ML12171A520
| ML12171A520 | |
| Person / Time | |
|---|---|
| Site: | Indian Point  |
| Issue date: | 12/19/2009 |
| From: | Gorman J, Hunt S, Riccardella P, White G A American Society of Mechanical Engineers (ASME) |
| To: | Atomic Safety and Licensing Board Panel |
| SECY RAS | |
| Shared Package | |
| ML12171A508 | List: |
| References | |
| RAS 22624, 50-247-LR, 50-286-LR, ASLBP 07-858-03-LR-BD01 | |
| Download: ML12171A520 (25) | |
Text
CHAPTER 44
44.1INTRODUCTION
Primary water stress corrosion cracking (PWSCC) of alloy 600nickel-chromium-iron base metal and related alloys 82 and 182
weld metal has become an increasing concern for commercial pressurized water reactor (PWR) plants. Cracks and leaks have been discovered in alloys 600/82/182 materials at numerous PWR plant primary coolant system locations, including at several loca-tions in the reactor vessels. The reactor vessel locations include top head control rod drive mechanism (CRDM) nozzles, top head ther-
mocouple nozzles, bottom head instrument nozzles, and reactor vessel outlet and inlet nozzle butt welds. The consequences of this PWSCC have been signicant worldwide with 72 leaks through May 2004 (56 CRDM nozzles, 13 reactor vessel closure head thermocouple nozzles, 2 reactor pressure vessel bottom-mounted instrument nozzles, and 1 piping butt weld), many cracked noz-zles and welds, expensive inspections, more than 60 heads replaced, several plants with several-month outage extensions to repair leaks, and a plant shutdown for more than 2 years due to extensive corrosion of the vessel head resulting from leak-age
from a PWSCC crack in a CRDM nozzle. This chapter addresses alloys 600/82/182 material locations in reactor vessels, operating experience, causes of PWSCC, inspection methods and ndings, safety considerations, degradation predictions, repair methods, remedial measures, and strategic planning to address PWSCC at the lowest possible net present value cost. Several example cases of PWSCC, and resulting boric acid cor-rosion, are described in the following paragraphs of this chapter
and, in some cases, the remedial or repair measures are described.
It is important to note that the repairs and remedial measures described may not apply to all situations. Accordingly, it is important to review each new incident on a case-by-case basis to ensure that the appropriate corrective measures are applied, including the need for inspections of other similar locations that may also be affected. 44.2ALLOY 600 APPLICATIONS Figure 44.1 shows locations where alloy 600 base metal andalloy 82 or 182 weld metal are used in PWR plant reactor ves-sels. It should be noted that not all PWR reactor vessels have alloys 600/82/182 materials at each of the locations shown in
Fig. 44.1. 44.2.1Alloy 600 Base Metal Alloy 600 is a nickel-based alloy (72% Ni minimum, 1417%Cr, 610% Fe) with high general corrosion resistance that has been widely used in light water reactor (LWR) power plants, i.e.,
in PWRs and boiling water reactors (BWRs). In PWR plants, alloy 600 has been used for steam generator tubes, CRDM nozzles, pressurizer heater sleeves, instrument nozzles, and simi-lar applications. The alloy was originally developed by the International Nickel Corporation (INCO) and is also known as Inconel 600, which is a trademark now held by the Special Metals Corporation [1]. The reasons that alloy 600 was selected for use in LWRs in the 1950s and 1960s include the following [27]:(a)It has good mechanical properties, similar to those of austenitic stainless steels. (b)It can be formed into tubes, pipes, bars, forgings, and cast-ings suitable for use in power plant equipment. (c)It is weldable to itself and can also be welded to carbon,low-alloy, and austenitic stainless steels. (d)It is a single-phase alloy that does not require postweld heat treatment. Also, when subjected to postweld heat treatments that are required for low-alloy steel parts to which it is weld-ed, the resulting sensitization (decreased chromium levels at
grain boundaries associated with deposition of chromium
carbides at the boundaries) does not result in the high sus-ceptibility to chloride attack exhibited by austenitic stain-less steels that are exposed to such heat treatments. (e)It has good general corrosion resistance in high temperaturewater environments, resulting in low levels of corrosion products entering the coolant and resulting in low rates of wall thinning. (f)It is highly resistant to chloride stress corrosion cracking (SCC), and has better resistance to caustic SCC than
austenitic stainless steels. (g)Its thermal expansion properties lie between those of car-bon/low-alloy steels and austenitic stainless steels, making
it a good transition metal between these materials.It was alloy 600s high resistance to SCC, especially chloride-induced SCC, that led to its selection for steam generator tubing in PWRs in the 1950s and 1960s. Several early PWRs had experi-
enced SCC of austenitic stainless steel steam generator tubing, variously attributed to chlorides and caustics, and this had led to a desire to use a tubing alloy with increased resistance to thesePWRREACTOR V ESSEL ALLOY 600 I SSUESJeff Gorman, Steve Hunt, Pete Riccardella, and Glenn A.White ASME_Ch44_p001-026.qxd 12/19/09 7:36 AM Page 1 2¥Chapter 44environments. Similarly, some early cases of SCC of stainless steel nozzle materials in BWRs during initial plant construction and startup, which was attributed to exposure to chlorides and uorides, led to the wide-scale adoption of alloy 600 and its relat-ed weld materials for use in BWR vessel nozzles and similar
applications [8].The rst report of SCC of alloy 600 in high-temperature pure orprimary water environments was that of Coriou and colleagues in
1959 [9] at a test temperature of 350C (662F). This type of crack-ing came to be known as pure water or primary water SCC (PWSCC) or, more recently, as low potential SCC (LPSCC). In response to Corious 1959 report of PWSCC, research was conduct-ed to assess alloy 600s susceptibility to SCC in high-temperature pure and primary water. Most of the results of this research in the 1960s indicated that alloy 600 was not susceptible unless specic
contaminants were present [1012]. The conditions leading to sus-ceptibility included the presence of crevices and the presence of
oxygen. Most of the test results of the 1960s did not indicatesusceptibility in noncontaminated PWR primary coolant environ-ments. However, by the early 1970s, it had been conrmed by sever-al organizations in addition to Coriou that PWSCC of highly stressed alloy 600 could occur in noncontaminated high-temperature pure and primary water environments after long periods of time
[1315]. Starting with Siemens in the late 1960s, some designers began to move away from use of alloy 600 to other alloys, such as alloy 800 for steam generator tubes and austenitic stainless steels for structural applications [15]. By the mid-1980s, alloy 690, an alternate nickel-based alloy with about twice as much chromium as alloy 600 (~30% vs. ~15%), had been developed and began to be used in lieu of alloy 600 for steam generator tubing [16]. By the early 1990s, alloy 690 began to be used for structural applications such as CRDM nozzles and steam generator divider plates.44.2.2Alloys 82 and 182 Weld Metal Weld alloys 82 and 182 have been commonly used to weldalloy 600 to itself and to other materials. These alloys are alsoFIG.44.1LOCATIONS WITH ALLOYS 600/82/182 MATERIALS IN TYPICAL PWR VESSEL ASME_Ch44_p001-026.qxd 12/19/09 7:36 AM Page 2 COMPANION GUIDE TO THE ASME BOILER & PRESSURE VESSEL CODE
¥3used for nickel-based alloy weld deposit (buttering) on weld preparations and for cladding on areas such as the insides of reac-tor vessel nozzles and steam generator tubesheets. Alloy 82 is bare
electrode material and is used for gas tungsten arc welding (GTAW), also known as tungsten inert gas (TIG) welding. Alloy
182 is a coated electrode material and is used in shielded metal arc welding (SMAW). The compositions of the two alloys are some-what different, leading to different susceptibilities to PWSCC.
Alloy 182 has lower chromium (1317%) than alloy 82 (1822%)
and has higher susceptibility to PWSCC, apparently as a result of the lower chromium content. Most welds, even if initiated or com-pleted with alloy 82 material, have some alloy 182 material.In recent years, alloys 52 and 152, which have about 30%chromium and are thus highly resistant to PWSCC, have been used in lieu of alloys 82 and 182, respectively, for repairs and for new parts such as replacement reactor vessel heads. 44.2.3RPV Top-Head Penetrations CRDMs in Westinghouse- and Babcock & WilcoxdesignedPWR plants and control element drive mechanisms (CEDMs) in Combustion Engineeringdesigned PWR plants are mounted on the top surface of the removable reactor vessel head. Figure 44.2 shows a typical CRDM nozzle in a Babcock & Wilcox-designed plant. Early vintage Westinghouse PWR plants have as few as 37 CRDM nozzles and later vintage Combustion Engineering plants have as many as 97 CEDM nozzles. These nozzles are machined from alloy 600 base metal with nished outside diameters ranging from 3.5 to 4.3 in. and with wall thicknesses ranging from about
0.4 to 0.8 in. In some cases, a stainless steel ange is welded to the alloy 600 nozzle with an alloy 82/182 butt weld. The nozzles are installed in the reactor vessel head with a small clearance or
interference t (0.004 in. maximum interference on the diameter) and are then welded to the vessel head by an alloy 82/182 J-groove weld. The surface of the J-groove weld preparation is coated with a thin butter layer of alloy 182 weld metal before stress relieving the vessel head so that the nozzles can be installed and the nal J-groove weld can be made after vessel stress relief.
This avoids possible distortion that could occur if the CRDM noz-zles were welded into the vessel head before vessel stress relief. Most vessels have a single 1.01.3 in. outside diameter alloy600 head vent nozzle welded to a point near the top of the head by a J-groove weld. Two of the early Babcock & Wilcoxdesignedvessels had eight 1.0-in. outside diameter alloy 600 thermocouplenozzles welded to the periphery of the head by J-groove welds. Most of the Combustion Engineering vessels have alloy 600 incore instrument (ICI) nozzles welded to the periphery of the top head by J-groove welds. These ICI nozzles are similar to CEDM nozzles except that they range from 4.5 to 6.6 in. outside diame-ter. Several Westinghouse plants have 3.5 to 5.4 in. outside diame-ter alloy 600 auxiliary head adapters and de-gas line nozzles attached to the top head by J-groove welds. Several Westinghouse plants have 5.3 to 6.5 in. outside diameter internals support hous-ings and auxiliary head adapters attached to the vessel top head surface by alloy 82/182 butt welds. In summary, PWR reactor vessels have 38 to 102 alloy 600 noz-zles welded to the top head, with most of these attached to the heads after stress relief of the head by alloy 82/182 J-groove welds. 44.2.4BMI Penetrations All of the Westinghouse and Babcock & Wilcoxdesigned reac-tor vessels in the United States and three of the Combustion Engineeringdesigned reactor vessels in the United States have alloy 600 instrument nozzles mounted to the vessel bottom heads.
These are often referred to as bottom-mounted instrument (BMI) nozzles. These nozzles range from 1.5 to 3.5 in. outside diameter.
