3.5. Materials Development and Test Introduction Materials R&D for ...

Review of the ITER Technology R&D 

3.5 Materials Development and Test 

3.5. Materials Development and Test 

Introduction 

Materials R&D for the vessel and in-vessel components has been performed in support of 

ITER design, and comprised the structural materials (SS 316L(N), CuCrZr, CuAl25, Inconel 

718 and Ti-alloys), plasma facing materials (Be, W and CFC) and joints between SS/SS, 

SS/Cu alloys, Cu alloys/Be, Cu/W and Cu/CFC, and ceramic materials, Al 2 O 3 , MgAl 2 O 4 [1, 

2]. 

3.5.1 Objectives 

The main objectives of the R&D are the following: 

- to provide rationale of the materials selection, 

- full characterisation of selected materials (including effect of components manufacturing 

cycle and irradiation effect), 

- providing required database for the components design. 

There was no R&D on standard materials (316L, 304B7, 430, Nimonic 80) with well 

established manufacturing techniques and known properties for the ITER relevant conditions. 

3.5.2 Structural Materials 

3.5.2.1 Stainless Steels 

Austenitic stainless steel (SS) of 316L type is the main structural materials for the ITER 

vacuum vessel and for in-vessel components (shielding blanket, divertor cassette body). This 

steel type is qualified in many national design codes and has adequate mechanical properties, 

good resistance to corrosion, weldability, forging and casting potential. It is industrially 

available in different forms, and can be manufactured by well-established techniques. For 

Fusion Engineering and Design Page 1



ITER, only minor modifications in chemical composition are required, in order to cope with 

the radiological safety limits and with the re-welding requirement. 

The safety implications of the cobalt (Co) and niobium (Nb) content show the following: 

- activated cobalt plays an important role in determining occupational dose level during 

maintenance and in the case of accidents, 

- reducing the Co content from 0.25% to 0.05% decreases the total decay heat in the 

vacuum vessel by ~20% and helps to reduce the activation of components; 

- further decrease of Co to 0.01% does not further reduce the decay heat because the 

Ni 60 (n,p) Co 60 reaction becomes dominant; 

- cobalt is one of the main components of activated corrosion products in the water 

cooling system; 

- niobium produces long-lived radioisotopes that could become important for the 

decommissioning and waste disposal of in-vessel components. 

In 316 type SS, niobium is present as a trace element picked up during the melting 

process from the ferroalloy addition: existing data indicate that it is possible to reduce 

Nb to ≤0.01%. 

Vacuum vessel, manifold and branch pipe connections must remain weldable throughout the 

machine lifetime in spite of the He generated by neutron irradiation. The helium content 

generated by gaseous transmutation should be kept as low as possible, preferably below 1 

appm. Boron gives the most significant helium generation during irradiation. Therefore, the 

boron concentration in the steel used for the vacuum vessel, manifold and water cooling 

branch pipes should be as low as feasible industrially. The reduction of boron from 20 wppm 

(typical content) to industry feasible level of about 10 wppm lowers the helium generation by 

~44%. 




For application in ITER, standard chemical composition of 316L type steel has been 

modified, in order to cope with radiological safety limits and with the re-weldability 

requirement. The modified stainless steel 316L(N)-IG has the suffix IG (ITER Grade) to 

indicate modification, and (N) gives an indication of controlled nitrogen content. The 

chemical composition of 316L(N)-IG steel is given in Table 1. 

The key R&D directions for the stainless steel are the following: 

- effect of components manufacturing cycle on the material properties, 

- effect of neutron irradiation, 

- re-weldability tests after irradiation, 

- corrosion behaviour of SS. 

(1) Effect of manufacturing cycle 

Hot isostatic pressing (HIP) process is one of the options for the shield blanket modules and 

for divertor cassett body manufacturing (see 3.3 and 3.4). A study of the effects of HIP cycles 

on the mechanical properties of solid and power HIPed steels has been performed in the 

frame of ITER R&D. It was shown that the tensile properties of HIPed steel are within the 

design allowables. In some cases the yield strength of solid-HIPed base metal is close to the 

minimum specified values [3]. Powder-HIP gives tensile properties slightly above the 

average of the wrought material [4, 5]. Summary of tensile strength tests of 316L(N) type 

steel after different manufacturing cycle is presented in Figure 1. In brackets are given 

reference to laboratory and HT where tests have been performed. 

Casting of SS was proposed as a manufacturing options for the divertor cassette body, and 

the shielding part of the first wall modules, because it is cheaper than conventional welding 




or HIPing (see §3.3 and §3.4). The precision casting technique for the divertor cassette has 

been developed by the US Home Team [6], and the material characterization has been 

performed by the US and RF Home Teams [6, 7]. The results of investigation show that nonmetallic 

inclusions from mould products have not been observed in the castings for the 

prototype of the divertor cassette body and are not present when employing good foundry 

practice. The strength of cast and HIPed steel is slightly lower than reference data for the 

wrought steel (see Figure 1). Cast SS exhibits relatively good fracture toughness, and very 

low crack growth rate indicative of high resistance to cracking. The Straus method, the 

Electrochemical Reactivation (EPR) method and crack growth rate measurements did not 

indicate any susceptibility of cast SS 316L(N)-IG to stress corrosion cracking [7, 8]. 

(2) Irradiation effect on SS. 

Neutron irradiation is another factor which affects the properties of stainless steel. Extensive 

irradiation tests have been provided within the ITER task activity. The overview of the results 

of irradiation R&D has been presented in several publications [2, 4, 9, 10]. The last results of 

irradiation tests of stainless steel with different manufacturing cycle are presented in Table 2. 

