amounts were calculated. With normal venting, the H pr sur is 2 s'cout 2 psi corresponding to only 0.11 ppm. Even if not vented, the maximum H 2 pressure is 80 psi, corresponding to a hydrogen concentration in the steel of 0.713 ppa. None of these conditions are cause for any concern with respect to degradation of mechanical properties of the steel or integrity of the pressuri-zer. p IV-1
V. FATIGUE LIFE CONSIDERATIONS An analysis has been made of the effect of the postulated cracks in the base metal on the fatigue life of the pressurizer during its subsequent operation. The fatigue evaluation has been conducted assuming various strength reduction factors to determine the maximum existing crack depth which can be tolerated in the vessel wall. The epherical and cylindrical portions of the vessel in both the steam and water phase region of the pressurizer were evaluated in this analysia. The manway flange, transition sections between the heads and shell, and the nozzle penetrations were given special consideration since they have a stress concentration factor associated with their geometry. From this evaluation it was found that the highest stressed area was around the nozzle penetration in the cylindrical portion of the vessel. This region was then used to complete the fatigue analysis. Tne following pressure and temperature transients occurring simultaneously were utilized in this investigation: No. of 6 Temperature Transient Cycles Duration Steam Region Water Region . Pressure (*F) (*F) (psi) 1. Heat-up(LOO *F/hr.) 60 5.66 hrs. -5 -5 2000 2. Cool-down(100*F/hr) 60 5 66 hrs. 5 5 -2000 3 +10% step change in 5,000 100 sec. 11 12 - 120 power 200 see. -6 23 80 4. -10% step change in 5,000 25 sec. -6 13 100 power 250 sec. 11 13 -100 5 Loss of load & loss lo 15 sec. -32 100 500 of flow 30 sec. 100 100 500 10 min. 27 27 -1070 6. Reactor scram 100 15 sec. 5 5 - 60 30 sec. -3 14 4c 7 Control rod position 22,000 1 min. -3 5 -3 5 50 changes 1 min. 35 35 - 50 8. On-off heater cycles infinite 2 hrs. + 25
= i The values listed in the "6 Temperature" and " 3 Pressure" columns are the maximums that occur during each transient. The "a Temperature" column contains the actual difference' between the mean temperature and inside surface temperature of the vessel vall. The above transients chosen for this analysis are conservative estimates covering all modes of normal and emergency operation believed to be significant in the operation of the plant. The specified number of occurrences for each transient is a concervative _ prediction of the number of occurrencer that could be anticipated in the remn.ining operational life of the pressurizer. These estimates are based on the number of i transients the plant has experienced to date. An added degree of conservatism has been introduced into the analysis by the assumption that the accident transients will occur at the end of core life. At this ti e during plant life there is a reduced plant operating temperature relative to the pressurizer temperature which results in larger transient temperatare changes. The stress fluctuation at the tip of_ the postulated crack which were assumed to exist in the highest. stressed areas of both the water and steam phase regions of the i ] vessel were then calculated for each transient condition. These stresses were computed
- 1a a function of the fatigue strength reduction factor (K) associated with the postulated l
crack depth, the pressure fluctuations, and the temperature changes. These stress i fluctuations are listed in the ?ollowing tabulation. l Pressure Stress-Thermal Stress l Transient . psi Steam Region Water Region Cycles l psi psi 1. -Heat-up 16,700 K -1330 K -1330 K 60 2. Cool-down -16,700 K 1330 K 1330 K 60 3 +10% power step - - 1,000 K-3480 K 3760 K 5,000 665 K -1885 K 7210 K L 4. -10% power step 835 K -1885 K 4080 K~ 5,000 -835 K 3330 K 4080 K '5 Loss of load & loss 4,180 K -10,000 K 31,500 K 10 of flow 4,180.K 31,500 K 31,500 K -8,950 K 8,480 K 28,000 K 6. Reactor scram- -500 K 1,570 K 1570 K 100 '30 K -950 K 4400 K f L7. -Control rod position hPO K -1,100 K -1100 K. 22,000 changes -420 K 1,100 K 1100 K 8. On-off heater cycles +210 K ~ infinite
Various values for the fatigue strength reduction factors were then applied on a trial and error basis to establish a peak combined stress for each transient coniition. For each value of assumed r+ ength reduction factor, the minimum and maximum values of fluctuati:3 stress were determined by combining the transient stress conditions in a manner to establish the maximum stress difference for the predicted aumber of cycles. In this manner, it was determined that a fatigue strength reduction factor of 9 25 may be tolerated at the tip of a crack and still maintain the cu=ulative fatigue usage factor of the water phase region of the vessel below the allowable value of 1. Likewise, a fatigue strength factor of 15 5 can be tolerated at the tip of a crack in the steam phase region of the pressurizer. From reference (1) it was determined that the depth of a crack on the wall surface required to develop the fatigue strength reduction factor of 9 25 in the s.ater phase region of the vessel is 0 35 inches. Similarly the s." face crack depth permitted in ~ the steam phase region of the pressurizer is in exces s of 0.6 inches. Therefore, it is concluded that a fatigue problem associated with the observed cracks in the Yankee Eowe presaurizer is unlikely. (1).Langer, B. F., " Design of Pressure Vessels for Low Cycle Fatigue", ASME Faper 61-WA-18 presented at the - ASME Winter Annual Meeting, New York City, November 26, 1961.
