E. R. Hodgson

4.22.1 Introduction

It is envisaged that early in the twenty-first century ITER (International Thermonuclear Experimental Reactor) will come into operation, and it is expected that this intermediate ‘technology’ machine will help to bridge the gap between the present-day large ‘physics’ machines and the precommercial DEMO power reactor, thus paving the way for commercial fusion reactors to become available by the end of the century. Although this ‘next-step’ device will undoubtedly help to solve many of the problems, which still remain in the field of plasma confinement, it will also present additional operational and experi­mental difficulties not found in present-day machines. These problems are related to the expected radia­tion damage effects as a result of the intense radiation field from the ‘burning’ plasma. This ignited plasma will give rise to high-energy neutron and gamma fluxes, penetrating well beyond the first wall, from which one foresees a serious materials problem that has to be solved. In the initial physics phase of opera­tion of such a machine, it is the radiation flux, which will be of concern, whereas in the later technology phase, both flux and fluence will play important roles as fluence (dose)-dependent radiation damage builds up in the materials. For structural metallic materials, radiation damage in ITER is expected to be severe, although tolerable, only near to the first wall. How­ever, the problem facing the numerous insulating components is far more serious because of the neces­sity to maintain not only the mechanical, but also the far more sensitive physical properties intact. An addi­tional concern arises from the need to carry out inspection, maintenance, and repair remotely because of the neutron-induced activation of the machine. This ‘remote handling’ activity will employ machin­ery, which requires the use of numerous standard components ranging from simple wires, connectors, and motors, to optical components such as windows, lenses, and fibers, as well as electronic devices such as cameras and various sophisticated sensors. All these components use insulating materials. It is clear, there­fore, that we face a situation in which insulating materials will be required to operate under a radiation field, in a number of key systems from plasma heating and current drive (H&CD), to diagnostics, as well as remote handling maintenance systems. All these sys­tems directly affect not only the operation, but also the safety, control, and long-term reliability of the machine. Even for ITER, the performance of some potential insulating materials appears marginal. In the long term, beyond ITER, the solution of the materials problem will determine the viability of fusion power.

The radiation field will modify to some degree all of the important material physical and mechanical properties. Some of the induced changes will be flux dependent, while others will be modified by the total fluence. Clearly, the former flux-dependent pro­cesses will be of concern from the onset of operation of future next-step devices. The fluence-dependent effects on the other hand are the important para­meters affecting the component or material lifetime. The properties of concern which need to be consid­ered for the many applications include electrical resistance, dielectric loss, optical absorption, and emission, as well as thermal and mechanical proper­ties. Numerous papers have been published discuss­ing both general, and more recently, specific aspects of radiation damage in insulating materials for fusion applications, and those most relevant to the present chapter are included.1-26

In recent years, because of the acute lack of data for insulators and the recognition of their high sensi­tivity to radiation, most work has concentrated on the immediate needs for ITER. A comprehensive cera­mics irradiation program was established to investi­gate radiation effects on a wide range of materials for essentially all components foreseen for H&CD and diagnostics in ITER, and to find solutions for the problems which have been identified. A large number of relevant components and candidate materials have been, and are being, studied systematically at gradu­ally increasing radiation dose rates and doses, under increasingly realistic conditions. A considerable vol­ume of the work discussed here was carried out within the ITER framework during the CDA, EDA, and EDA extension (Conceptual and Engineering Design Activities 1992-2002) as specific tasks assigned to the various Home Teams (T26/28 and T246; EU, JA, RF, US; T252/445 and T492; EU, JA, RF).27,28 Since these last ITER tasks, no new coordinated tasks related to insulators have been formulated. However, despite the lack of an official framework in which to develop and assign further common tasks following the end of the ITER-EDA extension, col­laborative work has continued between the EU, JA, RF, and US Home Teams on both basic and applied aspects of radiation damage in insulator materials. This has resulted in considerable progress being made on the understanding of the pertinent effects of radiation on in-vessel components and materials in particular for diagnostic applications. Problems which have been addressed and for which irradiation testing has been performed include comparison of absorption and luminescence for different optical fibers and win­dow materials, RIEMF (radiation-induced electro­motive force) and related effects for MI (mineral insulated) cables and coils, alternative bolometers to the reference JET type gold on mica, hot filament pressure gauges, and electric field effects in aluminas.