At the expected neutron fluence ~ 0.3 MWa/m 2 , the peak damage in the steel for the in-vessel 

components will be ~ 2 dpa with a peak He generation of about 55 appm. The irradiation 

temperature for these components is in the range 100-300°C. The result of investigation 

shows that hardening of the steel will occur, but predicted uniform and total elongation of the 

wrought and powder HIPed steel will remain higher than 10% and 20%, respectively, at the 

end of the components operation phase [10, 11]. For the solid-HIP steel, the estimated 

uniform and total elongation will be above 18% and 31%, respectively. Therefore, materials 

will be relatively ductile and retain work hardening capability during ITER life-time [1, 10]. 

Results of post irradiation tensile tests shows that even for displacement doses up to 4-5 dpa 




the material after all types of manufacturing cycle exhibits a relatively high ductility (see 

Table 2), and fracture of components due to flow localisation in irradiated steel can be 

avoided during the design. Irradiation has been performed at 175±15°C and 265±15°C [11]. 

Data given in Table 2 refer to irradiation temperature 265°C where irradiation effect was 

more significant. 

The fracture toughness of SS decreases up to 5-6 times after ~ 4 dpa irradiation and then 

slowly decreases with higher doses of irradiation [4]. For the powder HIPed steel showing 

most significant irradiation effect compared with other materials, the fracture toughness 

decreases with increasing dose from ~1000 kJ/m 2 in the unirradiated steel to ~ 250-255 kJ/m 2 

after 2-2.5 dpa at 77-290°C [4, 5]. 

(3) Re-welding of irradiated material 

The quality of re-welded steel depends on the welding procedure and on the He content in 

SS. The current assumption, based on existing data, is that re-welding of 316L(N)-IG can be 

successfully carried out when the He contents is less than 1 appm. Nevertheless, the reweldability 

study performed by the JA HT shows that there is a potential of sound re-welding 

of irradiated material with higher helium concentration (up to 3-10 appm) if precision 

tungsten inert gas arc (TIG) or yttrium-aluminium-garnet (YAG) laser welding technique 

with low energy is applied [12, 13]. 

To minimise the probability of crack formation during re-welding, it is recommended: 

- to minimise He content and radiation dose for the irradiated areas (as low as 

possible), 

- to minimise the heat input during welding, 

- to minimise (or avoid) tensile stresses and/or to provide compressive stresses during 

welding. 




(4) Corrosion behaviour 

The water chemistry is the key requirement to avoid or minimise corrosion damage of 

materials. Result of analysis lead to decision that neutral pH, hydrogen water chemistry 

(HWC) shall be used for the ITER cooling system. The main results of corrosion effect study 

at the ITER working conditions are given in refs. [14, 15]. 

The estimation of corrosion rate of steel in the ITER working conditions (neutral water 

chemistry and temperature ~ 100-200°C) shows that corrosion rate decreases with time, and 

varies from ~0.5 to ~1 µm/y [15]. 

Three methods have been used for the revealing steel to SCC: Straus method, electrochemical 

reactivation (EPR) and slow strain rate tests (SSRT). Results of R&D show the following: 

- Wrought 316L(N)-IG steel, base metal after one HIP cycle, powder HIPed, cast material, 

solid HIPed joints SS/SS, TIG welded joints are not susceptible to SCC. 

- The steel anneal at 550°C during 15h does not result in sensitisation of steel products. 

- Anneal at 675°C during 15h gives severe SCC sensitisation of cast SS (EPR parameter 

increases up to 1.12%). Wrought SS shows approximately twice better resistance to SCC 

than cast SS. Sensitisation behaviour of heat-affected zone of welded joints is similar to 

base metal. 

- Heat treatment at 675°C for 50 h results in sensitisation to SCC of the cast SS. This 

sensitisation was revealed by Strauss method, oxalic acid etching and by EPR tests (EPR 

parameter was about 10%). 




- No SCC of unirradiated 316L(N)-IG observed with SSRT and constant deformation 

testing (CDT) in pure water, water plus hydrogen, and water plus oxygen and/or 

hydrogen peroxide in the temperature range 90°C-320°C. 

- No SCC of 316L(N)-IG SS observed with in-pile testing consisting CDT specimens 

irradiated at 200°C in water plus oxygen (> 10 wppm) and hydrogen peroxide (>3 wppm) 

up to about 3 dpa. 

- No SCC of unirradiated 316L(N)-IG SS observed at 320°C with SSRT and CDT in 

neutral solution of chloride, sulphate and copper ions, containing about 2 wppm of 

hydrogen. However, with addition of 1-30 wppm of hydrogen peroxide to simulate 

possible water radiolysis, SCC occurred on sensitised specimens after 1500 hours of 

exposure. In simulated acidic hide-out environment containing chloride and sulphate as 

observed in BWRs, 316L(N)-IG SS is sensitive to SCC at 340°C. 

- No corrosion effect was observed on fatigue behaviour of 316L(N)-IG SS at 320°C after 

10 5 cycles at 0.5% strain amplitude. 

The estimation (by RF HT) of time to crack appearance during drying was based on the 

following relatively conservative assumption: 

- tension stress are close to yield stress, 

- presence of chloride or fluoride ions in the water (assisted corrosion cracking) and 

- presence of oxidizer (oxygen, hydrogen peroxide) in the water. 

The chloride concentration was assumed for the analysis at the saturation level, ~240g/kg. 

Time to appearance of crack during drying depends on impurities content in drying gas. For 

high purity gases (nitrogen and helium) time to cracking will be for about few thousand 

hours. For the gases with impurities (in particular oxygen) the time to cracking could be 




about a few hours. It was recommended to use for the drying only high purity gases and to 

perform drying at low temperature. 

