ML20214H956
| ML20214H956 | |
| Person / Time | |
|---|---|
| Site: | Sequoyah  |
| Issue date: | 07/20/1986 |
| From: | Cipolla R, Proctor R APTECH ENGINEERING SERVICES |
| To: | |
| Shared Package | |
| ML20214H850 | List:
|
| References | |
| NUDOCS 8612010102 | |
| Download: ML20214H956 (9) | |
Text
i APPLICATION OF LEAK 8EFORE BREAK ANALYSIS METHODS TO PRIMARY SYSTEM BOLTED CLOSURES R. C Cipolla, Principal Engineer and R. R. Proctor, Engineer Aptech Engineerine Services, Inc.
Feio Alto, California e
ABSTRACT basis for integrating appropriate mitigating measures, A strategy is proposed that will establish leak j such as preload control, nondestructive examination rate margins and nondestructive examination limits for (NDE), and leak detection capabilities, in order to bolting materials corsnonly used in primary pressure assure the integrity of the primary pressure boundary.
boundary closures. In the application of leak-before- A Bolted Joint Integrity Program has been break analysis methods to closures, an analogy is sponsored by the Electric Power Research Institute drawn between a welded joint and a bolted joint with (EpRI) with the main objective of obtaining a better regard to structural redundancy, load shedding understanding of the behavior of bolted closures.
behaylor, and early warning detection created by the Primary emphasis is placed on the safety acceptance of
~ presence of a leak. Analysis methods for determining the degraded bolted closure, but it is expected that the structural behavior and leakage of a bolted improvements in closure reliability will occur as closure for various amounts of bolt degradation are well . The purpose of this paper is to present a leak-presented. Calculations have been completed for two before-break strategy for resolving bolted closure steam generator / pressurizer manway cover designs, two integrity issues as the continuation of past work (5) check valve flanges, and a reactor enolant pump main and to show how this approach could be implemented flange. Results indicate that leak rates in excess of through the ASME Boller and Pressure Vessel Code (6).
1 GPH (0.042 kg/s) and as high as 10 GPH (0.42 kg/s) are possible without compromising the closure LEAK-BEFORE-BREAK EVALUATIONS FOR CLOSURES
~ '
integrity M gnificantly.
The leak-before-break criterlon was originally ,
.!NTRODUCTION proposed in the late 1960s as a means of estimating the necessary toughness of pressure vessel steels 50 Recent service expertence with primary pressure that a surface crack could grow through the wall, boundary bolting in pressurized water reactors (PWR) causing leakage of vessel cor. tents to detectable indicates carbon steel fasteners can become degraded levels before fracturing. As a result, this philo-as the result of prolonged contact with primary sophy has been effectively used in the assessment of integrity issues for welded pressure vessels and coolant water at elevated temperatures (1, 2 3). The piping components fabricated from ductile materials.
closures that have experienced bolting degra[fation include primary side manway covers of steam generators if a crmprnent exhibits a leakage failure mode prior and pressurizers, coolant pump main flanges, and some to the point where the actual integrity becomes l primary valve flangas. Of the closures listed, the questionable, then the demands on NDE methods other steam generator manway covers have been the most than leak detection can be reduced. Hence, the troublesome (4). objective of a9v leak-ttfcie-break analysis is to show Ipdividual fasteners have been observed to suffer that leaage will always precede failure by a suitably safe c.argin.
from general corrosion (wastage) at the shank or threaded sections or from stress corroston cracking The basic similarities between a bolted closure
, (SCC) at the thread root.' Although degradation of and a welded joint with respect to material selection individual fasteners has raised some questions with design requirements, control of fabrication processes,,
regard to closure integrity, operating experience also and preservice inspection suggest that an assessment suggests that only a small numbe- of closures have plan for closures could make use of a leak-before-break philosophy in much the same fashion as with actually degraded while in service. By focussing on these " service sensitive" closures, a generic plan for welded pipes or vessels. Since one of the principal addressing the integrity of a joint could be deve- design features of a bolted connection is its o
loped. Such a plan would also provide a rational structural redundancy it seems plausible that a bolted closure, even w,ith some degraded or failed 5
8612010102 861117 PDR P
ADOCK 05000327 PDR
$ _ - h. _ _ _ -
C * *
. ,. l I
=
l fasteners, could meet acceptance criteria consistent The parameters that govern bolt degradation and witti current industry practice provided that ample ultimately the integrity of a closure would naturally safety margins and closure reliability could be include the material condition, the closure loads, demonstrated. As an alternative to current emphasis and the environment being contained. Because the SCC within the ASME Code on individual fastener integrity, susceptibility of low alloy steels increases with
.an assessment strategy is proposed that will establish ' increasing strength, those parameters that affect the acceptance of a closure provided thet the variability in strength are the most important:
following conditions are met: specifically, material specification, heat treatment, ,
and nominal strength level. The stress-related e Leak-before-break of the closure is assured variables include preloading method, preload level, under the design basis conditions for the plant anticipated service loadings, and the joint stiffness and load redistribution characteristics of the e The safety consequences of closure leakage are closure. Given suitable numerical methods, the
, acceptable closure displacements and bolt stresses can be computed for a wide range of degraded bolt conditions.