One must however remember that ITER is only an intermediate ‘technology’ machine on the road to a precommercial power reactor. Such power reactors will face the same radiation flux problems as antici­pated in ITER, but the fluence problems will be far more severe. It is also important to note that the radiation flux and fluence levels will be different from one type of device to another depending on the design (e. g., in ITER and the Fusion Ignition Research Experiment (FIRE)26), and also on the spe­cific location within that device. Because of the gen­eral uncertainty in defining radiation levels, most radiation effects studies have taken this into account by providing where possible data as a function of dose rate (flux), dose (fluence), and irradiation tempera­ture. Although the task ahead is difficult, important advances are being made not only in the identifica­tion of potential problems and operational limita­tions, but also in the understanding of the relevant radiation effects, as well as materials selection and design accommodation to enable the materials lim­itations to be tolerated.

Following a brief introduction to the problem of radiation damage in both metals and insulators, the important aspect of simulating the operating envi­ronment for the component or material under exam­ination will be presented, with reference to present experimental procedures. The chapter will then con­centrate on the problems facing the use of insulators, with the radiation effects on the main physical prop­erties being discussed, concentrating in particular on the dielectric properties.

Application of nondestructive examination methods to NPP reinforced concrete structures presents chal­lenges: wall thicknesses can be in excess of 1 m; struc­tures often have increased steel reinforcement density with complex detailing; there can be a number of penetrations or cast-in-place items; accessibility may be limited because of the presence of liners or other components, harsh environments, or structures located below ground; experience with nondestruc­tive examinations of NPP concrete structures is somewhat limited; and methods utilized for the NPP structures are often based on equipment developed for other materials or technologies. Available methods are relatively good at identifying cracking, voids, and delaminations as well as indicating the relative quality ofconcrete. Methods for determining concrete prop­erties, however, generally are somewhat more quali­tative than quantitative because they tend to be indirect in that they often require the development ofcorrelation curves for relating a measured parame­ter (e. g., ultrasonic velocity or rebound number) to a property (e. g., concrete compressive strength). Infor­mation on identification and description of methods for determining the strength of concrete and evalua­tion of concrete structures is available.97-100 A practical guide related to nondestructive examination of con­crete, which not only identifies and describes the cap­abilities, limitations, and applications of the various methods that are available but also presents results from a number of examples, has been developed.101

The status of nondestructive examination meth­ods and priorities for its development with respect to examination and instrumentation and monitoring of concrete structures in nuclear plants was addressed by prior NEA/CSNIIAGE workshops.41,45,51 It was noted that although nondestructive examination techniques have been successfully used on a variety of reinforced and post-tensioned concrete structures, there has been somewhat limited experience in their use to evaluate typical NPP safety-related structures. With respect to these structures, three conditions exist where performing inspections or conduct of nondestructive examinations is not straightforward and requires development — inspection of thick — walled, heavily reinforced concrete sections, base — mats or foundations, and inaccessible portions of a containment metallic pressure boundary. Information summarizing the activities conducted addressing these conditions has been documented.1