3.5.2.2 Copper Alloys 

High thermal conductivity, high strength, satisfactory thermal and radiation resistance are 

required for the heat sink materials of the ITER first wall, divertor and limiter. Two Cubased 

alloys have been retained as candidate heat sink materials for the high heat flux 

components: CuCrZr and GlidCop ® Al25. Rationale for the selection of copper alloys is given 

in refs. [2, 16, 17, 18]. 

Starting from the original specification, a tighter limitation for composition and a specific 

heat treatment were proposed for the CuCrZr alloy, with the definition of an ITER grade (IG) 

specification, CuCrZr-IG. The proposed composition of CuCrZr alloy differs from the 

standard one, mainly by its narrower range of the Cr (0.6-0.9%), Zr (0.07-0.15%) content, 

limitation of oxygen (



creep-fatigue for both CuCrZr-IG and CuAl25-IG alloys, effect of manufacturing heat 

treatment on the properties degradation of CuCrZr-IG alloy and corrosion behaviour of 

copper alloys in the ITER working conditions. 

(1) Fracture toughness and creep/fatigue properties 

Low fracture toughness of CuAl25-IG is confirmed by various types of experiments, but the 

crack growth is stable and fracture is ductile [20, 21, 22, 23]. However, the initiation of crack 

nucleation does not appear to be instantaneous. Summarised data of fracture toughness tests 

are presented in Figure 2. Anisotropy of fracture toughness are observed for CuAl25-IG (see 

data in Figure 2 for T-L and L-T directions). Fracture toughness of CuCrZr-IG decreases with 

temperature, but remains considerably higher than in the case of CuAl25-IG. There is a big 

difference (approximately in two times) in fracture toughness of CuCuZr-IG measured by US 

and EU HTs (ORNL and VTT data). Further study is needed to determine the reason this 

difference in fracture toughness measurement. 

Fatigue behaviour of Cu alloy has been studied over the range 10 3 -2x10 5 cycles [24]. Results 

of creep-fatigue interactions and creep crack growth experiments suggest that the time 

dependent cyclic phenomenon may play an important role for the CuAl25 in determining the 

component’s lifetime. The number of cycles to failure decreases by about 25 times, if during 

pulse the hold time varies from 10 to 1000 second [25]. 

(2) Effects of heat treatment history 

Mechanical properties and thermal conductivity of CuCrZr alloy are sensitive to heat 

treatments (hence to joining technology). During component manufacturing, the base 

material is generally subjected to additional thermal cycles e.g. welding, brazing, HIPing, etc. 

Brazing, even at relatively low temperatures, results in a significant decrease of strength in 




the case of long-term exposure and/or slow cooling rates [26, 27]. Brazing or HIPing at high 

temperature can be combined with solution annealing. In this case the brazing/HIPing 

temperature should be about 950-980°C, followed by fast cooling and ageing at 475-480°C. 

The most critical step is the cooling rate from the brazing/HIPing temperature. If a fast 

cooling rate can be realised after high temperature brazing/HIPing, the loss of strength is 

minimised. About 80 to 90% of the strength of the "ideally" heat treated CuCrZr-IG can be 

achieved without deleterious effect of ductility, if the cooling rate (after the thermal exposure 

during HIP) is at least 70-180°C/min. 

CuAl25-IG shows much less sensitivity to heat treatment than CuCrZr-IG [28, 29]. Thermal 

stability is one of the main advantages of CuAl25-IG. This allows a wider flexibility in the 

use of different joining technologies for high heat flux components. 

(3) Irradiation effects on copper alloys 

Even low irradiation doses (~ 0.1-0.3 dpa) result in significant change of mechanical 

properties of Cu alloys. The irradiation results in increase of strength of Cu alloys at low 

temperatures, approximately below 300°C. Strength properties of unirradiated materials 

should be used for the design analysis as more conservative. At higher temperatures, 

approximately above 300°C, the strength of irradiated materials is lower, and those values 

(strength of irradiated materials) should be used [16, 30]. 

Ductility of CuCrZr-IG alloy increases with increase of irradiation temperature [16, 18]. At 

the irradiation temperatures < 230°C uniform elongation is below 2%, therefore criteria for 

immediate plastic strain localisation and fracture due to exhausting of ductility should be 




used for the design analysis [30]. At the temperatures exceeding 230°C material will be 

ductile (uniform elongation is above 2 %), and high temperature criteria should be used. 

The minimum value of uniform elongation of irradiated CuAl25-IG in the temperature range 

100-300°C is below 2% [30]. Therefore, in all temperature ranges, plastic strain localisation 

and local fracture criteria should be used for structural analysis. Combination of low ductility 

and decreasing of strength of material results in relatively low stress allowables that might be 

critical for the ITER components design. 

To take into account the embrittlement effect due to irradiation, two additional stress limit 

parameters, S e and S d , have been included in the ITER structural design criteria [30, 31]. 

These stress intensities can be used for the structural analysis of material to prevent fracture 

due to flow localization in materials with low work hardening capability (due to irradiation 

effect). 

(4) Corrosion behaviour 

Corrosion behaviour of copper alloys in the ITER working condition is one of critical issue 

for life-time and reliability of high heat flux components cooled by the water (divertor and 

baffle). Corrosion tests in static water (in autoclave with deaerated and demineralised water) 

have been performed by RF HT [15]. Corrosion rate varies from 0.3 up to 5 µm/y depending 

on exposure time and material. Test temperature was ~100°C. 

In the flowing water (up to 10 m/s) the corrosion rate increases. The measurements 

performed by the EU HT [14, 15] show the following: 

- Corrosion rate was observed ~ 17 µm/y for pure Cu and ~ 31 µm/y for CuAl25-IG at 

about 100 °C and under isothermal test conditions. 




