Neptunium is the first transuranic element. Its most stable isotope, 237Np (half-life = 2.144x 106 years), is a by-product ofnuclear reactors and plutonium pro­duction, formed either by a-decay of 241Am or from 235U by neutron capture:

235U + n 236 U + n-! 237U —^ 237 Np

It is also found in trace amounts in uranium ores due to transmutation reactions.

237Np is the most mobile actinide in the deep geological repository environment.194 Moreover, because of its long half-life, it becomes the major contributor of the total radiation in 10 000 years. This makes it and its predecessors such as 241Am candidates of interest for destruction by nuclear transmutation. Knowledge of Np carbide properties is therefore important for the management of carbide fuel waste, as well as for transmutation concepts, such as the ‘deep-burn,’ involving the production of fuel in contact with carbon.195

2.04.5.1 Preparation

Single-phase NpC0 94 was prepared by Lorenzelli196 by reacting Np hydride (obtained by reacting Np metal and water at 423 K) with carbon at 1673 K for 4h under vacuum. The final material had oxygen impurities of about 0.3 wt%. Lorenzelli
also observed that the solid state equilibrium between Np monocarbide and carbon at 1673 K under vacuum evolved after powdering and sintering with the for­mation of Np2C3. Sandenaw eta/.197 obtained purer NpC091 samples by arc-melting Np metal and carbon (0.024 wt% O2).

Neptunium dicarbide was prepared by heating NpO2 with graphite under H2 between 2930 and 3070 K.198 However, the preparation of neptunium dicarbide could not be repeated, not even by melting Np2C3 in a graphite crucible,196 and the identification of this compound therefore remains controversial.

4.14.4.1 Assessment of Irradiation Embrittlement Changes

The Master Curve methodology was originally developed for applications like RPV surveillance programs, for which the conventional methods are generally less accurate (based on Charpy V-notch energy shifts) or not suitable due to specimen size requirements (LEFM Kjc and Kja tests). As a direct measurement approach, the Master Curve approach is preferred over the correlative and indirect meth­ods, based mostly on the Charpy test, used in the past to assess irradiated RPV integrity. jt is therefore reasonable to expect that the future determination of plant-operating limits will be based on the Master Curve and related methods rather than on the

indirect methods or the trend curves based on the chemical composition of materials and their expected neutron fluence.

The advantages achievable with the ASTM E 1921 methodology especially for RPV applications are as follows:

• Direct fracture toughness estimation of the RPV using irradiated small SE(B) or C(T) type speci­mens and determination of a statistically correct mean behavior for ageing assessment and a realistic lower bound curve definition for integrity assess-

15,32-34

ments.

• Expansion of the analysis to cover issues related to material inhomogeneity and the quality of measured data are possible utilizing the proposed statistical methods.

• Expansion of the maximum likelihood estimation to take into account specimen-specific fluence data is possible when needed.

• Different data and those measured with different size and type specimens can be included in the analysis (including LEFM Kjc with caution). (Note that C(T) and SE(B) specimens may show a (usu­ally about 8 oC) bias due to different geometry so that C(T) specimens yield higher T0.)

• Utilization of correlations between the parameters characterizing different loading conditions like crack arrest and dynamic loading.

A typical situation especially for older RPVs is that the existing material data consists of very miscella­neous information on the properties of materials, such as test results measured with numerous different material conditions, test standards, specimen types, equipment, etc. jn such a situation, a method for handling and analyzing all available data in a syner­gistic way can provide significant savings especially if there are no archive test materials available for addi­tional testing. These aspects of characterization are described schematically in Figure 26.

jt is important to note that, based on present knowledge, the shift in Charpy transition temperature (e. g., AT41j) due to neutron irradiation on average is close to or less than the transition temperature shift in fracture toughness (AT0 from the Master Curve method); that is, the shift in Charpy data is generally unconservative in respect of the cor­responding shift in T0. However, there is large scatter in the relationship between these two shifts, and cau­tion is needed when assessing equivalence.

Although there are no specific requirements for Master Curve testing in RPV surveillance programs

Surveillance 3-PB specimens

Data evaluation using the Master Curve approach

Figure 26 Characterization of irradiated reactor pressure vessel materials and related parameters.

in the current Codes and Regulations, the methodol­ogy has been applied in national RPV surveillance programs, and numerous retroactive analyses have been made using data measured in the past in the surveillance and material characterization programs of NPPs. Today, there are also many publications and guidelines which can be used as guidance for using the Master Curve and the related methodologies effec­tively and in a proper manner. Applications for nuclear grade pressure vessel materials and irradiated materials are addressed in the IAEA publication Technical Report Series No. 429.33 Chapter 4.05, Radiation Damage of Reactor Pressure Vessel Steels, provides additional information and mecha­nistic details of ‘Radiation Damage of Reactor Pres­sure Vessel Steels.’

