The AMP can be developed in the following sequence:

1 Identification of degradation mechanisms and locations susceptible to ageing

2 Identification of the mitigation and preventive measures

3 Identification of the parameters to be controlled

4 Definition of the method for the detection of ageing effects

5 Definition of the monitoring, trending, condition evaluation

6 Definition of the acceptance criteria

7 Identification of the corrective actions

8 Organising the administrative control

9 Organising the operational experience feedback.

In reality, the development is some kind of iterative process and steps over­lap, as will be shown below.

All structural materials contain some types of flaw in them, the size of which can range from microscopic to mesoscopic in scale; these defects promote stress to concentrate locally around them leading to premature failure of the structure. The toughness of structures, in presence of inherent defects, is evaluated through a fracture mechanics approach. While most of the codes use a linear elastic fracture mechanics (LEFM) approach, small structures and ductile materials require elastic-plastic fracture mechanics (EPFM) formulations. The validity of LEFM compared to EPFM depends on the plastic zone size as shown in Fig. 1.4 and, in general, LEFM is not applica­ble when the plastic zone size is too large compared to either the crack size, the uncracked ligament or the member height.9 In very large structures and relatively brittle materials where LEFM is valid, the stress fields are charac­terized by stress intensity factor, Kh given by

K^YoJm, [1.9]

where a is half-crack length, a is applied nominal stress and Y isa geometry factor which is a function of the ratio of crack length to its width (a/w). As long as Kj is lower than the plane strain critical fracture toughness KIC, the structure with the crack can withstand the applied loads. In cases where LEFM is not valid (Fig. 1.4) either crack tip opening displacement (CTOD) or elastic-plastic fracture toughness (J-integral) can be conveniently adopted.

Although fracture toughness is a fundamental parameter characterizing the fracture behaviour of cracked bodies, it is often more convenient to use the ductile to brittle transition temperature (DBTT) measured using the relatively simple Charpy impact tests, to study the effect of neutron

radiation exposure (Fig. 1.5).10 These effects are well defined in BCC met­als, as against FCC metals, which exhibit a clear transition from ductile to brittle fracture behaviour as test or operating temperature decreases. It is common practice to consider a reference transition temperature corre­sponding to a specific Charpy impact energy of 41J (50 ft-lb) in lieu of actual nil-ductility transition (RTndt) such that brittle fracture is expected to take place below this reference temperature. As we will note later, exposure of ferritic steels to neutron irradiation leads to decreased fracture energy and increased RTndt commonly referred to as radiation embrittlement of RPV steels. Charpy impact tests are very useful and are conveniently adopted for reactor pressure vessel surveillance programmes (RVSPs). The transition temperature is a function of various factors such as the chemi­cal composition, the temperature, the neutron flux and fluence as well as the microstructure (such as base material, heat-affected zone (HAZ) or weld metal). Validation of thermal annealing of radiation defects in RPV steels is also often established using the Charpy test method. It has been well recognized that other fracture parameters such as crack arrest frac­ture toughness (KIa), dynamic critical stress intensity factor (KId), etc., need to be considered in detailed analyses involving strain-rate effects that become important during a loss of coolant accident (LOCA) condition. It has also been found applicable to high-temperature crack growth, pre­sumably because the plastic stress zone is often relatively small and linear elastic fracture mechanics are considered valid. Another fracture mechan­ics based parameter used to describe creep crack growth is C*. While there are many advantages in using C* analyses in creep of cracked bodies, these types of studies are confined more to scientific curiosities than to techno­logical applications.

In Japan, utilities inspect in accordance with JEAC-4205 (Japan Electric Association Code for ISI Requirements), and the inspection requirements are similar to those in the United States. For the RPV weld lines, a volu­metric examination is conducted on a regular basis. In France, utilities con­duct inspection according to their RSE-M (Rules for In-service Inspection of Nuclear Power Plant Components), and also undertake water pressure test­ing with acoustic emission monitoring, non-destructive inspection during the outages, loose-parts (noise) monitoring during operation, leak detection dur­ing operation and fatigue monitoring. The range of inspection covers the belt­line region of the shell, all welds, top and bottom heads, nozzles and safe end welds, penetrations, control rod drive housings, studs, threaded holes and sup­ports. In Germany, utilities conduct regular inspection using the non-destruc­tive inspection method in accordance with German Code KTA 3201.4.