- The corroded oxide layer was black, and there was no essential differences in the 

morphology and composition of oxide layer both for pure copper and CuAl25-IG alloy. 

- The erosion-corrosion was not clearly revealed during high flow water tests. 

- Hydrogen addition results in lower corrosion rate (and thin surface oxide layer). 

There was no evidence of galvanic corrosion effect in contact with stainless steel due to the 

very low electrical conductivity of the water. 

3.5.2.3 Ti Alloys Irradiation Effects 

For the blanket flexible supports, a flexible cartridge in the form of a "squirrel's wheel" is 

screwed into the vessel from one side and bolted through access holes in the blanket. The 

material should have a high strength, because of high axial loading force, and at the same 

time allow for a wide range of elastic deformation during bending. The estimated temperature 

range of the flexible cartridge is between 100 and 260˚C, and damage dose is ~ 0.1 dpa. 

Among the commercial Ti alloys, Ti-6Al-4V (α+β alloy) was proposed as reference grade. 

Since it is widely used, there is an industrial experience in many fields and the data base of 

unirradiated material is relatively complete. The R&D on Ti alloys included study of 

irradiation effect on the mechanical properties (tensile, fatigues, fracture toughness) and 

structural stability of alloys under irradiation [32, 33]. 

Ti alloy property change due to irradiation were studied for doses in the range 0.01-0.4 dpa 

and temperatures 40-350°C. The following general conclusion can be made from the results 

of R&D: 

- Strength of all materials is increased due to irradiation. 




- Ductility is decreased. However, in all experiments (even at dose ~0.4 dpa) the 

uniform elongation was above 2%. Therefore, structural criteria of ductile 

materials can be used for the design analysis. 

- There was no significant alteration of fatigue lifetime due to irradiation. 

- Fracture toughness decreases (2-5 times) due to irradiation. Nevertheless, fracture 

toughness remains at a relatively high level (~20-60 kJ/m 2 at 20°C). 

Alpha-based Ti alloy (Ti-5Al-2.4Sn) exhibited less embrittlement due to irradiation [32]. 

However, the strength of alpha-based alloy is lower than that of the alpha-beta alloy. Ti-5Al- 

2.4Sn alloy is therefore considered as a back-up option. 

It is known that Ti alloys might be susceptible to hydrogen embrittlement in certain 

conditions. The cartridge operates in the environment of the vacuum chamber, but it is 

shielded from the direct bombardment of energetic particles by the module itself. The 

possibility of hydrogen or helium implantation by energetic particles is, therefore, eliminated. 

Available data indicate that a relatively high hydrogen content of 3-5 wt. % is necessary to 

produce significant embrittlement, in terms of impact fracture energy, in Ti alloys. ITER 

R&D shows that hydrogen saturation of Ti-6Al-4V alloy up to 200 wppm of hydrogen did 

not result in changes of tensile properties both for unirradiated and irradiated material [33]. 

3.5.2.4 Ni-based Alloys Irradiation Effects 

High strength, fatigue and fracture toughness are required for the bolt attaching the shielding 

blanket modules through the flexible support to minimise the size of attached units and to 

provide reliability of the attachment. Inconel 718 is selected as the reference material due to 

its high strength and good ductility. Inconel 718 has satisfactory fracture toughness and 




fatigue lifetime. Data on the radiation behaviour of Inconel 718 is sparse. Results of 

irradiation tests show that irradiation up to 0.5 dpa will result in an increase of strength and in 

a slight decrease of ductility that will not affect the component structural integrity and 

lifetime [34]. 

Radiation induced stress relaxation is one of the critical issue for the bolts feasibility. Results 

of stress relaxation investigations performed by the EU HT [34, 35, 36] show that for the 

bolts of flexible cartridge fastenings where the irradiation dose is ~ 0.02-0.1 dpa, the stress 

relaxation will be ~ 80-90% from initial pre-load to the end of life, so, the required pre-stress 

should be ~ 800 MPa. 

3.5.3 Plasma Facing Materials 

The choice of plasma facing materials for different components is determined mainly by 

plasma compatibility and erosion lifetime (ion erosion, thermal erosion during disruption and 

transient events) [37]. 

- Beryllium has been chosen as the armour material for the first wall and port limiter, 

- Tungsten has been selected as the material for the divertor baffle area with high 

concentration of neutral particles due to lowest erosion rate, 

- Carbon fibre composites (CFC) have been proposed for the lower part of the divertor 

vertical target due to absence of melting during plasma disruptions. 

The key issues related to the plasma-wall interaction phenomena have been discussed 

elsewhere [37]. In this section the following R&D results will be highlighted: 

- evaluation of the different armour material grades and the selection of the reference 

grade for the ITER application, 




- justification of the material performance at ITER specific conditions (neutron 

irradiation effect, thermal fatigue/thermal shock resistance, etc.). 

3.5.3.1 Beryllium 

Based on the R&D results, S-65C VHP (vacuum hot pressed) grade (manufacturer Brush 

Wellman Inc., USA) has been selected as a reference grade. DShG-200, the RF Home Team 

and plasma sprayed (PS) Be are considered as a back-up option. 

The R&D activity has been focused on: 

- study of the thermal fatigue and thermal shock resistance of the different Be grades, 

- study of the neutron irradiation effect on the properties and thermal performance of 

Be, 

- development of the Be plasma spray technology and optimisation of the properties of 

PS Be. 

