ML17174A782

Sys Matls Analysis Dept Rept on Exam of Cracked Reactor Head Closure Stud from Dresden Unit 2

ML17174A782
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Site:	Dresden
Issue date:	12/28/1990
From:	COMMONWEALTH EDISON CO.
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ML17174A781	List: ML17174A780 ML17174A782 ML20082A693
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M-7058-90, NUDOCS 9107110216
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SYSTEM MATERIALS ANALYSTS DEPARTMENT REPORT ON M-7058-90 12-28-90 THE EXAMINATION OF A CRACKED REACTOR HEAD CLOSURE STUD FROM DRESDEN UNIT 2 BACKGROUND The Dresden 2 reactor head is held in place by 92 closure studs.

During a normal ISI inspection in January 1989, ultrasonic testing reported cracks in the lower threaded portions of head closure studs number 61-198-47 and 61-198-70.

The inspection technique used was a straight beam examination from the upper end of the studs.

Subsequent magnetic particle testing on the removed studs verified that cracks were present.

Additional ultrasonic testing using an ID bore probe estimated the maximum crack depths to be 35% through wall (0.88") for stud 61-198-47, and 83% through wall (2.09") for stud 61-198-70.

In an effort to determine the cause of the cracking, stud 61-198-47 was sectioned to allow for metallurgical evaluation.

Stud 61-198-70 was retained at Dresden Station for use as a calibration standard in future ultrasonic examinations.

The reactor closure studs measure approximately 65 inches long with a 6 inch diameter and 1.1 inch diameter bore hole.

A 27 inch long section containing the lower threaded region of stud 61-198-47 and 12 inches of the shank was removed for examination.

Because an unconditional release could not be obtained, the failure analysis was performed at Argonne National Laboratories under the direction of SMAD-Metallurgy Group personnel.

The stud material is specified as ASTM A320-Gr.L43, a quench and tempered low alloy steel that is 5.1.1 4334M

..------------~-

ATTACHMENT A CECO EXAMINATION REPORT ON REACTOR VESSEL HEAD CLOSURE STUD 61-198-47 ZNLD/899/38

similar to AISI Grade 4340.

EXAMINATION M-7058-90 Page 2 A visual inspection of the received stud section revealed pitting and corrosion of the OD surface.

The attack was most severe in the 12 threads located nearest to the shank (Figure 1).

Apparently, these threads were not engaged with the reactor vessel flange bushing during service.

Magnetic particle testing revealed cracking in the 7 thread roots located from 14 to 20 threads below the shank transition.

The cracking was intermittent and extended around approximately 40% of the stud circumference.

In several of the cracked threads, pits appeared to be aligned in the thread root (Figure 2).

Axial sectioning was performed at two locations that had been identified as sections 1 and 2 by the NDE testers.

The ID bore probe had reported maximum crack depths of 0.40" at section 1 and 0.88 11 at section 2.

Note that section 2 was reported to contain the deepest crack in the stud.

Metallographic examination of sections removed from the identified areas indicated that the cracks initiated at pits in the thread roots (Figures 3 and 4).

Cracks were observed in adjacent threads and oriented primarily at an angle between 65 and 75 degrees to the stud axis.

The maximum crack depth, as measured to the nearest thread root, was 0.40 11 in section 1 and 0.70" in section 2.

Crack propagation occurred in a branching manner, particularly near the crack tips (Figures 5 and 6).

Etching revealed the microstructure consisted of tempered martensite with a prior austenite grain size of ASTM #8.

This is considered normal for the 4334M-2

M-7058-90 Page *3 specified Grade L43 material.

Because of the small grain size and the fact that most of the cracks were oxidized, it was very difficult to determine if the crack propagation mode was intergranular or transgranular (Figure 7).

Metallographic examinations of several of the secondary cracks suggested a mixed propagation mode.

Scanning electron microscope examinations were performed on the section 2 sample after opening up the largest crack.

Energy dispersive X-ray analyses of the surface did not identify a corroding species.

In addition, because the surface was covered by a thick, black oxide, little information could be gained from the uncleaned surface.

After cleaning the surface with an inhibited acid solution, some evidence of intergranular attack could be observed (Figure 8).

Chemical and mechanical tests were performed on material removed from the shank region of the stud.

Chemical analyses, tensile, hardness and Charpy V-notch impact testing results are shown in Tables 1, 2 and 3 along with the data reported on the original CMTR.

The chemical analyses correlate well with the original CMTR.

However, the tensile test results are 10-20 ksi higher than reported on the CMTR.

In addition, the Charpy impact data indicates that a significant reduction in toughness has occurred.

For information purposes, the charpy results at the bolt up (S0°F) and pressurization test (150°F) temperatures are also listed.

Additional J-R curve toughness testing will be performed at the bolt-up and pressurization test temperatures in the future and the

~

results will be issued in a separate report.

4334M-3

• C~NCLUSION M-7058-90 Page 4 The cracking was the result of stress-corrosion cracking (SCC) which initiated at the base of pits located in the thread roots.