In addition to the integrity and safety requirements according to which a blanket component has to be designed in order to be licensed, there is another matter of economics: in case of transient events or accidents not affecting safety, they may prohibit fur­ther operation of the blanket and require replace­ment. Those blanket concepts that are more tolerant to design base and other accidents will be preferred by utilities. In this context, the use ofberyllium-based neutron multipliers, the type of coolant, and the tritium issues for the purge and coolant processing units (isotope separation, purification, steam generator operation, toxicity at accidental conditions) also appear to be very important aspects associated with the breeder material use.

4.15.8.5 Development of Tools

Thermomechanical codes to describe the interaction of pebble beds with the structural material have achieved significant progress in recent years. For complete TBMs and, even more, DEMO blanket modules, computational codes based on continuum mechanics will be the first choice in the near future. These codes should be quickly developed to enable fairly good predictions for the blanket behavior at the BOL of a DEMO plant or power reactor.

The database for irradiated material is still very small and the amount of relevant engineering data in the near future will also be very limited, as material development is still ongoing. Pebble-bed experiments that require a considerable amount ofmaterial will be costly and hardly possible for some time. This is the significant challenge for discrete-element methods (DEM). These codes must be improved in order to describe more realistically the interaction between nonirradiated pebbles (taking into account pebble shape, surface condition, and material properties, including thermal creep). If this goal is achieved, it should be possible to do this for irradiated materials, because a small pebble mass is required to make corresponding experiments. The results obtained then with DEM codes should be fed into formal correlations in the continuum codes in order to assess the blanket behavior toward EOL conditions.

4.15.8.6 Compatibility with Structure

Though experimental evidence is now accumulating for the irradiation behavior of current candidates for ceramic breeder and structure, there is no clear insight into the extent of chemical/physicochemical interactions taking place in long-term operations under a reducing atmosphere in a DEMO or power reactor. While these may affect the breeder proper­ties, they may also affect the blanket components structural integrity.

4.15.8.7 Waste Management and Reuse/Recycling

The large volume of ceramic breeder and multiplier required for breeding blankets in future power reac­tors necessitate ecological and sound economic solu­tions for intermediate storage and back end. Cost effectiveness and sound nuclear industrial practices will promote the selection and qualification of ceramic breeder technologies with full processing and recycling capabilities.

During thermal shock loads, steep temperature gradients of hundreds to several thousand degrees Celsius on a length scale of millimeter or even micrometer (depending on the pulse length) are formed, influencing only a limited volume near the loaded surface. While the heat load is applied, due to thermal expansion and the decreasing strength of the material at the surface compared to the bulk material, compressive stresses are formed in the surface plane. These stresses can lead to permanent plastic defor­mation that might, during cool down, generate tensile stresses sufficiently high to initiate crack formation perpendicular to the surface and thereby cause stress relaxation at the surface.

Depending on the mechanical properties in the surface plane, the amount and starting point of crack formation can be influenced. Based on this and on the fact that the mechanical properties are strongly dependent on the material’s microstructure (see Section 4.17.3.2.3), a grain orientation parallel to the surface and therefore high strength in the surface plane might be preferred.162 However, grains oriented parallel to the surface, such as in rolled materials or plasma-sprayed coatings, might result in delamination (see Figure 3(a)), which causes over­heating and subsequently surface melting if they have a lower strength in the depth direction and exhibit preferential cracking along the weak grain boundaries.

Therefore, a grain orientation perpendicular to the surface and parallel to the direction of the heat

flow is recommended.90 This will cause cracks to form along the grain boundaries toward the cooling structure (see Figure 3(b)) causing no degradation or only a negligible degradation of the material’s ther­mal transfer capabilities. Due to the lower mechani­cal properties in the surface plane, more or larger cracks will form during thermal shocks, running per­pendicular to the surface and following the grain orientation.