To justify the safety of LTO, the scope of TLAAs which must be recon­structed or newly performed covers Safety Class 1 and 2 mechanical com­ponents. Examples of the calculations/analyses follow.

For low-cycle fatigue analysis of Safety Class 1 and 2 piping and mechanical components, ASME BPVC was adapted for the calculations (Katona, Ratkai and Pammer, 2011). This task also includes identifica­tion of needs for fatigue monitoring. The most critical ones are the high stresses in the body and sealing block of the main circulating pumps. These, however, could be managed via focused non-destructive examina­tion programmes.

Analysis of thermal ageing of Class 1 and 2 components focuses on components manufactured from 15Ch2MFA, 22K, 08Ch18N9TL cast stainless steel materials and also on welds (Sv04Ch19H11M3, EA400/10T, Sv10ChMFT, IONI 13/55) which are sensitive to thermal embrittlement. Significant changes of material properties due to thermal embrittle­ment are to be expected above 220°C operational temperature in case of ferrit-pearlit materials or cast stainless steel. Only a few components match these conditions at the Paks NPP. According to fatigue analyses performed, there are no cases where crack propagation due to fatigue might be expected. The analysis performed for the main gate valve cast stainless steel body shows that crack propagation should not be expected even if the J-R curve for C8 steel is changing due to embrittlement and a crack is postulated.

For analysis of thermal stratification for Class 1 and 2 pipelines, a mea­suring system was operated at the Paks NPP Unit 1 pressurizer surge line in 2000-2001. Assessment of the measured data shows significant thermal stratification (110°C), which moved periodically from the pressurizer to the hot leg. This temperature swing was maintained by the swing of water level control in the pressurizer during the heat-up and cool-down. During normal operation, the temperature differences were decreased to a negligible level. A similar temperature monitoring system has been operating on both legs of the surge line at Unit 3 since 2007. Evaluation of the measured data and the subsequent fatigue analysis justify LTO for the pressurizer surge lines. Other pipelines have also been identified where thermal stratification might occur. These are the pipelines connecting coolant cleaning system No 1 to the pri­mary system; the pipeline of the passive emergency core cooling system and the feed-water system pipeline and also the auxiliary emergency feed-water pipelines. Experience gained at other VVER-440/213 plants (Mochovce and Dukovany NPP) has been taken into account in identifying the pipelines of interest. Implementation of monitoring programmes is ongoing for these pipelines with temperature and displacement measurements.

High-cycle fatigue analysis of flow-induced vibration of internal struc­tures of the steam generator tubes shows that the flow-induced vibration of the heat-exchange tubes does not cause significant stresses compared to those from operational loads. Taking into account 60 years of operation and 108% of reactor thermal power, the CUF is equal to 0.027 due to vibration even if a pipe wall thinning of 50% is assumed.

Analysis of the corrosion of piping wall must question whether the erosion-corrosion allowance applied in the design provides sufficient mar­gin for 50+10 years of operation. Only a few cases are expected where the existing corrosion-erosion monitoring programme using COMSY software will have to be extended.

In analysing for material property change of the steam generator tubes, the main finding of the study is that the thermal ageing of 08H18N10T material used for heat-exchange tubes is negligible at operating tempera­tures ~290°C. Results of laboratory tests show that there is no change in the fatigue crack propagation rate due to LTO at 288°C (NPO Hidropress, 2007). An operational time of 60 years is justified in this respect.

Various reports of incomplete control rod cluster assembly insertion trig­gered investigations to identify causes of the sticking problem. The root cause was understood to be excessive deformation (bowing) of fuel assem­blies. When the bowing exceeds the limit, it increases friction between the control rod and the guide thimble and can result in the breaking of the con­trol rod cluster assembly.72 The mechanism of bow, though not clearly under­stood yet, is believed to be caused by creep affecting the overall assembly and guide thimble. If the fuel assembly experiences a flux gradient, the tubes at the lower flux side will grow less than those in high flux side causing the fuel rod to bow. Provision is made in the design, to accommodate an increase in length of the fuel rod (due to creep and growth) on either end and any restriction in the free movement leads to bowing of the rod. Rods under cold-worked stress-relieved (CWSR) conditions show more elonga­tion than those under fully recrystallized conditions.72

The VVER-440/213 reactor pressure vessel

The design of the VVER-440 RPV is rather specific: the relatively small RPV diameter has to allow its transportation on rails. As a consequence of its limited diameter, the water gap between the RPV and the core is small, so the fast neutron flux (E>0.5 MeV) on the RPV is rather high at 1015 m-2s-1 and the RPV base material should therefore be more resistant to irradia­tion embrittlement. The RPV is assembled from forged rings without lon­gitudinal welds. The coolant from the low-pressure emergency core cooling systems and hydro-accumulators is directly injected into the RPV, and from the high-pressure system into the cold leg of the loops. The inlet and outlet nozzles of the loops are separated on different levels. The penetrations for the instrumentation for core control are on the RPV head.