Noninvasive techniques for characterization, inspection, and monitoring of thick-walled, heavily reinforced concrete sections to provide additional assurances of their continued structural integrity are desirable (e. g., as-built or current structural features determination, flaw detection and characterization, identification of honeycomb areas and embedded items, and location of voids adjacent to the liner). Methods that can be used to inspect the basemat without the requirement for removal of material and techniques that can detect and assess corrosion are of particular interest. Acoustic (e. g., ultrasonic pulse velocity, spectral analysis of surface waves, impact echo, and acoustic tomography), radar, and radiography appear to have potential for application to thick-walled, heavily reinforced concrete struc­tures in NPPs; however, additional development is required. The most commonly used type of founda­tion for both concrete and steel NPP containments is a mat foundation, which is a flat, thick slab support­ing the containment, its interior structures, and any shield building surrounding the containment.10 As such, the concrete foundation elements of NPPs are typically either partially or totally inaccessible for inspection unless adjacent soil, coatings, waterproof materials, or portions of neighboring components or structures are removed. As a result, indirect methods related to environmental qualification are often utilized to indicate the potential for degradation of the NPP concrete foundations.20 This is generally done through an evaluation of the surrounding medium (e. g., air, soil, humidity, groundwater, or cool­ing water). Methods employed are based primarily on chemical evaluations to assess the presence and con­centration of potentially aggressive ions (e. g., sulfates or chlorides). In addition to an assessment of the aggres­siveness of the surrounding environment, the CFR requires a complete description of the effects ofground — water levels and other hydrodynamic effects on the design bases of the plant foundations and other struc­tures, systems, and components important to safety.104

Inspection of inaccessible portions of metallic pressure boundary components ofNPP containments (e. g., fully embedded or inaccessible containment shell or liner portions, the sand pocket region in Mark I and II drywells, and portions of the shell obscured by obstacles such as platforms or floors) requires special attention. Embedded metallic por­tions of the containment pressure boundary may be subjected to corrosion resulting from groundwater permeation through the concrete; a breakdown of the sealant at the concrete-containment shell inter­face that permits entry of corrosive fluids from spills, leakage, or condensation; or in areas adjacent to floors where the gap contains a filler material that can retain fluids. NPP inspections have identified corrosion of the steel containment shell in the dry — well sand cushion region, shell corrosion in ice con­denser plants, corrosion of the torus of the steel containment shell, and concrete containment liner corrosion. Corrosion incidences such as these may challenge the containment integrity and, if through — wall, can provide a leak path to the outside environ­ment. Several techniques have been investigated that exhibit potential for performing inspections of inaccessible portions of NPP metallic pressure boundaries (i. e., ultrasonics, electromagnetic acoustic transducers, half-cell potential measurements, mag— netostrictive sensors, and multimode guide waves).1 However, these techniques tend to be time consum­ing and costly because they tend to examine only a small area at a time. A technique that can be applied remotely to perform global inspections and determines the overall condition of the contain­ment metallic pressure boundary in a cost — and performance-effective manner is desired.

4.15.4.1 Introduction

One of the issues that need to be addressed is the thermal and mechanical behavior of constrained peb­ble beds under (cyclic) nuclear loading conditions experienced in a tritium breeding blanket.

The way a pebble bed responds to a thermal load depends primarily on the thermal transport proper­ties of the bed, such as the packing factor of the bed, and the thermal conductivities of the pebble material and the surrounding purge gas. In addition, due to differences in the thermal expansion coefficients between the pebble-bed material and its surrounding structure, stresses will be induced in both the pebble bed and the structural material, and the contact
areas between pebbles and pebbles and walls become an important parameter. Also, during longer term operation in the neutron-irradiation environment, swelling of the breeder material will generate stresses. Furthermore, the pebble-bed thermal transfer prop­erties may deteriorate with irradiation dose and lithium burnup.

These induced stresses may directly or indirectly affect the functional operation if the mechanical integrity of the blanket element is endangered or if heat or tritium removal is significantly deterio­rated due to pebble fracture, sintering, or melting. Various creep phenomena will affect the actual evo­lution of stresses, like thermal creep and irradiation creep. In addition, chemical reactions in between pebbles and between structures and pebbles may be enhanced under high contact pressures and the compositional changes arising from long-term oper­ation by lithium burnup and other transmutation reactions.

When the (constrained) pebbles experience stres­ses, by either compression due to thermal expansion or irradiation-induced swelling, these stresses will be (partially) relieved through thermal creep, that is, irreversible deformation of the pebbles. Under high stresses and at high temperatures, typically 0.6-0.8 times the melting temperature TM, these relaxation processes include sintering of the material. This irre­versible strain and deformation in the material can lead to embrittlement and, ultimately, to fragmenta­tion of the pebbles.