The fact that the cracks propagated in a branching manner and were located in consecutive threads suggests an sec mechanism.

Based on the metallographic and scanning electron microscopic examinations, the p~opagation was mixed mode.

The mechanical testing results indicated the stud material has a higher strength and lower toughness than reported on the original CMTR.

The cause of this discrepancy is not known.

It is postulated that some long-term aging may have occurred at the operating temperature.

For stress-corrosion cracking to occur, a tensile stress must be applied to a susceptible material in a corrosive environment.

Based on the mechanical test results, the presence of stress-corrosion cracks in stud 61-198-47 is not entirely unexpected.

USAEC Regulatory Guide 1.65 (dated October 1973) states that closure stud materials such as Grade L43 are increasingly susceptible to sec above a tensile strength of 170 ksi.

Note that a tensile strength of up to 180.9 ksi was measured in the examined stud.

The environment that contributed to the sec is thought to be exposure to oxygenated water during.unit outages.

Maintenance personnel reported that the bushings in which the studs are engaged are slightly lower than the vessel flange sealing surface.

This creates a small reservoir where water can accumulate.

When the head is re-tensioned, the studs are in tension for up to two weeks prior 4334M-4

~-*

M-7058-90 Page 5 to the vessel heat-up, when the water would be driven away.

This could allow sec to occur.

The testers that performed the original straight beam ultrasonic examination from the top of the stud reported that the cracks in stud 61-198-47 were barely discernible.

Based on the sectioning performed, the detectability limit of the straight beam examination is a 0.70 11 deep crack.

The ultrasonic examination from the bore was able to size the cracks to within 0.18 11

• It should be noted that prior to the bore examination, the bore surface must be cleaned.

RECOMMENDATION Failure analysis indicated that the cracking in closure stud 61-198-47 was the result of stress-corrosion cracking.

Because other BWR closure studs are fabricated from similar materials and are exposed to a similar environment, it is possible that cracking may develop in other studs.

All 92 Dresden Unit 2 closure studs were inspected using the straight beam ultrasonic technique during the January 1989 inspection.

Additional inspections of other BWR studs should be performed in the future.

Because of the long lead time for delivery of new studs, it is recommended that several replacement studs be kept in stock when performing future inspections of the~uds.

Approved by:,J}~!Jh.W I

Reported Copies to: Station Manager, Dresden 4334M-5 Asst. Supt. Maintenance, Dresden Tech Staff Supervisor, Dresden BWR/PWR Projects Manager, DG BWR Syst. Design Supt., DG L. A. Sues, NSD, DG R. J. Tamminga. NSD, DG M. s. Horbaczewski, ISI, Dresden

Table 1 - Chemical Analysis Results- (Wt.

ASTM A320 stud 61-198-47

%)

Stud M-7058-90 Page 6 CMTR 61-198-47 For Heat Element Gr. L43 Near OD Near Bore 67-80278 Carbon 0.38-0.43 0.43 0.43 0.43 Manganese 0.60-0.85 0.72 0.67 0.72 Phosphorus 0.035 max.

0.010 0.010 0.010 Sulfur 0.040 max.

0.017 0.012 0.014 Silicon 0.15-0.35 0.30 0.29 0.29 Nickel 1.65-2.00

1. 68 1.67

1. 75 Chromium 0.70-0.90 0.80 0.74 0.80 Molybdenum 0.20-0.30 0.26 0.23 0.26 le 2 - Hardness and Tensile Test Results Tensile Yield Strength Strength

% Reduction Rockwell c Cksi)

Cksi)

% Elongation in Area Hardness A320, Gr.L43 125 min.

105 min.

16 min.

50 min.

No Requirement Stud 61-198-47 Near OD 180.2 167.8 17.4 53.5 38/39 1/2 Radius 173.0 155.2 18.0 56.5 34/35 Near Bore 164.0 145.7 17.4 54 32/33 Heat 67-80278 CMTR at 1/2 Radius Test 1 156.5 140.0 19.0 59.1 32/36 Test 2 160.0 145.0 18.5 56.9 36/38 Test 3 154.0 137.5 18.5 57.3 31/33 4334M-6

Table 3 -

Charpy V-Notch Impact Test Results (ft.-lbs.)

Room 100F Temperature 80°F Heat 67-80278 47,52,36 CMTR at 1/2 Radius Stud 61-198-47 Near OD 22,18 31,32 39,31 1/2 Radius 21,20 22,25 28,27 Near Bore 20,20 25,23 22,26 4334M-7 M-7058-90 Page 7 1so°F

.47,47-47,46 44,46

Figure 1 -

Figure 2 -

4334M-8 M-7058-90 Page 8 An as-received view of the closure stud near the lower shank-to-thread transition.

Note the heavy corrosion and pitting of the 12 threads nearest to the shank.

Cracks were present in the thread roots located from 14 to 21 threads beneath the transition.

An enlarged view of the pitting in the region of cracking.

The arrows highlight a region of aligned pits.

M-7058-90 Page 9 igure 3 - crack initiation at the thread roots in section 2.