VI. Integrity Analysis 4 A. Transition Temperature Approach As part of the engineering analysis of the Yankee plant, the hydrostatic test temperature was determined by use of the transition temperature approach. Essentially the hydrostatic test temperature is at the Fracture Transition Elastic temperature (FIE) which is accepted as being NUTT +60 F (nil ducti-lity transition temperature +60 F). The nil ductility t ransition tempera-i . ture is. determined by the drop weight test, ASTM-E-208. However, accepted engineering practice is to use a Charpy V-notch fix of 30 ft. lbs. to deter-mine the NDIT of SA302 Grade B because of the established correlation be-tween NUPf and Charpy V-notch as stated in Section III and PB 151987 (Tenta-tive Structural ~ Design Basis for Reacter Pressure Vessels at.1 sirectly Acsociated Components). Th -B & W resistance weld-clad was applied to SA302 Grade B plate material for the Yankee pressurizer shell course sections and the top and bottom heads, g. B & W cpecified a hydroetatic test tem.perature of TO F for the Yankee pres-curizer based on the transition temperature approach. This hydrostatic test temperature was determined adding 60 F to the temperature at which 30 ft. lbs. was obtained during notch toug'.iness teats, i.e FTE. Considering' t' ' same conservative criteria for possible increase in tran-cition temperature due to hydrogen in WCAP-2855, an additional 40 F should be added to the PTE before opereti;6 stresses are applied. The recommended minimum pressurization tempere.ture for the Yankee pressurizer is FIE +k0 F U or llc F. The actual operation of the pressurizer provides a minimum pres-surization temperature oof approximately 212 F. I B. Fracture Mechanics Analysis The likelihood of fast fracture originating at a postulated large crack -in the vessel wall has been evaluated. Based.on the evaluation below, it .y i has been concluded _that fast - fracture of the Yankee pressurizer is not pos-sible'. VI-l 7
4 Regardless of how the crack is postulated to form, even a crack h inches long, which is entirely through the 2 inch tv.ek vecsel head at the connec-tion to the cylindrical section, will not propagate at any of the stress and - temperature ccaditions that the vessel experiences. Irwin's/1-criterion for fast fracture from a through-crack has been applied in this analysis: o rr ' a (pgt)2 in (1) K = 1 a (dI )2 1 syg where K = the crack tip stress intensity factor (a material factor), psisOUI. a = half crack length (set a equal to the vall thickness), in.
- = tensile stress normal to the crack, psi ys, = yield st rength, p
r If the calculated value of K is less than the actual K, fast fracture vill 3 - not occur. For a through crack in a vessel vall of length twice tne wall thickness, the appropriate value of K corresponds to the condition of plane 9 stress. If the temperature is high enough, the fracture made is entirely by Shear, that is, the failure is ductile rather than brittle. For SA 302B Charpy V-notch data shows ductile behavior ab~;c 200 F. The temperature of the pressurizer is well above this while it is under load. Peak values of K, the plane stress intensity factor, are not available, but lower tempera-ture data /2, shown in Figure 1 (open circles) can be extrapolated to show the trend. The peak value (the curve levels off) is probably above 100 Ksi /Tr. The pressurizer functions such~ that saturation conditions exist at all temperatures. The' maximum saturation pressure of 2500 psia corresponds to a temperature of 668 F, and a vall stress of about 20,000 psi. Under reak conditions, the wall stress may-rise to 25,000 psi, but at the same tem-perature: 668 F.- VT-P
100 g i g. t g j i i i i g i 00$5H M6, rO-AL- = = 90 ,7r TEMPERATURE RANGE Eue cea[p.pg s,, ~ - yoX l OVER WHICH PRESSufE C;cq7g %\\ j LOADING TAKES PLACE / ~ USAEC 'C 1 Ic63> h 80 4 $9wnacm-p)j e u,..,, y D [ g d 2 ? OVER$ TRESS g 70 CONDITION E O O MEA 5URED VALWES FOR MINIMUM VALUE OF Ke 60 SA 302 B GIVEN IN TO STOP PROPAGAT!ON OF ANS TR ANS ACTIONS 4" LONG CRACK (2"WNL) AT ACTUAL STRESS N JUNE 1964 LEVEL. \\ 4 mw z o
- e-w E
o m @ 40 g o e .a 4UEM 0 m. o e U o e u O 20 FIG. I FRACTURE ANALYSISs YANKEE PRESSUR - 10 IZER. I l i I I I I O -400 -200
- 00 200 400 600 800
- 1000, TEMPERATURE, "F N
e.