e The margin against break at the point when the Finally, the environment variables include tempera-leakage becomes detectable exceeds an ture. humidity, and the presence of corrosive agents.
acceptance level These environmental effects are used to estimate the range of possible initial degraded bolting conditions A proposed assessment strategy for bolted clo- prior to the application of service loadings. Based sures that exploits the leak-before-break philosophy on these postulated conditions, requirements for is depicted in Figure 1. The suitability of this alternative MOE measures can be proposed once the strategy to closure evaluations will depend on resulting leak rates and available safety margins are available margins as dictated by the conditions established for a given closure design.
required for closure failure, the amount of external leakage from the closure, and the availability of leak SERVICE SENSITIVE CLOSURES detection instrumentation. Clearly, the character-1stics characteristics of joint behavior in terms of The focusing of inspection and maintenance load redistribution and gasket unloading followed by ictivities on service sensitive closures will allow flange separation must be quantified for a valid and for more effective resolution of equipment leakage accurate determination of safety margins. Load problems. During the investigation of primary changes within the joint are due to postulated bolt pressure boundary bolting problems, the AIF/MPC Task degradation (wastage or cracking due to corrosion) Group on Bolting and EPRI developed a Bolt Failures that will cause the degraded region of the closure to Data Base with a specific objective of identifying unload at the expense of neighboring regions which now troublesome closures. The failure data were compiled must carry a greater portion of the pressure loadings.' primarily from utility responses to IE Bulletin 82-02 and Licensee Event Reports up to September 1984. It was the intention of the AIF/MPC Task Group that this timre field infomation, along with historical data on plant
'"*'" specific closure performance from preservice and hydrotesting, will help define the service sensitive closures.
e The Bolt Failure Data Base was used to estimate rejection rates for fasteners used in five closures guir .t.i
- I t =t ri,i (4): .steem generator manways, pressurizer manways, ruters
- 8****'" tree,rties valves, reactor coolant pump (RCP) seals, and RCP flanges. A surriary of bolt rejection rates for all reported causes including boric acid corrosion, mechanical damage, cracking, etc., is given in
" Table 1. The rejection rates were computed on two bases: first on the total number of bolts at risk and Lua again on the total service years for the bolts at
"*j'"g'"
. risk. On either basis, the ranking of closure type is the same with the steam generator manways exhibiting the highest frequency of fastener replacements. The o RCP main flange, pressurizer manway, and valves i greater than six inches (15 cm) in diameter were also j n troublesome but all exhibited rejection rates less than half that for the steam generator manway.
The cause for rejectio of generator menway studs
,8 M', c,in r si.re r. was principally dye to boric acid corros. ion as shown
- of the causes for rejection. J Although one can argue on the overall conclusions-that can be reached from these limited data, the stu,ii7,.,at
, o' information does help to focus the types of components
=rgin requiring utility attention for improved maintenance
- practice as well as identifying candidate closure designs for evaluation by leak-before-break analysts Figure 1 - Closure Integrity Assessment Strategy. methods.
6 y k , +- .
