Elevated temperatures cause the polymers in the cable insulation to degrade through loss of elongation, embrittlement, and cracking (U. S. NRC, 2001). Cable polymers are primarily degraded by thermal oxidation in the pres­ence of oxygen, accelerating with increases in temperature as defined by the modified Arrhenius equation (IAEA, 2011):

к = A exp(-EA/RT) [6.1]

where EA is the activation energy, A is the frequency factor, and R is a con­stant. Temperature is the most important ageing stressor for most cables in a light water reactor (IAEA, 2011).

As a result of internal ohmic self-heating, power cables age uniquely, depending on how long the cable carries electric current, which current it carries, and the specific configuration of the cable installation itself. Treeing (the appearance of small tree-shaped cracks in the insulation caused by electrochemical reactions) and the loss of the dielectric properties of cable insulation are characteristic results of power cable ageing (IAEA, 2011; U. S. NRC, 2001 ).

Exposure to moisture can also degrade cables that have been installed directly in the ground or in ducts or conduits where water has access. ‘Wetting ’ describes conditions in which a cable is exposed to moisture or high humidity for extended periods of time, including limited periods of complete submergence. Submersion describes conditions when the cable is completely submerged in water for extended periods. So long as the insula­tion and outer jacket are not damaged, intermittent wetting will not damage most cables, but extended submersion is beyond the qualified operating conditions for most cables (U. S. NRC, 2010a).

Moisture can cause water treeing where voids or contamination are pre­sent in the cable. This combination of water and electrical stress degrades the insulation’s dielectric properties (U. S. NRC, 2001). In fact, the U. S. NRC (NUREG 6704) identified wetting as the primary ageing-related cause of failure (specifically, short circuit) for medium-voltage cables, in particu­lar the insulation (U. S. NRC, 2001). Such failure could allow currents and voltages to spread into the adjacent power distribution system, potentially causing other degraded power cables to fail too (U. S. NRC, 2010b). For this reason, cables in hard-to-access underground ducts and conduits, covered trenches, bunkers, and manhole vaults are the subject of special concern (U. S. NRC, 2010b).

Both power cables and I&C cables are directly affected by mechanical stress including bending, abrasion, cutting, contact, deformation, and per­foration, as a result of installation and maintenance, for example. Cables connected to vibrating machines are also subjected to stress, leading to chafing, cutting, or cracking of the cable insulation material (AMS Corp., 2011). Cable jacket and insulation material as well as cable conductors can be damaged by electromechanical forces caused by high levels of short cir­cuit current passing through a power cable (U. S. NRC, 2010a).

Radiation is another significant cause of cable degradation. During nor­mal operation, gamma and neutron radiation cause oxidative degradation in increasing (nonlinear) relation to the radiation dosage absorbed by the cable. During accidents, beta radiation may also affect cables unprotected by a conduit (IAEA, 2011).

The time dependence of plastic strain is described by plots of strain against time, also known as creep curves. A typical creep curve is shown in Fig. 3.1,3 and consists of three different regions: the primary, secondary and tertiary creep regions. Usually the primary creep region commences only after the material has experienced an instantaneous strain, e0 which is a result of sudden loading of the material and corresponds to period ‘a’ observed by Andrade.1 The instantaneous strain is composed of elastic (recoverable on release of load), anelastic (recovers with time) and plastic (non-recover­able) components. Though the applied stresses for creep are smaller than

the yield strength of the material, the instantaneous strain is composed of a plastic strain component.

The primary creep (also known as transient creep) region, as the name suggests, describes the first or initial stage of creep deformation and corre­sponds to period ‘b’ in Andrade’s work.1 Such a region is characterized by a strain rate decreasing with time. The decrease in strain rate continues until the secondary stage (also known as steady-state creep) is attained. In the sec­ondary creep region (period ‘c’ in Andrade’s study) the strain rate of defor­mation remains constant. This is evident in Fig. 3.1 with the secondary creep region described by a straight line indicating a constant slope. The secondary stage strain rate is the minimum strain rate of the creep curve. The useful creep life of most engineering materials is generally estimated from second­ary stage creep strain rate values. However such a methodology might not be applicable for materials which have a large primary creep region or where the tertiary creep region completely dominates the primary and secondary creep stages. The tertiary creep region is the last stage of creep deformation and concludes with the failure of the material. In the tertiary creep regime (as identified by Hanson and Wheeler2) the material undergoes deformation at very high strain rates. The tertiary stage of the creep curve usually occurs over significantly smaller time periods in comparison to the primary and the secondary stage, and is often regarded as ‘fracture’ mode.

Since dry storage of spent fuel (SF) is gaining importance, it is necessary to assure the fuel rod integrity during interim storage for relatively long times. A clad with high burnup is likely to contain large amount of hydrogen (1000 ppm). The initial level of hydrogen is kept very low in order to reduce the in-reactor hydride-related problems and hydrogen pickup during service is controlled by employing new alloys.129 Short term creep tests in Zircaloy-4 reveal that after a burnup of 64 MWd/kgU, hydrogen did not pose any dele­terious effect and the material possessed sufficiently good ductility.130 But it is interesting to note that hydrogen affects the creep rate in zirconium alloys differently as atomic hydrogen and as hydride.

