Report on creep and LCF-Data on ODS-Materials for GEN IV-Reactors

A Hobt1, A. Klenk

1

M. Serrano2, R. Hernandez

2

1 MPA University of Stuttgart

2 CIEMAT, Technoligy Department

Abstract

For the next generation of high temperature nuclear reactors new structural materials

are necessary to meet the requirements of the loading and sufficient life time.

Therefore several Oxide-Dispersion-Strengthened (ODS) ferritic-martensitic materials

have been development. Within the GetMat-project there have been testes different

materials, differing in chromium content, under creep, fatigue and creep-fatigue

loading at temperatures between 600°C and 750°C. The materials were made from

sheets and rods. The work was then continued in the MaTiSSE-project for tube

materials. This contribution deals with the evaluation of the uniaxial creep tests for

rod material with specimens extracted in different orientations and plate material.

Especially for structural materials the properties and material behaviour in different

orientations and different products are essential to know within the life-time

assessment and the evaluation of the operational loadings in terms of stress and

strain field analysis. This can be achieved by providing appropriate material laws for

an optimized description of the material behaviour.

Introduction

In order to keep up the competitiveness of the European nuclear power stations in the long run,

innovative solutions must be found, guaranteeing a more efficient and at the same time safe

operation of the power plants. The increase in efficiency normally goes along with an increase in the

operating temperature and pressure. At the same time the power generation is on the step to a new

demand on the operation due to a rising portion of renewable energies forcing a flexible operation of

the power plants. To meet the demand, new materials must be developed with increased strength.

Beside the material development, the behaviour of the materials under operating conditions must be

described using adequate material models.

For this purpose the GETMAT project was launched to address the cross-cutting aspects of classical

ferritic-martensitic materials and the new Oxide-Dispersion-Strengthened(ODS) materials for core

and primary circuit components. This envelopes the issues of qualification of the materials by

mechanical testing, investigation of joining and fabrications techniques and the development of

modelling techniques based on a fundamental understanding of the material. The present

publication deals with the mechanical characterisation of ODS alloys with different chromium

contents and the link to the microstructure which enables an interpretation of the material

behaviour.

Material characterisation and specimen fabrication

Two ferritic ODS alloys and one martensitic ODS alloy are characterised. A 14Cr ODS extruded bar,

supplied by CEA (J27 heat) and a 12Cr ODS and 9Cr ODS plates supplied by Kobelco in the framework

of the GETMAT 7FWP project [2]. The 14Cr ODS bar was produced by CEA by mechanical alloying f a

master alloy of composition (wt. %) 13.98Cr, 1.03W, 0.39Ti, 0.29Mn, 0.32Si, 0.17Ni with 0.3 wt%

Y2O3 under hydrogen atmosphere in a vertical attritor. Subsequently, the material was hot extruded

in the form of bar at 1100 ºC and annealed for 1.5 h at 1050 ºC [3]. By this process, there are

generated two regions with different mechanical properties and also material properties, which can

be clearly distinguished, see Figure 1a). The 12Cr ODS plate were produced with a pre alloyed metal

powder of composition (wt. %) 11.59Cr, 1.87W, 0.22Ti and 0.1Si was prepared by argon gas

atomisation method. The powders were mechanically alloyed in dry type attrition ball mill with 0.23

wt. % Y2O3 and subsequently extruded at 1150°C, hot forged at 1150°C and annealed at 1150°C for

1h. Finally, the forged plates were cold rolled with 40 % reduction and annealed to re-crystallisation

at 1200°C for 1 hour [4], see Figure 1b). The 9Cr ODS plate was produced my mechanical alloying T91

powders with 0.3 wt. % Y2O3. These powders were extruded at 1160ºC, hot forging at 1150ºC and

heat treatment at 780ºC for 1 hour and air cooled afterwards. Finally a cold working up to 40%

reduction were performed and heat treated for normalization at 1050ºC for 1 hour and tempered at

750ºC for 1 hour, [5].

The fabrication process leads to inhomogeneous grain size distribution comparing the longitudinal

and transversal direction. The microstructure of the 14Cr ODS bar show an elongated-grained

structure parallel to the extrusion direction, see Figure 3a.