The ferritic steel reactor pressure vessel is clad internally with austen­itic stainless steel. The RPVs are made from low alloy steel (15Cr2MVA; at Loviisa NPP 12Cr2MFA) and the circumferential submerged arc weld­ing was made using Sv-10CrMoVTi wire. The RPV was covered inter­nally by a welded clad of two stainless steel layers. The inner layer is a non-stabilized stainless steel (Sv-07Cr25Ni13, similar to AISI 309) and that, when in contact with the coolant, is a niobium stabilized stainless steel (Sv-08Cr19Ni10Mn2Nb; Sv-07Cr19Ni10Nb at Loviisa; both equivalent to

AISI 347). Components of the primary circuit in contact with the primary coolant, other than the RPV, are also made of austenitic stainless steel, that is the piping of the primary loop, the main circulating pumps, gate valves and the emergency and auxiliary systems pipework.

From the point of view of longer-term operation, the main deficiency of VVER-440/230 was the high irradiation exposure of the reactor pressure vessel wall by fast neutrons, and the relatively quick embrittlement of the RPV material. The issue had been aggravated by the lack of a proper RPV surveillance programme at these plants. Several attempts have been made to assess the embrittlement of the base and weld material of those RPVs. For the first generation RPVs, essential data for RPV materials were absent, for example transition temperature, concentration of copper and phosphorus; the archive metal of the RPVs was not available. The phosphorus and cop­per contents in the welds of VVER-440/230 are in the range 0.030-0.048% and 0.10-0.18%, respectively. In the case of VVER-440/213, the same con­centrations are in the range 0.010-0.028% for P and 0.03-0.18% for Cu (Brumovsky et al. 2005; Vasiliev and Kopiev, 2007).

Reactor pressure vessel surveillance programmes became obligatory in all VVER plants that had been commissioned after Units 1 and 2 at Loviisa. Proper RPV surveillance programmes have been implemented at VVER-440/213 plants outside of the former Soviet Union from the com­mencement of plant operation.

An ‘ Extended Surveillance Specimen Programme ’ was prepared with the objective of validating the results of the standard programme (Kupca, 2006). It aimed to increase the accuracy of the neutron fluence measure­ment, make a substantial improvement in the determination of the actual temperature of irradiation, fix the orientation of RPV samples to the centre of the reactor core, minimize the differences in neutron dose between the Charpy-V notch and crack-opening-displacement specimens and evaluate any dose-rate effects. For Units 1 and 2 of the Mochovce NPP, a completely new surveillance programme was prepared, based on the philosophy that the results of the programme must be available during the entire service life of the NPP. The new, advanced surveillance programme deals with the irradiation embrittlement of both the weld area heat affected zone and the austenitic stainless steel cladding of the RPV, which were not previously evaluated in surveillance programmes.

Several measures were implemented for the resolution of the RPV embrittlement issue: [11]

• Decreasing the stressors, for example, heating up the water in the emer­gency core cooling system (ECCS) to lessen thermal shock in a pressur­ized thermal shock (PTS) situation; steam-line isolation; system solu­tions interlocks.

• Introduction of volumetric non-destructive testing for in-service inspection.

Annealing of RPV has been implemented at Loviisa NPP and Kola NPP (also at the shut down plant Bochunice V1). Annealing in the case of the VVER-440 reactor vessel weld was performed at a temperature of 475±15°C and the holding time was 150 hours. Assessment of annealing effectiveness (level of properties recovering after annealing), determination of re-irradiation re-embrittlement rates after annealing, and the behaviour of VVER-440 weld materials, showed the real possibility of recovering RPV toughness properties of irradiated VVER-440 RPV materials. Measures were also taken to improve the knowledge of the vessel material by ves­sel sampling. A more detailed description of the RPV neutron irradiation embrittlement issue is provided by Erak et al. (2007), for example. Based on the results of the VVER-440/213 plants, annealing of the RPV has been implemented at Rivne NPP.