As shown in Fig. 44.3, a typical BMI nozzle is welded to the bot-tom head by a J-groove weld. In the case of the Westinghouse and Combustion Engineering plants, the J-groove welds were made after stress relieving the vessel. In the case of the Babcock &
Wilcoxdesigned plants, the J-groove welds were made prior to vessel stress relief. Early test experience at a Babcock & Wilcox-designed plant showed a ow vibration concern with the portions of the BMI nozzles inside the bottom head plenum. Accordingly, all of the Babcock & Wilcock plant BMI nozzles were modied
after initial installation to increase the diameter of the portion of the nozzle extending into the lower plenum. The new extension was alloy 600 and the modication weld was made using alloy
82/182 weld metal, with no subsequent stress relief heat treatment. 44.2.5Butt Welds Many Westinghouse reactor vessels have alloy 82/182 buttwelds between the low-alloy steel reactor vessel inlet and outlet nozzles and the stainless steel reactor coolant pipe, as shown in Fig. 44.4. In most cases, these welds include alloy 182 cladding on the inside of the nozzle and an alloy 182 butter layer applied to the end of the low-alloy steel nozzle prior to vessel stress relief.FIG.44.2TYPICAL CONTROL ROD DRIVE MECHANISM (CRDM) NOZZLEFIG.44.3TYPICAL BOTTOM-MOUNTED INSTRUMENT (BMI) NOZZLE ASME_Ch44_p001-026.qxd 12/19/09 7:36 AM Page 3 4¥Chapter 44Babcock & Wilcoxdesigned plants, and all but oneCombustionEngineering-designed plant, do not have alloy 82/182 butt welds at reactor vessel inlet and outlet nozzles since the reac-tor coolant piping is low-alloy steel as opposed to stainless steel. Reactor vessel core ood line nozzles in Babcock & Wilcoxdesigned plants have alloy 182 cladding and alloy 82/182 butt welds between the low-alloy steel nozzle and stainless steel core
ood pipe.44.2.6Core Support Attachments Most PWR vessels have alloy 600 lugs attached to the insidesurface of the vessel, as shown in Fig. 44.5, to guide the reactor internals laterally or to support the reactor internals in the event of structural failure of the internals. These lugs are attached to cladding on the inside of the vessel by full penetration alloy82/182 welds. In most cases, the vessel cladding in the area of thelugs is also alloy 182 weld metal.44.2.7Miscellaneous Alloy 600 Parts Most reactor vessel lower closure anges have alloy 600 leak-age monitor tubes welded to the ange surface by alloys 82/182
weld metal. These are not discussed further since the leakage monitor tubes are not normally lled with water and, therefore, are not normally subjected to conditions that contribute to
PWSCC.44.3PWSCC 44.3.1Description of PWSCC PWSCC is the initiation and propagation of intergranularcracks through the material in a seemingly brittle manner, with little or no plastic deformation of the bulk material and without the need for cyclic loading. It generally occurs at stress levels close to the yield strength of the bulk material, but does not involve signicant material yielding. PWSCC occurs when three controlling factors, material sus-ceptibility, tensile stress, and the environment, are sufciently severe. Increasing the severity of any one or two of the three factors can result in PWSCC occurring, even if the severity of the remaining factor or factors is not especially high. The three factors are discussed separately in the following sections. While mechanistic theories for PWSCC have been proposed, a rm understanding of the underlying mechanism of PWSCC has not been developed. Accordingly, the inuence of material susceptibility, stresses, and environment must be treated on an
empirical basis, without much support from theoretical models.44.3.2Causes of PWSCC: Material Susceptibility Based on laboratory test data and plant experience, the follow-ing main factors inuence the susceptibility of alloy 600 base metal and its weld alloys to PWSCC: (a)Microstructure.Resistance to PWSCC tends to increase as the fraction of the grain boundaries that are decorated by chromium carbides increases. Various models have been proposed to explain this effect such as one where the car-
bides act as dislocation sources and enhance plastic defor-
mation at crack tips, thereby blunting the cracks and imped-ing their growth [17]. The absence of carbides in the matrix
of grains also correlates with higher resistance to PWSCC, as does larger grain size [18]. (b)Yield Strength.Susceptibility to PWSCC appears to increaseas the yield strength increases. However, this is considered to
be a result of higher yield strength material supporting high-er residual stress levels and is, therefore, more of a stress than a material effect. As discussed in para. 44.3.3, tests indi-cate that the time to PWSCC initiation varies inversely with the fourth to seventh power of the total (applied plus resid-
ual) tensile stress [1921].(c)Chromium Concentration.Tests of wrought materials andweld materials in the nickelchromiumiron alloy group of
materials consistently indicate that susceptibility to PWSCC
decreases as the chromium content increases [22,23].
Materials with 30% chromium or more are highly resistant to PWSCC. The improved resistance of alloy 82 vs. alloy 182 weld metal is attributed to the higher chromiumFIG.44.4TYPICAL REACTOR VESSEL INLET/OUTLET NOZZLEFIG.44.5TYPICAL CORE SUPPORT LUG ASME_Ch44_p001-026.qxd 12/19/09 7:36 AM Page 4 COMPANION GUIDE TO THE ASME BOILER & PRESSURE VESSEL CODE
¥5concentration of alloy 82 (1822%) vs. that of alloy 182(1317%). Alloy 690 base metal and alloys, 52 and 152 weld metal, with about 30% chromium, have been found to
be highly resistant to PWSCC in numerous tests. (d)Concentrations of Other Species and Weld Flaws. No cleartrends in PWSCC susceptibility have been observed as a function of the concentration of other species in the alloy such as carbon, boron, sulfur, phosphorous, or niobium.
However, to the extent that these species, in combination
with the thermomechanical processing to which the part is subjected, affect the carbide microstructure, they can have
an indirect inuence on susceptibility to PWSCC. Also, hot
cracks caused by some of these species (e.g., sulfur and
phosphorous) can act as PWSCC initiators and, thus, increase PWSCC susceptibility.44.3.3Causes of PWSCC: Tensile Stresses Industry design requirements, such as ASME BPVC SectionIII, specify the allowable stresses for reactor vessel components
and attachments. The requirements typically apply to operating condition loadings such as internal pressure, differential thermal expansion, dead weight, and seismic conditions. However, the
industry design standards do not typically address residual stress-es that can be induced in the parts during fabrication. These resid-
ual stresses are often much higher than the operating condition stresses and are ignored by the standards since they are secondary (self-relieving) in nature. It is the combination of operating condi-
tion stresses and residual stresses that lead to PWSCC. For the case of penetrations attached to the vessel heads by par-tial penetration J-groove welds, high residual stresses are caused by two main factors. Firstly, the surfaces of nozzles are typically machined prior to installation in the vessel. This machining cold works a thin layer (up to about 0.005 in. thick) on the surface, thereby signicantly increasing the material yield and tensile strength near the surface. Secondly, weld shrinkage, which occurs when welding the nozzle into the high restraint vessel shell, pulls the nozzle wall outward, thereby creating yield strength level
residual hoop stresses in the nozzle base metal and higher strength cold-worked surface layers. These high residual hoop stresses contribute to the initiation of axial PWSCC cracks in the cold-worked surface layer and to the subsequent growth of the axial cracks in the lower strength nozzle base material. The lower frequency of cracking in weld metal relative to base metal may result from the fact that welds tend not to be cold worked and then subjected to high strains after the cold work.
Residual stresses in the nozzles and welds can lead to crack ini-tiation from the inside surface of the nozzle opposite from the weld, from the outside surface of the nozzle near the J-groove weld, or from the surface of the J-groove weld. Most PWSCC cracks have been axially oriented. This is consis-tent with results of nite element stress analyses, which predict that the hoop stresses exceed the axial stresses at most locations.However, axial stresses can also be high and circumferentialcracks have occurred in a few cases. For the case of butt welds, the weld shrinkage that occurs asprogressive passes are applied from the outside surface produces
tensile hoop stresses throughout the weld, axial tensile stresses on the outside weld surface (and often also the inside weld surface),
and a region of axial compressive stress near midwall thickness.
The hoop stresses can contribute to axial PWSCC cracks in the weld and the axial stresses can contribute to circumferential cracks. Finite element analyses show that the hoop stresses on the wetted inside surface of a butt weld are typically higher than the
axial stresses at high stress locations, such that cracks are predict-ed to be primarily axial in orientation. However, if welds are repaired on the inside surface, or subjected to deep repairs from the outside surface, the residual hoop and axial stresses on the wetted inside surface can both approach the yield strength of the
weld metal and can cause circumferential as well as axial cracks. 44.3.4Causes of PWSCC: Environment Several environmental parameters affect the rate of PWSCCinitiation and growth. Temperature has a very strong effect. The effects of water chemistry variations are not very strong, assum-ing that the range of chemistry variables is limited to those that
are practical for PWR primary coolant, i.e., with the coolant con-taining an alkali to raise pH above neutral and hydrogen to scav-
enge oxygen.(a)Temperature. PWSCC is strongly temperature dependent.The activation energy for crack initiation is about 44
kcal/mole for thick section nozzle materials [24] and 50 kcal/mole for thinner cold-worked steam generator tubing material [25]. The activation energy for crack growth is about 31 kcal/mole [26]. Using these values, the relative factors for crack initiation and growth at typical pressuriz-er and cold leg temperatures of 653F and 555F relative toan assumed hot leg temperature of 600F are given inTable 44.1. (b)Hydrogen Concentration. Tests using crack growth ratespecimens have shown that crack growth tends to be a max-
imum when the hydrogen concentration results in the elec-
trochemical potential being at or close to the potential where the Ni/NiO phase transition occurs [27]. Higher or lower values of hydrogen decrease crack growth rates. This effect can be substantial, with peak crack growth rates in some cases being up to four times faster when the hydrogen con-centration is at the value causing peak growth rate as com-pared to conditions with hydrogen values well away from the peak growth rate value, as shown in Fig. 44.6 [27]. Tests at various temperatures show that the hydrogen concentra-tion for the Ni/NiO transition varies systematically with
temperature, and that the hydrogen concentration causing the peak growth rate exhibits a similar trend, with the ASME_Ch44_p001-026.qxd 12/19/09 7:36 AM Page 5 6¥Chapter 44concentration causing the peak crack growth rate becominglower as temperature decreases (e.g., 10 cc/kg at 320C, 17 cc/kg at 3301/4C, 24 cc/kg at 338C, and 27.5 cc/kg at 360C).Crack initiation may depend on hydrogen concentration in a similar manner. However, enough testing to determine the effect of hydrogen on time to crack initiation has only been
performed at 330C, where it resulted in the most rapidcrack initiation in alloy 600 tubing at about 32 cc/kg vs.
about 17 cc/kg for peak crack growth rate. Reported data regarding effects of hydrogen concentration on PWSCC ini-tiation and growth are shown in Fig. 44.7 [28]. The reasons that the hydrogen concentration for peak aggressivity
appears to be about twice as high for crack initiation vs.
crack growth rate (32 cc/kg vs. 17 cc/kg) are not known; the difference may be real or may be an artifact of data scatter
or imprecision.