In Zircaloy-4, the creep rate was reported to depend on the condition of the material — whether in cold-worked stress-relieved (CWSR) or annealed con­dition; CWSR alloy shows a significant strengthening on addition of hydrogen. The reason for this behavior is attributed to hydrogen influencing the strain hardening rate and static recovery of the material. Biaxial tests in Zircaloy-4 show that the presence of hydrides in the cladding will help to prevent the cold work microstructure from being annealed out of dislocations and thereby lower creep rates are maintained in the spent fuel cladding.131 The same alloy in annealed condition shows a decrease in creep rate when hydrogen is in solu­tion and an increase when part of the hydrogen is precipitated as hydrides. This behavior is attributed to the reduction in the stacking fault energy of Zr caused by diffusion of hydrogen to the core of the screw dislocations and an increase in their mobility. On the other hand when hydrogen is present in the form of hydrides, it increases the matrix strength and reduces the creep rate. The creep rate of Zircaloy-4 at a temperature of 693K and a stress of 150 MPa, a mar­ginal increase in creep rate is noted for a hydrogen content of 200 wt. ppm. The increase in the creep rate is believed to be brought out by the reduction in the modulus value when hydrogen is added.132 In a Zr-2.5wt.%Nb alloy, the creep rate at 723K is reported to increase by 2-2.5 times for a hydrogen content of 160 wt. ppm and the stress exponent reduces from 2.41 to 1.59, indicating the change in creep mechanism (Fig. 3.30).133

The results above indicate that hydrogen in dissolved state increases the creep rate and this is pertinent to SF which remains in this temperature range (~573K) and has sufficient hydrogen dissolved in it.

PWR/VVER fuel assembly bowing may occur due to excessive guide tube (GT) growth that will result in larger holding down forces (Fig. 5.1) (Strasser et al, 2010a). The bowing is caused by the complex interaction of a variety of parameters that include the bowing of the skeleton assembly. The param­eters include:

• GT irradiation growth as a function of fluence and temperature, see Section 4.6.1 on irradiation growth.

• GT creep as a function of fluence and temperature, see Section 4.6.2 on irradiation creep.

• GT stiffness and buckling strength as a function of temperature.

• The effect of hydrogen pickup and irradiation on the GT properties, see Section 4.5 on corrosion of zirconium alloys.

• Hold-down force.

• Thermal expansion of the skeleton components, the core plate spacing and their interaction.

• Fuel rod/grid friction force and relaxation over time.

BWR fuel channel bowing was studied by Cantonwine et al. (2009). According to them, channel-control blade interference had been a
challenging issue over the previous 8 years for operating BWR plants where ~2-year cycles are normal and Zircaloy-2 is the standard channel material. The primary reason for this was the unaccounted channel dis­tortion caused by differential hydrogen across the channel resulting from shadow corrosion on the blade side (known as shadow corrosion-induced bow). Zircaloy-2 is particularly susceptible to this distortion mechanism because it has a high hydrogen pickup fraction (HPUF) that increases with exposure.

Several strategies have been developed to combat bow. As an interme­diate resolution to this issue Zircaloy-4 has been reintroduced because it is effectively resistant to shadow corrosion-induced bow and has similar irradiation growth and creep performance to Zircaloy-2. The one disadvan­tage of Zircaloy-4 is that it has less corrosion resistance than Zircaloy-2. However, based on the extensive experience with Zircaloy-4 channels both in the United States and Japan (plus processing improvements have been made specifically to enhance corrosion resistance), the corrosion perfor­mance of Zircaloy-4 is claimed to be adequate for channel applications. Other examples of global nuclear fuel (GNF) publications on channel bow are described by Mahmood et al. (2007) and Cantonwine et al. (2009). Other reasons for BWR fuel channel bowing are (Strasser et al, 2010a):

• Fast neutron flux gradients from a variety of causes including the flux gradient at the core periphery (see Fig. 5.2).

• Non-uniform metallurgical structure (e. g. texture difference between the two opposing channel sides leading to difference in irradiation growth rate) or composition.

• Non-uniform wall thickness.

• As-fabricated bow.

The bowing may result in difficulties in inserting the control rods (a safety issue) and/or in a decrease in thermal margins, the latter from two possible causes. First a departure from nucleate boiling (DNB) value: if the fuel rod surface heat flux becomes large enough, the water film adjacent to the fuel rod will convert into a steam film with a much lower thermal conductivity resulting in a rapid large increase in the fuel cladding temperature which, in turn, will accelerate the oxidation and embrittlement of the fuel cladding. The maximum heat flux at which the water is converted into a steam film is referred to as the DNB value. Second, a loss of coolant accident (LOCA) could, for example be caused by a coolant pipe break in the primary circula­tion system since larger water gaps between assemblies may exist in the core than is accounted for in the core nuclear design. To ensure that the LOCA licensing criteria are met, the fuel rod surface heat flux must be limited.

Core periphery

5.2 Schematics showing fuel outer channel bowing at core periphery due to large fast neutron fluence. Largest degree of bowing in BWRs occurs at the core periphery due to the flux profile. Also the type of FA bow seems to be very dependent on core location (Strasser et al., 2010a).

DOI: 10.1533/9780857097453.3.315

Abstract: This chapter discusses management strategies in terms of mitigation and repair techniques for degradation in pressurized water reactors (PWRs). We begin by introducing PWR materials management strategies followed by details of ageing and life management of the PWR components. International cooperation activities for ageing management are also included. Development and application of mitigation techniques for reactor pressure vessels, reactor internals, steam generators, pressurizer nozzles, control rod drive mechanisms, and secondary piping are described next. Mitigation and repair methods for degradation of PWR components include: material changes, isolation techniques, weld material changes, design changes, weld overlays, stress improvements, environment improvement, mechanical repair and component replacement or removal. Finally, cracking history of components around the world and countermeasures are introduced.

Key words: PWR, life management, reactor pressure vessel, reactor internals, steam generators, pressurizer nozzles, control rod drive mechanisms, primary and secondary pipings.

7.1 Introduction

In the case of PWR power plants, Korea, China and Europe continue to build nuclear power plants which show advanced performance, whereas the [8]