,, . y 2%p ys %H e hs x
l Table I !
the gasket. The finite element mesh for the 16-stud REJEC[10N RATES FOR BOLTING IN PRIMARY cover is shown in perspective view in Figure 2. The ,
PRESSURE BOUNDARY CLO$URES (ALL CAUSES) studs were modelled by beam elements which were
. connected directly to the solid elements. Orthogonal 1 % rigid links were connected to the beam element end Closure Type Total Bolts Total Bolt nodes to induce stud bending when cover and flange At Risk Service Years surfaces do not remain perpendicular to the stud 4 1
during loading. To simpif fy the analysis, the gasket Steam Generator Manways 5.811 was codelled as an elastic foundation represented by 3.98% discrete uniaxial elastic springs. The elastic Reactor Coolant Pump 2.85% loading and unloading behavior of a spiral wound Main Flange 2.48%
asbestos filled gasket was inferred from experimental Pressurizer Manways 2.28%
cyclic stress deflection curves (8) and used to 1.20% define the spring element stiffnesses. The Valves (>6 inches 2.10%*
(15.2cm) Diameter) deflections from the tests were matched to the actual manway by relating the gasket properties through Reactor Coolant Pump 0.82% ratios involving stress area and gasket thickness.
Seal Flange 0.851 Because of the massiveness of the vessel flanges, the flange surface was assumed to be rigid. Stud preload was established in the model by imposing a
- Estimated (Note: For All Bolts at-risk is Primary Valves ByofS'tatistical the total population bolts inser-Analysis (4) differential temperature between the stu cover. A nominal preload of approximately 30 ksi vice for the given closure during the reporting period)
(207 HPa) and internal pressure of 2235 psi (15.4 MPa) were used in the study. Stud degradation was Table 2 modelled by changing the area of individual bolt STEAM GENERATOR MANWAY elements to simulate partial wastage or by removin STUD REJECTIONS BY CAUSE bolt elements to model complete fastener failure. g Cover separation was predicted in the 20-stud' Bolts ! Of model when approximately two studs were assumed to Cause For Rejection Rejected have failed; whereas, in the 16-stud manway, the Total cover first lifted away from the gasket when one stud Boric acid corrosion 116 37.1 was assumed to have failed. When increased amounts Galled / mechanical damage / 65 21.3 of degradation were permitted, including multiple thread damage / removal damage stud f ailures, a redistribution in both gasket and Pitting / removal damage 65* 21.0 stud loads was observed. The change in gasket load Stress corrosion cracking 32 10.3- in a 20-stud manway from the "preload only" case Linear indications 16 5.5 through to various degrees of stud failures under Cracks internal3.pressure of 2235 psig (15.4 MPa) is shown in 5 1.6 Figure Corrosion / erosion / steam cut 4 1.2 The uniform gasket load becomes nonlinear Corrosion / mechanical damage 3 1.0 as the studs degrade and eccentric pressure Icading Other causes gasket compression to shift. The angular J 1.0 position at zero gasket load indicates the extent of TOTAL cover separation.
310 100.0 Stud load redistribution was most significant for
- 61 at one facility for one event the five studs nearest to the degraded region.
Figure 4 illustrates the load shedding and redistri-bution characteristics of the 20-stud manway for a ANALYSIS OF TYP! CAL CLOSURES
_ Primary Manway Cover Although there are more than 300 primary manway '
covers in use in steam generators and pressurizers of United States plants, the basic design is very similar .U.o in all applications. fiost covers are typically i 27-inch (69 cm) diameter circular plates covering a '
I 16-inch (40.6 cm) opening. The cover is 5.75 inches (14.6 cm) thick and held to the vessel by 16 studs.
The 16 studs are fabricated from AISI 4340 steel according to either ASTM A193-87 or A320-L43 speci- /
fications and are 1.875 inches (4.76 cm) in diameter.
A 20-stud manway cover of similar geometry is also used by one PWR vendor. The 20 studs are smaller in size, typically 1.3 inches (3.3 cm) in diameter and y 'e jd}"*"
fabricated from sfrd'ar materials, A three-dimensional finite element model was s-in eia developed to study the deformation behsvlor of both cover' designs as a function of stud preload and ' J different degrees of stud degradation. A general -s.ii.
ei-m purpose finite element computer program called ANSYS Q) was used to solve for cover displacements as a e function of circumferential position, and the --
conditions under which the cover would separate from Figure 2 - Sixteen-Stud Hanway Cover Model.
7
_ ,_ w dc
e
- e
- ,t s
I 4 4 4 8 4 5 & 4 4 4 gg - 20*stv4 N***1 Cever _
\
16 - .