In order to determine the time limit of operation of the RPV, it is neces­sary to consider and analyse the neutron irradiation damage, thermal age­ing and low-cycle fatigue in decreasing the fracture toughness of the RPV materials.

Pressurized thermal shock (PTS) is the most critical lifetime lim­iting event for the RPV. Since the PTS screening requirement (pressure-temperature-loading limits) is the lifetime limiting process for the RPV of VVERs, the methodology of PTS evaluation has to be established in the national regulations. This will take into account the applicable best practices, features of the RPV and the thermal-hydraulic peculiarities of the VVERs. The assumptions of renewed PTS analyses have been confirmed with mixing tests. International research projects supported the effort of VVER plants in the evaluation of PTS for the RPV (IAEA, 2005).

The results of the PTS calculations, based on the analysis of postulated embedded flaws, endorse the possibility of 50 years of operation for all of the units, without annealing of the 5/6 welds. At the Paks NPP, the assump­tion of the embedded postulated crack (under-cladding semi elliptical type) was justified by the results of qualified in-service inspections, which followed the procedure of European Network for Inspection Qualification (ENIQ). Two types of inspection were applied to the full cladding area: (1) ultra­sonic inspections from the inner surface and (2) Eddy current inspection, overlapping the first 5 mm thickness of the RPV inner-wall area. There is

no generic need for the heating up of the emergency core cooling water. It was introduced as an example at Rivne NPP in Ukraine, however it seems unnecessary at Paks NPP in Hungary.

The critical locations when considering fatigue are the welds of the inner tubes of the control rod drive nozzles.

Under current operating exposure times imposed by the five weight percent U-235 enrichment limit, the behavior of zirconium based fuel cladding alloys is reasonably understood on an empirical, and even somewhat phenomeno­logical, basis. This understanding is based on predictions of the oxide layer and hydride content of the cladding, both of which affect its ability to with­stand stresses due to normal and accident conditions. This does not mean that there are not unknown issues with the current zirconium alloy systems. Prediction of the cladding resistance to RIAs is not well understood; here large amounts of heat are almost instantaneously imposed on the fuel which causes rapid expansion of the pellets into the cladding resulting in clad­ding failure. As the exposure of the fuel increases and these rapid expan­sion effects become more pronounced and less well understood, the nuclear fuel manufacturers are left with the choice of understanding the mechanical behavior of highly exposed (and oxidized and hydrided) cladding to rapid stresses imposed by the fuel either on an empirical or phenomenological basis. The empirical basis used so far relies on the collection of immense amounts of data covering almost any potential event during fuel operation, a very time consuming and expensive approach. The phenomenological approach would be much better, but there is practically no such basis for understanding except at the very rudimentary level. Previous attempts to model and predict these effects have been unsuccessful due to the com­plexity of the interactions between the models of the various effects which lead to severe non-convergence problems. Major programs are currently underway funded by the U. S. Department of Energy (DOE) and industry to try to understand and model fuel behavior on a phenomenological basis using improved convergence algorithms (for instance the Consortium for Advanced Simulation of Light Water Reactors (CASL) and the Nuclear Energy Advanced Modeling and Simulation (NEAMS) Program).

Besides operating under extended burnup, the known operating limits of nuclear fuel cladding are being exceeded by high thermal duty condi­tions. This arises from the desire of nuclear utilities to get the most electrical generating capacity out of their current plants while still using a minimal amount of fuel. These higher thermal operating limits (in kilowatts per meter) increase the temperature of the cladding which, in turn, increases the build-up of crud on the surfaces, further increasing the cladding temper­ature and its oxidation rate and hydride content. Higher thermal rates also increase fretting issues between the fuel rods and their support structures causing unintended wear and fuel cladding failures. The added vibration is due not only to increased stress due to flow variations, but also to relaxation of the grid structures that support the fuel rods. In addition, there is a trend to move to slightly higher pH values by adding more lithium to the primary coolant. These higher lithium levels affect the corrosion rate of the fuel cladding though they also seem to inhibit stress corrosion cracking (SCC) of steam generator alloys in the primary circuit (EPRI, 2012).