(c)Lithium Concentration and pH. Tests indicate that theeffects of changes in pH on crack growth rate, once the pH is well above neutral, are minimal and cannot be distin-guished from the effects of data scatter [28]. However, when
considering the full pH range from acid to neutral to caus-tic, several tests indicate that crack growth rates decrease as pH is lowered to the neutral range and below, but is essen-
tially constant for pH Tof about 6 to 8 [29,30]. While tests of crack growth rate indicate increases in pH and lithium concentration within the normal ranges used for PWRs have minimal effects on crack growth rate, some evaluations of
crack initiation data indicate that increases in pH and lithium
cause moderate increases in the rate of crack initiation, e.g., in the range of 1015% for increases in cycle pH T from 6.9 to 7.2 [29].However, recent tests sponsored by the Westinghouse Owners Group (WOG) indicate that the effect may be stronger, such as an increase by a factor of two for an increase in cycle pH T from 6.9to 7.2. Further tests under EPRI sponsorship are underway (as of
2004) to clarify this situation.44.4OPERATING EXPERIENCE44.4.1Precursor PWSCC at Other RCS Locations PWSCC of alloy 600 material has been an increasing concernin PWR plants since cracks were discovered in the U-bend region
of the original Obrigheim steam generators in 1971. The history
of PWSCC occurrences around the full reactor coolant system up though 1993, i.e., not limited to the reactor vessel, is documented
in an EPRI report [31]. Between 1971 and 1981, PWSCC cracks
were detected at additional locations in steam generator tubes (e.g., at dents and at roll transitions), and in an increasing number of tubes. This experience showed that alloy 600 in the metallurgi-cal condition used for steam generator tubes was quite susceptible to PWSCC, with susceptibility increasing as stress, cold work, and temperature increase. It was found that susceptibility was also strongly affected by the microstructure of the material, with sus-
ceptibility tending to decrease as the density of carbides on the
grain boundaries increases. The rst case of PWSCC of alloy 600 in a nonsteam generatortube application was reported in 1982. This incident involved PWSCC of an alloy 600 pressurizer heater sleeve [31]. Swelling of a failed electric heater element inside this sleeve was identied as a contributing cause. Subsequent to this occurrence, an increas-ing number of alloy 600 instrument nozzles and heater sleeves in pres-surizers have been detected with PWSCC. Also, increasing numbers of instrument nozzles in reactor coolant system hot legs and steam generator heads have also been detected with PWSCC.
Many of the susceptible nozzles and sleeves have (as of May 2005) been repaired or replaced on a corrective or preventive
basis [31].FIG.44.6ALLOY 600 CRACK GROWTH RATE AT 338C PLOTTED VS.HYDROGEN CONCENTRATION [27]FIG.44.7HYDROGEN CONCENTRATION VS.TEMPERA-TURE FOR N2/N2O PHASE TRANSITION,PEAK PWSCC SUSCEPTIBILITY,AND PEAK CRACK GROWTH RATE [28]
ASME_Ch44_p001-026.qxd 12/19/09 7:36 AM Page 6 COMPANION GUIDE TO THE ASME BOILER & PRESSURE VESSEL CODE
¥7PWSCC in alloys 182 and 82 weld metals was rst detected inOctober 2000 in a reactor vessel hot leg nozzle weld [32]. This was only a month before the rst detection of PWSCC in a reac-tor vessel head penetration weld, as discussed in para. 44.4.2. 44.4.2RPV Top-Head Penetrations The rst reported occurrence of PWSCC in a PWR reactorvessel application involved a leak from a CRDM nozzle at Bugey 3 in France that was detected during a 10-year inservice inspec-
tion program hydrostatic test conducted in 1991 [33]. This initial occurrence, and the occurrences detected during the next few years, involved PWSCC of alloy 600 base material at locations with high residual stresses resulting from fabrication. The high
residual stresses were mainly the result of weld-induced defor-mation being imposed on nozzles with cold-worked machined surfaces.
Subsequent to the initial detection of PWSCC in a CRDM nozzle in 1991, increasing numbers of plants detected similar types of PWSCC, typically resulting in small volumes of leak-age and boric acid deposits on the head surface as shown in
Fig. 44.8. In 2000, circumferential cracks were detected on the outside diameter of some CRDM nozzles. In 2002, significant wastage of the low-alloy steel Davis-Besse reactor vessel head occurred adjacent to an axial PWSCC crack in an alloy 600 CRDM nozzle. The wastage was attributed to corrosion by boric
acid in the leaking primary coolant that concentrated on the vessel head. Figure 44.9 shows a photograph of the corroded surface at Davis-Besse. The Davis-Besse plant was shut down for approximately 2 years for installation of a new head and
incorporation of changes to preclude similar corrosion in the future. The NRC issued several bulletins describing these events
and requiring utilities to document their inspection plans for this
type of cracking [3436]. The cracking discussed above was mainly related to PWSCC ofalloy 600 base materials. Starting in November 2000, some plants found PWSCC primarily in the J-groove weld metal, e.g., in CRDM nozzle-to-vessel alloy 182 J-groove welds [37]. Since that time, several other cases of PWSCC of CRDM nozzle-to-head welds have been detected. Also, detection of PWSCC in alloys 182 and 82 welds appears to be increasing in frequency at other nonreactor vessel locations around the reactor coolant system.
However, the frequency of PWSCC in welds remains lower than in alloy 600 base material. For example, after the detection of
PWSCC in the weld metal of a CRDM nozzle at a PWR in the United States in November 2000, and the detection of PWSCC in the alloy 182 weld metal at reactor vessel outlet nozzles in the
United States and Sweden in late 2000, EDF inspected 754 welds in 11 replaced reactor vessel heads without detecting any cracks
[24]. 44.4.3RPV Nozzle Butt Welds In October 2000, a visual inspection showed a leak from analloys 82/182 butt weld between a low-alloy steel reactor vessel hot-leg outlet nozzle and stainless steel hot-leg pipe at the V.C.
Summer plant. Destructive failure analysis showed that the leak was from a through-wall axial crack in the alloys 82/182 butt weld, as shown in Fig. 44.10. The axial crack arrested when it reached the low-alloy steel nozzle on one side and stainless steel
pipe on the other side, since PWSCC does not occur in these materials. The axial crack can propagate into the low-alloy steel and stainless steel by fatigue, but the fatigue crack growth rates will be low due to the small number of fatigue cycles. The destructive examination also showed a short-shallow circumferen-tial crack intersecting the through-wall axial crack that grew through alloy 182 cladding and terminated when it reached the low-alloy steel nozzle base metal. Examination of fabricationFIG.44.8TYPICAL SMALL VOLUME OF LEAKAGE FROM CRDM NOZZLE ASME_Ch44_p001-026.qxd 12/19/09 7:36 AM Page 7 8¥Chapter 44records showed that the leaking butt weld had been extensivelyrepaired during fabrication, including repairs made from the inside surface. Nondestructive examinations of other reactor ves-sel outlet and inlet nozzles at V.C. Summer showed some addi-tional shallow axial cracks.Shortly before the leak was discovered at V.C. Summer, part-depth axial cracks were discovered in alloys 82/182 reactor vessel outlet nozzle butt welds at Ringhals 3 and 4. Some of these cracks were removed and two were left in place to allow a determination of the crack growth rate. The crack growth rate is discussed in
para. 44.7.2. In addition to the PWSCC cracks in alloys 82 and 182 weldmetal in reactor vessel CRDM nozzles and inlet and outlet nozzle butt welds, a leak was found from a pressurizer nozzle butt weld at Tsuruga 2 in Japan and a part-depth crack was detected in a hot-leg pressurizer surge line nozzle butt weld at TMI-1. Both of these cases occurred in 2003. Cracks were also detected in alloys
82 and 182 cladding in steam generator heads that had been ham-mered and cold worked by a loose part [24]. In the 20052008 time period, the industry has begun imple-menting a massive inspection program for PWSCC in primary coolant loop Alloy 82/182 butt welds (In accordance with Industry Guideline MRP-139 [58] see Section 44.5.6 below for complete discussion). Considering the temperature sensitivi-ty of the PWSCC phenomenon discussed above, this program
started with the highest temperature welds in the system: those at pressurizer nozzles. To date, essentially all pressurizer nozzle dissimilar metal butt welds (typically five or six per plant) have
been inspected, mitigated, or both. Approximately 50 nozzles were inspected (many more were mitigated using weld overlays with no pre-inspections), resulting in PWSCC-like indications being detected in nine nozzles, as documented in Table 44.2 below.Through mid-2008, inspections of reactor vessel nozzle buttwelds have not yet been performed; hot leg nozzle inspections under MRP-139 are slated to begin in Fall 2008. Given the above pressurizer nozzle experience, it would not be surprising if at least some welds with PWSCC-like indications are discovered.FIG.44.9LARGE VOLUME OF WASTAGE ON DAVIS-BESSE REACTOR VESSEL HEADFIG.44.10THROUGH-WALL CRACK AND PART-DEPTH CIRCUMFERENTIAL CRACK IN V.C.SUMMER REACTOR VESSEL HOT-LEG OUTLET NOZZLE ASME_Ch44_p001-026.qxd 12/19/09 7:36 AM Page 8 COMPANION GUIDE TO THE ASME BOILER & PRESSURE VESSEL CODE
¥944.4.4RPV Bottom-Head Penetrations In 2003, bare metal visual inspections of the reactor vessel bot-tom head at South Texas 1 showed small leaks from two BMI noz-zles, as shown in Fig. 44.11. These leaks were traced to PWSCC cracks in the nozzles that initiated at small regions of lack-of-fusion in the J-groove welds between the nozzles and vessel
head [38]. The nozzles were repaired. Examinations of the other BMI nozzles at South Texas 1 showed no additional cracks.
Essentially all other U.S. plants have performed bare metal visual inspections of RPV bottom-head nozzles without any evidence of leaks. At least a dozen U.S. plants have completed volumetric examinations of the BMI nozzles, representing more than 20% of
the total population of RPV bottom-head nozzles in the U.S., with no reported cracking. Similarly, no indications of in-service degra-dation have been identied in volumetric inspections of RPV bot-
tom-head nozzles performed in other countries. PWSCC of BMI nozzles that operate at the plant cold-leg temperature is generally considered to be less likely than PWSCC at locations operating at hot-leg or pressurizer temperatures. The earlier-than-expectedPWSCC in BMI nozzles at South Texas 1 may be related to a com-bination of high material susceptibility and welding aws.44.5INSPECTION METHODS ANDREQUIREMENTS As a result of the increasing frequency of PWSCC cracks andleaks identied in important PWR reactor vessel alloys 600, 82, and 182 materials since 2000, signicant efforts are in progress by the nuclear industry and the NRC to improve inspection capabilities and develop appropriate long-term inspection requirements. The following summarizes the status of inspection methods and require-
ments as of May 2005. It is recommended that users check with the
NRC and industry programs to remain abreast of the latest changes
in inspection methods and requirements. 44.5.1Visual Inspections Bare metal visual inspections have proven to be an effectiveway of detecting very small leaks, as shown by Figs. 44.8 and 44.11, and, therefore, should play an important role in any inspec-tion program. A key prerequisite for these inspections is that the surface should be free of preexisting boric acid deposits from other sources, because the presence of preexisting boric acid deposits can mask the small volumes of deposits shown in Figs. 44.8 and 44.11. Visual inspections with insulation in place can provide a useful backup to bare metal visual inspections but will be inca-pable of detecting small volumes of leakage, as shown in Figs.