Figure 1: Product forms from which the specimens have been extracted

a) 14Cr b) 14Cr c) 12Cr / 9Cr

The mean gran size in the transverse orientation is 432 nm length and 415 nm width, while the size in

the longitudinal orientation is 826 nm length and 496 nm width. A preferential crystallographic

orientation of the grains along <110>// to the extrusion direction is observed. A detailed

microstructure description of the 14Cr ODS bar can be found in [6]. Regarding the microstructure of

the 12Cr ODS plate, in all the sections studied it is possible to distinguish a bimodal grain size

distribution, existing small grains with sizes usually lower than ten micron and strange bent shapes

and some big and elongated grains along the extrusion direction which sizes are bigger than tens of

micron up to two hundred micron, see Figure 3b. This could be an indication of an incomplete

recrystallization on the fabrication process. EBSD maps show that the smallest grains seem to have

an orientation along <110> in the ED, while larger grains have not got a preferential orientation. No

detailed microstructure of the 9Cr ODS were performed but information can be found in [7].

Figure 2: Microstructure in the as-received state

a)

14

Cr

b)

12

Cr

Results from conducted tests and first derivation of material laws

Within the GETMAT project the 9-14 %Cr-ODS materials have been analysed by basic experiments,

comprising, both in air atmosphere and Argon atmosphere at high temperatures:

• Tensile tests at different strain rates

• Creep tests

• Fatigue tests

Based on these tests it is possible to describe the material behaviour with macroscopic material

models which then can be used in the analyses of components under high temperature operation.

For the design of the creep tests and the fatigue tests, plastic flow behaviour of the materials was

derived by tensile tests. To quantify the influence of the loading speed, the strain rate of the tensile

tests was changed for the 9Cr ODS and the 14Cr ODS materials. Due to the high temperature of

750°C, all tests at MPA Stuttgart on the 14Cr ODS material were conducted in Argon atmosphere at

ambient pressure. The tests on the other two material were conducted in ambient air.

The basic material characterisation was done by tensile tests. At MPA Stuttgart, the specimens were

all extracted in longitudinal direction. The testing temperature was the main temperature also for

the further investigations, see Table 1, together with the main results. At CIEMAT, also specimens in

transversal direction have been analysed. At CIEMAT also tensile tests were performed, using with

dog-bone specimens (gage section 15x3x2 mm) mechanised in longitudinal (L) and transverse (T)

orientation for the 14Cr ODS and 12Cr ODS steel, defined as the direction of the loading during the

test, for the bar and plate. Tensile tests were performed at 22°C, 400°C, 600°C and 700°C a

displacement rate of 0.1 mm/min (corresponding to a strain rate of 1x10-4

/s) in a servo-hydraulic

MTS testing machine. Total elongation and reduction in area measurements were performed on the

broken specimens. A summary is given in Table 2.

Table 1: Summary of conducted tensile tests, MPA Stuttgart

Specimen Material Temp / °C Strain rate / 1/s UTS / MPa Strain at failure /%

2AB4

14Cr ODS 750

10-4 315 1,8

2AB1 2·10-5 315 1,34

2AB2 2·10-5 306 1,31

2AB3 10-6 283 0,54

X2C/2D 12Cr ODS 650 10-5 402 13,8

9CB2

9Cr ODS 600

10-3 592 37

9CA2 10-4 488 35,5

9CB1 10-5 394 27,9

9CA1 10-6 354 16,4

Table 2: Summary of conducted tensile tests, CIEMAT

Specimen Material Temp / °C Strain rate / 1/s UTS / MPa Strain at failure /%

14B13-L1

14Cr ODS

24 1.00E-04 1112 15 14B13-L4 24 1.00E-04 1131 20 14B22-L2 27 1.00E-04 1123 22 14B22-L1 29 1.00E-04 1127 22 14B22-L6 400 1.00E-04 910 14

14B13-L23 400 1.00E-04 905 17 14B13-L5 400 1.00E-04 893 20 14B22-L3 600 1.00E-04 461 12 14B22-L4 600 1.00E-04 471 11

14B13-L24 600 1.00E-04 423 22 14B22-L7 700 1.00E-04 396 7 14B22-L8 700 1.00E-04 356 9 14B22-L8 700 1.00E-04 356 9

14B13-L30 700 1.00E-04 400 12 14B13-L2 700 1.00E-04 374 7 14B13-L3 700 1.00E-04 379 6

14B22-L10 800 1.00E-04 300 2 14B22-L11 800 1.00E-04 298 7 14B22-L12 800 1.00E-04 307 6 14B13-L6 800 1.00E-04 292 2 14B13-L7 800 1.00E-04 269 8

