1

Creep and Elevated Temperature Properties of Aluminum

1.0 Experimental Procedure

1.1 Material

Two types of aluminum alloy were supplied by Virginia Tech for the purposes of determining

high temperature static and creep properties.

One type of aluminum alloy was formed via extrusion, and is referred herein as Type A. This

material was supplied in the form of extruded strips of dimensions 6.60 mm, 50 mm (width)

and 600 mm (length).

The other type of aluminum alloy had the same composite as Type A, but was formed by

strain hardening (i.e. hot rolled) and in this report is refereed to as Type B. This material was

supplied as rectangular plates of dimensions 600 mm × 300 mm and two thicknesses. One

plate had a thickness of 7.60 mm (referred to herein as B-Thin) and the other with a thickness

of 9.90 mm (referred to herein as B-Thick).

1.2 Methodology

1.2.1 Specimen Dimensions

All specimens for high temperature tests were machined into dog-bone shaped coupons.

Extruded samples (Type A) had a necked gauge section of 10 mm wide, 6 mm thick and 300

mm long. The strain-hardened aluminum specimen; B-Thin had a necked gauge section of 10

mm wide, 7.60 mm thick and 300 mm long. Specimen from sample B-Thick had a necked

gauge section of 10.0 mm wide, 9.90 mm thick and 300 mm long. Figure 1 shows the dog-

bone shaped test specimens used in this study. The extrusion direction (for Type A) and

rolling direction (for Type B) was aligned in the length-wise direction of the specimens.

Figure 1: Dog-bone specimen used for static tests.

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1.2.2 High Temperature Static Tests

Elevated temperature static tests were performed in tension loading for temperatures ranging

between 20 and 460°C. The specimens were centrally heated over a 100 mm long section of

the gauge region using a temperature-controllable heating device. Prior to loading in tension

at a loading rate of 0.5 mm/min, the specimens were allowed to equilibrate at the set test

temperature for 10 min. Figure 2 shows the experimental set-up of the high temperature tests.

The ends of the specimen show dimples made by the grips of the testing machine. From the

calculated stress-strain relationships, the Young’s modulus and 0.2% proof strength of the

aluminum alloys were determined.

Figure 2: Experimental set-up for tensile and creep testing.

1.2.3 Creep Tests

Using the same test set-up as the elevated temperature static test shown in Figure 2, high

temperature creep tests were performed at constant temperature and constant stress on the

dog-bone shaped specimens. The specimens were held at the creep test temperature until they

had reached thermal equilibrium, after which they were pre-loaded at a rate of 3 kN/min to the

pre-determined applied creep stress. Creep tests were performed for constant stress levels

ranging between 15 and 200 MPa at constant temperatures ranging from 20 to 460°C for both

types of aluminum alloy. The creep behavior (i.e. strain evolution as a function of time) was

measured for the entire test until the specimen failed via stress rupture. For Type A

specimens, the creep behaviour was measured at temperature ranges and stress levels that

allowed the specimens to fail within a short period of time realistic of the duration of a fire on

naval vessel (i.e. less than several hours). However, creep tests for sample B-Thick were

performed in two sets: (1) constant temperature with varying applied stress and (2) constant

applied stress and varying temperature. This procedure was adopted since it allows the

3

extraction of creep parameters for use in thermo-mechanical (analytical and numerical)

models.

2.0 Results

2.1 Elevated Temperature Static Properties

2.1.1 Sample A: Extruded Aluminum Alloy

Figure 3 shows the tensile stress-strain curves for selected elevated temperature tests

performed for sample A; i.e. the extruded aluminum alloy. For this sample, the test

temperatures were varied from 25 to 430°C.

Figure 3: Stress-strain curves for sample A.

The elastic moduli and 0.2% proof strength of specimens from sample A calculated for

different test temperatures are shown in Figures 4 and 5. The mechanical properties decrease

with increasing test temperatures. A fitting tanh function was used to mathematically fit the

elastic moduli and proof strength to test temperatures using equations (1) and (2),

respectively.

( ) ( )( )%50tanh22

TTkEE

TE RTRT −−= (1)

( ) ( ) ( ) ( )( )%50

%2.0%2.0

%2.0 tanh22

TTkTRTRT

−−=σσ

σ (2)

where ( )TE and ( )RT%2.0σ are the Young’s modulus and the 0.2% proof strength at room

temperature, respectively; k is a material fitting constant describing the breath of the property-

0.000 0.002 0.004 0.006 0.008 0.0100

50

100

150

200

250

300

430 oC

363 oC

326 oC

290 oC

202 oC

110 oC

25 oC

Str

ess (

MP

a)

Strain (mm/mm)

4

temperature dependency and %50T is the temperature at which the room temperature measured

property value has dropped by 50%. These values are given for respective property-

temperature dependencies in Figures 4 and 5.

Figure 4: The effect of temperature on the Young’s modulus of the extruded sample (A).

Figure 5: The effect of temperature on the 0.2% proof strength of the extruded sample (A).