(1) Thermal performance 

The comparative studies of the thermal fatigue and thermal shock resistance have been 

performed by the US, EU and RF Home Teams [38, 39]. More than 15 different industrial Be 

grades were tested. The results are briefly summarised as follows: 

- S-65C VHP (vacuum hot pressed) and DShG-200 Be grades have been determined as 

materials with the best thermal fatigue resistance, 

- thermal shock tests (0.2-3 GW/m 2 , 5 ms) of the different grades show that S-65C 

VHP has generally better performance (lower thermal erosion and crack formation), 

- simulation tests of VDE demonstrated the structural integrity (absence of the 

macroscopic destruction) of the Be S-65C, 




- thermal shock tests of the neutron irradiated Be revealed the slight increase of the 

thermal erosion, He bubble formation has been observed in the re-solidified layers. 

(2) Irradiation effects on Be 

The comprehensive studies of the neutron irradiation effect on the beryllium properties have 

been carried out in all Home Teams, [40, 41, 42, 43] and these results were summarised in 

[44]. For ITER operational conditions there is no significant influence on the physical 

properties (including swelling). Generally, neutron irradiation causes an increase of tensile 

strength and a decrease of ductility and fracture toughness due to combined effects of neutron 

damage and helium generation. These are typical for all beryllium grades. The quantitative 

data for the recent grades have been generated. It was shown that for expected ITER 

conditions (damage ~ 1-1.4 dpa, irradiation temperature ~ 200-400°C) the behaviour of the 

recent Be grades (S-65C VHP and S-200F) is very similar and the ductility will remain in the 

level of ~ 1%. Experimental results of ductility change due to irradiation are presented in 

Figure 3. 

The thermal performance of the brittle beryllium (after irradiation) is a key issue. However, 

there are some positive experiments showing that high heat flux tests of neutron irradiated 

mock-ups did not reveal any damage in Be and the high heat flux tests of the cracked 

unirradiated Be did not reveal any detrimental behaviour due to cracking. 

Still more tests of neutron irradiated Be at relevant conditions (1000 appm He, 1-1.4 dpa) are 

needed, especially to simulate the VDE conditions and further thermal cycling behaviour. 

(3) Be plasma spray technology 




US Home Team has performed R&D on the development of the Be plasma spray technology, 

[45]. In the frame of this program, the significant improvements of the Be PS properties has 

been achieved, e.g. the thermal conductivity of the PS Be could be in the range 100-160 

W/mK at density ~ 92%. However, in comparison with S-65C, plasma-sprayed Be has lower 

thermal shock resistance and lower mechanical properties. 

3.5.3.2 Tungsten 

The R&D activity has been focused on the selection of the W grades based on the behaviour 

during thermal loads, simulating normal heat flux, disruption and thermal shocks (including 

neutron irradiation). Several W grades from different suppliers (pure sintered W, cast W 

alloys, W-1%La 2 O 3 , chemical vapour deposited W, monocrystal W, plasma sprayed W etc.) 

have been studied. The summary of the results could be found in [46]. 

Comparative tests of the low cycle fatigue resistance of the 10 different W grades have been 

performed by US HT and it was concluded that monocrystal W with proper orientation has 

the best resistance to thermal fatigue cracking, the next grade is the pure rolled W. 

The results of the thermal shock testing of the different W grades shows that weight loss of 

the W-1%La 2 O 3 is significantly higher than for pure W (JA HT) [47]. The thermal shock 

tests performed in EU HT, FZ Juelich have shown that the crack formation in W significantly 

depends on temperature and the neutron irradiation leads to increasing of the cracks 

formation, however, the thermal erosion was not increased, [48]. Thermal fatigue tests at 

15MW/m 2 revealed some surface crack formation in the W -1%La 2 O 3 tiles. 




The study of the different W grades at thermal fatigue and disruption simulation has been 

performed by the RF HT [49]. All tested tungsten grades show good thermal fatigue 

performance at proper armour design up to heat flux of 15 MW/m 2 . Rolled tungsten with 

texture orientation perpendicular to heat flow direction shows tendency of delamination at 

heat fluxes higher than about 5 MW/m 2 . Disruption simulation tests show that nearly all 

tungsten grades suffer from surface cracking except for single crystal tungsten. Thermal 

fatigue test at 20 MW/m 2 of mock-up after disruption simulation demonstrated that there is 

severe crack formation in almost all materials except for W-5Re-0.1ZrC alloy and W-0.02 Re 

single crystal alloy. 

Typically for bcc metals, neutron irradiation leads to an increase in the ductile-to-brittle 

transition temperature (DBTT) [44]. Based on the available data it could be concluded that all 

W grades have the problem of brittleness, at the expected fluence of 0.1-0.5 dpa in the region 

near the heat sink, for an irradiation temperature less than ~ 500°C. 

One of the findings from the R&D results is that the key issue is the proper design of the W 

armour (size of tiles and grain orientation). The recommended grain orientation is parallel to 

the direction of the heat flow, in which case possible delamination is not so critical. Based on 

the results of R&D the pure sintered W is recommended as reference material for divertor 

components. 

3.5.3.3 Carbon Fibre Composites 

The R&D activity has been performed in EU and JA HTs and has been focused on the 

following directions [50]: 

- development of the new advanced CFCs (including doped) with improved properties, 

- characterisation of these new materials, 




- study of the critical issues as neutron irradiation and thermal shock/disruption. 

Significant progress has been achieved by the EU HT in the development of the new CFCs. 

Several new grades as Dunlop, 3D CFC Sepcarb ® NB31, silicon doped CFC with improved 

resistance to chemical erosion Sepcarb-inox NS31 have been developed [51, 52]. The key 

property for CFCs is thermal conductivity, for Sepcarb ® NB31 it is in a range 320 W/mK at 

room temperature. 