The fracture of pebbles results in an inhomoge­neous pebble-bed density and contact area and would lead to an inhomogeneous temperature distribu­tion that is hard to model or predict. The lack of predictability of the temperature distribution can be a major safety issue. A mechanically stable pebble bed with a high packing factor is desired.

4.16.3.2.5.1 Aluminum

The permeability of hydrogen isotopes through alu­minum is smaller than that through most metals. This is due to a very low intrinsic solubility and a moderate hydrogen diffusivity. Young and Scully99 reviewed the diffusivity of hydrogen in aluminum and discussed the reasons for the wide range of reported values. A majority of hydrogen present in aluminum alloys at lower temperatures is trapped at defects and alloying content tends to increase the solubility for hydrogen in aluminum slightly.99,170,171

Application of aluminum alloys in fusion reactors is limited by their neutron activation, low melting temperature (933 K for pure aluminum or signifi­cantly lower for some alloys), the formation of embrittling intermetallics when layered on other metals, and the difficulty of joining aluminum alloys (because of the native oxide layer). The thermal limitation is significant because precipitates in con­ventional aluminum alloys coarsen or dissolve above 450 K and most conventional alloys are not very resistant to creep deformation. Few studies of alumi­num alloys have been made to date for use at tem­peratures greater than about 450 K. Because many aluminum-containing intermetallics and aluminum oxide also have low hydrogen permeability and do not have the same thermal limitations as metallic aluminum, it has been proposed that they be used as hydrogen isotope barriers.123

4.16.3.2.5.2 Nickel

Nickel alloys are described in Chapter 2.08, Nickel Alloys: Properties and Characteristics. The per­meation of hydrogen and its isotopes in nickel has been extensively studied. The results from Louthan et al.10 provide a good estimate of the transport properties. They found that cold-worked nickel dis­plays higher permeability to hydrogen, speculating that diffusion is enhanced along dislocation networks. Louthan4 also measured permeation in high-pressure gaseous hydrogen isotopes, showing that the trans­port is not dependent on concentration (and that Sievert’s law is appropriate at elevated pressure). Normalization of nickel’s permeation4 and diffusion3 by the mass ratio (eqn [3]) has been shown to provide good agreement for all the three isotopes.

4.16.3.2.5.3 Copper

There are numerous reports on the permeability of hydrogen in copper; the gas permeation results of Begeal103 are suggested here. Caskey and co — workers172 found that the diffusivity of hydrogen in copper depends on the oxygen content, and conse­quently, that the effective diffusivity can be much lower than the lattice diffusivity. Nevertheless, the reported diffusivity is consistent among a number of reports.3,103,172-174 Copper is also described in Chapter 4.20, Physical and Mechanical Properties of Copper and Copper Alloys.

Besides He, hydrogen isotopes, particularly the fuel elements deuterium and tritium, are the main inci­dent ion species contacting PFMs and PFCs. The energy of these particles corresponds with the plasma temperature at the edge, which is in the range of some eV, but also includes highly energetic particles (<10keV) escaping from the inner core of the

(g)

Figure 10 Cross-sectional scanning electron microscopic images for nine different grades of W relevant to fusion engineering practice. All target specimens were exposed to consistent pure He plasmas at 1120 K for 1 h. The He+ impact energy was ~40eV; (a) PLANSEE stress-relieved W, (b) single crystal (100) W, (c) ITER ASTM B760 compliant W,

(d) PLANSEE W-5% Re, (e) PLANSEE W-1% La2O3, (f) ultrafine-grained W-1.5% TiC, (g) ULTRAMET CVD W-10% Re,

(h) VPS W (EAST), and (i) recrystallized W. Reproduced from Baldwin, M. J.; Doerner, R. P. J. Nucl. Mater. 2010,404,165-173, with permission from Elsevier.

plasma. The impact of the energetic hydrogen ions is influenced by the incident ion energy, the ion flu — ence, the temperature, and the material’s composition and microstructure. The resulting damage, that is, vacancy formation, vacancy clustering, bubble formation, and blistering, determine not only the amount of material degradation and erosion but also the hydrogen (tritium) retention in the material. For active control of the hydrogen retention, short-term thermal treatments of the surface are being investi­gated. However, the short thermal load required to effectively remove the deuterium and tritium may also destroy the thin material layer (nm to low-pm range) that is responsible for the majority of the

259

retention.