In contrast to deformed materials, crack formation and crack orientation in materials with isotropic or almost isotropic grain structures, for example, MIM-W or recrystallized W, is rather unstable and is strongly enhanced for the weakened recrystallized material. Depending on the applied power densities, the formed temperature gradient, and the resultant stress fields within the material, cracks initially running perpen­dicular to the surface might deflect at zones with compressive stresses and keep running parallel to the surface (see Figure 4).

4.17.4.1.1 Power density and pulse duration

The material’s response is strongly related to the applied temperature fields and by this to the absorbed power density and the pulse duration. This results in a material-related surface temperature increase and heat penetration depth.163 A classification of the impact of the temperature field is made by establishing three parameters: the damage, the cracking, and the melting threshold. While the latter depends on the ther­mal conductivity and the melting temperature (for alloys or mixed materials formed during tokamak
operation) of the material, the damage and cracking threshold are determined mainly by the material’s mechanical properties. Damage here means that the material’s surface has undergone a visible and measur­able modification, for example, by surface roughening, recrystallization, or pore/void formation.

The limiting temperature for graphite use in fusion systems is defined by thermal sublimation 1500— 2000 °C). However, a process that is very similar to thermal sublimation (in cause and in effect) appears to define the current temperature limit. This phe­nomenon, which is known as radiation-enhanced sublimation (RES), is not clearly understood yet, but dominates above a temperature of about 1000 °C, and increases exponentially with increasing temperature.

One theory says that the process responsible for initiating RES follows from the earlier discussion of radiation damage in graphite. Specifically, in a displacement event, a Frenkel pair is created. The interstitial has a low (~0.5 eV) migration energy, is quite mobile between the basal planes, and thus diffuses readily. Some fraction of these interstitials are condensed at vacancy sites, which are essentially immo­bile below about 700 °C (migration energy ^4eV). Other migrating interstitials can be trapped by microstructural defects or can coalesce into simple clusters, which limits their mobility. However, some fraction of the interstitials diffuse to the surface of the graphite and thermally sublime. The thermal sublimation of radiation-induced interstitials is RES, and must be distinguished from both physical and chemical sputtering. Time of flight measurements have shown that the thermal energy of RES ions has a Maxwellian energy distribution, which is directly coupled to the mean surface temperature.83 This clearly distinguishes RES atoms from physically sput­tered atoms, which exhibit highly anisotropic energy distributions. RES atoms are also distinguished from thermally sublimed species in that only single carbon atoms are detected, whereas single atoms and atom complexes (C2, C3, …) are found during thermal sublimation. Another theory for the explanation of RES is simply that the bombarding hydrogen ions turn the very near-surface region into a low-density amorphous zone. A very large fraction of the carbon atoms in this zone are now edge atoms with weak bonding to the connecting atoms. These edge atoms are much more easily thermally volatized into the plasma.

The effect of RES in the next generation of high surface particle flux fusion systems is presently unclear. Evidence suggests that the erosion yield does not scale linearly with flux, as physical sput­tering does, but may in fact decrease significantly with increasing flux.84 Moreover, as with chemical erosion, the inclusion of interstitial boron into the crystal lattice can decrease RES and shift the thresh­old to higher temperatures. Boron will volatilize above 1500 °C, thus limiting the PFM temperature to <1500 °C.

4.18.4.3 Erosion of Graphite in Simulated Disruption Events

Finally, the effect of plasma disruptions needs to be considered. Section 4.18.2 discussed the thermo­mechanical response of the PFCs to the excessive plasma energy in a disruption. This large thermal energy dump can additionally cause enhanced ero­sion due to the increased particle flux, elevated sur­face temperature, or simply by exfoliation of the surface due to thermal shock. The latter two material losses are reduced for materials with high thermal conductivity. This has been demonstrated experi­mentally, and is shown in Figure 27,3 which gives weight loss as a function of thermal conductivity for a number of graphites and composites of varying thermal conductivities subjected to one electron beam pulse at 4.1 MW m~ . As discussed in Section 4.18.2, and as seen in the data of Figure 27, high thermal conductivity materials reduce the surface tempera­ture, and hence the overall erosion yield, during a disruption.

Although not a concern in present day tokamaks, in-vessel dust and tritium inventories have been recognized as a safety and operational issue for next step devices such as ITER.190-193

In particular, accident scenarios that result in water or steam exposure of hot plasma-facing mate­rials are one of the greatest concerns for ITER, because steam interacts with hot beryllium leading to the production of hydrogen, and hydrogen in the presence of air can lead to an explosion.