JA HT has developed new advanced 3D CFC NIC-01 with thermal conductivity ~ 540 

W/mK at room temperature [53]. 

These new developed materials have been widely studied in comparison with existing grades 

with regards to effect of the neutron irradiation and thermal shock effect. For ITER condition 

(damage ~ 0.1-0.3 dpa and irradiation temperature 150-1200°C) the changes in properties are 

not expected to be crucial except for the thermal conductivity. CFCs with high initial thermal 

conductivity retain high conductivity after irradiation, but at low irradiation temperatures 

(less than 300°C) the thermal conductivity could be ~ 3-5 times lower than the unirradiated 

CFC [44]. 

The performances of mock-ups using the Sepcarb ® NB 31 and NIC-01 armour tiles have been 

studied including irradiation effects. Generally, the performance was as expected (higher 

temperature due to the loss of thermal conductivity). The decrease in the thermal conductivity 

due to neutron irradiation leads to an increase in thermal erosion under disruption conditions 

and this may have to be taken into account during the erosion lifetime assessment. 




The results of the thermal shock testing of the unirradiated and irradiated CFCs have been 

reported: it was shown that neutron irradiation leads to increasing the thermal erosion of 

~1.5-2 times for CFC [54, 55] 

For ITER application Sepcarb ® NB 31 and NIC-01 have been selected as a references grades. 

The back-up option is 2D felt-type CFC CX 2002U and Sepcarb-inox NS31. 

3.5.4 Characteristics of Joints of Different Materials 

The design of the ITER Plasma Facing Components (PFCs) involves various combinations of 

joints between armour materials and Cu alloy heat sinks and between Cu alloys and austenitic 

steels. The joints must withstand the thermal, mechanical and neutron loads and the cyclic 

mode of operation and operate under vacuum, while providing an acceptable design lifetime 

and high reliability. This paragraph summarizes the results of the R&D program on the 

development of joining technologies and properties of joints. 

3.5.4.1 Characteristics of Be/Cu Joints 

The main problem of bonding Be to Cu alloys is that Be reacts with almost all metals at 

moderate and high temperatures and forms brittle intermetallic phases that are detrimental for 

the joint reliability and the fatigue lifetime. To solve this problem and to provide good quality 

Be/Cu joints, different approaches have been studied by the ITER Home Teams. These 

approaches have been summarized in [56] as follows. 

- Use of materials as fillers or interlayers between Be and Cu alloy which do not form 

intermetallic phases with Be (e.g. Al, Ge, Si, AlSi or AlBeMet). The joining 

temperature for these materials is typically equal to or less than the melting 

temperature of these filler materials. The room temperature strength of the joint with 




pure Al interlayer is 120 MPa with ductile fracture inside the Al layer. For the 

AlBeMet-150 interlayer, the RT tensile strength increased to 195 MPa and 100 MPa 

at 300°C. 

- Use of diffusion barrier materials with less affinity for the formation of beryllide 

intermetallics. Different types of these barriers have been studied: Ti, Cr, Ti/Ni, Al/Ni 

Ti/Cu, Al/Ti/Cu, and Cr/Cu. Typically for this type of joints, HIP is used as a joining 

procedure with the temperature range 500 – 850°C. The shear strength of Be/Cu joints 

with Ti interlayer is ~ 108 MPa at room temperature. 

- Brazing with Cu-based brazing alloys. At a brazing temperature more than 650°C 

(typical temperature for these types of brazing alloys) intermetallic formation occurs, 

but using a "fast" brazing technique, this formation could be limited. 

- Direct bonding of Be to Cu alloy (with or without a soft intermediate layer of Cu) at 

moderate temperatures (~ 500-700°C). Direct Be plasma spray can also be carried out 

under this condition. 

The results of the high heat flux testing of the small-scale mock-ups are described in §3.3. As 

a part of the materials R&D program, several small scale Be/Cu mock-ups [57] have been 

irradiated (0.3 dpa, 350°C) and tested (up to 7 MW/m 2 ) after irradiation and Be/Cu brazed 

mock-up has been tested in the in-pile (simultaneous neutron 0.2 dpa and heat flux 3 MW/m 2 

irradiation) [58]. Based on these preliminary results, the conclusion could be made that 

neutron irradiation is not so critical for the durability of the Be/Cu joints. 

3.5.4.2 Characteristics of W/Cu Joints 

The main problem in the development of W/Cu joints is the large difference in the coefficient 

of thermal expansion (CTE) and of elastic modulus. From engineering point of view the 




solution is to use brush-like (rectangular or rod) or lamella type of the W armour design and 

use the soft intermediate layer between W and heat sink. 

The problem of the joining of W and Cu could be solved by different methods (see summary 

in [56]. One of the studied techniques is the casting of pure Cu onto W and further joining of 

the W/Cu to heat sink by different methods. The yield tensile strength of W/cast Cu joint is 

typically higher than yield strength of pure Cu and equal ~ 100-120 MPa. The brazing 

technology with CuMn also gives good results. The use of CuMn base brazing alloy 

demonstrated a good tensile strength of joint ~ 200 MPa with failure in Cu alloy near the 

brazed joint. Direct diffusion bonding of the W tile in the form of rods seems also promising. 

The use of W plasma spray gives good joining with Cu alloy heat sink. The use of chemical 

vapour deposition (CVD) of W onto Cu produces also good joints, but has limitations in 

thickness. Both these methods have been successfully applied for manufacturing of the 

components with curved surfaces. However, these technologies could be applied only for 

components with low and moderate heat flux. 