As deuterium retention in plasma-exposed beryllium targets saturates after a given ion fluence (see Section 4.19.3.2.1), it is apparent that retention in codeposits will eventually be the dominant accumulation mech­anism with respect to beryllium PFCs. This is pri­marily due to the fact that the thickness of a codeposit will continue to grow linearly with time. It is, therefore, critical to understand both the

retention amounts and the release behavior of hydro­gen isotopes from beryllium codeposits. In this sec­tion, a ‘codeposit’ includes both the codeposition (where a BeD or BeD2 molecule is deposited on a surface) and co-implantation (where deposited layers of beryllium are bombarded with energetic hydrogen isotopes) processes.

Initial interpretation of studies of beryllium code­posits were made difficult by relatively high oxygen impurity content within the codepositing surface.102,103 Subsequent measurements104 with lower oxygen con­tent seemed to indicate that the oxygen level within the codeposit was correlated to the level of hydrogen isotope retention in the codeposit. The other variable that was identified to impact the retention level in these studies was the temperature ofthe codepositing surface.

Measurements seriously questioning the impor­tance of oxygen on the retention level in beryllium codeposits were made by Baldwin eta/.105 In this data set, the oxygen content throughout the codeposit was measured by depth profiled X-ray photoelectron spectroscopy and the oxygen content did not corre­late with the deuterium retention level (Figure 8), although the temperature of the codepositing surface was still a dominating term in determining the deu­terium retention level. Later, more detailed measure­ments confirmed that the presence of a beryllium
oxide surface layer was not correlated with an increase in retention in beryllium.106

A systematic study of beryllium codeposition fol­lowed,107 identifying three experimental parameters that seemed to impact the retention level in a code­posit. Along with the surface temperature, the inci­dent deuterium energy and the beryllium deposition rate were determined to be influential scaling para­meters. The previously reported data in the literature was also evaluated using the derived scaling and found to agree with the predictions of the retention levels measured under the various experimental conditions present in the different machines. Later the derived scaling was revised108 to use the ratio of the fluxes of the codepositing species, rather than the deposition rate to permit more accurate extrapo­lation to conditions expected in the edge of confine­ment devices.

The ability to predict the level oftritium retention in beryllium codeposits is an important aspect of a safety program; however, developing techniques to remove the trapped tritium from codeposits is a more important issue. The deuterium release behavior during thermal heating of beryllium code­posits has been investigated.1 9 The results show that the maximum temperature achieved during a bake — out is the figure of merit for determining the amount of deuterium release from beryllium. Increasing

the time spent at lower baking temperatures did not increase the amount of deuterium released from the beryllium codeposits. These results, along with the retention level predictions, should make it possi­ble to design baking systems for different areas of a confinement device to control the accumulation rate of tritium to a desired level.

Physical properties of pure copper and copper alloys are quite similar in terms of the melting point, the density, the Young’s modulus, and the thermal expan­sion coefficient. Table 1 compares the room tem­perature physical properties of pure copper, PH CuCrZr, and DS CuAl25.2,27-29 Because PH copper alloys and DS copper alloys contain only a small amount of fine second-phase particles, the physical properties of these copper alloys closely resemble those of pure copper.

The conductivity of copper and copper alloys is the most important physical property for their applications. The electrical conductivity of copper can be reduced by thermal vibration of atoms and crystal imperfections, for example, solute atoms, vacancies, dislocations, and grain boundaries. These different mechanisms have additive contributions to the increase in resistivity. As with other metals, the thermal conductivity of copper, kth, is proportional to the electrical conductivity, l, described by the Wiedemann-Franz law, that is,

kth = 1LT [1]

where T is the absolute temperature and L is the Lorentz number. The electrical conductivity of pure copper is sensitive to temperature, and less sensitive to the amount of cold work and the grain size. The linear temperature coefficient for electrical

resistivity in copper is dp/dT = 6.7×10-11 OmK-1.30 Severe cold work can reduce the electrical conductiv­ity of copper by only 2-3% IACS.