Besides the more recent concern with RIA-type accidents and the ability to store used nuclear fuel for long times under both wet (spent fuel stor­age pool) and dry (spent fuel storage casks and perhaps internment in a final repository) conditions, there is the continuing concern of how fuel which has experienced high burnups will respond to such classic accident scenarios as large and small break loss of coolant accidents (LOCAs) due to loss in ductility and strength. Current knowledge of both oxidation layer thickness and hydride content is gained entirely through empirical testing. Development of new alloys is based entirely on an empirical approach with very little theoretical guidance.

Areas of research for zirconium (and any other metal) alloy claddings include the phenomenological understanding of: [22]

• The effect of oxide level and hydride content on the mechanical behav­ior of zirconium alloys.

• The effect of rapid stress levels on the mechanical behavior of zirconium alloys.

• Effects of long-term wet and dry storage as well as environmental condi­tions in any potential disposal site on the integrity of the zirconium alloy cladding.

• Effect of mechanical stresses induced by flow vibration on the fuel rod cladding as well as the grid support structures.

• Effect of flow rate on corrosion and erosion of both the fuel and the fuel structure.

• Modeling of these effects to allow performance prediction in extended burnup and in transient operating conditions.

The development of AMPs has to begin with the identification of the ageing mechanisms, critical locations and effect of ageing on the intended safety function. When an AMP is developed for a complex structure or compo­nent, like the reactor or steam generator, several mechanisms and critical locations can be identified. The material, conditions and stressors are con­sidered at this step of the AMP development. Examples for the mechanisms are listed in the Table 8.4.

As a matter of fact, the structuring of the AMPs together with the identifi­cation of the commodities is not independent from the identification of ageing mechanisms. For example, a commodity group can be defined as follows, see Table 8.7 : Safety Class 3 + Piping and pipe elements + working in prepared water (e. g. feed-water line) + carbon steel. From experience, the dominating ageing mechanism of this group is flow-accelerated corrosion (FAC), a deg­radation process resulting in wall thinning of piping, vessels, heat exchanger and other equipment made of carbon and low alloy steel. This degradation mechanism of the identified commodity group should be addressed by proper AMP, which can be developed for example via application of the COMSY system (Zander, Nopper, Roessner, 2007) used by several VVER operators.

The premature fracture of materials under fluctuating load (stress/strain/ temperature) is known as fatigue. Fatigue is a sudden failure exhibiting no overall ductility in the component and is known to be the cause in 90% of the total failures of structures. During each fatigue cycle the material absorbs part of the applied energy and, when the accumulated strain energy reaches the value of the surface energy of the material (in that environ­ment), a surface forms (i. e. a crack appears). The accumulation of strain energy is facilitated by the presence of a notch or scratch and the surface energy is the minimum for the exposed crack than the embedded one. Often, the fracture surface is perpendicular to the direction of the applied stress and a compressive residual stress is beneficial in delaying the fatigue failure. Fatigue life is represented by a plot of applied stress (S) against the number of cycles to failure (Nf) known as the S-N curves. Figure 1.6 depicts the S-N curves for various metals8 and we note that ferrous metals exhibit a distinct ‘endurance’ limit below which fatigue failures do not occur whereas non­ferrous metals do not seem to exhibit such a limit, albeit the slope of the S-N curve decreases at very high cycles. The stress axis can also be either the stress amplitude (omax — omin)/2, the stress range (omax — omin) or mean stress (omax + omin)/2 and it is generally seen that the fatigue life depends weakly on the R ratio (R = omin/omax), where omax and omin represent the maximum and minimum stresses, respectively. Depending on the number of cycles to fail­ure the fatigue curve is classified as low cycle fatigue (LCF) and high cycle

fatigue (HCF) regions, corresponding to the plastic and elastic deforma­tion ranges, respectively. LCF is characterized by macroscopic cyclic plastic strains and is generally limited to less than 104 cycles. LCF is controlled by the ductility and HCF by the strength of the material, and thus, cold-work and radiation hardening (both of which result in reduced ductility) result in decreased fatigue life in the LCF range while being beneficial in the HCF range, especially at low stresses/strains. Figure 1.7 shows a typical fatigue life plot as strain range (Де) against number of failure cycles (Nf) along with the corresponding stress-strain loops (broad in LCF and narrow in HCF). In the high cycle region corresponding to HCF, the Basquin equation relates the applied stress (Да) to the number of cycles:

Nf (Да)р = C or in terms of strains Nf (ЕДе)р = C, [1.10]

where C and p are material constants. LCF with inelastic strains is often described by the Coffin-Manson equation