Preload and pressure 64 . (all stves latect) ,
!r
-t f
/
10 - .
One sted removed -
6 -
4
- Three studs se.ee studs .
r - .
I i i/ I I I I I I f
,0 30 to ,0 120 460 480
. aa, vier resition. e toe,rees Figure 3 - Gasket L'oad Redistribution For Different Stud Failure Conditions.
(NOTE: 1 KIP = 4.45 KN) range of conditions including a worst case of seven The finite element model representing 180' adjacent or contiguous studs completely failed (1005 segment of the pump casing, flange, and cover is shown degraded). It is observed that the two studs nearest 3
to the failed region receives the greatest increase in load, while the second and third nearest neighbors 20-st.4 m=.f ce.ee 38 -
receive a smaller fraction. The load in the fourth ,,,,,,,,,,,,,,
and fifth closest stud decreases with the unloading e
- 2235 pit its.e nas) 2.s -
caused by the reduced stiffness of the cover / flange s joint. The applied pressure loading performs a ***=
greater amount of work in deforming the more flexible r.s -
(degraded) portions of the closure, while slightly less work is done on the greater stiffnesses of the 2.4 -
undergraded portions of the closure.
A similar trend in load change is observed in the z.: -
16-stud design except that the load increase in the nearest stud is greater due to the fewer number of r.o -
studs and greater angular distance between the faste-ners. Here, load redistribution was most significant s.a -
for.the three nearest studs to the degraded region, as !
r shown in Figure 5. Only the first two studs share an , i,, _
ste* marest to
'*"d""
increased load whereas the third nearest is observed E g,,,,,
to unload. Percentage-wise, the stud stress increases : ,,, _ merest f aster in the 16-stud manway for a given amount of ::
closure damage, but larger amounts of leakaga would E 3,, _
ntre =. rest also be expected.
t0 Reactor Coolant pump Main Flange g The main flange and cover of a Type E RCP was fourtn warest evaluated in the same manner as the manway closure. ~ '
The pump cover is composed of ai insert plate and ~
bolting ring with a bolt circle diameter of ap.proxi-mately E8 inches (147 cm). The insert and ring is 0*
- held to the pump casing by 16 studs 4.75 inches (12 -
cm) in diameter and approximately 36 inches (91 cm) long. The opening of the pump casing is 48 inches 82 -
(122 cm) in diameter and the outside diameter of the ringisapproximately80 inches (203cm). The studs 88, ) , $ .. [ l l g are fabricated from A!SI 4340 steel. Because the %, ,, c.,, , ,,,,, , , n ,, g, g, mating flange on the casing is of comparable site to e the cover, the pump casing was also modelled so that Figure 4 - Load Redistribution in Five Nearest the compliance of the mating flange is well repre- Studs to Degraded Region in a sented. 20-Stud Manway.
8 O- --
- -- - 1 -
- 8. 6 -
studs (beams) and the co..*/ body to give an approxl.
- t. s -
st.s 4'e mate stretch of 25 mils (0.64 m) translating to a 0
o 000 stud stress of about 25 kst (172 MPa). Internal pump f.: -
, 9.
pressure was assumed to be 2250 psi (15.5 HPa).
. ortpui sirni -
The unloading of he flange as studs were removed ca. 4eirasatical t.o -
h I.Nie.e was similar to that uuserved for the manway cover ei . ::35 esas (15.4 *al - , except that gasket unloading was more untform with little or no increase in gasket compressive load.
- a. ear
~ Ring / insert plate is expected to separate when only one stud is assumea '.o have failed. The increase in g*88
~
~ stud stress for vartous degrees of stud degradation is i
shown in Figure 7. lhe four closest studs to the 2 8* -
degraded region are observed to carry increased 3 sneasanent amounts of load above their original level of approxi-
- 1.2 - mately 35 kst (240 IN.). As with the manway cover.
e f 1- the two studs adjace.it to the degraded regica receive
- t.e the largest increase in load. The load ratios are greater than the manways because the pressure load is
- 0. 8
- about nine times greater for the pump cover.