All alloying elements in copper reduce the elec­trical conductivity, and the amount of degradation depends on the type of element, the concentration, and microstructural form (e. g., solid solution, pre­cipitation, or dispersion). Figure 2 compares the strength and conductivity of copper and several types of copper alloys.31

An advantage of the Master Curve method is that it makes possible the use of Charpy-size and even smaller specimens for a valid determination of frac­ture toughness and T0. The number of tests always has to be determined so that the minimum required confidence level is achieved for the estimation.

ASTM E1921 describes a special weighting system to ensure a sufficient confidence level for the analysis. The final check can be made only after the value of T0 is rather well known, when the ade­quacy of tests can be determined from the condition У]ni ri > 1, where Гі is the number of valid data in the valid temperature range і and ni is the weight factor of this range. Normally, the required number can be tentatively determined only after some tests have been conducted, because the optimal test tempera­ture range is not known beforehand. With very small specimens, the final number of tests needed for a

.

valid estimate will likely be larger than the given minimum of six because of the smaller validity win­dow, which is reduced with decreasing specimen size and the material yield strength (see Figure 7). Exam­ples of possible small and miniature SE(B) specimen geometries for fracture toughness testing are com­pared in Figure 12.

As mentioned previously, a commonly used speci­men configuration for irradiated RPV steels is the full Charpy-size geometry with 10 x 10 mm cross­section. For most applications, this geometry provides a sufficiently large validity window (Figure 9), because as few as six specimens may be sufficient for a valid estimate. From the present experience for irradiated steels with different size specimens, Charpy square or half Charpy rectangular SE(B) or 0.4 T or 0.5 T C(T) geometries are generally optimum when the amount or form of the test material is limited. In some cases, even smaller test specimens may be required due to very small amounts of test material.

A comparison made between T0 and scatter estimates from test results measured with miniature specimens (i. e., smaller than the Charpy-size) shows that the definitions of scatter and the measuring capacity (specimen size) criterion (eqn [14]) apply even to miniature specimens that are of 3 x 4 mm and

3.3 x 3.3 mm cross-section SE(B).15 The results indi­cate no bias between the T0 estimates measured with the miniature specimens compared to the overall
mean values of T0, which is shown in Figure 13 for the Charpy-size specimen and three subsize SE(B) specimens. The comparison demonstrates, in all respects, applicability ofthe miniature size specimens for the fracture toughness estimation using the Mas­ter Curve approach. In some datasets (e. g., on grade 15 Kh 2 MFA and on the International Atomic Energy Agency (IAEA) reference material JRQ), the smallest specimens (3 x 4 and 5 x 5 mm) produced some low T0 values, which most likely were caused by macroscopic inhomogeneity encountered with such small specimen dimensions. An example of miniature specimen results for the A508 Cl. 3 steel FFA (a French steel grade) is presented in Figure 14, demonstrating nearly consistent fracture toughness data independent of the specimen size.

Another benefit of using the Master Curve approach is that the uncertainty associated with the T0 estimation can be determined and taken into account for assessing a conservative, realistic estimate for the lower bound fracture toughness. As discussed previously, ASTM E 1921 does not set limits for the specimen size or configuration; however, the mini­mum number of test results, dependent on test tem­perature relative to T0, is predefined to ensure an acceptable minimum confidence level for the esti­mate. If a larger uncertainty for the T0 estimation is accepted, the minimum number of specimens required can be reduced from that given in the stan­dard. Correspondingly, having less than the minimum

number of valid data in the test series does not necessarily invalidate the dataset, but it does result in a lower confidence level of the estimate. It is essential that the most realistic confidence level is estimated and taken into account in final integrity assessments.

Very small specimens (around 3 x 3 mm) tend to give slightly (1-3 °C) higher values of T0 compared to 10 x 10 mm specimens.15 This trend is likely due to the censoring procedure, which screens out pro­portionally more data from the upper tail of the dataset than from the lower tail. This screening affects both the scatter caused by possible material inhomogeneity and that from statistical outliers. The optimal test temperature range for miniature specimens has been proposed to be —50 °C < T — T0 <—20 °C.