Де = 2A(2Nf)c [1.11]

where A, a function of the ductility, and c (-0.5 to -0.7) are material con­stants and Nf is the number of stress/strain reversals. The Coffin-Manson equation is seen to be valid for many materials over a broad range of tem­perature, environment, stress history and microstructural conditions. The complete fatigue curve can be described by combining the LCF and the HCF

formulations by either the universal slopes equation (Equation [1.12a]) or the characteristic slopes equation (Equation [1.12b]):

where Su is ultimate tensile strength, ef is true fracture strain, af true frac­ture stress, and b and c are material constants. In terms of the characteristic slopes (Equation [1.12b]) the value of fatigue life at which the transition from low cycle (plastic) to high cycle (elastic) occurs is given by

[1.12c ]

Fatigue crack growth rate (FCGR, da/dN) is determined by measuring the extension of a pre-crack using visual, potential drop, unloading compliance or other techniques over the elapsed number of load cycles from stress con­trol tests conducted on either compact tension (CT) or three-point bend specimens and is related to the range of stress intensity factor (AK). Typical crack extension curves at two different starting stress ranges (Ac) versus number of cycles are shown in Fig. 1.8a and the slopes of the curve yields da/dN. The plot on logarithmic scale of (da/dN) versus AK (Fig. 1.8b) clearly reveals three stages. Stage I is associated with crack blunting with very little crack growth, while crack growth in stage II can be related using Paris’ law:

[1.13 ]

where p is the Paris parameter/constant with values ranging from 2 to 4; this covers the majority of the crack growth event before entering the final stage (stage III) where plastic fracture occurs as crack length reaches a crit­ical value (af) corresponding to the plane strain fracture toughness (KIC) value:

In Equation [1.14], Y is the geometric factor which is a function of a/w (a is crack length and w is specimen width).

Stage I corresponds to formation of a fine crack from surface defects (such as scratches, key ways, stress concentrations) with slow initial propagation along specific crystallographic directions covering few grains before the growth enters stage II where the crack propagates at a relatively faster rate and on a plane perpendicular to the loading direction. In general, persistent slip bands (PSBs), beach marks and fatigue striations (Fig. 1.9a and 1.9b)

are characteristics of stage II crack propagation and the separation between striations depends on the stress range and frequency of loading. The total number of cycles to failure can be estimated as follows from Equations 1.15 and 1.16:

and

Another important aspect of considering the crack growth versus AK is to examine the effects of superimposed environment such as corrosion and radiation. The variation of da/dN with AK in these cases would shift the threshold stress intensity range to lower values and the critical crack length at fracture would be indicated by KISCC instead of KIC.

In strain controlled fatigue tests for life evaluation, it may be noted that the cyclic stress-strain curve leads to a hysteresis loop as depicted in Fig. 1.10a where O-A-B is the initial loading curve11 and, on unloading, the yielding occurs at lower stress (point C as compared to A) which is known as the Bauschinger effect. The material may undergo cyclic hardening or softening; in rare cases it remains stable (Fig. 1.10b). This behaviour depends on the initial metallurgical condition of the material. According to Fig. 1.10b, as the number of cycles increases cyclic hardening leads to decreasing peak strains while the peak strains increase in the case of cyclic softening. In general, the hysteresis loop stabilizes after about 100 cycles and the stress-strain curve obtained from cyclic loading will be different from that of monotonic load­ing (Fig. 1.10c), but the stress-strain follows a power law relationship similar to that in monotonic loading (Equation [1.3]):

Ac = K'(Ae)n

where the cyclic hardening coefficient n’ ranges from 0.1 to 0.2 for many metals and is given by the ratio of the parameters (b/c) (Equation [1.12b]). In some cases fatigue ratchetting occurs resulting in an increase in strain as a function of time when tested under a constant strain range (Fig. 1.11); this is often referred to as cyclic creep. i2 In a stress controlled test with non-zero mean stress, the shift in the hysteresis loop along the strain axis, as depicted in Fig. i. ii, is attributed to thermally activated dislocation movement at stresses well below the yield stress and/or due to dislocation pile up result­ing in stress enhancement. Fatigue ratchetting may also occur in the pres­ence of residual stress and in cases where microstructural inhomogeneities exist such as in welded joints.