3.5.4.3 Characteristics of CFC/Cu Joints 

The problem in the development of CFC/Cu joints is worse than for the W/Cu joints due to 

the even larger difference in the CTE of the materials [56]. To provide a high quality CFC/Cu 

joint the surface of the CFC has to be activated to increase the wetting and a compliant layers 

between the CFC and Cu alloy heat sink is needed to relieve the residual stresses. Several 

joining technologies have been developed and studied. 

- Active metal casting (AMC ® ) technology, which was originally developed by Plansee 

AG for the Tore Supra limiters. This technology includes the special laser treatment 




of the CFC surface, followed by casting of pure Cu onto CFC, machining and final 

joining with the Cu alloy heat sink. In the monoblock configuration, CFC tiles with 

AMC ® Cu can be joined to Cu alloys by brazing or HIP. For flat tiles, the same joint 

could be obtained by e-beam welding. AMC ® technology was applied to different 

CFC grades (SEP N11, SEP N31, etc.). It was shown that tensile strength of joints 

produced by AMC ® technology is ~ 30 MPa which is less than strength of CFCs, 

whereas the shear strength is ~ 40-50 MPa. 

- Brazing with silver-free alloys (CuMn, CuSiAlTi) and HIP-assisted brazing with 

CuMn and CuTi alloys. The best high heat flux results have been observed for CuMn 

brazes 

As part of the materials R&D program, several small scale CFC/Cu mock-ups have been 

irradiated up to ~ 0.3 dpa at 350°C and 700°C. After irradiation this mock-up was tested in 

the JUDITH facility and survived 1000 cycles at 15MW/m 2 , [57]. Similar experiments with 

irradiated CFC-Cu mock-ups (irradiated up to fluence ~0.3-0.5dpa at irradiation temperature 

~300-340°C) have been performed at the OHBIS – facility by the JA Home Team [59]. No 

degradation of the heat flux performance (except of the expected change of the thermal 

conductivity) has been observed. Based on these preliminary results it could be concluded 

that neutron irradiation seems not so critical for the CFC/Cu joints. 

3.5.4.4 Characteristics of SS/Cu Joints 

Plasma facing components include the following joints for Cu alloys and stainless steel: for 

the first wall - copper plates with steel cooling tubes and plates; for limiter and divertor - 

copper structures with steel plate and/or copper and steel tubes. 

Generally, good bonding of both CuCrZr and CuAl25 alloys to SS was achieved. 




For HIP joining the use of CuAl25-IG alloy is preferable due to its excellent resistance to 

high temperature heating. Optimised SS/CuAl25-IG joints have good radiation resistance. 

Strength of the HIP bonded joints after 0.4 dpa irradiation dose at 150 and 300°C is higher 

than or equal to that of the base metal, with the failure location within the Cu part close to the 

interface [60]. 

Properties of CuCrZr/SS joints depend on the properties degradation of Cu alloy due to heat 

treatment used for joining technology. Providing fast cooling (~ 1-2°C/s) and applying 

further ageing, the properties of CuCrZr-IG after HIP are acceptable [27, 61]. This method 

gives high quality bonding and tensile properties of joints close to the strength of Cu alloys. 

Other joining methods such as explosion bonding seem promising and provides a good 

quality of joint, but it is applicable only for simple geometric shapes of the components. For 

Cu tube to stainless steel tube joining, friction welding, diffusion bonding (for DS Cu) and 

TIG welding with Ni interlayer (for CuCrZr-IG) can be used. The casting of CuCrZr onto 

stainless steel also gives a good quality of joint, that is also resistant to neutron irradiation 

[62]. 

The R&D results obtained for the joints lead to the following relationship of fracture 

resistance: 

- base material > joint sample > irradiated joint sample, 

- CuCrZr joints >>CuAl25 joints, 

- lower temperature data > same at high temperature. 

However, results are geometry dependent, and the samples do not generally fail at the 

interface. 




Thermal fatigue tests of the first wall mock-ups performed by the EU HT and JA HT did not 

reveal any failure of SS/Cu joints under the following test conditions: 30,000 cycles at 0.75 

MW/m 2 , and 1000 cycles at 5 MW/m 2 plus 1500 cycles at 7 MW/m 2 [65, 66]. 

Last generation of HIPed CuAl25/316L(N) and CuCrZr/316L(N) joints demonstrate high 

level of strength and satisfactory (non-zero) ductility after irradiation to 0.4 dpa at 150°C and 

300°C and to 2 dpa at 200°C [64]. The tensile strength of CuAl25/316L(N) and 

CuCrZr/316L(N) joints after irradiation (and test) at 200°C is about 350 MPa and 260 MPa, 

respectively, and the total elongation of CuCrZr/316L(N) joint is about 13% [67]. 

3.5.4.5 Characteristics of SS/SS Joints 

All austenitic stainless steels are readily weldable with a wide range of well-established 

methods, such as narrow gap TIG welding, electron beam welding, laser welding. The 

properties of the weld are well defined and compatible with the load requirement of these 

components. For TIG joints, the unirradiated welded material shows in general a higher yield 

stress, lower ultimate tensile strength and lower ductility than plate material. The effect of 

neutron irradiation on the tensile properties is similar to that for plate material, i.e. significant 

hardening and loss of ductility occurs [3]. It is expected that the worst ductility reduction 

occurs in the irradiation temperature range of 275-375°C, well above that appropriate for 

ITER. 

The joining of stainless steel to steel by HIP is considered as the main method for the 

manufacturing of the shield blanket modules. The quality of joint strongly depends on the 

quality of the base metal and on the surface preparation during the HIPing process. Most of 




the tensile data for SS/SS HIP joints in unirradiated state are within the design allowable of 

the base metal but close to the lower bound of the specified strength [2, 3]. 