The cracks initiated at pits and propagated at an angle between 65 to 75 degrees to the stud axis.

Higher magnification views revealed the cracks were branching.

(15X Magnification, Unetched)

I Figure 4 - Crack initiation at a pit in section 2.

The gray areas are oxide fill ed pits.

(50X Magnification, Unetched) 4334M-9

M-7058-90 Page 10 igure 5 -

Branching near the tip of the largest section 2 crack.

(15X Magnification, Unetched)

• *~

/ >

• ,.<~ *

~

I

. r.

..;:r

-~

• )f

• t....

igure 6 -

Branching near the tip of one of the section 1 cracks.

(lOOX Magnification, Unetched) 4334M-10

M-7058-90 Page 11 Figure 7 -

igure 8 -

A scanning electron microscope view of a section 2 crack surface showing evidence of intergranular cracking.

(lOOOX Magnification) 4334M-ll

ATTACHMENT B STRUCTURAL EVALUATION OF COMMONWEAL TH EDISON BWR REACTOR PRESSURE VESSEL HEAD STUD CRACKING ZNLD/899/39

ORF #137-0010 SASR #91-03 STRUCTURAL EVALUATION OF COMMONWEALTH EDISON BWR REACTOR PRESSURE VESSEL HEAD STUD CRACKING Prepared for Co11111onwealth Edison Co.

Prepared by GE Nuclear Energy 175 Curtner Avenue San Jose, CA 95125

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STRUCTURAL EVALUATION OF COMMONWEALTH EDISON BWR REACTOR PRESSURE VESSEL HEAD STUD CRACKING H.S. Mehta, Principal Engineer Materials Monitoring & Structural Analysis Services Prepared by: ~

~.L. Herrera, Senior Enginee~

Materials Monitoring & Structural Analysis Services Verified by:

S. Ranganath' anager Materials Monitoring & Structural Analysis Services Approved by: _;~

0_--....,....~+-i"-e_* _ 4_'= ____ _

S. Ranganath, Manager Materials Monitoring & Structural Analysis Services

1.0 2.0 3.0 4.0 5.0 6.0 TABLE OF CONTENTS BACKGROUND CRACK GROWTH ASSESSMENT FRACTURE TOUGHNESS ASSESSMENT

3. 1 CVN Energy 3.2 Fracture Toughness Calculation 3.3 References FRACTURE MECHANICS ASSESSMENT 4.1 Applied Stress Intensity Factor Ca~culation 4.2 Fracture Margin Assessment

4. 3 References ASHE CODE MARGIN ASSESSMENT 1-1 2-1 3-1 3-1 3-1 3-2 4-1 4-1 4-3 4-4 5-1 5.1 Minimum Stud Area Required for Dresden 2 & 3 5-2 5.2 Minimum Stud Area Required for Quad Cities and 5-2 Las a 11 e Pl ants 5.3 Acceptable Number of Partially Cracked Studs 5-3 5.4 References 5-4

SUMMARY

AND CONCLUSIONS 6-1 APPENDIX Theoretical Crack Propagation Rate/Stress Intensity Factor Relationships for Low Alloy Steels

1.0 BACKGROUND

During routine in-service inspection (ISi) in January 1989, at the Dresden 2 plant, suspected crack indications were observed by UT in two RPV head studs below the first engaged thread in the RPV flange.

The indications were detected with a Section XI Code UT technique where a straight beam (OO L wave) introduced from the upper end traverses the length of the stud.

The calibration reflector in this case was a 3/8 inch diameter flat bottomed hole.

The reflections received from the indications had amplitudes of 10% to 25% of the calibration reflector, and therefore were not required by Code to be recorded.

However, these indications were evaluated as crack indications.

For this reason, studs with indications were removed for additional testing.

Two available spare studs were installed in place of the two with indications. All other studs ( 90 out of a total of 92) were UT inspected without observation of other crack indications.

Magnetic particle examination was performed on both studs to verify the presence of cracks.

One stud showed cracks in seven threads extending to 40%

of the circumference.

The other stud contained cracks in two threads extending 50% of the circumference.

Depth measurements were made with a 700 shear wave probe from the center bore hole.

Maximum depths were estimated to be 0.88 inches in one stud and 2.09 inches in the other.

Subsequently, a metallurgical examination was performed on the stud with the lesser crack depth and it was determined that 'the actual depth was 0.7 inches.

The metallurgical evaluation reported the cause of the cracking to be stress corrosion cracking (SCC) initiating at pits.

Fig 1-1 and 1-2 show the approximate shapes of the cracks based on UT, and the remaining cross sections.

1-1

he studs in which crack indications were observed were fabricated from material specified as SA-193 Class 3, with Code Case 1335-1.

The outside diameter of the stud is approximately 6 in. and the inside diameter is one inch.

During refueling, the stud, including the lower portion which is threaded, is exposed to the stagnant air saturated water environment following vessel flooding.

The stud threads near the vessel flange surface may remain wet until plant startup, when the flange reaches a temperature high enough to evaporate the water.