In real situations stresses change at random frequencies and, in gen­eral, the percentage of life consumed in one cyclic loading depends on

1.10 ( a) Cyclic stress-strain curve illustrating hysteresis loop.11

(b) Hysteresis loops during cyclic hardening and cyclic softening.12

(c)

Comparison of cyclic stress-strain curve for cyclic hardening and stress-strain curve under monotonic loading.11

1.11 An example of ratcheting fatigue.12

the magnitude of stress in subsequent cycles. However, the linear cumu­lative damage rule, known as Miner’s rule, assumes that the total life of a component can be estimated by adding up the life fraction consumed by each of the loading cycles. If Nfi is the number of cycles to failure at the ith cyclic loading and Nt is the number of cycles experienced by the structure then

although Miner’s rule is too simplistic and fails to predict the life when notches are present. Further, it fails to predict the life when mean stress and temperature are high or cyclic frequency is low where creep deformation dominates over fatigue loading. In such situations a better approximation is given by combining Robinson’s rule for creep fracture with Minor’s rule;

[1.19]

where (fi) and fracture time (t) corresponding to the ith creep conditions.

It turns out that many materials exhibit deviations from this linear addi­tion depending on whether it is cyclically hardening or softening.13 In par­ticular the predictions tend to be highly non-conservative for cyclically softening materials.

Fatigue strength or life of structures can be improved by reducing the mean positive stress, through appropriate design with no stress raisers and by surface finish and modifications. In particular, case hardening by carbu­rizing and nitriding as well as shot-peening, which increase surface residual compressive stresses, result in distinct improvements in fatigue life. In com­parison to pure metals, solid solution has been found to improve fatigue strength. Other factors such as interstitials inducing strain ageing could also improve fatigue life.

Environmental effects on creep-fatigue are quite complex and each case needs to be considered separately. While Equation [1.19] gives an approxi­mate assessment, the mechanistic explanations of high-temperature fatigue effects are corrosion — or creep-related. Coffin considered the time dependent fatigue to be essentially SCC and formulated frequency-modified fatigue life-time correlations for crack initiation and propagation.12 Manson pro­posed a plasticity oriented fatigue model using a strain-range partitioning method.13 Fatigue crack growth assisted by creep cavitation at grain bound­aries was considered by Majumdar and Maiya14 to model high-temperature fatigue crack growth.

As described in Section 1.2.1, exposure of materials and structures to high energy neutrons leads to the creation of microscopic defects such as vacan­cies, interstitials, Frenkel defects, dislocations and faulted loops, as well as voids and cavities. Figure 1.12a depicts voids and precipitates in irradiated stainless steel1 5 while large Frank loops are shown in Fig. 1.12b. 1 6 Similar faulted Frank loops are noted in irradiated aluminium and copper as well as iron (Fig. 1.12c).17

Materials undergo many changes on exposure to neutron radiation: defect concentration increases, neutron transmutation occurs, chemical reactivity changes (generally gets enhanced), diffusion of the elements increases and new phases (both equilibrium and non-equilibrium) form. The extent of change in properties is, in general, proportional to radiation flux, particle energy and irradiation time, while it decreases with an increase in irradia­tion temperature. The creation of voids, cavities and depleted zones leads to decreased density of the material with a corresponding increase in vol­ume known as radiation swelling. Increased defect concentration leads to increased electrical resistivity and decreased thermal conductivity while magnetic susceptibility decreases. The threshold neutron fluence or dpa that leads to extensive degradation in a material depends on the crystal structure and nature of atomic bonding — semiconductors and polymers degrade at much lower neutron fluences compared to ceramics and metals. The reader is referred to various monographs on nuclear materials and radiation effects for more details.18 These defects result in hardening and embrittlement of the material with an increase in strength and accompanying decrease in ductility commonly referred to as radiation hardening and radiation embrittlement; strain hardening in the material decreases accompanied by a decreased uni­form elongation and an increase in DBTT (or RTndt), which decreases the fracture toughness. The increased defect density enhances the diffusivity in the material which in turn increases the creep rates and reduces the rupture time. These various phenomena will be discussed in detail in the following sections.

In this section we look at applied management practice around the world.

74.1 Degradation of reactor vessels

There have been no cases of ageing degradation of reactor vessels. Thermal annealing at high temperature (475 ± 15°C) for 100-150 h has been applied to plants operating in Russia since 1987 as a preventative measure (Badanin, 1989; Cole and Friderichs, 1991).