m Thus, the vall stress as a function of tenperature is given by: 20,000 2, YA sat., psi (2) P where P = saturation pressure at a given tenperature, psia The stress fron equation (2) and an assuned crack length of 4 inches (a = 2") are substituted in equation -(l) to give the calculated curve of E versus.tenperature shown in Figure 1, c Considering Irwin's criterion t'or rapid fracture, it has been shcwn that a leak-before-rapid-fracture situation exists since the calculated values of K are so much less than the actual value of over 100 psi in. O b VI.
References 1. Irwin, G. R., " Relation of Crack Tcughness Measurements to Practical Appli-cations," ASME paper 62-MET-15 2. Innderman, E. and S. E. Yanichko, " Selection of Pressure Vessel Materials Based Upon a Fracture Mechanics Analysis," Transactions _ , Vol. 1, No. 1, June 1964 e VI-4
Appendix A Visual and Penetrant Exa=ination of Pressurizer Cladding Visual and penetrant examinations were performed to determine the condition of i the resistance and weld deposited cladding in the upper head section of the pres-surizer. The visual examination was perfomed at approximately 3x magnification using a conventional reading glass and suitable lipting. The penetrant examination was perfomed using Type II penetrant caterials and technique. The sequence of operations and the results are itemized below: a. Cleaning: The entire head surface and approximately 1 foot of the shell surface, in-cluding the head-shell girth veld and the 2 shell seam velds, were vigor-i ously dry-wire brushed (manually). These same surfaces were thoroughly viped with cloths saturated with Stoddard type solvent. Drying was ac-complished using dry rags an. 27 evaporation. i b. Baking: Two incandescent heat lamps were placed in the vessel and directed onto j the clad surface. The lamps were oriented 180 apart. They were rotated approximately 90 each 4 hour period. The baking time was approximately 30 hours. The surface temperature was approximately 125 F. c. Visual Examination: Prior to apply penetrant, the entire curface was visually inspected. There was no evidence whatsoever of moisture, rust or any other type of effluent on the surfaces. d. Penetrant Application: Type II penetrant, SKL-5, was applied sparingly by brushing using an alter-nating motion to work the penetrant into the surface. As soon as the entire j surface was covered with a thin layer of penetrant, a clean dry rag was lightly wiped across the surface to remove excessive penetrant so that it did not run down the lover shell surface and into the pressurizer water. The penetrant was allowed to remain for 20 minutes. Appendix A, 1
5-e. Penetrant Removal: The initial removal was done using dry rags. The final removal was done using rags saturated with Stoddard type solvent followed by final viping with dry rags. Removal was continued until the drying rag was faintly pink. i. f. Development: A very thin uniform coat of SKL-5 developer was applied by spmying from a pressurized can. g. Penetrant Examination: The indications located at three positions appeared in: mediately, i.e., as identified below: positions 20 South of West, 20 West of North, and 25 South of West. All of the remaining indications appeared within 30 minutes..(Note belov for descrinH ;n and location of all indications.) There were no additional indica. ions appearing, nor was there excessive diffusion of the penetrant -.em any of the previously disclosed indica-9 tions approximately 3 hours after developing. The orientation of all pene-trant indications is shown in Figure 1. Dye Penetrant Indications on Head Forging of Yankee Pressurizer - 1. 20 South of West 4 a):- One crack.600 long on the third blister running south to vest - crack 'is located on the south side of the blister at the base of the radius - the crack is not located in the nugget. Four inches above veld clad-top edge..
- 2.. 25 South of West
-a) _ 'One pit '.050 in diameter (depth unknovr) bleed freely. Located nine ~ inches.- above veld clad top edge. 3 -3 20 West of North a) L0ne vertical indication.100'long located in the ligament (raised Appendix A.
po* tion) located 3/4" above veld clad top edge. Does not appear to be cracked. 4. 10 East of North 2) Two cracks .100 and.050 long - crack begins two (2) inches above veld clad top edge. Crack runs parallel to a veld repair bead. In-dication is 1/16 to the' East of weld bead. Located in ligament. .i. 5 30 East of North 1 a) Indication.600 long. ' Located two inches above veld clad top edge. Located in ligament..
- 6. '20 North of East a).