The metallurgical examination of the stud with the 0.7 inch deep crack showed mixed mode transgranular cracking and branching, which, in this alloy, is typical of stress corrosion cracking.

The probable cause of the sec is the exposure of the stud~ in the preloaded condition to oxygenated water during outages since extensive oxidation and pitting was observed on the studs.

The

.metallurgical evaluation indicates that cracking originated on the outer edge of the stud at pitted locations near the thread roots.

The cracked studs had een in service for about 18 years, including 11 refueling outages.

The specification for the studs sets a minimum tensile strength of 145 ksi.

At the time of design, no specifications on maximum stud tensile or yield strength were defined.

USNRC Regulatory Guide 1.65, issued in October 1973, requires a maximum tensile strength of 170 ksi on RPV studs, because stud material with greater than 170 ksi (with the corresponding hardness of Re 38) had shown susceptibility to sec. The metallurgical evaluation reports a tensile strength value of 180 ksi and a hardness of Re 38 for the material in the outside threaded area of one cracked stud. The metallurgical evaluation also reports a Charpy V-notch (CVN) impact toughness of 21 ft-lbs at +1o*F, while the CMTR for that heat had reported 36 to 52 ft-lbs at +1o*F.

The reasons for the drop in toughness are being investigated.

However, for the purpose of the structural evaluation described here, the current measured CVN values are used.

This report describes the structural evaluation of the remaining vessel studs at Dresden-2 assuming postulated cracks in the remaining ninety original studs.

Also included is the structural evaluation in terms of allowable number of cracked studs for the other CECO BWR units.

1-2

Since these studs were examined in 1989 and found to be free of. indications, it is conservative to postulate crack depths equal to the threshold of detection for the UT procedure used.

With this initial flaw size assumption and a crack growth rate based on the maximum depth, the crack depth at the end of the current outage is estimated. Fracture toughness values are determined based on the current measured CVN data.

Fracture margins are determined for the limiting condition for the stud-bolt up.

Finally, the minimum stud area required to meet the ASME code stress limits is determined.

This will provide information on the number of cracked studs that can be tolerated while still maintaining the ASME code requirements for joint integrity.

1-3

Figure 1-1 Cracked Cross Section of Stud # 61-198-47 1-4

\\

Figure 1.;2 Cracked Cross Section of Stud I 61-198-70 STu,D :IF"~/- 198-70 1-5

2.0 CRACK GROWTH ASSESSMENT The vessel studs are not exposed to the water environment when the pressure vessel is at temperature since the studs are dry.

The only time when the studs are in tension and exposed to the water environment is after refueling and retensioning when the vessel head is in place but the flange temperature is not high enough to allow evaporation of the water. This time is estimated to be on the average 3 weeks for each refueling outage.

During this period the lower portion of the stud is exposed to stagnant air saturated water, but at temperature below 212*F.

Stress corrosion cracking can occur under these conditions since the applied stud loads are high and the environment is stagnant.

The Dresden 2 head has experienced 11 bolt up cycles when the cracked studs were found in 1989.

It is assumed that crack initiation occurs in half this time i.e., 6 cycles.

The total period for crack growth is (6 bolt up cycles) (3 weeks/bolt up) (7x24 hours/week).. 3024 hours0.035 days <br />0.84 hours <br />0.005 weeks <br />0.00115 months <br />.

Assuming 2 in.

growth in this time the average crack growth rate is 6.6x10-4 in/hour. This is within the range of predictions (Appendix A) from analytical models for low alloy steel at high temperatures.

The estimated* crack growth rate of 6.6x10-4/hr. is higher than the bounding low sulfur line but is below the model predictions for the high sulfur line for stagnant conditions.

The effect of the lower temperature of the studs and the moderate sulfur content could explain why the crack growth rate is between the low sulfur and the high sulfur predictions.

Nevertheless, the comparison with the model predictions does support the premise that the exposure to the stagnant water environment during refueling *could have caused observed cracking. If the crack growth rate stayed at this level for the 1989 and 1991 bolt up cycles the increment in crack depth is 2x(3x7x24) x6.6 x 10-04 = 0.6 in. with the assumed initial flaw size of 0.7 in. corresponding to the threshold for UT inspection.

The estimated crack depth following the current bolt up is 1.3 in. This crack depth value will be used in the fracture margin assessment.

2-1

3.0 FRACTURE TOUGHNESS ASSESSMENT 3.1 CVN Energy The fracture toughness measured from one heat of the stud material ranged from 36 to 52 ft-lb at IOOF (Table 3-1). However, Table 3-1 shows that the measured CVN energies of specimens taken.from one of the cracked studs are considerably lower.

It is not clear whether this variation in CVN values is due to heat to heat variations or an aging phenomenon.

Nevertheless, for this evaluation the CVN values based on the specimens removed from the cracked stud will be used conservatively. Table 3-2 shows the hardness and tensile data.

The most limiting condition for the vessel stud from the fracture mechanics viewpoint is the bolt up condition.