12L1

12Cr ODS

24 1.00E-04 1160 12 12L2 24 1.00E-04 1138 10 12L3 400 1.00E-04 861 7 12L4 400 1.00E-04 896 11 12L5 600 1.00E-04 482 20 12L6 600 1.00E-04 478 21 12L7 700 1.00E-04 317 9 12L8 700 1.00E-04 335 11

12L10 800 1.00E-04 258 7 12L9 850 1.00E-04 229 1 9L1

9Cr ODS

24 1.00E-04 1098 15 9L4 25 1.00E-04 1098 17 9L2 400 1.00E-04 864 6 9L3 400 1.00E-04 888 12 9L5 600 1.00E-04 397 22 9L6 600 1.00E-04 393 21 9L7 700 1.00E-04 229 9 9L8 700 1.00E-04 237 9 9L9 800 1.00E-04 170 10

In Table 3 and Table 4 the parameters and the main results of the conducted creep tests are

summarised. Again, the creep test specimens at MPA Stuttgart were all tested in longitudinal

material direction. Some of the specimens have been dismounted at the end of the GETMAT project.

The tests were stopped because the results were sufficient for the characterisation in correlation

with the commonly used ferritic-martensitic steels such as the P92. And also the remaining testing

time could not be estimated precisely.

Table 3: Summary of conducted creep tests, MPA Stuttgart

Specimen Material Temperature / °C Stress / MPa Time to rupture / h Strain at failure /%

3A1/B1

14Cr ODS 750

250 10.3 2.89

3A2/B2 230 15.6 1.36

4A3/B3 170 227.3 1.68

4A2/B2 150 4,859.4 0.734

4A1/B1 130 dism. after 6.190 -

X3A/X3B 12Cr ODS 650 180 dism. After 4369 -

9CB3 9Cr ODS 600

170 2301 1,7

9CA4 250 199 4,2

Table 4: Summary of conducted creep tests, CIEMAT

Sepecimen Material Temperature T / °C Stress / MPa Time to rupture / h

CIEMAT-J27 Longitud.

14Cr ODS

600 205 7416

600 281 8088

600 326 216

600 387 528

CIEMAT-J27 Tranverse

600 382 48

600 278 72

600 205 216

CIEMAT-J27 Longitud.

650 182 4248

650 184 2856

650 204 2712

650 225 888

650 227 2856

650 253 192

650 255 216

650 255 48

650 281 48

650 281 96

650 281 24

650 324 24

650 324 48

After the tensile tests, there have been also conducted fatigue tests at MPA Stuttgart according to

the summary in Table 5. For the 9Cr ODS and the 12Cr ODS there have been conducted creep-fatigue

tests with a hold time of 10 minutes in the tension section. The creep fatigue tests for the 14Cr ODS

could not be finished due to buckling of the specimens.

Table 5: Summary of conducted fatigue tests

Specimen Material Temperature / °C Total strain range / % Cycles to failure / -

A1/B1

14Cr ODS 750

0,5 9101

A3/B3 0,8 397

C1/D1 0,6 1458

E3/F3 1,5 9101

X1A/X1B 12Cr ODS 650

0,8 778

X1C/X1D 0,8 (10 Min HT in tension) 379

9CB6

9Cr ODS 600

0,5 17108

9CA6 0,6 14692

9CB5 0,8 1953

9CA5 1 1040

9CB7 0.5 (10 Min HT in tension) 3097

9CA7 0.8 (10 Min HT in tension) 246

In the following section, the results from the described tests above are shown in more detail. In

Figure 3 the results on the 14Cr ODS material are shown. It can be seen, that at different strain rates

at the highest temperature the strength is on almost the same level, but the elongation at failure is

decreasing with decreasing testing speed, Figure 3a). When the transverse and the longitudinal

direction is compared, it can be seen, that the longitudinal direction has a higher strength, tested for

different temperatures, Figure 3b).