Third erarest 3.2 o.s - -
e 3.0 .g p,,, p g ,,,, -
38 -
$0 p $s
- e. - -
I 1.6 -
, I E t f 0 l ! 3 4 $ j E4 -
=v-eer et coati . . ra sies st.4 -
Stud taoes aeares..t re, . to t.t -
Figure 5 - Load Redistribution in the Three -
Nearest Studs Due to Stud Degradation 88 -
in a 16-Stud Hanway. sneas auent -
.s - l in Figure 6. The model is comprised of 1200 solid -
elements with the studs being represented by beam t, I.6 -
elements and attached to the solid body in the same
- manner as the manway model. Two pressure retaining i.4 -
rni,, ,,,,n t gaskets are used in the actual assembly of this pump; 2 -
however, a single line of gasket (spring) elements of equivalent area and location is employed in the model j .t -
to simplify the model geometry. The studs are rerteaurnt preloaded by a differential temperature between the i ., -
0.s -
j/ rYag a i rt plate) l 0.4 -
0.1 -
,, I f f I I O + 3 2 3 a 5 e nueer of Centtivow Faded Sted.
Figure 7 - Load Redistribution in Reactor Coolant Pump Main Flange.
Reck Valves D'" Two bolted flange check ulves, one small six-inch (15 cm) swing check and another larger ten-inch (25 cm) check, were analyzed in similar manner as the previous closures. The valve flanges complied with ANSI B16.5 steel pipe flenge design, 1500 lb class.
Check valves were selected for evaluation becaus/ they
'"'*' exhibited the most flange leakage problems as docu-mented in the Bolts Data Base.
The six-inch (15 cm) valve has a 14.25 inch (36.2
. Figure 6 - Reactor Coolant Pump Model. cn) diameter cover with a neck diameter of 7.825 inches (19.9 cm). Twelve 1.25 inch (3.2 cm) diameter 9
i k mLmi x_;~ * ^^ ^ ^ ~ ~
studs hold the cover to the body with a specified The unloading of the flange due to stud degrada-preload torque of 500 ft-lbs (680 J). The ten-inch tion was similar to the manway cover analysis except (25 cm) valve has a 19.875-inch (50.5 cm) bonnet that the ten-inch (25 cm) valve was less untform, c:vering a 11-inch (28 cm) diameter opning. Sixteen probably because of the nonsymetric valve flange body studs,1.625 inches (4.1 cm) in diameter, are used in geometry. The redistribntion of the original 37 ksi this design. The stud material is the same for both . (255 HPa) bolt stress is shown in Figures 10 and 11.
valves, specifically ASTM Alg3-87. Cover separation is expected to occur at some point A three-dimensional finite element nodel of each after two contiguous studs have failed although the .
valve is shown in Figures 8 and 9. Both models are specific analysis to show conditions for bonnet lift-180' sprietric representations containing approxi- off was not performed. The load redistribution in the c:stely 700 elements in each. Because of the impor- ten-inch (25 cm) valve was relatively uneven between tance of flange stiffness on stud load, the valve two and four contiguous stud failures; however.
bodies were also modelled. The basic modelling of the because of the greater density of studs. significant studs and gasket follow that of the previous analyses. load is carried by the two studs nearest the degraoed A uniform preload of approximately 35 ksi (240 HPa) region of the closure as compared to the smaller was applied to the studs and the internal pressure was valve.
2250 psi (15.5 HPa).
a.: , , , , ,
3.o si.. .o cnui ,,in -
e, t.s -
N l'i$oYiE. Eld [
r.. -
a n.
2.4 -
t2 -
searest it.e to failed regt.a
,. i.s
, 2 esi.e.n
[ I.4 *
.a ,,,,,, j sneae . .it
. s.t -
,/ -
E Figure 8 - Six-Inch (15 cm) Check Valve flodel.
- 1.0 m
Thlre nearest 0.6 - ' -
o.4 _
( sena.: o.: -
t i f f f "o
jj ;;; '+-
a r 2 e s 6 lI1 l
9 f -
u., .c c..ii, . r.n. si.e.
_\
l / llll
/ / ll 7* Figure .0 - Load Redistribution in Three Nearest I ~( Studs to Degraded Region in a 6-Inch C-f (15 cm) Check Valve.
i ---s 'Ml, LEAK RATE PREDICTIONS 1
l l _
l -_I p/
F Model Description i
3 Selection of an appropriale model fcr predicting
'g flow through a slit will depend on the fluid conditions p d ^a p g ,.. '
p% , . ,
, 'h,'
and geometric characteristics of the crack (Figure 12). In this case, trie slit is represented by,
'* the gap between the unloaded portion of the gasket and l the previously mating flange surface. The ratio of j
' flow path length to characteristic dimension (i.e..
hydraulic diameter) defined as L/D. is used to specify the degree of thermal nonequilibrium of the escaping fluid. A leak rate model following this approach Figure 9 - Ten-Inch (25 cm) Check Valve Model. based on Henry's horaogeneous nonequilibrium critrical 10 m -
-- L d --
1 s.