74.2 Degradation of reactor internals

There is a report that a guide tube support pin made of Alloy 750 was dam­aged in Mihama Nuclear plant in Japan in 1978. The damaged pin was mov­ing around as a foreign body and it was discovered in the steam generator chamber. It was determined that the damage was from high stress in pri­mary water at high temperature. After this incident, research studies were conducted over many years to understand the damage mechanisms. It was found that Alloy 750 was a material sensitive to PWSCC, and that sensitivity increased largely depending on heat treatments. Since it was replaced with a new cold worked material (CW316 SS), no additional cracks have been found.

Regarding baffle former bolts, some cracks were found in an old French power plant (Fessenheim and Bugey) in 1988. Some bolts had 10-25 dpa fluence after being in operation for 10-20 years. They were made of a 316 cold worked stainless steel and showed intergranular stress corrosion crack­ing (IGSCC). The ageing mechanism was assumed to be IASCC. The cracks had spread from the shank zone of the head to the lower part of the head. The material was found to be hardened and radiation-induced segregation was found in the grain boundary. According to the hardness profile mea­surement, it was approximately 5-10 dpa, and there was no evidence of swelling. The bolts with cracks were detected at rows 2 and 3 from the lower part of the nuclear reactor where a considerable amount of neutron radi­ation had accumulated. According to the report, until that point, cracks in the baffle former bolts had been found in the ‘down-flow’ design in which inlet coolant flows downward (Gerard, 2009). From 1989 to 1993, the flow of the coolant in the nuclear reactor of the CP0 (name of French PWR 900 MW pre-series units) plant was changed to up-flow and, between 2000 and 2003, one third of the bolts were replaced. The cracking rate of baffle bolts increases slowly depending on dose. Based on the information on all the baffle bolts researched during in-service inspection (ISI) for all CPO units, the dose threshold is estimated to be approximately 3-4 dpa. The number of cracked bolts increases slightly at higher doses.

In Japan, two kinds of approach have been applied to the PWR plants since 1998 in light of the baffle former bolt issues experienced in France and also in the United States. The first approach was to replace the baffle former bolts. From 2001 to 2002, type 347 stainless steel bolts were replaced with cold worked 316CW stainless steel bolts in Mihama Unit 1 and 2. The second approach was to replace the internal structure of the nuclear reac­tor. In this case, the lower zone including the baffle former bolts and the upper zone were replaced. Since 2004, the internals of the nuclear reactors in three PWR plants have been replaced entirely. The measures described above were applied to the baffle former bolts in a 2-loop PWR plant which was built in the early 1970s.

A research programme aimed at preventing defects of the internals of nuclear reactors has been ongoing since 2000. For example, there has been research on IASCC of austenitic stainless steel used in PWR inter­nals such as in baffle former bolts. Valuable data was collected through this research, which showed that IASCC initiation is closely related to stress and exposed neutron fluence. In other words, the threshold stress which deter­mines IASCC occurrence is dependent on the amount of neutron fluence. The threshold stress value tends to decrease as neutron fluence increases.

Based on the data and experiments, it was decided that a guideline would be published with detailed interpretation. Japan published a guideline on management activities such as inspection of the baffle former bolts of actual power plants in 2002. This guideline will be revised in the future to reflect up-to-date knowledge obtained through international collaboration.

In Korea in 2007, defects were found in the control rod guide tubes made of Inconel X-750 in the internals of a nuclear reactor. The control rod guide tubes were replaced with CW 316 stainless steel tubes. Since then, no cracks have been found in the baffle former bolts (Hwang et al, 2010).

In the United States, the Westinghouse Owners Group (WOG) has inspected cracks in baffle former bolts from PWRs in other countries. They have also provided information on activities planned for Westinghouse power plants with potential cracking. The WOG has clearly demonstrated that destruction of minor bolts would not have a serious impact on safety because a number of baffle former bolts would still support the structure. The WOG activities are as follows:

• Development of analytical methods and acceptance criteria for bolt analysis.

• Performance of risk-informed evaluations.

The Nuclear Energy Institute (NEI) which consists of the Materials Technical Advisory Group (MTAG) of the United States is formed of energy company representatives. The MTAG has received support from EPRI to prepare a guideline for In-Service Inspection (ISI) of RVI equipment which has signif­icant impact on continued and safe operation of power plants. In preparing an inspection guideline, the damage to equipment inside the nuclear reactor by inspection, fatigue, abrasion and corrosion were considered. Recently, Westinghouse and AREVA have published a report using screening based on various significant damage mechanisms as a part of EPRI MRP on reac­tor internals.