The temperature for bolt up can be as low as sooF and the loading is essentially the maximum value corresponding to stud tensioning. Other conditions such as hydrotest and normal operation are not as severe as bolt up since the temperature (and therefore, the toughness) is significantly higher.

Furthermore, the applied load on the stud during tensioning is higher than that under the hydrotest and operating conditions.

Thus, the vessel bolt up represents a 'proof test' and after.a successful bolt up, the likelihood of a fracture problem in the operating condition is negligible.

For the purpose of this fracture assessment an average CVN value of 27.5 ft-lb corresponding to the mid-wall location and so*F temperature will be used.

3.2 Fracture Toughness Calculation Fracture mechanics assessments require the conversion of the CVN values to K1c values. Several empirical relationships (Table 3-3) are available relating K1c to CVN.

These fall into two broad areas - transition temperature range and upp~r shelf range. Table 3-4 shows the K1c-CVN relationships and the 3-1

calculated fracture toughness values.

The transition range predictions probably represent lower bound values and the upper shelf correlation may be an upper bound.

Amongst the transition range relationships the Barsom-Rolfe correlation predicts K1d values and may be overly conservative for fracture margin calculations since the bolt loading is virtually static.

The Carten-Sailors correlation is more widely used and probably represents the best transition range prediction.

Thus for the purpose of the fracture assessment it is concluded the fracture toughness of the stud material is in the range of 81.3 ksi/in (Carten-Sailors transition range) to 123.7 ksi Jin (Rolfe-Novak-Barsom upper shelf) at the ao*F bolt up.condition.

3.3 References 3-1 J.M. Barsom and S.T. Rolfe, "Correlations Between Kic and Charpy V-Notch Test Results in the Transition-Temperature Range," Impact Testing of Metals, ASTM STP 466, American Society for Testing and Materials, Philadelphia, 1970, pp. 281-302.

3-2 J.T. Carten and R.H. Sailors, "Relationship Between Material Fracture Toughness Using Fracture Mechanics and Transition Temperature Tests," T.

& A.H. Report No. 346, University of Illinois, Urbana, Aug. 1971.

3-3 R. Roberts and C. Newton, "Report on Small-Scale Test Correlations with Kie Data," WRC Bulletin 299, Nov. 1984.

3-4 S.T. Rolfe and S.R. Novak, "Slow-Bend Kic Testing of Medium-Strength High-Toughness Steels," Review of Developments in Plane Strain Fracture-Toughness Testing, ASTM STP 463, American SOciety for Testing and Materials, Philadelphia, 1970, pp. 124-159.

3-2

Heat 67-80278 CMTR at 1/2 Radius TABLE 3-1 Charpy V-Notch Results (ft-lbs) 47, 52, 36 Room Temperature 3-3

I SA193 Cl.3, per Code Case 1335-1 Stud 61-198-47 Near OD 1/2 Radius Near Bore Heat 67-80278 CMTR at 1/2 Radius Test 1 Test 2 Test 3 Tensile Strength Cksil 145 min 180.2 173.0 164.0 156.5 160.0 154.0 TABLE 3-2 Hardness and Tensile Test, Results Yield Strength

% Reduction Rockwell C (ksil 3 Elongation in Area Hardness 130 min 12 min 40 min No Require.

167.8 17.4 53.5 38/39 155.2 18.0 56.5 34/35 145.7 17.4 54 32/33 140.0 19.0 59.l 32/36 145.0

. 18.5 56.9 36/38 137.5 18.5 57.3 31/33 3-4

TABLE 3-3 Typical Fracture Toughness-CVN Relationships Transition Range (1)

Barsom and Rolfe (Reference 3-1)

(Kid)2/E.. 5 CVN (Kid in psiJin, E in psi and CVN in ft-lb)

(2)

Carten & Sailors (Reference 3-2)

K1c = 15.5 J(CVN)

(K1c in ksiJin, CVN in ft-lb)

(3)

Roberts and Newton (Reference 3-3)

K1c

• 9.35 (CVN)0.63 (K1c in ksi Jin, CVN in ft-lb)

Upoer Shelf (4)

Rolfe-Novak-Barsom (Reference 3-4)

(K1dSy)2

• 5.(CVN/Sy - 0.05)

(K1c in ksi Jin, Sy in ksi, CVN in ft-lb) 3-5

(1)

(2)

(3)

(4)

TABLE 3-4 Predicted Fracture Toughness as a Function of the Measured CVN Temp (°F}

CVN (half rad;us) average K1d (Barsom & Rolfe)

(E = 30 x 106 psi)

Transition Range K1c (Carten & Sailors)

Transition Range K1c (Roberts-Newton)

Transition Range K1c (Rolfe-Novak-Barsom)

Sy

• 155 ksi Upper Shelf 20.5 55.4 70.20 62.70 3-6.

23.5 27.5 46.5 59.47 64.20 83.50 75.10 81.30 105.70 68.30 75.40 105.0 123.7 173.3

4.0 FRACTURE MECHANICS ASSESSMENT The fracture mechanics assessment was conducted in two steps.

First, the applied value of stress intensity factor, K, was calculated as a function of *crack depth for the various operating conditions.