44.8and 44.11. In many cases, it has been necessary to modify insulation pack-ages on the vessel top and bottom heads to facilitate performing
bare metal visual inspections. As of May 2005, most of these modications have been completed for PWR plants in the United
States. ASME Code Case N-722, Additional Examinations for PWRPressure-Retaining Welds in Class 1 Components Fabricated with Alloys 600/82/182 Materials,Section XI, Division 1, was approved in 2005 to provide for increased visual inspections of
potentially susceptible welds for boric acid leakage. TABLE 44.2CRACKING INDICATIONS DETECTED IN REACTOR COOLANT LOOP ALLOY 82/182 BUTT WELDS,2005 THROUGH MID-2008Inspection Type ofIndication OD Indicationa / l / PlantDateNozzleIndicationDepth (a, in)Length (l, in)thicknesscircumferenceCalvert Cliffs 22005CL DrainCirc0.0560.62810%10%Calvert Cliffs 22005HL DrainAxial0.3920.00070%0%
DC Cook2005SafetyAxial1.2320.00088%0%
Calvert Cliffs 12006HL DrainCirc0.1000.45019%5%
Calvert Cliffs 12006ReliefAxial0.1000.0008%0%
Calvert Cliffs 12006SurgeCirc0.4002.40025%6%
Davis Besse2006CL DrainAxial0.0560.0007%0%
San Onofre 22006SafetyAxial0.4200.00030%0%
San Onofre 22006SafetyAxial0.4200.00030%0%
Wolf Creek2006ReliefCirc0.34011.50025.8%46%
Wolf Creek2006SafetyCirc0.2972.50022.5%10%
Wolf Creek2006SurgeCirc0.4658.75032.1%19%
Farley 22007SurgeCirc0.5003.00033%6%
Davis Besse2008AxialCrystal River 32008CircFIG.44.11LEAK FROM SOUTH TEXAS 1 BMI NOZZLE ASME_Ch44_p001-026.qxd 12/19/09 7:36 AM Page 9 10¥Chapter 4444.5.2Nondestructive Examinations Technology exists as of May 2005 to nondestructively examineall of the alloys 600, 82, and 182 locations in the reactor vessel. Partial penetration nozzles (CRDM, CEDM, ICI) are typicallyexamined using one of two methods. The nozzle base metal can be examined volumetrically from the inside surface by ultrasonics
to conrm that the nozzle base material is free of internal axial or circumferential cracks. Alternatively, the wetted surfaces of the alloy 600 base metal and alloys 82 and 182 weld metal can be examined by eddy current probes to ensure that there are no sur-face cracks. If there are no surface cracks on wetted alloy 600 sur-faces, then it can be inferred that there will also be no internal cracks. Nozzles in the reactor vessel top head can be examined
when the head is on the storage stand during refueling. Nozzles in the reactor vessel bottom head can be examined ultrasonically or by eddy current when the lower internals are removed from the vessel during a 10-year in-service inspection outage. In some cases, the inside surfaces of BMI instrument nozzles can be examined by tooling inserted through holes in the lower internals. Reactor vessel inlet and outlet nozzle butt welds are normallyinspected ultrasonically from the inside surface using automated
equipment. These inspections are typically performed during 10-year in-service inspection outages when the lower internals are removed from the reactor vessel. Eddy current methods are also being used in some cases for examining the inside surfaces of these welds for cracks, although eddy current inspection sensitivi-ty is a function of the condition of the weld surface. For example, discontinuities in the weld prole can cause the eddy current probes to lift off of the surface being examined and, thereby, adversely affect the inspection sensitivity. CRDM nozzle butt welds can be examined from the outsidesurface by standard ultrasonic methods. A key to obtaining good nondestructive examinations is to have the process and the operators qualied on mockups containing prototypical axial and circumferential aws. The EPRI NDE Center in Charlotte, NC, is coordinating qualication efforts for
inspection methods and inspectors in the United States. 44.5.3ASME BPVC Reactor Vessel InspectionRequirements ASME BPVC Section XI species inservice inspection require-ments for operating nuclear power plants in the United States.
Portions of these requirements that apply to PWSCC susceptible components in the RPV are summarized as follows:(a)Table IWB-2500-1, Examination Category B-P, requires aVT-2 visual examination of the reactor vessel pressure-retaining boundary during the system leak test after every
refueling outage. No leakage is permitted. (b)Table IWB-2500-1, Examination Category B-O, requires that 10% of the CRDM nozzle-to-ange welds be inspected by volumetric or surface methods each inspection interval. (c)Table IWB-2500-1, Examination Category B-N-1, requiresthat attachment welds to the inside surface of the reactor vessel be examined visually each inspection interval. Welds in the beltline region must be inspected by VT-1 methods while welds outside the beltline region must be inspected by VT-3 methods. (d)Table IWB-2500-1, Examination Category B-F, speciesexamination requirements for dissimilar metal welds in reactor vessels. Nozzle-tosafe end socket welds must be examined by surface methods every inspection interval.
Nozzle-to-safe end butt welds less than NPS 4 must be exam-ined by surface methods every inspection interval. Nozzle-to-safe end butt welds NPS 4 and larger must be examined by volumetric and surface examination methods every inspection interval. Some deferrals of these inspections are permitted. (e)As of May 2005, the ASME Code did not require nonde-structive examination of the partial penetration welds for the CRDM and BMI nozzles. However, Code Case N-729-1
[63] was published later in 2005 that contained alternative examination requirements for PWR closure heads with noz-zles having pressure-retaining partial-penetration welds.
This Code Case included visual, surface and volumetric examinations for PWR closure heads with Alloy 600 noz-zles and Alloy 82/182 partial-penetration welds at inspec-tion intervals that are based on the temperature dependence
of the PWSCC phenomenon described in para. 44.3.4.
(Since RPV closure heads operate at varying temperatures, there are signicant head-to-head temperature differences
between plants.) Code Case N-729-1 also contains inspec-
tion requirements for PWR closure head with nozzles and
partial-penetration welds of PWSCC resistant materials to address new and replacement heads.(f)As noted in para. 44.5.1, Code Case N-722 [64] for visualinspections of alloys 82/182 welds was approved in 2005. (g)As of May 2008, the ASME Code is working on a new Section XI Code Case that contains alternate inspection requirements Alloys 82/182 welds butt welds. ASME Code actions are also in progress addressing various repair and
mitigation options for dealing with PWSCC. These are discussed below in para. 44.9.44.5.4NRC Inspection Requirements for RPV Top-Head Nozzles Subsequent to the discovery of signicant corrosion to theDavis-Besse reactor vessel head, the NRC issued NRC Order
EA-03-009 [39]. This order species inspection requirements for RPV head nozzles based on the effective degradation years of operation. Effective degradation years (EDYs) are the effective full-power years (EFPYs) adjusted to a common 600F tempera-ture using an activation energy model. For plants with 600F headtemperatures, the EDYs are the same as the EFPYs. For plants
with head temperatures, greater than 600F, the EDYs are greaterthan the EFPYs. For plants with head temperatures less than
600F, the EDYs are less than the EFPYs. The NRC orderspecies two types of inspections: (a)bare metal visual inspections of the RPV head surface including 360around each RPV head penetration nozzle (b)nondestructive examinations of the RPV nozzles by one ofthe two following methods: (1)ultrasonic testing of each RPV head penetration nozzle(i.e., base metal material) from 2 in. above the J-groove
weld to the bottom of the nozzle plus an assessment to
determine if leakage has occurred through the interfer-
ence t zone (2)eddy current testing or dye penetrant testing of the wettedsurface of each J-groove weld and RPV head penetration nozzle base material to at least 2 in. above the J-groove weld The rst of the nondestructive examinations is to show that there are no axial or circumferential cracks in the nozzle base ASME_Ch44_p001-026.qxd 12/19/09 7:36 AM Page 10 COMPANION GUIDE TO THE ASME BOILER & PRESSURE VESSEL CODE
¥11metal or leak paths past the J-groove weld. The second of thenondestructive examinations is to show that there are no axial or
circumferential cracks in the nozzle base metal by conrming the absence of surface breaking indications on the nozzle and weld wetted surfaces. The order species inspection intervals for three categories ofplants: high susceptibility plants with greater than 12 EDY or where PWSCC cracks have already been detected, moderate sus-ceptibility plants less than or equal to 12 EDY and greater than or equal to 8 EDY, and low susceptibility plants with less than 8 EDY. As of June 2008, the U.S. NRC is expected shortly to transition the requirements for inspection of RPV top-head nozzles based on
NRC Order EA-03-009 [39] to a set based on ASME Code Case N-729-1 [63], with caveats. The inspection schedules in this code
case are generally based on the RIY (reinspection years) concept, which normalizes operating time between inspections for the effect of head operating temperature using the thermal activation energy appropriate to crack growth in thick-wall alloy 600 material
(31 kcal/mol (130 kJ/mol)). The basis for this approach to nor-malizing for the effect of head temperature is that the time for a aw just below detectable size to grow to through-wall (and leak-age) is dependent on the crack growth process. The requirements in ASME Code Case N-729-1 [63] were developed by ASME, with extensive technical input provided by a U.S. industry group (Materials Reliability Program) managed by EPRI [68].44.5.5NRC Inspection Requirements for RPV BMI Nozzles NRC Bulletin 2003-02, Leakage from Reactor Pressure VesselLower Head Penetrations and Reactor Coolant Pressure Boundary Integrity
[40], summarizes the leakage from BMI noz-zles at South Texas 1 and requires utilities to describe the results of BMI nozzle inspections that have been performed at their plants in the past and that will be performed during the next and following refueling outages. If it is not possible to perform bare metal visual examinations, utilities should describe actions that are being made to allow bare metal visual inspections during sub-
sequent outages. If no plans are being made for bare metal visual or nonvisual surface or volumetric examinations, then utilities must provide the bases for concluding that the inspections that have been performed will ensure that applicable regulatory
requirements are met and will continue to be met. On September 5, 2003, the NRC issued Temporary Instruction 2515/152 [41],
which provides guidance for NRC staff in reviewing utility sub-mittals relative to Bulletin 2003-02. While the Temporary
Instruction does not represent NRC requirements, it does indicate the type of information that the NRC is expecting to receive in response to the bulletin. 44.5.6Industry Inspection Requirements forDissimilar Metal Butt Welds The industry in the United States has developed a set of manda-tory inspection guidelines for PWSCC susceptible. Alloy 82/182 butt welds, which are documented in the report MRP-139 [58].
MRP-139 denes examination requirements in terms of categories
of weldments that are based on 1) the IGSCC resistance of the
materials in the original weldment, 2) whether or not mitigation
has been performed on the original weldment, 3) whether or not a pre-mitigation UT examination has been performed, 4) the exis-
tence (or not) of cracking in the original weldment, and 5) the likelihood of undetected cracking in the original weldment. The categories range from A through K, with the higher letter categories requiring augmented inspection intervals and/or samplesize. Category A is the lowest category, consisting of piping that has been replaced (or originally fabricated) with PWSCC resistant
material. These weldments are to be inspected at their normal ASME Code frequency, as dened in ASME Section XI, Table IWB-2500-1. Category D refers to unmitigated PWSCC suscepti-ble weld in high temperature locations (e.g. pressurizer or hot leg
nozzles). These require an early initial inspection (before end of 2008 for pressurizer nozzles and before 2010 for hot leg nozzles),
followed by more frequent inspections if they are not treated with some form of mitigation. Other categories (thru Category K) address susceptible welds that have been mitigated (B and C),
welds that have been inspected and found cracked, with or with-
out mitigation, and welds for which geometric or material condi-tions limit volumetric inspectability. For the latter group, by the time the examination is due, plant owners are required to have a
plan in place to address either the susceptibility of the weld or the
inspectability of the weld.