Inclusion indication located 3/4" above veld clad top edge. .160 in diameter - Located in nugget. 1 4 T. Due East . a); One crack [" long located both in the ligament and nugget of stitch veld. Located 8"'above veld clad top edg'e. 8. 10 South of East a) One crack.250 long located in the ligament. Six inches above veld clad top edge.
- 9. 10 South of.. East L a)'
Three cracks'having total length of.800 very possibly these cracks i jointoformonecrack. Crack runs in both the nugget and ligament
- located two.- three inches above veld 'elad top edge.
[ -, 1 i i -Appendix A' -
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Appendix B STAINLECS STEEL CLADDING CRACKING Stainleso Steel Cladding Cracking Cracking of the stainless steel cladding on water coobed reactor components has only recently been reported. The only case reported in the open literature has been that of EBWR. Westinghouse has investigated this incident insofar as infor-mation has been made available. Because this case has been publicized, and because it bears on the Yankee vessel and pressurizer cladding integrity question, a brief summary of the facts and Westinghouse conclusions is included here. Reported Facts:
- 1. The EEWR reactor vessel was clad by the 1&W spot velding process with Type 304 stainless steel plate.
- 2. Cracking of the cladding was found and reported by Argonne.
s
- 3. The cracks were intergranular, more extensive above than below the water line.
- 4. The material was sensitized, as would be expected from the manufacturing processing and stress relief cycles.
- 5. A portion of the clad plate used for the head, and a portion of a test plate, both clad similarly by B&W vere examined and found to be sensitized and cracked intergranularly.
- 6. A recent ANL report ' indicated that the cracks were caused by stress rupture.
l Westinghouse Investigation Results: i
- 1. Samples of the head plate, test plate, and EEWR cladding were given to Westing-l house for study, as the problem might be related to all B&W spot velded cladding, or even to austenitic cladding in general.
Appendix B _1
Metallographic studies showed that the material was sensitized in all cases, and intergranular attack and cracking was much more evident in the samples of EEWR cladding. The type of cracking--not associated with noticeable deformation--is precisely what is expected of sensitized stainless steel subject to corrosive conditions. This type of failure is well docunented in the literature, and one of the clascie papers on the subject, shew 3ng the influence of stress on intergranular corrosion is that of Hoyt and Scheil This is not to be confused with transgranular stress corrosion cracking of stainless steels usually caused by chloride or other halide ions. It is known that cracking probicms were encountered by B&W during the develop-ment work in applying this technique to cladding with austenitic stainless steels. Further, severe pickling had to be resorted to,in removing iron and 9 other contanination from the clad surface. As pickling is known to cause intergranular attack of sensitized stainless steel, it is very probable that J some veld cracking and intergranular attack from pickling was present in components as manufactured. From the limited information at present available to Westinghouse, it appears that there has been service--related cracking of the EEWR cladding, that the primary cause was intergranular corrosion, possibly accelerated by stress, although it is probable that some cracking or intergranular attack was present at the time the vessel was placed in service. Mechanical properties of sensitized stainless steel have been videly studied and reported, showing no significant reduction of ductility as the result of In a particular severe case (3) sensitization. , severely carburized and sensitized Type 304 stainless steel was tested at lov strain rates at elevated temperatures. Ruptures were intergranalar, but photomicrographs showed the usual necking at failure. This is in contrast to the lack of deformation associated with the E3WR cladding cracking. Appendix B.
Fossible further evidence for a corrosion--related failure mode is that the cracking was more severe above the water line, where relatively high amounts of oxygen vould be expected. The failures cannot be explained by thermal cycling in service because the strains will be too small to cause plastic deformation after the first cycle, and if they were large enough to cause failure, local deformation would be obvious in a photocicrograph. In summary, the most logical reason for the crac';ing of the EF4R cladding is that of intergranular corrosion, probably accelurated by stress, although it is not unlikely that sote veld cracking and intergranular ateack was present originally. Yankee Pressurized Cladding Cracking The cracking found in the Yarkee pressurizer is probably related to the EFJR experience. The cladding was manufactured in the same way, and probably had sone cracking in its original condition. It is possible that further cracking occurred by stress accelerated intergranular corrosion, as the oxygen levels in ) the pressurizer were high on occasions early in life. Steps have been taken to reduce the oxygen level during operations, and this will minimize further deterioration of the cladding. 4 i Appendix B
References:
1. ANL 7071, Progress Report, June 1965 2. Hoyt, S. L. and M. A. Scheil, " Stress-Corrosion Cracking in Austenitic Stain-less Steels," AShi Trans., Vol. 27,1939, pp.191-226. 3 Stone, D. H. and A. D. Schwartz, "Effect of Carburization on the Lov Strain Rate Behavior of Type 304 Stainless Steel," NAA-SR-10487, August 31, 1965 3 J o-s Appendix B k_ --}}