Figure 3: Tensile test results on 14Cr ODS

a) longitudinal b) longitudinal and transverse

The tensile test results for the 12Cr ODS are shown in Figure 4a). It was conducted one tensile test

only, to have a comparison of the strength with the 9Cr- and 14Cr material to be able to design the

loads for the fatigue tests and the creep tests. The results for the conducted tests on the 9Cr ODS

0,0 0,5 1,0 1,5 2,0 2,5

0

50

100

150

200

250

300

350

2AB3

2AB2

2AB1

2AB4 [dε/dt = 10-4/s]

2AB1 [dε/dt = 10-5/s]

2AB2 [dε/dt = 10-5/s]

2AB3 [dε/dt = 10-6/s]

14Cr-ODS, T = 750°C,

Argon-Athmosphere

Str

ess σ

/ M

Pa

Strain ε / %

2AB4

material are shown in Figure 4b). For the 9Cr material a strong influence of the strain rate can be

seen, not only on the strength properties but also on the ductility. The 9Cr ODS and 12Cr ODS exhibit

larger ductility than the 14Cr ODS, although the temperatures are lower. This behaviour is not known

from the classical 9-12Cr ODS martensitic materials.

Figure 4: Tensile test results on 12Cr and 9Cr ODS

a) 12Cr b) 9Cr

For the analysis of the long-term behaviour of the materials, creep tests were conducted. For the

tests in Argon atmosphere, the specimen was put in a elastic, soft bellow, which was filled with

Argon at a slight overpressure against ambient conditions to ensure that the specimen is not getting

in contact with air, see Figure 5.

Figure 5: Test setup for the creep tests in Argon atmosphere

Digester with soft bellows

The tests at MPA Stuttgart where conducted on specimens out of the longitudinal direction. At

CIEMAT there have been conducted several specimens in transverse direction, see Figure 6. It can be

seen, that the ODS alloys have a higher creep strength than the classical P92 martensitic material, in

case of the 14Cr ODS even at a higher temperatures. Also shown are results obtained at JRC at 650°C

in air, [2].The 9Cr ODS and the 12Cr ODS alloys are on the level of the 14Cr ODS at 650°C. It has to be

mentioned, that the 9Cr ODS alloy was tested at 600°C. Comparing only different testing

temperatures of the 14Cr ODS experiments, the expected increase in creep strength with decreasing

temperature is evident. What also can be seen, that beside the high scatter of the material itself, the

specimens from transverse direction exhibiting a lower creep strength than the specimens form

longitudinal direction. This will be discussed in connection with the microstructural investigations.

Beside the high creep strength also the ductility must be considered. Looking at the creep strain

0 2 4 6 8 10 12 14

0

50

100

150

200

250

300

350

400

450

500

12CrODS

T = 650 °C

= 10-5/s

Str

ess σ

/ M

Pa

Strain ε / %

End of clip gauge

control

ε&

0 10 20 30 40

0

50

100

150

200

250

300

350

400

450

500

550

600

650

700

750

9CB2 [dε/dt = 10-3/s]

9CA2 [dε/dt = 10-4/s]

9CB1 [dε/dt = 10-5/s]

9CA1 [dε/dt = 10-6/s]

9Cr-ODS, T = 600°C

Str

ess σ

/ M

Pa

Strain ε / %

evolution, see Figure 7, it can be seen, that with increasing testing time, the creep ductility decreases

and is generally on a level between 2-4% which is classically defined as creep brittle. This must be

taken into consideration when assessing these material parameters in a structural component design.

Figure 6: Creep strength diagram, ODS materials

a) Comparison different ODS alloys b) 14Cr ODS, diff. temperatures / directions

Figure 7: Creep deformation diagrams

a) 14Cr ODS b) 9Cr ODS

The fracture anisotropy of the 14Cr ODS and 12Cr ODS steels has been characterised by tensile,

impact and small punch tests, see [9]. The longitudinal orientation, parallel to the extrusion/rolling

direction, shows in general better mechanical properties than the transverse orientation. This

anisotropy found in the fracture behaviour could be attributed to the elongated grained structure

and to the grain boundary weakness due to the presence of micron size oxide particles. In this paper

different behaviour is found for the creep properties of the 14Cr ODS and 12Cr ODS steels. For the