At the time of this writing, inspections are well under theMRP-139 guidelines are well underway in U.S. plants. Essentially all pressurizer nozzles have been inspected and or mitigated, and
plans are in place to perform the other initial inspections required
by MRP-169. Plans include mitigation of most susceptible weld-ments in high temperature locations, thus moving the weldments into Categories A, B or C. Work is also currently underway to develop an ASME Section XI Code Case (N-790, alternative examination requirements for PWSCC pressure-retaining butt welds in PWRs) which will eventually replace MRP-139 and place the augmented examination requirements for PWSCC sus-ceptible butt welds back under the ASME Section XI Code.44.6SAFETY CONSIDERATIONS 44.6.1Small Leaks Small leaks due to axial cracks such as shown in Figs. 44.8 and44.11 do not pose signicant safety risk. The leak rates are low enough that the leaking primary coolant water will quickly evapo-rate leaving behind a residue of dry boric acid. Most of the leaks detected to date have resulted in these relatively benign condi-tions. As shown in the gures, very small leaks are easily detected by visual inspections of the bare metal surfaces provided that the surfaces are free from boric acid deposits from other sources. One explanation for the extremely low leak rates is that short tight
PWSCC cracks can become plugged with crud in the primary coolant, thereby preventing leakage under normal operating con-
ditions. It is hypothesized that distortions, which occur during plant transients, allow small amounts of leakage through the crack before it becomes plugged again. Regardless, these small leaks do
not pose a signicant safety concern.44.6.2Rupture of Critical Size Flaws Initially, leaking RPV top-head nozzles were thought to beexclusively the result of axial cracks in the nozzles, and it was thus believed that they did not represent a signicant safety con-cern. However, as more examinations were performed, ndings
arose that called this hypothesis into question. (a)Relatively long circumferential cracks were observed in twonozzles in the Oconee Unit 2 RPV head, and several other plants also discovered shorter circumferentially oriented
cracks. ASME_Ch44_p001-026.qxd 12/19/09 7:36 AM Page 11 12¥Chapter 44(b)Circumferential NDE indications were discovered in theNorth Anna Unit 2 head in nozzles that showed no apparent
signs of boric acid deposits due to leakage.
Figure 44.12 presents a schematic of a top-head CRDM nozzleand J-groove weld and the nature of the cracking that has been observed. There is some uncertainty as to whether circumferential cracks arise as a result of axial cracks growing through the weld or nozzle and causing leakage into the annular region between the
nozzle and head, ultimately leading to reinitiation of circumferen-tial cracking on the outside surface of the tube, or if they are due to the axial cracks branching and reorienting themselves in a
circumferential direction, as depicted on the right-hand side of Fig. 44.12. A destructive examination program has been per-formed on several of the North Anna Unit 2 nozzles, indicating that the circumferential nozzle defects found there were in fact the result of grinding during fabrication rather than service-related cracking. Nevertheless, the occurrence of circumferential crack-ing adds a new safety perspective to the RPV top-head nozzle
cracking problem, because of the potential for such cracks to grow to a critical length and ultimately lead to ejection of a nozzle from the vessel, although a large circumferential aw covering more than 90% of the wall cross section is typically calculated for nozzle ejection to occur because of the relatively thick wall typical
of RPV top-head nozzles.
PWSCC in PWR RPV inlet/outlet nozzles could also potentiallydevelop circumferentially oriented aws, which could lead to pipe rupture. To date, observed cracking has been primarily axial with only very small circumferential components. With time, however, PWSCC in large piping butt welds might be expected to follow
trends similar to the IGSCC cracking issue in BWRs [42]. In the
BWR case, cracking and leakage were initially seen only as axial-ly oriented cracks in smaller diameter piping. With time, however, axial and circumferential cracking were observed in pipe sizes up to and including the largest diameter pipes in the system.
Considering the potential existence of weld repairs during initial
construction of the plants and the associated high residual stresses that they produce in both axial and circumferential directions, signicant circumferential cracking may eventually be observed in large-diameter PWR pipe-to-nozzle butt welds.
Because of the concern for PWSCC in PWR piping dissimilarmetal butt welds, methods for predicting the critical crack size for rupture in such welds have received recent attention [59]. Axial PWSCC aws in these welds are limited to the width of the alloy
82/182/132 weld material. Experience has conrmed that the PWSCC cracks arrest when they reach the PWSCC-resistant low-alloy steel and stainless steel materials [50]. Therefore, the maxi-mum axial crack lengths are limited to a few inches at most (much less than the critical axial aw length), except for the small number of cases involving alloy 600 safe ends or alloy 600
pipe/tube (CRDM and BMI nozzles), where axial cracks initiating in the weld could potentially propagate into the alloy 600 base
metal. Thus, critical crack size calculations for PWR piping dis-similar metal butt welds typically assume one or more circumfer-entially oriented PWSCC aws.In 2007, EPRI sponsored a detailed investigation of the growthof circumferential PWSCC aws in PWR pressurizer nozzle dis-similar metal butt welds [59]. Using nite-element methods, this study examined the effect of an arbitrary crack prole on crack growth and subsequent crack stability and leak rate versus the
standard assumption of a semi-elliptical crack prole. The crack stability (i.e., critical crack size) modeling of the EPRI study was
based on a standard limit load (i.e., net section collapse)
approach as applied to an arbitrary crack prole around the weld circumference [65]. The potential for an EPFM failure mode was considered using a Z-factor approach specic to piping dissimilar metal welds [66]. Finally, the role of secondary piping thermal constraint stresses in the rupture process was investigated on the basis of available experimental pipe rupture data [67], elastic-
plastic nite-element analyses of a pipe with an idealized through-thickness crack [59], and pressurizer surge line piping models applied to evaluate the maximum capacity of the secondary loads to produce rotation at a cracked pressurizer surge nozzle [59].44.6.3Boric Acid Wastage Due to Larger Leaks Small concentrations of boron are added to the primary coolantwater in PWR plants in the form of boric acid to aid in controlling core reactivity. At the start of an operating cycle with new fuel, the boron concentration is typically about 2,000 ppm or less. The concentration of boron is reduced with fuel burnup to about 0 ppm at the end of an operating cycle when fuel is ready to be replaced. Work by EPRI and others to determine the probable rate of corrosion of low-alloy steel by boric acid is documented in the
EPRI Boric Acid Corrosion Guidebook [43]. This document shows that the corrosion rate of low-alloy steel by deareated pri-mary coolant (inside the pressure vessel and piping) with 2,000 ppm boron is negligible. The corrosion rate for low concentration (2,000 ppm) aerated boric acid is also very low. However, when high-temperature borated water leaks onto a hot surface, the water can boil off leaving concentrated aerated boric acid. The corro-sion rate of low-alloy steel by hot concentrated aerated boric acid
can be as high as 10 in./year under some conditions. As evidenced by the signicant volume of material corrodedfrom the Davis-Besse reactor vessel head, boric acid corrosion
has the potential to create signicant safety risk. Figure 44.13 shows cross-section and plan views of the corroded region of the Davis-Besse head shown in Fig. 44.9. As indicated, a large vol-ume of the low-alloy head material was corroded, leaving the stainless steel cladding on the inside of the vessel head to resist the internal pressure. Part-depth cracks were discovered in the
unsupported section of cladding. FIG.44.12SCHEMATIC OF RPV TOP-HEAD NOZZLEGEOMETRY AND NATURE OF OBSERVED CRACKING ASME_Ch44_p001-026.qxd 12/19/09 7:36 AM Page 12 COMPANION GUIDE TO THE ASME BOILER & PRESSURE VESSEL CODE
¥13Based on available evidence, it was determined that the leakage that caused the corrosion had been occurring for at least 6 years.
While it was known that boric acid deposits were accumulating on the vessel top head surface, the utility attributed the accumula-tions to leakage from spiral-wound gaskets at the anged joints
between the CRDM nozzles and the CRDMs. The accumulations of boric acid had not been removed due to poor access to the enclosed plenum between the top of the vessel head and the bot-tom of the insulation, as shown in Fig. 44.14. The transition from relatively benign conditions, such as shownin Figs. 44.8 and 44.11, to severe conditions, which created the cav-ity shown in Figs. 44.9 and 44.13, is believed to be a function of the
leakage rate. A PWSCC crack that rst breaks through the nozzle wall or weld will initially be small (short), resulting in a low leak rate. It is believed that the small leak rate will not lower the metal surface temperature enough to allow liquid conditions to exist. As the crack grows in length above the J-groove weld, the leak rate is expected to increase to the point where boric acid on the surface near the leak remains moist or where the leaking borated water locally cools the low-alloy steel material to the point where the sur-face will remain wetted and allow boric acid to concentrate.
Preliminary models of these conditions have been developed, and test work was started by EPRI in 2004 to more accurately deter-
mine the conditions where the leakage produces wetted conditions
that can cause high boric acid corrosion rates and where the leakage
results in essentially benign dry boric acid deposits. Conditions such as occurred at Davis-Besse can be prevented bya three-step approach. Firstly, perform nondestructive examinations of the nozzles frequently enough to catch PWSCC cracks beforethey grow through wall. Secondly, clean the external surfaces of preexisting boric acid deposits from other sources and perform bare metal visual inspections at frequent enough intervals to detect leaks at an early benign stage. Thirdly, if the risk is believed high orFIG.44.13PLAN AND CROSS-SECTION THROUGH CORRODED PART OFDAVIS-BESSE REACTOR VESSEL HEADFIG.44.14CROSS-SECTION THROUGH DAVIS-BESSE REACTOR VESSEL HEAD ASME_Ch44_p001-026.qxd 12/19/09 7:37 AM Page 13 14¥Chapter 44inspections are difcult or costly, replace the susceptible parts or apply a remedial measure to reduce the risk of PWSCC leaks. 44.7DEGRADATION PREDICTIONS 44.7.1Crack Initiation Initiation of PWSCC in laboratory test samples and in PWRsteam generator tubing has been found to follow standard statisti-cal distributions such as Weibull and log-normal distributions
[4447]. These distributions have been widely used for modeling
and predicting the occurrence of PWSCC in PWRs since about
1988, and continue to be used for this purpose.The parameters of a statistical distribution used to model agiven mode of PWSCC, such as axial cracks in CRDM nozzles, only apply to the homogeneous set of similar items that are exposed to the same environmental and stress conditions, and only to the given crack orientation being modeled. For example, axial and circumferential cracking are modeled separately since the stresses acting on the two crack orientations are different. In general, two parameter Weibull or log-normal models are used to model and predict the future occurrence of PWSCC. An initia-tion time, which sometimes is used as a third parameter, is not gen-
erally modeled, because use of a third parameter has been found to result in too much exibility and uncertainty in the predictions.
PWSCC predictions are most reliable when the mode of crack-ing is well developed with results for detected cracking available
for three or more inspections. In this situation, the tted parameters
to the inspection data are used to project into the future. When no
cracking has been detected in a plant, rough predictions can still be developed using industry data. This is generally done using a two-step process. The rst step involves developing a statistical distribu-tion of times to occurrence of PWSCC at a selected threshold level (such as 0.1%) for a set of plants with similar designs. Data for plants with different temperatures are adjusted to a common tem-perature using the Arrhenius equation (see Table 44.1). The distrib-ution of times to the threshold level is used to determine a best esti-mate time for the plant being modeled to develop PWSCC at that threshold level. Techniques are available to adjust the prediction to
account for the time already passed at the plant without detecting the mode being evaluated. Once the best estimate time for occur-rence at the threshold level is determined, future cracking is pro-jected from that point forward using the median rate of increase (Weibull slope or log-normal standard deviation) in the industry for the mode of PWSCC being evaluated. 44.7.2Crack Growth Numerous PWSCC crack growth studies have been performedon thick-wall alloy 600 material in PWR environments at test tem-
peratures that span the range of typical PWR operating tempera-tures. In 2002, these tests were reviewed and summarized under
sponsorship of EPRI [26,48]. The EPRI study (MRP-55) conclud-ed that PWSCC crack growth rates for thick-wall alloy 600 base metal behave in accordance with the following relationship:
wherecrack growth rate at temperature T in m/sec (or in./hr)
Q gthermal activation energy for crack growth 130 kJ/mole (31.0 kcal/mole) a..a=exp c-Q g R a 1 T-1 T ref bd a (K-K th)b R universal gas constant 8.314 103 kJ/mole ¥ K (1.103 103 kcal/mole ¥ R)
T absolute operating temperature at location of crack, K (or R)Trefabsolute reference temperature used to normalize data 325C 598.15 K (617F 1076.67 R) crack growth amplitude K crack tip stress intensity factor, Mpa m (or ksi in)K thcrack tip stress intensity factor threshold 9 Mpa m (8.19 ksi in) exponent 1.16Temperature dependence is modeled in this crack growth rate equation via an Arrhenius-type relationship using the aforemen-tioned activation energy of 31 kcal/mole. The stress intensity factor dependence is of power law form with exponent 1.16.