The steam-chemical reactivity of different grades of Be has been studied extensively in the past.194-200 The amount of hydrogen produced depends on the specific material, temperature, exposure time, and especially the effective surface area. Because of the large surface area of dust, its chemical reactivity is an issue.

Dust is expected to be produced inside the vac­uum vessel of a tokamak by interaction of the plasma with the components of the first wall and the divertor. A detailed discussion of the mechanisms of dust pro­duction and of the influence of parameter variations is beyond the scope of this contribution, but it should be noted that the processes and the production rate of dust are not fully understood and the extrapolation of knowledge from existing tokamaks to ITER is difficult. Research into dust production mechanisms and rates, the appropriate dosimetric limits for per­sonnel exposure, and methods of removal has only

recently begun.201,202

The location where the dust settles will determine its temperature, and consequently, its chemical reac­tivity. At the moment about 6 kg of C, 6 kg of W, and 6 kg of Be dust are allowed ‘on hot surfaces’ in ITER, with these limits set by the H production risk. This corresponds to the maximum allowable quantity of H (2.5 kg) for the vessel integrity to be guaranteed in case of explosion. A complete oxidation of Be at 400 °C and C at 600 °C is assumed for the calculation. If no C is present in the machine, the limits are relaxed to 11 kg for Be, or 230 kg for W. These quan­tities are set such that the overall hydrogen combus­tion limit is not exceeded.9

It must be recognized that a limit on the order of ~10 kg for beryllium dust on ‘hot-surfaces’ is very restrictive, and in particular, the development of diagnostics techniques that can determine from local measurements the global inventory in the machine could prove to be very challenging.203 How­ever, it is also likely that dust in ITER produced by Be eroded from the wall and deposited on the divertor will not survive on plasma-facing surfaces exposed to heat fluxes and will tend to accumulate in grooves or castellations in the armors of PFCs. They are an essential feature of the design of PFCs to relieve stresses during cyclic high heat flux load­ing, thus maximizing the fatigue lifetime of the armor to heat-sink joint. Some reduction in reaction rates is expected because the steam supply is not unlimited and steam must diffuse through the dust in the grooves. Experiments have been carried out in the Russian Federation, both in the Bochvar Institute of Moscow and the Efremov Institute of St. Petersburg.204 Although not conclusive, the main results summarized in Figure 20, show a reduction of the measured Be steam reactivity, particularly at high temperatures (more than a factor of 20). How­ever, further experimental and modeling work is needed to clarify if the observed slower kinetics at high temperatures (800-900 °C) eliminates the risk of explosion in the event of an accident.

Radiation-induced conductivity (RIC) is the loss of insulation only during irradiation, which is a common issue for insulator ceramics in irradiation environ­ments. Historically, evaluation of RIC has been car­ried out mostly for Al2O333 (see also Chapter 4.22, Radiation Effects on the Physical Properties of Dielectric Insulators for Fusion Reactors).

Figure 10 shows RIC as a function of the dose rate for Er2O3, Y2O3, and CaZrO3 bulk specimens for 14 MeV neutron, fission neutron, and g-ray irradiation, in comparison with data on Al2O3.34,35 For Y2O3 and AlN, results for the coating are also shown. The RIC of the candidate materials of Er2O3, Y2O3, and CaZrO3 are comparable with Al2O3. According to these results and the expected dose rate in an Li—V fusion blanket36, the expected induced conductivity in the fusion blanket condition is much lower than the maximum allowable value of ^10~2 S m-1.37

The effects of the nuclear properties of Er on radio­activity and tritium breeding ratio (TBR) of V—Li blankets were also investigated. Figure 11 shows the contact dose rate of a V—Li blanket with and without neutron multiplier Be in the cases of (1) no coating, (2) 10 pm, and(3) 1 pm coatingwithEr2O3. Withoutthe coating, the dose rate was dominated by V-4Cr-4Ti substrates reaching the hands-on recycle limit after several decades of cooling. (Note that impurities in

V-4Cr-4Ti were not considered in this calculation.) With the coating, the dose rate increases because of the contribution from Er but still satisfies the remote recycling limit.38 Er is a neutron-absorbing element and can reduce the TBR especially for an in situ coating where Er is doped into Li. However, because only

0. 15% Er is needed in Li for the in situ coating, 9 the impact of Er in Li on the TBR is not an issue.39,40

Operating experience has demonstrated that periodic inspection, maintenance, and repair are the essential elements of an overall program to maintain an acceptable level of reliability for structures over their service life. Assessment and management of aging in NPP concrete structures requires a more systematic approach than simple reliance on existing code margins of safety.82 What is required is the integration of structural component function, poten­tial degradation mechanisms, and appropriate control programs into a quantitative evaluation procedure. A methodology for demonstrating the continued reli­able and safe performance of these structures should include (1) identification of structures important to public health and safety; (2) identification of environ­mental stressors, aging mechanisms and their signifi­cance, and likely sites for occurrence; (3) a monitoring — or in-service-inspection-based methodol­ogy that includes criteria for resolution of existing conditions; and (4) a remedial measures program.