The fracture margin assessment was then performed considering the measured fracture toughness properties of the closure studs.

4.1 Applied Stress Intensity Factor Calculation The general expression for calculating the stress intensity factor, K for any cracked geometry is the following:

K

• a (F) J(wa)

(4-1) where, a factor.

depth.

is the nominal stress, a is the crack depth and F is the amplification The magnitude of F is a function of the component geometry and crack Based on the observed crack geometries in the two closure studs removed from service, the crack geometry shown in Figure 4-1 was used in calculating the K values.

A literature search was conducted to obtain the value of F for various crack depths. Reference 4-1 gives the experimentally determined values of K for a round bar with the crack geometry shown in Figure 4-1, subjected to three point bending.

Based on the experimental values reported in Reference 4-1, a mathematical expression for K was derived in Reference 4-2.

The corresponding mathematical expression for the amplification factor, F, is the following:

F=(/w/8)[3.75-10.93(a/D)+20.0S(a/D)2

-19.93(a/D)3+7.56(a/D)4]/(l-a/D)2 4-1 (4-2)

he closure stud is subjected to tension loading rather than bending.

However, due to the constraints at the each end, the closure studs are likely to experience a loading that may be somewhere in between the three point bending and pure tension.

Since there are no K solutions available in the literature for an edge cracked hollow stud subjected to tension load, the K values were obtained by multiplying the bending case K values by a ratio, R given as the following:

R = Kp,t/Kp,b

{4-3)

Where, Kp,t

• Stress intensity factor for a single-edge cracked plate subjected to tension with crack length to plate width ratio the same as the crack depth to stud diameter ratio.

Kp,b - Stress intensity factor for a single-edge cracked plate subjected to bending with crack length to plate width ratio the same as the crack depth to stud diameter ratio.

The Kp,t and Kp,b value were obtained using the mathematical expressions given in Reference 4-3.

The other key information n~eded to calculate the K value in equation (4-1) is the stress, a.

The.stresses in the studs do not change significantly after the initial bolt up.

Also, the temperature (correspondingly, the material fracture toughness) is lowest during the bolt up compared to any other plant operating conditions.

Therefore, it is reasonable to conclude that the bolt up condition

• represents the most limiting condition from the fracture mechanics assessment considerations.

Reference 4-4 indicates that a pressure of 6600 psi is applied in the stud tensioner during the bolt up.

This is equivalent to a stud tensioning load of 1,442,100 lbs. The corresponding stud nominal stress, a is 52780 psi. This value 4-2

of nominal stress was used in calculating the K values.

Thus, the K values were evaluated using the following expression:

K =a {F){R)J{wa)

{4-4)

Table 4-1 shows the calculated values of K as a function of crack depth.

The last column in Table 4-1 shows the K values for the tension case and the second from last column shows the values for bending case.

The actual K value at any crack depth is expected to be somewhere between the bending and tension case values.

4.2 Fracture Margin Assessment As indicated in Section 2, the estimated crack depth in any cracked stud following the current bolt up is 1.3 inch.

An examination of Table 4-1 indicates that the applied value of K at this crack depth ranges from 82.3 ksiJin to 117.2 ksiJin.

In Section 3, the fracture toughness of the stud material at the bolt up condition was estimated to range from 81.3 ksiJin to 123.7 ksiJin.

Thus, the available material toughness is in the range of applied K values.

Thus, even with the conservative assessment, crack extension is not predicted.

Furthermore, a stud failure or a pop up type crack extension during stud tensioning would definitely

• be noticed and recorded. Since a review of the stud tensioning records indicated no such event, it is concluded that the toughness of each of the studs is at least equal to or greater than the applied value of K.

Thus, the fact that a stud did not experience failure in a 'proof test' assures that the stud will maintain its structural integrity at least till the next bolt up.

4-3

4.3 References 4-l Bush, A.J., "Experimentally Determined Stress-Intensity* Factors for Single-Edge-Crack Round Bars Loaded in Bending," Experimental Mechanics, 16, 249-257 (1976).

4-2 Underwood, J.H. and Woodward, R.L., "Wide Range Stress Intensity Factor Expression for an Edge-Cracked Round Bar Bend Specimen," Experimental Mechanics, 29, 166-168 (1989).

4-3 Tada, H., Paris, P.C. and Irwin, G.R., The Stress Analysis of Crack Handbook, Del Res. Corp., Hellertown, PA. (1973).

4-4 Telephone conversation between Tom Spry of CECO and H. Mehta of GE, January 12, 1991.