14Cr ODS the anisotropy on the creep properties is clear, being the transversal specimens less

resistant to creep strength. Whereas the 12Cr ODS do not show an appreciable effect of specimen

orientation. This different behaviour can be due a very different grain morphology and a different

grain aspect ratio, being higher for the 14Cr ODS than the one for the 12Cr ODS, were the grains are

0,1 1 10 100 1000 10000 100000

0

50

100

150

200

250

300

350

400

450

dism.( )

running at the End of GETMAT

14CrODS [750 °C, Argon]

12CrODS [650 °C]

9CrODS [600 °C]

14CrODS [650 °C, Data from JRC, air]

all longitudinal direction

P92 [650°C]

Mean Data

Cre

ep r

uptu

re s

tre

ss R

u/t

/T /

MP

a

Time to rupture tU / h

*

* Stress increased from

120 Mpa to 180 MPa

dism.( )

0,1 1 10 100 1000 10000 100000

0

50

100

150

200

250

300

350

400

450

dism.( )

running at the End of GETMAT

14CrODS [750 °C, Argon]

14CrODS [650 °C, Data from JRC, air]

14CrODS [650 °C, Data from CIEMAT, air]

14CrODS [600 °C, Data from CIEMAT, air]

14CrODS [600 °C, Data from CIEMAT, air, Trans.]*

*rest in longitudinal direction

P92 [650°C]

Mean data

Cre

ep r

uptu

re s

tre

ss R

u/t

/T /

MP

a

Time to rupture tU / h

*

* Stress increased from

120 Mpa to 180 MPa

0,1 1 10 100 1000 10000

0,01

0,1

1

10

B BB

3A1/3B1 [σn = 250 MPa]

3A2/3B2 [σn = 230 MPa]

4A3/4B3 [σn = 170 MPa]

4A2/4B2 [σn = 170 MPa]

3A3/3B3 [σn = 150 MPa]

4A1/4B1 [σn = 130 MPa]

Cre

ep s

train

εc /

%

Time t / h

14CrODS

T = 750 °C B

test dur.

4502 hrs.

0,1 1 10 100 1000 10000

0,01

0,1

1

10

9CB3 [σn = 170 MPa]

9CA4 [σn = 250 MPa]

Cre

ep s

train

εc / %

Time t / h

9CrODS

T = 600 °C

larger but equiaxed and thus can enhance creep strength because grain boundary sliding is inhibited,

see [10].

Figure 8: Determination of anisotropy

a) b)

c) d)

The conducted low-cycle-fatigue tests performed at MPA are shown in Figure 9, together with the

results obtained at JRC in air medium and from ENEA in GFR atmosphere, [2].

Figure 9: Low-cycle-fatigue

It can be seen, that the results in Argon are creating a common scatter band with the results in GFR

at lower strain ranges. Compared the results in air from JRC, there can be seen no influence of the

temperature. What is worth mentioning is that the results from the 12Cr ODS, test in air, matching

the results from the 14Cr ODS in Argon but at a lower testing temperature and the 9Cr ODS results

come together with the results in air of the 14Cr ODS alloy, but the temperature is also markedly

lower. The effect of a hold time seems to be as expected based on the knowledge, see for example

[3], [4].

First modelling approach

Based on the results obtained from the basic material tests, preliminary modelling approaches could

be derived. Based on this modelling it is possible to adapt assessment procedures, e.g. according to

the standards, for the design of structural components in high temperatures power plants. For the

numerical analysis of transient loading conditions, the inelastic strains must be calculated to evaluate

stress and strain ranges during cyclic operation. Therefore the dependence on the strain rate is

necessary to describe the materials’ behaviour correctly. A very simple approach was chosen, based

on the tensile tests, conducted at different strain rates. The modelled stress-strain curves for the 9Cr-

ODS material are shown in Figure 10. To model the influence of the strain rate, a visco-plastic model

based on a formulation from Chaboche was used, [3], [4], see Eq. (1)-(3).