Because significant quantities of ceramics will be needed in the near future for the fabrication of ITER TBMs and for a potential ITER driver blanket, various efforts have been initiated to evaluate fabri­cation process development. One of the fabrication issues is the hygroscopic nature of several candidate

Table 1 Some design and loading parameters for a number of fusion reactor blanket concepts utilizing ceramic breeder materials. For comparison values for the EU ITER Test Blanket Modules are given also. In addition to the references mentioned in the table, reader can turn to references in the text like e. g. 16, 21, 24 and 38. Design BIT BOT HCPB WCSB SSTR A-SSTR2 PPCS-B HCPB-TBM in (HCPB) ITER EU EU EU Japan Japan Japan EU EU References 6,7,8 2,3,4,8 15 20 217 218 16 17-19 Ceramic breeder (CB) LiAlO2 Li4SiO4 Li4SiO4 or Li2TiO3 or Li2O Li2TiO3 Li4SiO4 Li4SiO4 Li2TiO3 other Li-6 enrichment % 90 25 30% (Li4SiO4), 90 30-40 90 60% (Li2TiO3) alternative ceramic (Li2ZrO3; Li2TiO3 Li2TiO3 Li2TiO3 breeder Li2TiO3) Shape Pellet Pebble Bed Pebble Bed Pebble Bed Pebble Bed Pebble Bed Pebble Bed Pebble Bed Coolant water He He water H2O orSCW He He He Neutron multiplier Be-pellets Be-pebbles Be-pebbles Be-pebbles, Be-pebbles Be-pebbles Be-pebbles Be-pebbles or Be12Ti Minimum ceramic breeder C 410 300 400 400 450 700 400 250 operation temperature Maximum ceramic breeder C 590 660 890 900 750 800 920 900 operation temperature Structural material 316 SS MANET Eurofer F82H F82H SiCf/SiC Eurofer Eurofer MW/m2 (RAFS) Surface heat flux 0.5 1 1 <1 0.4 / 0.5 (peak) 0.27/0.5 (peak) Neutron wall loading MW/m2 2.4 3.2 3-5 50 2.2 /3.5(peak) 0.78 Lifetime fluence MWa/m2 >10 10 7-8 0.1

Figure 10 View of a typical test blanket module port cell arrangement in ITER. Reproduced from http://www. iter. org/mach/ tritiumbreeding.

lithium ceramics. Sensitivity to moisture increases as the lithium oxide content increases and as the specific surface area increases.

The research activity initially involved g-LiAlO2, Li2O, Li2SiO3, and Li2ZrO3; see, for example, Johnson eta/.25 and Roux eta/.52 Later work concerned Li4SiO4, Li8ZrO6, and Li2TiO3.26’33’53-55’185’189 Cur­rently, most blanket concepts are based on Li4SiO4 or Li2TiO3, though recently, work on other systems such as Li3TaO456 and Li8PbO657 as well as compo­sites of Li2TiO3 with Li2O or Li4TiO4 additives58 has been reported.57 Breeder development has also started in Korea and India.59-61

There is significant interest in reduced activation ferritic/martensitic (RAFM) steels to replace nickel­bearing austenitic stainless steels in reactor applica — tions117. There are many RAFM steels that have been proposed and investigated in the literature specifi­cally for fusion applications; these typically contain between 7 and 12 wt% chromium, relatively low carbon (<0.15 wt% C), and controlled alloying additions to bolster structural properties, while min­imizing activation (e. g., additions of W, Ta, and vanadium and reductions of nickel, molybdenum,

and niobium content). The transport of hydrogen and its isotopes has been extensively studied in MANET (MArtensitic for NET, including the so-called MANET II) and modified F82H (generally referred to as F82H-mod). Some of the other desig­nations ofRAFM steels that can be found in literature include EUROFER 97, Batman, OPTIFER-IVb, HT — 9, JLF-1, and CLAM steel.