. . i . . . . . Flow fills
. , ,ii.
u -
.... y c . . .. . .. - ca..in, ig m m .smssso,,,,, .i.. . - , , Jy O'utathisT0:1 '
s.s
..J i.* -
7
_m,,_ . .
%.,,m_, p t.t - _ saco.iea_ _,,,,,,,,- % ,a. e lieutaJet[ ,, =tature l ;
s.0 -
~
st.e ... : t.
.s -
H'8 as'a (a) Two-Phase Flow Through a Long, Narrow
_ 511t.
1.4 -
4
- . l.4 -
, Length E
seceae nearest
/
. #ag talt plane 0.0 - Ih 8 " t . # Al (area at t/0 = 12) 0.4 - # *"" "" 8 8 '
L- / Ag intet plane
~
a.1 -
0~0' ' ' I I I 1 i l
' Direction of flow 0 i a s s i 4 nu.eer et cut. e : rest e st ' (b) Critical Flow Model For Leakage Through a Slit.
Figure 11 - Load Redistribution in a Ten-Inch (25 cm) Check Valve. Figure 12 - Critical Flow Model For Leakage Through a Slit.
flow model (9) has been developed by Collier (10 and subsequently modified by Abdo11ahlan and ChexaT()1 where a' is the momentum density equal to mixture The general features of the discharge of initially],). density (p for the homogeneous flow assumption.
subcooled or saturated liquid through a slit is shown Eque. ion (f)) can be integrated along the flow pat in Figure 12. In the region 0 < L/D < ), a liquid jet evaluate the overall pressure drop across the slit as surrounded by a vapor annulus is forne t For lengths the sum of individual drop components to give:
between L/D = 3 and L/D = 12, the liquid jet breaks up AP =
into droplets at the surface and small bubbles are total APe + APat + aPaa + APf (3) entrained within the jet. It is assumed that no mass ,
or heat transfer takes place between entrance and where AP is the entrance pressure loss, aP and AP L/D = 12 and also the friction pressure drop in this are the Icceleration pressure drops due to fTuid phaf$
region is negligible. change and area change, respectively, and AP is the l The fluw is assumed to be isenthalpfc and homo- frict9 s re ss f geneous, and all nonequilibrium effects are introduced through a single parameter whieb is a function of iterative process for a given set of stagnation equilibrium quality and flow path length to di,ameter conditions and s'ft geometry. The details of the ratio L/D. The one-dimensional mixture mass and 11 12 . For l momentian centervation equations are used to evaluste numerical situations where procedure the flow isare notgiven choked, elsewhere (Id Tia)k 1
the pressurs drop components. The continuity equation is calculated from single phase relati(ns v1th 15: friction included:
l h + kh = 0 (1) G =
29, 0
8 (4)
I o where G 15 the mass flux, A is the slit openirg area' and Z is the direction of flow coordir. ate. The momentum equation is:
where g is gravitatfoaal acceleration, P and pressure and specific volume at stagnatioR. andh are is the back pressure. Calculated leak rates by the $bove I
dP 11 d GA - fC 2 methods have agreed well with experirrental studieg (j 0,) .
! ~H*{ . A H (~7}.
- 26 T" I2}
l Computed Closure Leakages The leak rate for each closure was calculated by I PICEP (12) which was modified to acconnodate the 11 l
l
-f . .