Figure 44.15 presents the distribution of the coefcient () in thepower law relationship at constant temperature (617F). The datain this gure exhibit considerable scatter, with the highest and lowest data points deviating by more than an order of magnitude from the mean. The 75th percentile curve (see Figure 44.15a) was recommended for use in deterministic crack growth analyses
[26,48], and this curve is now included in Section XI for disposi-tion of PWSCC aws in RPV top-head nozzles. In addition, prob-abilistic crack growth rate studies have been performed of top head nozzles using the complete distribution [49]. An additional factor of 2 has been applied to the 75th percentile value when analyzing crack growth exposed to leakage in the annular gap between the nozzle and the head, to allow for possible abnormal water chemistry conditions that might exist there [26,48]. Similar crack growth rate testing has been conducted foralloys 82 and 182 weld metals. The weld metal crack growth data are sparser and exhibit similar statistical variability. A review of weld metal PWSCC crack growth data has also been
completed under EPRI sponsorship [61,62]. This study (MRP-115) showed that Alloy 182/132 weld metal crack growth obeys a similar relationship to that shown above for alloy 600 base metal, but with crack growth rates about four times higher than the alloy 600 curve for stress intensity factors greater than about 20 ksi in (see Figure 44.15a). Similar to the heat-by-heat analy-sis for the wrought material, a weld-by-weld analysis was per-formed on the available worldwide laboratory crack growth rate
data for the weld materials (see Figure 44.15b). The EPRI study (MRP-115) concluded that PWSCC crack growth rates for alloy 82/182/132 weld metal behave in accordance with the following relationship, where no credit for a stress intensity factor thresh-old greater than zero was taken because of insufficient data on
this parameter:
where:crack growth rate at temperature T in m/s (or in/h)
Q gthermal activation energy for crack growth130 kJ/mole (31.0 kcal/mole)
Runiversal gas constant8.314 103 kJ/mole-K (1.103 103 kcal/mole-R)
Tabsolute operating temperature at location of crack, K (or R)a..a=exp c-Q g R a 1 T-1 T ref bd a f alloy f orient K b ASME_Ch44_p001-026.qxd 12/19/09 7:37 AM Page 14 COMPANION GUIDE TO THE ASME BOILER & PRESSURE VESSEL CODE
¥15 Trefabsolute reference temperature used to normalize data598.15 K (1076.67R)power-law constant1.5 1012at 325C for in units of m/s and K inunits of MPa m (2.47 107at 617F for in units of in/h and K in units of ksi in)f alloy1.0 for Alloy 182 or 132 and 1/2.6 0.385 for Alloy 82 f orient1.0 except 0.5 for crack propagation that is clearly perpendicular to the dendrite solidication direction Kcrack-tip stress intensity factor, MPa m (or ksi in)exponent1.6Deterministic crack growth rate predictions have been per-formed for axial and circumferential cracking in RPV top- and bottom-head nozzles and in large-diameter butt welds [49,50].
Welding residual stresses are a primary factor contributing to crack growth in all these analyses. Stress intensity factors versus
crack size, considering residual stresses plus operating pressure
and thermal stresses are rst computed in these studies. These are a.a.then inserted into the appropriate crack growth relationship (alloy 600, 82, or 182) at the component operating temperature and inte-grated with time to predict crack size versus operating time at the
applicable temperature. Figure 44.16 shows typical crack growth predictions for a cir-cumferential crack in a steep angle RPV top-head (CRDM) noz-zle. (Nozzles in the outer rings of vessel heads having the steepest angles between the nozzle and the head have been found to be controlling in terms of predicted growth rates for circumferential
cracks). The analysis depicted in Fig. 44.16 assumed a through-wall, 30of circumference crack in the most limiting azimuthallocation of the nozzle at time zero, and predicted the operating time for it to grow to a size that would violate ASME Section XI aw evaluation margins with respect to nozzle ejection (~300). It isseen that, even for relatively high RPV temperatures, operating times on the order of 8 years or greater are predicted for circumfer-ential nozzle cracks to propagate to a size that would violate ASME Section XI safety margins.Figure 44.17 shows similar crack growth predictions for apostulated circumferential crack in a large-diameter nozzle butt weld. Stress intensity factors were computed in this analysis forFIGURE 44.15ADETERMINISTIC CRACK GROWTH RATE CURVES FOR THICK-WALL ALLOY 600 WROUGHT MATERIAL AND FOR ALLOY 182/132 AND ALLOY 82 WELD MATERIALS [61,62]FIGURE 44.15BLOG-NORMAL FIT TO 19 WELD FACTORS FOR SCREENED MRP DATABASE OF CGR DATA FOR ALLOY 82/182/132 [61,62]
ASME_Ch44_p001-026.qxd 12/19/09 7:37 AM Page 15 16¥Chapter 44a 6-to-1 aspect ratio crack in a large-diameter RPV inlet/outletnozzle, ranging in depths from 0.1 in. to 2.2 in. The nozzle was conservatively assumed to have a large, inside surface repair, and the crack was assumed to reside in the center of that repair (i.e., in the most unfavorable residual stress region of the weld).
The predicted crack growth in this case is fairly rapid for a typi-
cal outlet nozzle temperature, 602F, propagating to 75%through-wall (the upper bound of ASME Section XI allowable aw sizes in piping) in about 3 years. Conversely, if no weldrepair were assumed, little or no crack growth would be predict-ed over the plant lifetime. For this same crack, including the effect of the repair, the predicted time for a 0.1 in. deep crack to grow to 75% through-wall at a typical inlet nozzle temperature
(555F) is about 11 years. The strong effect of operating temperature is apparent in bothcrack growth analyses. The outlet nozzle analysis also demon-strates the detrimental effect of weld repairs that were performed
during construction at some plants. FIG.44.17CRACK GROWTH RATE PREDICTIONS FOR CIRCUMFERENTIALCRACKS IN RPV MAIN COOLANT LOOP DISSIMILAR METAL NOZZLE BUTT WELD AT OPERATING TEMPERATURES TYPICAL OF REACTOR INLET AND OUTLET NOZZLES INITIAL CRACK ASSUMPTION 0.10.6INSIDESURFACE CRACK AT MAXIMUM STRESS AZIMUTH IN NOZZLE WITH ASSUMED INSIDE SURFACE FIELD REPAIR.FIG.44.16CRACK GROWTH RATE PREDICTIONS FOR CIRCUMFER-ENTIALCRACKS IN RPV TOP-HEAD NOZZLE AT VARIOUS ASSUMED OPERATING TEMPERATURES INITIAL CRACK ASSUMPTION 30THROUGH-WALLCRACK AT MAXIMUM STRESS AZIMUTH IN HIGH ANGLE NOZZLE.
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¥1744.7.3Probabilistic Analysis Because of the large degree of statistical scatter in both thecrack initiation and crack growth behavior of PWSCC in alloy
600 base metal and associated weld metals, probabilistic fracture mechanics (PFM) analyses have been used to characterize the phenomenon in terms of the probabilities of leakage and failure
[49] for RPV top head nozzles. The analysis incorporates the fol-lowing major elements:(a)computation of applied stress intensity factors for circum-ferential cracks in various nozzle geometries as a function
of crack length and stresses (b)determination of critical circumferential aw sizes for noz-zle failure (c)an empirical (Weibull) analysis of the probability of nozzle cracking or leakage as a function of operating time and tem-
perature of the RPV head(d)statistical analysis of PWSCC crack growth rates in thePWR primary water environment as a function of applied stress intensity factor and service temperature (e)modeling of the effects of inspections, including inspectiontype, frequency, and effectivenessA series of PFM analysis results is presented in [49], which cov-ers a wide variety of conditions and assumptions. These include base cases, with and without inspections, and sensitivity studies to evaluate the effects of various statistical and deterministic assump-tions. The model was benchmarked with respect to eld experience, considering the occurrence of cracking and leakage and of circum-ferential cracks of various sizes. The benchmarked parameters were then used to evaluate the effects of various assumed inspection pro-grams on probability of nozzle failure and leakage in actual plants.
A sample of the results is presented in Figs. 44.18 and 44.19. Figure 44.18 shows the effect of inspections on probability ofnozzle failure (Net Section Collapse, or ejection of a nozzle) for head operating temperatures ranging from 580F to 600F. A no-inspection curve is shown for each temperature. Runs were then
made assuming NDE inspections of the nozzles. Inspections were assumed to be performed at intervals related to head operating tem-
perature (more frequent inspections for higher head temperatures,FIG.44.18PROBABILITY OF NOZZLE FAILURE (NSC) AS A FUNCTION OFVARIATIONS IN TOP-HEAD TEMPERATURE AND INSPECTION INTERVALSFIG.44.19PROBABILITY OF NOZZLE LEAKAGE AS A FUNCTION OF VARIATIONS IN TOP-HEAD TEMPERATURE AND INSPECTION INTERVALS ASME_Ch44_p001-026.qxd 12/19/09 7:37 AM Page 17 18¥Chapter 44less frequent for lower temperatures). It is seen from the gurethat the assumed inspection regimen is sufcient to maintain the nozzle failure probability (per plant year) below a generally accepted target value of 1 103 for loss of coolant accidents due to nozzle ejection. Figure 44.19 shows similar results for the probability of leak-age from a top-head nozzle. It is seen from this gure that the same assumed inspection regimen maintains the probability of
leakage at or about 6% for the cases analyzed. Analyses similar to those reported in Figs. 44.18 and 44.19 have been used, in conjunc-
tion with deterministic analyses, to dene an industry-recommended inspection and corrective action program for PWR top heads thatresults in acceptable probabilities of leakage and failure. This work also constituted the basis for the inspection requirements
incorporated in ASME Code Case N-729-1 [63].Similar probabilistic analyses have been performed for PWSCCsusceptible butt welds in pressurizer nozzles, as part of the effort
documented in MRP-216 [59]. Analyses established the current expected aw distribution based on pressurizer nozzle DMW inspections to date, (Table 44.1), estimates were made of the prob-ability of cracking versus aw size, and of crack growth rate ver-sus time. A plot of the aw indications found to date, in terms of crack length as percentage of circumference (abscissa) versus crack depth as percentage of wall thickness (ordinate) is illustrated in Figure 44.19a. Axial indications plot along the vertical axis (l/circumference = 0) in this plot, with leaking aws plotted at a/t
= 100%. Circumferential indications plot at non-zero values of
l/circumference, at the appropriate a/t. Clean inspections are plot-ted randomly in a 10% box near the origin, to give some indication of inspection uncertainty. Also shown on this plot are loci of criti-cal aw sizes based on an evaluation of critical aw sizes present-ed in Ref. [59]. 50th and 99.9th percentile plots are shown. It isseen from this gure that all of the aw indications detected werefar short of the sizes needed to cause a rupture. The probabilistic analysis also addressed the small but nite probability that larger aws may exist in uninspected nozzles, plus the potential for crack growth during future operating time until all the nozzles are
inspected (or mitigated) under MRP-139 [58] guidelines.44.8 REPAIRS When cracking or leakage is detected in operating nuclearpower plant pressure boundary components, including the reactor vessel, repair or replacement may be performed in accordance
with ASME BPVC Section XI [51].Section XI species that the aws must be removed or reduced to an acceptable size in accor-dance with Code-accepted procedures. For PWSCC in RPV alloy 600 components, several approaches have been used. 44.8.1Flaw Removal For relatively shallow or minor cracking, aws may beremoved by minor machining or grinding. This approach is per-mitted by the ASME Code to eliminate aws and return the com-ponent to ASME Code compliance. However, this approach gen-
erally does not eliminate the underlying cause of the cracking.