In general, studies of RAFM steels report rela­tively consistent transport properties of hydrogen and its isotopes; some of these studies are reviewed in Serra et a/.118 Despite the consistency of the data available in literature from several research groups, few studies verify the expected pressure dependence of the transport properties that is expected for diffusion- controlled transport. Pisarev and coworkers119,120 have suggested that the literature data may underestimate diffusivity and solubility due to surface limited trans­port. Similar suggestions have been presented to explain some of the data for the austenitic stain­less steels1; however, the work on austenitic stainless steels has been cognizant of the issues with surface effects; generally surface effects are mitigated by coating specimens with palladium or other surface catalyst. Such precautions have not been systemati­cally employed for permeation studies of the RAFM steels, although the need to control the surface
condition (and confirm the square root dependence on pressure) has been widely acknowl — edged.29,30,11 ,121 While the apparent transport prop­erties in the absence of trapping are relatively consistent for all the RAFM steels, the issue of sur­face effects and the suggestions of Pisarev et al. need further validation in the literature because the trans­port of tritium is less likely to be affected by surface conditions compared to deuterium and protium.

The diffusivity of hydrogen is shown in Figure 12 along with an average relationship (Table 1). The literature data are generally within a factor of 2 of the average relationship. The MANET alloys tend to have lower diffusivity of hydrogen and its isotopes than F82H-mod. Differences in permeability between these two alloys has been attributed to Chromium content;29,30 however, a clear correlation of transport properties with Chromium content cannot be estab­lished on the basis of existing data.122 At temperatures less than about 573 K, the apparent diffusivity is signif­icantly less than the exponential relationship extrapo­lated from higher temperatures. This is attributed to the effect of trapping on the transport of hydrogen and its isotopes.

The reported values of apparent solubility of hydrogen and its isotopes in RAFM varies very little in the temperature range from 573 to 873 K. Pisarev

and coworkers report values that are three to four times higher on the basis of their assessment of sur­face effects. Here we recommend a relationship for the apparent solubility (Table 1) that is consistent with the majority of the literature data with AHs = 28.6 kJ mol — , which is based on a simple curve fitting of the data shown in Figure 13. The values of the solubility are about an order of magnitude less than the austenitic stainless steels in the temperature range between 500 and 1000 K, although the solubil­ity of hydrogen is more sensitive to temperature for the RAFM steels since AHs is four times the value for the austenitic stainless steels.

The trapping characteristics of the RAFM steels have been estimated for several alloys.19,118,121’123-126 Although binding energies and densities of hydrogen traps vary substantially, the majority of reported values for RAFM steels are in the range 40-60 kJ mol-1 and 10 3—10 5 traps per metal atom, respectively. The traps are attributed primarily to boundaries118 and result in a significant reduction in the apparent diffu- sivity at temperatures less than about 573 K. At higher temperatures, the traps are essentially unoccupied and do not affect diffusion.20

The measured recombination coefficient is many orders of magnitude lower than theoretical predic­tions; moreover, the measured values can also vary
substantially from one study to another.118,127,128 Measured values for the recombination coefficient for deuterium on MANET alloys are approximately in the range 10-2—10-4m4s-1 per mol of H2 for the temperature range 573—773 K.127,128 Oxidation of MANET was shown to induce surface-limited trans­port of deuterium and reduce the recombination coefficient kr « 10-6 m4 s-1 per mol of H2.128 Furthermore, it is suggested that structure and composition of the oxide may also affect the recom­bination coefficient and that oxidation can increase the energy barrier associated with dissociation of the gaseous diatomic hydrogen isotopes.128

In summary, the diffusivity and the solubility of hydrogen and its isotopes are consistently similar for all the RAFM steels that have been tested for fusion applications. RAFM steels show a relatively rapid diffusion and low solubility of hydrogen and its isotopes at ambient temperature. The diffusivity is six orders of magnitude greater than that of the austenitic stainless steels at 300 K, while the solubility is more than three orders of magnitude lower than that of the austenitic stainless steels. The diffusivity of hydrogen and its isotopes is not strongly sensitive to temperature compared to most other metals. On the other hand, the heat of solution (AHs) for the RAFM steels is quite large,

and thus the solubility of hydrogen approaches that of austenitic stainless steels at temperatures >1000 K. Consequently, at elevated temperatures (e. g., >700 K), the permeability is less than an

573 K, and thus the apparent diffusivity is much lower than expected from tests that are performed at higher temperatures.