4-4

TABLE 4-1 K Calculation for Cracked Studs k* a(F)(R) Jfl a t

I D

j a

a/D F

R K,t K,b 0.10 0.017 0.819 1.024 24.8 24.2 0.20 0.033 0.808 1.050 35.5 33.8 0.30 0.050 0.798 1.080 44.1 40.9 0.40 0.067 0.790 1.111 51.9 46.7 0.50 0.083 0.782 1.143 59.2 51.8 0.60 0.100 0.776 1.177

_66.2 56.3 0.70 0.117 0.772 1.212 73.2 60.4 0.80 0.133 0.768 1.247 80.2 64.3 0.90 0.150 0.766 1.283 87.2 68.0 1.00 0.167 0.765 1.318 94.4 71.6 1.* 10 0.183 0.766 1.354 101.8 75.2 1.20 0.200 0.768

• . 1. 389 109.3 78.7 1.30 0.217 0.772 1.424 117.2 82.3 1.40 0.233 0.777 1.458 125.4 86.0 1.50 0.250 o.783 1.492 133.9 89.7 1.60 0.267 0.792 1.526 142.9 93.7 1.70 0.283 0.802 1.558 152.4 97.8 1.80 0.300 0.813 1.591 162.4 102.1 1.90 0.317 0.827 1.623 173.1 106.6 2.00 0.333 0.842 1.656 184.5 111.4 4-5

(a)

~ross Section of Cracked Stud D

{b)

Assumed Geometry for K calculation Figure 4-1 4-6 I

i I

I I

5.0 ASHE CODE MARGIN ASSESSMENT The minimum stud area required to maintain ASHE Code margins is evaluated in this section.

The calculations are based on the criteria presented in the ASHE Code Section III Appendix E, "Minimum Cross-Sectional Area".

E-1200 gives the criteria to determine the minimum stud area.

The design load is given by; Wml = H + Hp

{5-1)

= 0.785G2P + (2b x 3.14GmP)

Wm2

• 3.14bGy

{5-2) where Wm1* minimum required stud load for the Design Pressure Wm2.. minimum required stud load for gasket seating H* total hydrostatic end force G* Diameter at location of gasket load reaction P* Design pressure b* effective gasket or joint contact surface seating width m* Gasket Factor (from Table E-1210-1)

The minimum required stud area is given by; where (5-3)

{5-4)

Sa* allowable stud stress at atmospheric temperature Sb* allowable stud stress at design temperature 5-1

Note that the actual stud area must be greater than either Ami or Am2*

Based on the above criteria for minimum stud area, the minimum number of studs is given by the greater of; or, (5-5) where A is the cross-sectional area of one stud.

5.1 Minimum Stud Area Required For Dresden 2 & 3 Based on the criteria given in Section 5.0, the minimum stud area and number of studs can be calculated.

For self energizing gasket types, Table E-1210-1 of the ASHE Code states that the gasket factor, m, is 0.

However, for this calculation, it was conservatively assumed that the gasket factor is 6 which is a high number in Table E-1210-1.

Substituting the appropriate numbers into equations 5-1 through 5-5 the required stud area is 2030.3 in2.

The minimum number of studs is 79.

The actual number of studs is 92.

Therefore, there is significant margin since the actual number of studs is significantly greater than the minimum number of studs.

5.2 Minimum Stud Area For Quad Cities and La Salle 1 and 2 Plants The minimum stud area and the number of studs was also calculated for the Quad Cities and La Salle 1 and 2 plants.

The calculation of required stud area and number of studs is given in Reference 5-1 and 5-2 for La Salle 1 and La Salle 2 respectively.

The results for Quad Cities 1 & 2 are the same as for Dresden 2 based on identical geometries.

Results of these calculations are shown in Table 5-1. Also shown in Table 5-1 are the actual number of studs.

Since the cracked studs considered here are assumed to carry no loads, conservative design practice requires that the distribution be uniform (i.e., no clustered cracked studs)

• 5-2

.5.3 Acceptable Number of Partially Cracked Studs Another measure by which the structural margin of the reactor head closure can be assessed is the maximum number of partially cracked studs (with the maximum estimated crack depth) that can be present while still maintaining the code required minimum total stud area.

It was assumed in this calculation that the remaining area at a cracked stud contributes to the total stud area.

Table 5-2 shows the results of this calculation. The calculation for Oresden-2 is somewhat conservative since a higher stud load used in the vessel design report was considered.

The results in Table 5-2 clearly demonstrate that a significant number of cracked studs can be tolerated while still monitoring the required code structural margin.

Experience with the evaluation of preloaded studs shows that the compliance of a stud does not increase significantly for moderate crack depths.

Thus preload loss is not significant for the range of crack depths evaluate-0 here.

Therefore, there is no specific requirement on the distribution of the partially cracked studs (i.e., they can be adjacent to each other).

Furthermore, for the self sealing 0-rings the load needed for the sealing action 1s small.

5-3

-~-------

5.4 References 5-1 "Analytical Report For LaSalle County Station Unit No.I", Report No.

CENC-1250, Combustion Engineering, Chattanooga, Tennessee.

5-2 La Salle II Vessel Stress Report, CBIN Contract No. 72-2046, Chicago Bridge and Iron Nuclear Co~pany.

5-4

Table 5-1 Required Stud Area and the Number of Studsl La Salle 1 La Salle 2 Oyad Cities 1&2 Dresden 2&3 Required Stud Area(in2) 1810.0 1787.7 2030.3 2030.3 Required # of Studs2 66 73 79 Actual # of Studs 68 76 92

1.