. . _ _ N ._ - - - - - --
2 , F, expected slot openings for the bolted flange relatively high margins at the 1 GPM (0.042 kg/s) leak
' connections-as detehnined from the finite element rate.
results. The subcooled fluid conditions for a pressure of 2235 psi (15.4 HPa) and 2250 psi.(15.5 IMPACT OF CLOSURE INTEGRITY ASSESSMENTS ON MPa) of a temperature of 600*F (316*C) were assumed. MONDESTRUCTIVE EXANINATION
. Leak rate estimates for all the closures analyzed previously are presented in Figure 13. The pump main With reference to the requirements under
. flange showed the greatest capacity for producing -
Section XI of the ASME Code, two areas where closure large leak rates owing to the large diameter of the integrity assessments would affect NDE are the extent .
sealing surface and smaller number of studs per arc of examinations (IWB-2000) and the flaw acceptance length. The manway covers and valve bonnets exhibit similar leak rates and trends. Being a smaller standards (ll*3-3000). The extent of examination for pressure retaining bolting is divided into two.
closure, the six-inch (15 cm) valve is predicted to categories as dictated by bolt size. Category B-G-1 produce smaller leak rates at lesser levels of stud covers principally volumetric examination of bolting degradation; however. significant leakages are whose diameter is greater than two inches (5.1 cm).
possible once degradation has extended to a larger Category B-G-2 is for bolting two inches (5.1 cm) in percentage of the bolting. diameter or less with visual surface examination specified only. These NDE requirements were developed
,,2 from conventional bolted joint fabrication t
- I I I I
- - applications; however, nuclear power plant field
- experience presented earlier suggests that the
- ,,,. % - volumetric / visual examination cutoff at two inches eneci ..i (5.1 cm) may require reassessment. If these fleid 06 $td58 data provide a statistically representative measure of
- ~: ' primary pressure boundary closure performance, service
- "Ha 'laae' sensitive closures could be identified and
T "'"**'8 -
appropriately ranked and NOF requirements established based on known closure performance and on likely
} N go.,,a ,,,,,, -
failure modes. The NDE requirements developed from esuch an approach would not necessarily be the same as C,"' '
those in the present 1983 edition of the Code. It j , /
/ /
/ --
15'8 would be expected that any alternative approach would emphasize volumetric examination with supplemental i visual /volunetric NDE for those situations when d a
/ / : a leakage from the closure has occurred during service.
, rf /
". Category B-G-1 acceptance standards for nonaxial 3 :/ j' 2 -
" flaws are 0.250 inch (0.64 cm) and one inch (2.5 cm) 1
~ /
for axially oriented flaws. Closure assessment based
- . / I on leak-before-break will provide a relat1onship
/
to-r :
{ between leak rate and closure integrity as measured in si..i.cn en.ca ,,i., - terms of bolt degradation. By selecting a minimus I[ :
j / ::itws: i required safety margin, which may vary for different 3 : 3 service loading levels, the results from a closure no-i -
- - assessment wov1d give the basis for establishing NDE requirements. The logic of integrating a leak-before.
/ breal; philosophy into a determination of requirements f
f -
and criteria for NDE is shown in Figure 14. From an
/
- established set of safety margins, a range of degraded f f ".3 conditions would be postulated that maintains a f ~
constant level of closure safety. Leak rates are f I computed for the range of postulated conditions and
, l, ,
I'#,
, , the minimum leak rate used to establish detecta-
, 3 . , , bility limits. Likewise, the type and extent of n, r et c,ati, reites stwi . degradation used in the analysis for leak rates provides the basis for selecting NDE requirements and Figure 13 - Leak Rate Predictions For Different levels of acceptance. Clearly, the example analyses presented herein provide sufficient bases for Primary System Closures. Initiating Code revisions.
For the closures analyzed, a leak rate of one SUtg4ARY Ale CONCLUSIONS gallon per miiute or 1 GP't (0.042 kg/s) is achieved when approximately one to three studs have failed.
The available margins at 1 GPM (0.042 kg/s) are shown Closure integrity assessments will provide a
.in Table 3 where the safety factor is based on load rational basis for recorsnending revisions to presert Code NDE requirements. Satisfying a leak-befue break required to fall the stud nearest the degraded region criterion is an effective strategy for assuring by net section tensile overload. In this determina-tion, the direct (o and bending (o ) stresses were closure integrity while at the same time reducing conservativelyadde%)andcomparedw1Ihspecified demands on NDE. Preliminary analysis of various .f primary pressure boundary closurss (steam geherator sinimum strength properties. Thesix-inch (15cm) and pressurizer manways, RCP main flanges, and check valve exhibits the smallest margin for the condition valves) suggests that integrity can be assured by where 28% of the studs are gone, but because of the monitoring closure leakage in excess of operational
, smaller pressure load, a safety factor of 2.2 still limits. Large leak rates are predicted when a few exists. The pump and manway covers all exhibit fasteners are assumed to have failed. Adequate 12
D*.