There will still be susceptible material exposed to the PWR envi-ronment that caused the cracking originally, and in some cases the susceptibility might be aggravated by surface residual stresses caused by the machining or grinding process. Simple aw removal is thus not considered to be a long-term repair, unless combined with some other form of mitigation. However, in the short term, for example, where future component replacement is
planned, it may be a viable approach for interim operation.FIGURE 44.19APRESSURIZER DISSIMILAR METAL BUTT WELD FLAW INDICATIONSCOMPARED TO CRITICAL FLAW SIZE PROBABILITY ESTIMATES ASME_Ch44_p001-026.qxd 12/19/09 7:37 AM Page 18 COMPANION GUIDE TO THE ASME BOILER & PRESSURE VESSEL CODE
¥1944.8.2Flaw Embedment Surface aws are much more signicant than embedded awsfrom a PWSCC perspective, because they continue to be exposed to the PWR primary water environment that caused the crack and that can lead to continued PWSCC aw growth after initiation.
Accordingly, one form of repair is to embed the aw under a PWSCC-resistant material. Figure 44.20 shows an embedment approach used by one vendor to repair PWSCC cracks or leaks in
top-head nozzles and welds. The PWSCC-susceptible material, shown as the cross-hatched region in the gure, is assumed to be entirely cracked (or just about to crack). PWSCC-resistant material, typically alloy 52 weld metal, is deposited over the susceptible material. The assumed crack is shown to satisfy all ASME BPVC Section XI aw evaluation requirements, in the absence of any growth due to PWSCC, since the crack is completely isolated from the PWR environment by the resistant material. Note that the resistant material in this repair must overlap the susceptible material by enough length in all directions to preclude new cracks growing around the repair and causing the original crack to be reexposed to the PWR environment. Although this repair approach has been used successfully in several plants, there have been many incidents in which nozzles repaired by this approach during one refueling outage have been found to be leaking at the subsequent outage. These occurrences were attributed to lack of sufcient overlap of the repair, because it is sometimes difcult to distinguish the exact point at which the susceptible material ends (for instance the end of the J-groove weld butter and the begin-
ning of the RPV cladding in Fig. 44.20).44.8.3Weld Overlay Another form of repair that has been used extensively to repaircracked and leaking pipe welds is the weld overlay (WOL).
Illustrated schematically in Fig. 44.21, WOLs were rst con-ceived in the early 1970s as a repair for IGSCC cracking and leakage in BWR main coolant piping. Over 500 such repairs have
been applied in BWR piping ranging from 4 in. to 28 in. in diam-eter, and some weld overlay repairs have been in service for over 20 years, with no evidence of any resumption of the IGSCCproblem. Although WOLs, shown in Fig. 44.21, do not eliminatethe PWSCC environment from the aw as in the aw embedment process, the repair has been shown to offer multiple improve-ments to the original pipe welds, including the following: (a)structural reinforcement (b)resistant material (c)favorable residual stress reversalWeld overlays also offer a signicant improvement in inspec-tion capability, because once a weld overlay is applied, the required inspection coverage reduces to just the weld overlay material plus the outer 25% of the original pipe wall, often a
much easier inspection than the original dissimilar metal weld (DMW) inspection. Weld overlay repairs have been recognized as a Code-accept-able repair in an ASME Section XI Code Case [52] and accepted by the U.S. NRC as a long-term repair. They have also been used, albeit less extensively, to repair dissimilar metal welds at nozzles
in BWRs. The weld overlay repair process was rst applied to a PWRlarge-diameter pipe weld (on the Three Mile Island 1 pressurizer to hot-leg nozzle) in the fall of 2003. Since that time, as part of the MRP-139 inspection effort described in para. 44.5.6, over 200 weld overlays have been applied to pressurizer nozzle dissimilar metal butt welds. Part of the reason for this trend is that many pressurizer nozzles were unable to be volumetrically inspected to achieve the required examination coverage in their original con-guration. By applying weld overlays, in addition to mitigating the welds, their inspectability was enhanced such that post over-lay ultrasonic exams could be performed in accordance with applicable requirements. Technical justication for the WOL
process as a long-term repair is documented in Ref. [53].
Requirements for weld overlays in PWR systems, including their use as mitigation as well as repair, is documented in Ref. [60].44.8.4Weld Replacement Finally, the awed weld may be replaced in its entirety. In PWRtop-head nozzles, this process has been implemented extensively by
relocating the pressure boundary from the original PWSCC-
susceptible J-groove weld at the inside surface to a new weld at themidwall of the RPV head (see Fig. 44.22). With this repair approach, the PWSCC-susceptible portion of the original J-groove weld and buttering is left in the vessel, but it is no longer part of the pressure-retaining load path for the nozzle. The lower portion of the original nozzle is rst removed by machining to a horizontal ele-vation above the J-groove weld (left-hand side of Fig. 44.22).A
weld prep is produced on the bottom of the remaining portion ofFIG.44.20SCHEMATIC OF RPV TOP-HEAD NOZZLEFLAW EMBEDMENT REPAIRFIG.44.21SCHEMATIC OF WELD OVERLAY REPAIR APPLIED TO RPV OUTLET NOZZLE ASME_Ch44_p001-026.qxd 12/19/09 7:37 AM Page 19 20¥Chapter 44 the nozzle, and a new, horizontal weld is made between the original nozzle and the bore of the RPV head (righthand side of Fig. 44.22).
The new weld is made with PWSCC-resistant material (generally alloy 52 weld metal), and the surface of the weld is machined for NDE. The repair process still leaves some portion of the original PWSCC-susceptible alloy 600 nozzle in place, potentially in a high residual stress region at the interface with the new weld. However, a surface treatment process, such as roll peening or burnishing, has been applied to this interface in many applications to reduce poten-
tial PWSCC concerns. Experience with this repair process has been
good, in terms of subsequent leakage from repaired nozzles, and in most cases the repair need only survive for one or two fuel cycles, because, once PWSCC leakage is detected in an RPV head, com-
mon industry practice has been to schedule a future head replace-ment (not because of the repaired nozzle, but because of concerns that other nozzles are likely to be affected by the problem leading to costly future inspections, repairs, and outage extensions).44.9REMEDIAL MEASURES 44.9.1Water Chemistry Changes Three types of water chemistry changes that could affect the rate of PWSCC are zinc additions to the reactor coolant, adjust-
ments to hydrogen concentration, and adjustments to lithium concentration and pH. The factors are described below. (a)Zinc Additions to Reactor Coolant.
Laboratory tests indicatethat the addition of zinc to reactor coolant signicantly slows down the rate of PWSCC initiation, with the improvement factor increasing as the zinc concentration increases [29].
The improvement factor (slowdown in rate of new crack ini-tiation) shown by tests varies from a factor of two for 20 ppb zinc in the coolant to over a factor of ten for 120 ppb zinc.
The effect of zinc on crack growth rate is not as certain, with some tests indicating a signicant reduction in crack growth rate but others indicating no change. Further testing is under-way under EPRI sponsorship (as of 2004) to clarify the effects of zinc on crack growth rate. As of mid-2004, evalu-ation of plant data, especially the data for a two-unit station
with PWSCC at dented steam generator tube support plates, is encouraging but not conclusive with regard to whether use
of zinc is reducing the rate of PWSCC. The uncertainty is the
result of changes in inspection methods simultaneously with
changes in zinc concentration. (b)Adjustments of Hydrogen Concentration.
The EPRI PWRPrimary Water Chemistry Guidelines require the hydrogen concentration in the primary coolant to be kept between 25
and 50 cc/kg [28]. As discussed in the Guidelines and sum-marized above in para. 44.3.4, the rate of PWSCC initiation and rate of PWSCC crack growth both seem to be affected by the hydrogen concentration, with lower concentrations being more aggressive at lower temperature and higherFIG.44.22SCHEMATIC OF RPV TOP-HEAD NOZZLE WELD REPLACEMENT REPAIR ASME_Ch44_p001-026.qxd 12/19/09 7:37 AM Page 20 COMPANION GUIDE TO THE ASME BOILER & PRESSURE VESSEL CODE
¥21 concentrations at higher temperature. Depending on theplant situation as far as which parts are at most risk of
PWSCC, and depending on the temperature at those parts, there may be some benet, such as an improvement factor
of about 1.2, in operating at hydrogen concentrations at either end of the allowed range. In the longer term, increased benet may be achieved by operating slightly outside of the allowed range (e.g., at 60 cc/kg), although
this will require conrmation that the change does not result in some other undesirable effects. (c)Adjustments of Lithium Concentration and pH.