Assumes uniform distribution i.e., no clustered cracked studs.

2.

Assumes that the remaining studs are fully cracked and have no load carrying capability.

5-5 79 92

Table 5-2 Allowable Maximum Number of Partially Cracked Studs to Code Required Margin Plant

1. of Studs Dresden 2 & 3 43 Quad Cities 1 & 2 43 La Salle 1 15 La Salle 2 17 5-6

6.0

SUMMARY

AND CONCLUSIONS Th;s report describes the structural evaluation of the remain;ng vessel studs at Dresden-2 assuming postulated cracks in the remaining ninety original studs.

Since these studs were examined 1989 and found to be free of indications, it is conservative to postulate crack depths equal to the threshold of detection (0.7 in.) for the UT procedure used.

With this initial flaw size assumption and a crack growth rate based on the.maximum depth, the crack depth at the end of the current outage is estimated. Fracture toughness values are determined based on the current measured CVN data.

Fracture marg;ns are determined for the limiting condition for the stud-bolt up.

Finally, the minimum stud area required to meet the ASME code stress limits is determined.

This provides information on the number of cracked studs that can be tolerated while still maintaining the ASME code requirements *for joint integrity.

The results of the analysis confirm that crack extension is unlikely to occur for the limiting bolt up condition.

This w~s confirmed by the stud tension experience where no unusual load drop or compliance changes indicative of crack extension was observed.

Since the studs have undergone the 'proof test' condition during bolt up, no fracture concerns arise for the pressure test and normal operation conditions.

The analysis of the minimum required stud area for Dresden 2 pl ant shows that up to 13 could be fully cracked without violating the code stress limits.

If partial cracking (1.e., Postulated Crack depth of 1.3 in.) is assumed, cracking of up to 43 bolts can be tolerated. Similar analyses were also performed for other CECO BWRs.

The results of these analyses indicate that there is significant structural margin relative to the code requirements even with a large number of cracked studs.

The margin to failure is substantially higher and confirms the overall structural margin in the flanged joint.

6-1

APPENDIX Theoretical Crack Propagation Rate/Stress Intensity Factor Relationships for Low Alloy Steels

Appendix Stress corrosion cracking in ferritic steels is a function of several parameters: applied stress intensity factor, sulfur content, temperature, water conductivity, EC P, fl ow rate.

Severa 1 mode 1 s have been developed to mechanistically address these variables [A-1].

These models are in large part applicable to 2aa*c water environment.

The conditions for sec in the stud occur at lower temperature <lOo*c and for a short period of time.

The applied stresses are high.

And the environment is stagnant.* The justification for the applicability of the high temperature models to the stud environment is therefore not clear.

Nevertheless it is useful to compare the crack growth rate deduced from the experience with the worst cracked stud with the limiting crack growth rates based on analytical models.

Figure A-1 from Reference A-1 shows the theoretical crack propagation rate/stress intensity relationships for low allow steel in low flow rate water at sso*F.

The analytical equations corresponding to the two relationships shown in Figure A-1 are the following:

For high sulfur, VT* 1.23x10-5 Kl.4 in/hr.

for low sulfur, VT

• 5.32x10-12 K4 in/hr.

where K is in ksi Jin.

An average value of 40 ksi Jin was assumed for the crack growth rate calculation using the above relationships.

Accordingly, the predicted crack growth rates are 2.2x10-3 in/hr and l.4xlo-5 in/hr, respectively.

A-1

I~*

"\\

he estimated crack growth rate of 6.6x10-4 in/hr lies in between the two pred;cted crack growth rates.

Note that the analytical model predictions are based on 550 9F water, whereas the water environment experienced by the closure studs is at less than 212*F, which is expected to result in lower crack growth rate.

Therefore the crack growth rate of 6.6x10-4 in/hr used in the structural evaluation is a reasonable value.

Reference A-1 Ford, F.P., "Status of Research on Environmentally Assisted Cracking in LWR Pressure Vessel Steel," Trans. of ASHE, Journal of Pressure Vessel Technology, 1988.

A-2

20 JO 10..

20 30 40 MPa./n 50 60 70 80 90 100 110 M~

CRAOC ~"'°'TOI H'i'[/S'TRES1-'N'TtNSl'n'

~~S fall AS.Uli/A.SOI IC 2Cap?!01*&'1tlt AT 201i"C:

l 0-1

• 0. t ?ylcl:

• 1.JT ACM&NT,11.0! 'lOW IUi 'It J

~

"31 l*Z,, * ~- S'Tm.)

aMGt OT CCll90SIQ F C1Dt1'A&.S C0'4tlt!HC: 2.00INiea.:I 40 so 10 70 80 90 100 STRESS INTENSITY Ksi.r.;

'11t E

Theoretical cr:ack propaptlOD ra~/stnu IDtemltJ rel:atlomhlps ror low-111101 steel lD low now r:ate water ;at ~

ror wrioas combla:atlons or sulphur content (la steel) ~d corrosion potential.

Figure A-1 A-1