i, 6
,. Table 3 ASSESSMENT Of MARGINS AT ONE GPM (0.042 kg/s) LEAKAGE
. Percentage Computed Assumed Failed Studs ' Factor Of Bolting For One GPM safety At One GPM Closure Type Material Leak Leakage 16 Stud Manway Cover SA3?0-L43 15.9% 3.2 20 Stud Manway Cover A540-824 14.5% 3.0 RC Pump Flange A193-87 7.8% 3.3 6 Inch (15 cm) Check Valve A193-87 27.5% 2.2 10 Incn (25 cm) Check Valve A193-B7 17.8% 2.6 safety margins can be demonstrated provided that . 3. Hall, J. F., "A Survey of the Literature on Low-closure damage is local and that boltjng materials are Allow Steel Fastener Corrosion in PWR Power suf ficiently ductile as to tolerate heavy damage Plants," Electric Power Research Institute, induced by corrosion. Topical Report NP-3784 (December 1984).
4 Capener E. L., and R. C. Cipolla, " Evaluation of
'* Bolting Service Experiences in Primary Pressure
,,,',',','l',l,,,,'l'lt, Boundary Closures," Aptech Engineering Services, Inc., Report AES 8111290-3, EPRI RP2055-5 (December 1984).
- 5. Nickell, R. E., R. C. Cipolla, and E. A. Herrick,
,,ii.i.i. o.. . e cesiu m "The Use of Leak-Before-Break Criteria and Assessment of Margins in Addressing Closure Integrity Issues," Paper D6/3, SMIRT-8 Conference, Brussels, Belgium (August 1985).
- 6. American Scciety of Mechanical Engineers, Boller cios.t. i., tea and Pressure Yessel Code,Section XI, " Rules for the Ir rvice inspection of Nuclear Power Plant Compc' ts," 1983 Edition.
- 7. DeSalvo, G. J., and J. A. Swanson, "ANSYS -
Engineering Analysis Systems, User's Manual,"
titauin soc titauin au.,im, Revis1on 4.1, Swanson Analysis Systems, Houston, o,t s,t.a.nin u ,t.a .u e. ir ati c r i t'
PA (March 1.1983).
- 8. Bazergut, A., "Short Tenn Creep and Relaxation Figure 14 - Flowchart Showing the Detemination -
Behavior of Gaskets," Welding Research Council, of NDE Performance Requirements and Acceptance Criterion. Bulletin 294 (1985).
- 9. Henry, R. E., "The Two-Phase Critical Discharge ACKNOWLEDGEMENTS of Initially Saturated or Subcooled Liquid,"
Nuclear Science and Engineerinc, Vol. 41 (1970).
This work was performed under the sponshorship of the Electric Power Research Institute, Palo Alto, 10. Collier, R. P., et al., "Two-Phase Flow Through California (USA), Research Project 2055-5. The Intergranular Stress Corrosion Cracks and authors also wish to acknowledge the assistance ResuIting Acoustic Emission," Electric Power provided by the AIF/MPC Task Group on Bolting with Research insitute, Report (in publication),
special thanks extended to Mr. Kenneth Moore of the Babcock and Wilcox Company Hr. Edgar Landerman of 11. Abdollahtan, D., and R. Chexal, " Calculation of Westinghouse Electric Corporation, and Mr. Walter Bak Leak Rates Through Cracks in Pipes ar:d Tubes,"
of Combustion Engineering, Irc. Electric Power Research Instit'ite, Rep;rt NP-3395 REFERENCES ' December 1983).
- 1. 12. Norris, D., A. Okamoto, B. Cl.exal, and T.
Merrick, E. A., and T. U. Marston, " Background Griesbach, "PICEP: Pipe track Evaluation
- and industry Response to the issue of Bolting Program " Electric Power Research Insitute, Degradation or failure in U.S. Ccenercial Nuclear Special Report NP-3596-SR ( August 1984).
Power Plants," Paper-06/1 SMIRT-8 Conference, '.,
Brvssels, Belgium ( August 1985).
- 2. Anderson, W., and P. Sterner, " Evaluation of Responses to IE Bulletin 82-02," NUREG-1095 (May 1985).
13
- r -