Even though more specimens are needed when smaller specimens are tested, the consumption of test material becomes smaller, even if more than the minimum number of specimens were tested
for the T0 estimate. In this respect, the 5 x 5 mm specimen is the least material consuming SE(B) size (~12 specimens are needed for the standard estimate if sys = 500 MPa).15 Using the 5 x 5 or 3 x 4 mm specimen geometry, it is possible to prepare up to 8 or 12 subsize specimens from the halves of one tested 10 x 10 mm specimen (see Figure 15). When selecting specimen configuration, it should be noted that with deeply cracked specimens having the liga­ment size equal to or less than the thickness, the ligament is the primary dimension limiting the specimen-measuring capacity, not thickness. Also, with slim (reduced thickness) and very small speci­mens (like the 3 x 4 mm cross-section), it is recom­mended to side-groove the specimens to increase stress triaxiality near the surfaces.

4.15.5.1 Tritium Release

A wide range of mechanisms play a role in the tritium transport and release processes of the lithium — containing ceramics’ of which an impression is given in Figure 48.1 2 Tritium generated from neutron cap­ture is first transported to the grain boundary by bulk diffusion. The bulk diffusion and trapping inside the grains are affected by the neutron radiation-induced defects. Via the intergranular diffusion, the tritium is then delivered to the grain surfaces’ which are exposed to open and closed porosity. The closed porosity frac­tion provides another means to build up inventory in

the material. At the surface isotope exchange with hydrogen (H2) and water (H2O) lead to desorption of tritium in molecular forms of HT and HTO, respec­tively. Further, the tritium in molecular form is trans­ported through the interconnected pores and enters the flow of the purge gas. In order to assess the tritium retention in the candidate ceramic breeder material, one needs to know which of the steps are rate deter­mining and which operation parameters are the most relevant for facilitation of the tritium release (Table 3).

The tritium release characteristics oflithium cera­mics are typically studied in two parameter ranges:

1. Out-of-pile. Tritium production through exposure to neutron irradiation, followed by out-of-pile tri­tium desorption through stepwise isothermal or ramp annealing tests in laboratory setups, also known as temperature programmed desorption (TPD). If irradiation doses are very low, such activity is typically called ‘tritium doping,’ and tritium transport parameters reflect beginning of life (BOL) conditions only because irradiation damage and lithium burnup remain negligible.

2. In-pile: In case of in-pile experiments, typically steady-state tritium production and release condi­tions. In general, such parameters are closer to breeding blanket conditions, as they allow the application of a wide range of temperatures and purge gas conditions, and the study of long-term performance issues such as irradiation damage and lithium burnup. At present, such data are limited in terms of fast neutron damage doses (thermal and mixed spectra materials test reactor (MTR) only).

For fusion plasma-facing applications, the most essential properties are thermal conductivity, strength and ductility, thermal shock and thermal fatigue resistance, structural stability at elevated tem­perature, and stability of the properties under neu­tron irradiation. The advantages and disadvantages of tungsten for these conditions are manifold and opposed to each other as shown in Table 2. While the advantages of the material are mainly related to its high temperature-handling capability, the limitations are associated with manufacturing and handling at low temperatures (below ductile to brittle transition temperature, DBTT61-63), plasma com­patibility including neutron irradiation, and radio­logical issues.

However, with regard to other potential PFMs, for example, Be (see Chapter 4.19, Beryllium as a Plasma-Facing Material for Near-Term Fusion Devices), CFC (see Chapter 4.18, Carbon as a Fusion Plasma-Facing Material), and Mo, tungsten is still the most promising, offering an advantageous combination of physical properties and, therefore, has become the material of choice for ITER and DEMO. Since this decision was made, R&D efforts for investigating newly developed tungsten grades

Table 2 Features of W armor materials

Advantages Disadvantages

50 mm

(b)

and alloys that are able to overcome or at least miti­gate some of the above-mentioned disadvantages have significantly increased.