Neutron irradiation has a slight effect on the mechanical properties of the joints, 

demonstrating a similar behaviour as for the base metal [3]. The last result of irradiation tests 

are presented in Table 2 and, in more detail, is given in ref. [11]. 

HIP bonded SS/SS joints and powder–HIPed SS seems more sensitive to irradiation 

degradation of the fatigue resistance in comparison with wrought SS [3, 5]. 

3.5.5 Ceramic Insulators 

Ceramic insulators are proposed to use for the limiter plates insulation and to provide 

electrical insulation of the first wall modules. The electrical resistivity changes due to 

irradiation have been studied within the R&D for diagnostic components. The objective of 

this investigation was to study the structural integrity of irradiated coatings under impact and 

compression loads and ability to maintain the electrical resistance after these loads. 

Specimens with ceramic insulator coating have been manufactured and tested before 

irradiation. These are Al 2 O 3 and MgAl 2 O 4 with different substrates. The breakdown voltage 

of coating is above 2500V, electrical resistance ~ 10 10 Ω and compression strength >200 MPa 

that exceeds ITER requirements. The impact tests performed by the JA HT show that the 

coating withstands without visible defects 10000 impact cycles with energy ~7-13 kJ/m 2 (by 

dropping steel hammer, [68]). Specimens with ceramic coatings have been irradiated in 

JMTR reactor. Neutron fluence was ~3x10 20 n/cm 2 (E>1 MeV), irradiation temperatures were 

150°C and 250°C. Post irradiation mechanical impact tests with impact energy ~ 7-13 kJ/m 2 

did not result in failure of coating up to 30000 impact cycles. 




3.5.6 Summary 

ITER R&D activity provides basis for materials selection and justification of materials 

behaviour for different components. The effort was focused on the optimisation of materials 

to meet ITER requirement, study of materials behaviour under ITER working conditions and 

study the effect of component manufacturing on the properties of materials. 

On the basis of analysis of available data and results of R&D, materials for the ITER 

components have been prescribed. These are both standard and industrially available 

materials (steels, Ti and Ni based alloys, tungsten, beryllium, etc.) and materials modified for 

better performance of the ITER components (316L(N) steel, Cu alloys and CFC). 

The results of the ITER R&D programme support this selection and provide the justification 

of the material performance under operational conditions. The key critical issues for each 

material have been analysed and assessed, and the operational windows have been 

determined. 




List of Tables 

Table 1 

Table 2 

Chemical composition (wt. %) of 316L(N)-IG stainless steel 

Tensile properties of 316L(N)-IG SS changes due to irradiation 

List of Figures 

Fig. 1 

Fig. 2 

Summary of tensile tests performed in different laboratories 

Fracture toughness of unirradiated Cu alloys. (T and L is transverse and 

longitudinal directions, respectively, indicating tension and crack propagation 

relative to the rolling direction of plate. References in figure: OMG [19], ORNL 

[21], VTT [22, 23]) 

Fig. 3 

Change of the total elongation of the Be S-65C VHP after neutron irradiation 




Table 1. Chemical composition (wt. %) of 316L(N)-IG stainless steel 

Element Reference grade: 316L(N)-IG 

min 

max 

C 0.015 0.030 

Mn 1.6 2.0 

Si 0.50 

P 0.025 

S 0.005 0.01 

Cr 17.0 18.0 

Ni 12.0 12.5 

Mo 2.30 2.70 

Nb+Ta+Ti, 

and Ta, 

and Nb 

0.15 

0.01 

0.01-0.10 * 

Cu 0.3 

B 0.001-0.002 * 

Co 0.05-0.25 * 

N 0.060 0.080 

* Notes: 

- Low cobalt,



700 

600 

316L(N)-IG, min 

s-HIP (entek, RF) 

s-HIP (JAERI, JA) 

p-HIP (entek, RF) 

p-HIP (Studsvik, EU) 

316L(N)-IG, av 

s-HIP (SEA, EU) 

s-HIP (SGM/GRE, EU) 

p-HIP (CEA) 

cast+HIP (Boeing, US) 

500 

400 

300 

UTS 

YS 

200 

100 

0 

0 100 200 300 400 500 600 

Temperature, °C 

Fig. 1. Summary of tensile tests performed in different laboratories (UTS - ultimate tensile 

strength, YS - yield strength) 




600 

500 

CuAl25-IG, T-L (OMG) 

CuAl25-IG, T-L (ORNL) 

CuAl25-IG, T-L (VTT) 

CuCrZr-IG, T-L (ORNL) 

CuCrZr-IG, T-L (VTT) 

CuAl25-IG, L-T (OMG) 

CuAl25-IG, L-T (ORNL) 

CuAl25-IG, L-T (VTT) 

CuCrZr-IG, L-T (ORNL) 

CuCrZr-IG, L-T (VTT) 

400 

300 

200 

100 

0 

0 100 200 300 400 


Fig. 2. Fracture toughness (J Q value) of unirradiated Cu alloys. (T and L is transverse and 

longitudinal directions, respectively, indicating tension and crack propagation relative to the 

rolling direction of plate. References in figure: OMG [19], ORNL [21], VTT [22, 23]). 

50 

40 

30 

20 

10 

0 

Average, unirradiated 

R. Chaouadi, 1-2.5 dpa 

L.Snead, 0.34 dpa 

M.Roedig, 0.35 dpa 

Ttest=Tirrad 

0 200 400 600 800 


50 

45 

40 

35 

30 

25 

20 

15 

10 

5 

0 

Fig. 3. Change of the total elongation of the Be S-65C VHP after neutron irradiation [39, 40, 

42] 




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