As dis-cussed in para. 44.3.4, some tests indicate that the rate of
PWSCC initiation is increased by increases in lithium con-centration and pH, e.g., by factors ranging from about 1.15
to 2.0. On the other hand, increases in lithium and pH pro-vide proven benets for reducing the potential harmful deposit buildup on fuel cladding surfaces and for reducing shutdown dose rates [28]. Based on these opposing trends, plants can select a lithium/pH regime that best suits their needs, i.e., does not involve substantial risks of aggravating PWSCC, while still providing benets for reducing fuel deposits and shutdown dose rates. When evaluating the pos-
sible risks to PWSCC of increasing lithium and pH, it should be noted that crack growth rate tests show no harm-ful effect while crack initiation tests do. The data from crack growth rate tests are considered to be more reliable, and it is recommended that they be given greater weight than the
results from crack initiation tests. An additional considera-tion is that the use of zinc can provide a stronger benet
than the possible decit associated with increases in lithium and pH, and, thus, can make use of a combined zinc adjust-ment and increase in lithium and pH attractive. 44.9.2Temperature Reduction To date, a main remedial measure applied in the eld for RPV top-head PWSCC has been modication of the reactor internals package to increase bypass ow through the internals ange region and, thereby, reduce the head temperature. The lower head
temperature is predicted to reduce the rates of crack initiation and growth based on the thermal activation energy model, as shown in Table 44.1. However, experience in France suggests that PWSCC may occur at head temperatures close to the reactor cold-leg tem-perature. This is especially signicant given PWSCC of two South Texas Project Unit 1 BMI nozzles at a temperature of about
565F. The South Texas Project experience shows that materialsand fabrication-related factors can result in PWSCC at tempera-tures lower than otherwise expected.44.9.3Surface Treatment EPRI has sponsored tests of a range of mechanical remedialmeasures for PWSCC of alloy 600 nozzles. Results of these tests were reported by Rao at the Fontevraud 5 Symposium [54]. The
remedial measures test program consisted of soliciting remedial measures from vendors, fabricating full-diameter and wall-thick nessring specimens from archive CRDM nozzle material, installing specimens in rings that locked in high residual stresses on the specimen inside surface, applying the remedial measures to the stressed surface, and then testing the specimens in doped steam with hydrogen overpressure at 400C (750F). The specimenswere removed from the autoclave at intervals and inspected for SCC. A complicating factor in interpreting the test results is thatnot all of the specimens were fabricated from the same heat ofmaterial. Therefore, there were differences in material PWSCC susceptibility in addition to differences in remedial measure effec-tiveness. The methods used to correct for differences in specimen PWSCC susceptibility are discussed in the paper. The remedial measures fell into three main effectiveness groups. (a) most effective (1)waterjet conditioning (2) electro mechanical nickel brush plating
(3) shot peening (b) intermediate effectiveness (1) electroless nickel plating (2) GTAW weld repair
(3) laser weld repair(c) least effective (1) EDM skim cutting (2) laser cladding (3)apper wheel surface polishingAs of May 2005, it is not believed that any of these remedialmeasures had actually been applied to a reactor vessel in the eld. 44.9.4Stress Improvement To mitigate against the IGSCC problem in BWR piping, manyplants implemented residual stress improvement processes. These were performed both thermally (induction heating stress improve-
ment or IHSI) and by mechanical means (mechanical stress improvement process or MSIP). As described above, residual
stresses play a major role in susceptibility to both IGSCC and PWSCC, because large piping butt welds tend to leave signicant residual stresses at the inside surfaces of the pipes, especially
when eld repairs were performed during construction. Both stress improvement processes have been demonstrated to reverse the unfavorable residual stresses, leaving compressive stresses on the inside surface of the pipe, which is exposed to the reactor environment. MSIP has also been applied to PWSCC-susceptible butt welds in PWR piping, primarily dissimilar metal welds at vessel nozzles, such as the V.C. Summer outlet nozzle cracking problem described above. As long as the stress improvement process is applied relatively early in life, when cracking has not initiated or grown to signicant depths, it clearly constitutes a useful remedial measure that can be applied to vessel nozzles, eliminating one of the major factors that contribute to PWSCC. One of the benets of the weld overlay process described aboveto repair PWSCC-cracked butt welds is that it reverses the resid-ual stress pattern in the weld, resulting in compressive stresses on the inside surface. Thus, a novel mitigation approach that is being explored at several plants is the application of weld overlays pre-emptively, before cracking is discovered. Applying a preemptive WOL in this manner produces the same remedial benets described above for the stress improvement processes, but also
places a PWSCC-resistant structural reinforcement on the outer surface of the pipe. So, if the favorable residual stresses were to relax in service, or for some reason be ineffective in arresting the PWSCC phenomenon, the PWSCC-resistant overlay would still provide an effective barrier against leakage and potential pipe rupture. Moreover, the revised inspection coverage requirements specied for WOLs apply to such preemptive overlays, providing
the added benet of enhanced inspectability [52].
ASME_Ch44_p001-026.qxd 12/19/09 7:37 AM Page 21 22¥Chapter 4444.9.5Head Replacement The most obvious way to address RPV top-head cracking issues is head replacement. Approximately one-third of operating PWRs in the United States have replaced their heads or have
scheduled head replacements in the near future. Such head replacements take advantage of the lessons learned to date regard-ing the PWSCC phenomenon, and the new heads are generally
produced so as to eliminate all PWSCC-susceptible materials, replacing them with resistant materials (alloy 690 and associated weld metals alloys 52 and 152). RPV head replacement is a key aspect of strategic planning to address the alloy 600 problem in PWRs, and is performed as part of a coordinated alloy 600 main-tenance program that addresses steam generator, pressurizer, and piping issues as well as the RPV. 44.10STRATEGIC PLANNING Within constraints posed by regulatory requirements, utilitiesare free to develop a strategic plan that ensures a low risk of leak-age, ensures an extremely low risk of core damage, and results in the lowest net present value (NPV) cost consistent with the rst two criteria. Development of a strategic plan for RPV top-head nozzles was described by White, Hunt, and Nordmann at the 2004 ICONE-12 conference [55]. The strategic planning process was based on life cycle management approaches and NPV economic modeling software developed by EPRI [56,57]. The main steps in the strategic planning process are as follows:(a)predicting time to PWSCC (b)assessing risk of leaks (c)assessing risk of rupture and core damage due to nozzle ejection (d)assessing risk of rupture and core damage due to boric acidwastage (e)identifying alternative life cycle management approaches (f)evaluating economically the alternative management approaches While the paper and following discussion are based on RPV top-head nozzles, the same basic approach can be applied to BMI nozzles and butt welds. 44.10.1Predicting Time to PWSCC Predictions of the time to PWSCC crack initiation are described in para. 44.7.1. The predictions are typically based on a statistical distribution such as a two-parameter Weibull or log-
normal model. Predictions are most accurate if based on plant-specic repeat inspections showing increasing numbers of cracked nozzles. If such data are not available, then predictions
are typically based on data for other similar plants (e.g., design, material, operating conditions) with corrections for differences in
operating time and temperature. 44.10.2Assessing Risk of Leaks The risk of leakage at a particular point in time (typically refu-eling outage number) is typically determined by a probabilistic (Monte-Carlo) analysis using the distribution of predicted time to crack initiation (para. 44.7.1), crack growth (para. 44.7.2), and
other probabilistic modeling techniques (para. 44.7.3). The proba-bilistic analysis should include a sensitivity study to identify the
most important analysis input parameters, and these important parameters should be reviewed to ensure that they can be substan-tiated by available data. 44.10.3Assessing Risk of Rupture and Core Damage Due to Nozzle Ejection The risk of nozzle ejection (net section collapse) is determined using methods such as described in para. 44.6.2. 44.10.4Assessing Risk of Rupture and Core DamageDue to Boric Acid Wastage The risk of failure of the carbon or low-alloy steel reactor ves-sel head by boric acid wastage is determined using methods such
as described in para. 44.6.3. 44.10.5 Identifying Alternative Life CycleManagement Approaches An important step in developing a life cycle management planis to identify the alternative approaches that can be considered.
These alternatives can include the following: (a)continue to inspect and repair indenitely without applying remedial measures. (b)apply remedial measures, such as lowering the vessel headtemperature by increasing bypass ow through the vessel
internals ange, adding zinc to the primary coolant, and water-jet conditioning the wetted surface of nozzles and welds to remove small aws and leave the material surface with a compressive residual stress. (c)replace the vessel head as quickly as possible.
(d)replace the vessel head shortly after detecting the rst PWSCC cracks. (e)use other approaches identied.Each of these alternatives must be studied to determine thedifculty of application, the likely effectiveness, and the effect of the change on required inspections. For example, head replace-ment may involve the need to cut an access opening in the con-tainment structure or to procure a new set of CRDMs to allow the head changeout to be performed quickly, so as to not adversely affect the refueling outage time. If openings must be cut in con-tainment, consideration should also be given to the possible need
to cut other openings in the future, such as for steam generator or pressurizer replacements. Consideration must also be given to the
disposal of a head after it is replaced. 44.10.6Economic Evaluations of AlternativeManagement Approaches Most life cycle management evaluations include economicanalyses to determine the NPV cost of each alternative. The NPV cost is the amount of money that is required today to pay all pre-dicted future costs, including the effects of ination and the dis-
count rate. Inputs to an LCM economic analysis typically include the following: (a)costs of planned preventive activities including inspections, remedial measures, and replacements.(b)predicted failure mechanisms (e.g., cracks, leaks, and rup-ture) and failure rates. (c)costs for corrective maintenance in the event of a failureincluding the cost to make the repair, the estimated value of lost production, and an allowance for consequential costs such as increased regulatory scrutiny. Consideration should be given to the fact that a major incident such as the Davis-Besse RPV head wastage can result in lost production and conse-quential costs far higher than the cost to replace the affected
component.
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¥23Figure 44.23 shows typical results of a strategic planning analysis with economic modeling. The nal step in the economic evaluation is to review the pre-dictions in light of other plant constraints, such as planned plant life, potential power uprates, budget constraints, and the availability of replacement heads. In many cases, the alternative with the low-
est predicted NPV cost may not represent the best choice.44.11REFERENCES 1.SMC 027, Inconel Alloy 600. In: Special Metals Corporation Handbook. 2000. 2.White DE. Evaluation of Materials for Steam Generator Tubing.Bettis Technical Review, report WAPD-BT-16, December 1959. 3.Howells E, McNary TA, White DE. Boiler Model Tests of Materialsfor Steam Generators in Pressurized Water Reactors. Corrosion 1960;16:241t245t. 4.Copson HR, Berry WE. Qualication of Inconel for Nuclear Power Plant Applications. Corrosion 1960;16:79t85t.5.Copson HR. Effect of Nickel Content on the Resistance to Stress-Corrosion Cracking of Iron-Nickel-Chromium Alloys in Chloride Environments. First International Congress on Metallic CorrosionLondon, 1961, p328333; Butterworths, 1962.6.LaQue FL, Cordovi MA. The Corrosion of Pressure Circuit Materialsin Boiling and Pressurized-Water Reactors (Special Report 69).
London: The Iron and Steel Institute; 1961: 157178. 7.Copson HR, Berry WE. Corrosion of Inconel Nickel-ChromiumAlloy in Primary Coolants of Pressurized Water Reactors. Corrosion 1962;18:21t26t. 8.Bush SH, Dillon RL. Stress Corrosion in Nuclear Systems. StressCorrosion Cracking and Hydrogen Embrittlement of Iron Base Alloys
,Conference held at Unieux-Firminy, France, June 1216, 1973, pp.
61-79, Case 3, NACE, 1977. 9.Coriou MM, et al. Corrosion Fissurante suos Contrainte de LInconeldans LEau `a Haute Temprature. 3e Colloque de Mtallurgie Corrosion (S`eche et Aqueuse), Organis `
a Saclay les 29s30 juin et 1er juillet 1959,North Holland Publishing Cy, Amsterdam, Pays-Bas, 1960.10.Copson HR, Berry WE. Corrosion of Inconel Nickel-ChromiumAlloy in Primary Coolants of Pressurized Water Reactors. Corrosion 1962;18:21t26t.11.Copson HR, Dean SW. Effect of Contaminants on Resistance to StressCorrosion Cracking of Ni-Cr Alloy 600 in Pressurized Water.
Corrosion 1965;21(1):18.12.Copson HR, Economy G. Effect of Some Environmental Variables onStress Corrosion Behavior of Ni-Cr-Fe Alloys in Pressurized Water.
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Proceedings of Fontevraud 5 International Symposium , September 2327, 2002. 55.White GA, Hunt ES, Nordmann NS. Strategic Planning for RPV Head Nozzle PWSCC. Proceedings of ICONE12, 12 thInternational Conference on Nuclear Engineering, April 2529, 2004, Arlington, Virginia. 56.Demonstration of Life Cycle Management Planning for Systems,Structures and Components: With Applications at Oconee and Prairie Island Nuclear Stations, Palo Alto, CA: Electric Power Research Institute; Charlotte, NC: Duke Power; East Welch, MN: Northern States Power (now Xcel Energy); 2001.57.Demonstration of Life Cycle Management Planning for Systems,Structures and Components Lcm VALUE User Manual and Tutorial.
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Conference, San Antonio, TX, 2007.67.G. M. Wilkowski, et al., Degraded Piping Program-Phase II,Summary of Technical Results and Their Signicance to Leak-Before-Break and In-Service Flaw Acceptance Criteria, NUREG/CR-4082, Vol. 18, January 1985 to March 1989.68.Materials Reliability Program Reactor Vessel Closure Head Penetration Safety Assessment for U.S. PWR Plants (MRP-110NP):
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