Figure 10: Modelling approach to describe strain rate dependency

�� �� = ����

�� ����� (1)

�� = � ���������� ��

(2)

�� = �� ���� �� − ���� �� = ��� − ���� (3)

Here inε& is the total inelastic strain rate tensor and p& is

the derivation in time of the accumulated inelastic strain

10 100 1000 10000

0,1

1

GETMAT results

running at the

End of GETMAT

Tota

l str

ain

ran

ge

∆ε

A / %

Cycles to failure NA / -

14Cr [650°C, data from JRC, Air]

14Cr [data WP2 (750 °C, GFR env.)]

14Cr [750°C, data from JRC, Air]

14Cr [750°C, Argon]

12Cr [650°C, Air]

12Cr [650°C, Air, tH = 10min]

9Cr [600°C, Air]

9Cr [600°C, Air, tH = 10min]

Manson-Coffin 14Cr

Manson-Coffin 9Cr

*( )* not evaluated, thread failure

10 100 1000 10000

0,1

1

To

tal str

ain

ra

ng

e ∆

ε A / %

Cycles to failure Nf / -

14Cr [650°C, data from JRC, air]

14Cr [data WP2 (750 °C, GFR env.)]

14Cr [750°C, data from JRC, air]

14Cr [750°C, Argon]

Manson-Coffin-Fit

εel

εpl

The next step in describing the material behaviour was to describe cyclic loading by modelling the S-

N curves to gain the relevant number of cycles to crack initiation. This can be done by using a

Manson-Coffin equation, see Eq. (4)-(5).

( ) ( )c'

f

b'f

pl,ael,aa N2´N2E

⋅+⋅

=+= ε

σεεε (4)

c

b'n =

( )c

b

'

f

'f'K

=

ε

σ (5)

The parameters have been fitted based on the stable hysteresis loop at mid-life. The strain amplitude

was splitted in the elastic part εa,el and plastic part εa,el. Within the equation the yield strength σf and

the Young’s Modulus E, gained from the tensile tests was adopted. The results for the 14Cr ODS and

9Cr ODS alloy is shown in Figure 11. Since for the 12Cr ODS only one specimen was tested, it was not

possible to gain the equations for this alloy.

Figure 11: Possible modelling of the fatigue behaviour

In the last step, the creep behaviour was modelled using a modified creep law based on a Graham

and Walles formulation, [5], see Eq (6) and Eq. (7). This formulation was developed at MPA Stuttgart,

[6], the creep strain rateε& is dependent on the creep strain e and the von Mises stress σvM. The

tertiary creep stage is described using a damage parameter D. To account for the dependency of the

damage evolution on the multiaxiality of the stress state, the parameter q according to [7] was

introduced.

( ) ( )

2

2

21 m

v

n

Am

v

1n

1A

D1vM10

D1vM10 ε

σε

σε ⋅

−⋅+⋅

−⋅=&

(6)

2

2

21

2

1 ~~ 310

310 mD

nD

ADmD

nD

AD

vMqvMqD εσεσ ⋅

⋅

α

⋅⋅

⋅

α

⋅= +

& (7)

This was applied on the creep strain curves gained from the 14Cr ODS material, since there were

available the most creep data. The results is shown in Figure 12. As it can be seen, there is a strong

sensitivity to the stress. Therefore the material law is not capable of describing the full range of test

stresses. The material parameters, the exponents of the creep equations, have been found to fit best

the creep strain curves in the long term regime.

10 100 1000 10000

0,1

1

( )* not evaluated, thread failure

T

ota

l str

ain

ra

ng

e ∆

ε A / %

Cycles to failure Nf / -

14Cr [650°C, Data from JRC, air]

14Cr [750°C, Data from JRC, air]

9Cr [600°C, air]

9Cr [600°C, th = 10 min, air]

Manson-Coffin-Fit

εel

εpl

( )*

Figure 12: Modelling of the creep deformation

Microstructural analysis and link to experimental results

Based on the microstructural investigations, the behaviour in the experimental results shown above

can be explained. This will comprise the creep behaviour and the material behaviour under cyclic

loading and it is necessary for a safe design for these materials.

In Figure 13 the specimen 4A2/B2 is shown after the rupture. As it can be seen, the fracture was

brittle- there is less deformation in the ruptured cross-section. On the microsection cavities are

visible and secondary cracking along the grain boundaries in longitudinal direction. This is linked with

the low creep ductility after rupture. What must be mentioned are the damage phenomena which

might be influenced by the Argon atmosphere.

Figure 13: Example of microstructure of 14cr ODS after creep loading

As it can be seen from TEM analysis, Figure 14, the microstructure does show relevant changes

during creep loading. The example is taken from the creep test specimen on 14Cr ODS alloy, tested at

750°C. The pictures are showing the longitudinal direction exemplarily, since the findings in the

transverse direction are similar. When evaluating the dislocation density and the subgrain size there

can be measured only small changes. Again this validates the brittle behaviour without the softening

0,01 0,1 1 10 100 1000 10000

0,01

0,1

1

10

3A1/3B1 [σn = 250 MPa]

3A2/3B2 [σn = 230 MPa]

4A3/4B3 [σn = 170 MPa]

4A2/4B2 [σn = 170 MPa]

3A3/3B3 [σn = 150 MPa]

4A1/4B1 [σn = 130 MPa]

FIT G&W

Cre

ep

str

ain

εc / %

Time t / h

14CrODS

T = 750 °C

of the material through classical damage mechanisms such as precipitation coarsening or other

changes of the microstructure.

Figure 14: TEM analysis of microstructure, 14Cr ODS, longitudinal direction,

As

rece

ive

d

Aft

er

cre

ep

loa

din

g

t u =

4

85

9 h

, T

= 7

50

°C

Regarding the cyclic behaviour of the ODS materials, the high strength of the material will cause a

very strong dependency of the microstructure. At a strain range of 1.5 % which includes plastic

portions, there can be seen, that the fracture is not homogeneously perpendicular to the loading

direction, but along the grain microstructures of the material, see Figure 15. On the fracture surface

the striations can be seen clearly. But there are also secondary cracking visible which might be due to

the stress concentrations at the grain boundaries due to their small amount of deformability.

Figure 15: Microstructural analysis of 14Cr ODS LCF specimen, T = 750°C, ∆ε∆ε∆ε∆ε = 1.5 %,

The effect of the fracture along the laths of the grain structure is more evident for the 9Cr ODS and

the 12Cr ODS material, although their deformability in the first glance at the tensile tests is larger

than for the 14Cr ODS. In Figure 16, the micrographs are showing a LCF specimen of 9Cr ODS alloy

showing a crack path where the angle between the fracture surface and the loading direction is

changing. On the fracture surface there can be seen also secondary cracking. In Figure 17

micrographs of a specimen, also from 9Cr ODS, with the same strain range but with a hold time in

tension section is shown. Here the secondary cracking along the grain structure in longitudinal

direction is more evident.

Figure 16: Microstructural analysis of 9Cr ODS LCF specimen, T = 600°C, ∆ε∆ε∆ε∆ε = 0.8 %

Figure 17: Microstructural analysis of 9Cr ODS LCF specimen, T = 600°C, ∆ε∆ε∆ε∆ε = 0.8 %, th = 10 min

Conclusions

On different ODS alloys at their probable operating temperature different mechanical

characterisations have been conducted from which the fundamental yield properties, fatigue

behaviour and the creep properties could be derived. Based on these results component design

according to the standards became possible. However, the results are showing that the material

behaviour exhibits higher strength at high temperatures than the classical martensitic steels. This

was the aim when producing the ODS alloys. But there are some issues that must be considered

within the component design. The first point is the brittle material behaviour which influences safety

factors and the design of the component. The second point is due to the anisotropic material

structure. It must be considered, that the strength of the material is not isotropic. Therefore

especially for components multiaxial loading must be closely analysed.

Based on this the behaviour under multiaxial loading in creep and creep fatigue regime must be

analysed and the material behaviour must be characterised.

Outlook to the next phase

The next steps is the mechanical characterisation of cladding tubes for Generation IV nuclear

reactors made from ODS materials with the focus on multiaxial loading. This will be the work within

the MATISSE project which is ongoing. The testing of the cladding tubes is done under internal

pressure generating a two dimensional state of stress at the minimum. Since joining by welding is

difficult for those materials and the thin walled tubes in particular, first a mounting device must be

constructed. A first design was provided by VTT, which is the leader of the workpackage for the

mechanical characterisation.

Acknowledgements

The presented work was funded by the EU within the GETMAT-Project, GA No. FP7-212175.

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