The experimental data obtained from elevated static tests for sample A (elastic moduli and

0.2% proof strength) were compared to a tanh fitting functioning describing the Eurocode 9

data [1,2] and Langhelle and co-workers derived data for aluminum alloy 6082-T651 which is

contained in an article published by Fogle et al. [3]. These results are shown in Figures 6 and

7. The degradation in elastic moduli as a function of temperature for the extruded sample is

similar to that observed for 6082-T651. However, these data are significantly different from

0 100 200 300 400 500 6000

10

20

30

40

50

60

70

80

ERT

=78.0 GPa

k = 0.007207

T50%

= 373.61 oC

Ela

stic

Modulu

s (

GP

a)

Temperature (oC)

0 100 200 300 400 500 6000

50

100

150

200

250

300

σ0.2%,RT

=285.1 MPa

k = 0.008444

T50%

= 283.79 oC

0.2

% O

ffset Y

ield

Str

ess (

MP

a)

Temparature (oC)

5

0 100 200 300 400 500 6000.0

0.2

0.4

0.6

0.8

1.0

Eurocode 9 data

Aluminum Type A data

No

rma

lise

d E

lastic M

od

ulu

s

Temperature (oC)

0 100 200 300 400 500 6000.0

0.2

0.4

0.6

0.8

1.0

Aluminum Type A data

Eurocode 9 data

No

rma

lise

d 0

.2%

Off

se

t Y

ield

Str

ess

Temperature (oC)

observations made for 5083-H116 alloy specimens previously tested in our research facilities.

The proof strength comparison (see Figure 7) of the extruded sample and 6082-T651,

however, reveals significant differences in some of their mechanical properties as function of

temperature.

Figure 6: Comparison of experimental Type A Young’s modulus data from this study to

Eurocode 9-derived data for the 6082-T651 aluminum alloy [1,2].

Figure 7: Comparison of experimental proof strength data from this study to Eurocode 9-

derived data for the 6082-T651 aluminum alloy [1,2].

2.1.2 Sample B: Strain-Hardened (Hot-Rolled) Aluminum Alloy

Tensile stress-strain curves for sample B (both thick and thin) are shown in Figure 8. The

elevated temperature static property data for sample B was collected from tests performed at

temperatures in the range of 20 to 460°C. The elastic moduli and 0.2% proof strength of

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0.000 0.002 0.004 0.006 0.008 0.0100

50

100

150

200

250

460 oC

340 oC

260 oC

220 oC

150 oC

20 oC

Str

ess (

MP

a)

Strain (mm/mm)

sample B-Thin and B-Thick tested are shown in Figures 9 and 10, respectively. Similar to

observations made for the extruded sample, the mechanical properties (elastic modulus and

0.2% proof strength) decrease with increase in temperature. The same tanh functions used to

mathematically describe the relationship between the elastic moduli or the proof strength with

temperature (equations 1 & 2) were also used for this data. The rate of mechanical property

(elastic modulus and proof strength) degradation with temperature for both B-Thin and B-

Thick specimens is the same, Figure 9.

Figure 8: Stress-strain curves for sample B-Thin.

Figure 9: Effect of temperature on Young’s modulus of samples B-Thick and B-Thin.

0 100 200 300 400 500 6000

10

20

30

40

50

60

70

ERT

=70.41 GPa

k=0.005350

T50%

=285.41 oC

B Thin

B Thick

Ela

stic

Modulu

s (

GP

a)

Temperature (oC)

7

0 100 200 300 400 500 6000

50

100

150

200

250

σ0.2%,RT

=234 MPa

k=0.009133

T50%

=271.97 oC

B-Thin

B-Thick

0.2

% O

ffse

t Y

ield

Str

ess (

MP

a)

Temperature (oC)

0 100 200 300 400 500 6000

10

20

30

40

50

60

70

80

Sample B (Thin)

Sample A (Extruded)

Sample B (Thick)Ela

stic

Modulu

s (

GP

a)

Temperature (oC)

Figure 10: Effect of temperature on 0.2% proof strength of sample-Thick and B-Thin.

The experimental data for the elastic modulus and 0.2% proof strength for the strain-hardened

sample (B) are compared to that of the extruded sample (A) in Figures 11 and 12,

respectively. In both cases, the mechanical properties of the extruded sample (A) are superior

to those of the strain-hardened sample (B) over the entire temperature range over which

elevated temperature static tests were conducted.

Figure 11: Comparison of effect of temperature on Young’s modulus of aluminum alloys A, B-

Thin and B-Thick.

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0 100 200 300 400 500 6000

50

100

150

200

250

300

Sample B (Thick)

Sample A (Extruded)

Sample B (Thin)0.2

% O

ffset Y

ield

Str

ess (

MP

a)

Temperature (oC)

1 10 100 1000 10000

0.00

0.15

0.30

0.45

0.60

50 M

Pa-3

63 C

15 M

Pa-4

32 C25 M

Pa-4

00 C

50

MP

a-3

41

C

17

.5 M

Pa-4

37 C

100

MP

a-2

90

C

40

MP

a-3

63

C

30 M

Pa-3

90 C

20

MP

a-4

30

C75

MP

a-3

01 C

20

0 M

Pa

-200

C

Str

ain

(m

m/m

m)

Log (Time)

Figure 12: Comparison of effect of temperature on 0.2% proof strength of aluminum alloys A,

B-Thin and B-Thick.

2.2 High Temperature Creep Properties

2.2.1 Sample A: Extruded Aluminum Alloy

Selected creep strain curves for the extruded sample (A) at different temperature and stress

conditions are shown in Figure 13. The creep strain is plotted against the logarithm of time.

Figure 13: Creep curves for the extruded sample (Type A).

The rupture times were converted into Larson-Miller parameters by using the relationship:

T(log(t) + C) = LMP (3)

where T is the absolute isothermal exposure temperature, t the exposure time in hours, and C a

material constant; set at a value of 18 in this report. Figure 14 shows the variation of the

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0 5000 10000 15000 20000 250000.0

0.1

0.2

0.3

0.4

125 M

Pa

95 M

Pa

70 M

Pa

105 M

Pa

115 M

Pa

Cre

ep

str

ain

Time (s)

8000 9000 10000 11000 12000 13000 140001.0

1.2

1.4

1.6

1.8

2.0

2.2

2.4

Log(σ) = 0.32 + 9.90exp(-LMP/5488.76)

Lo

g (

σ)

Larson-Miller Parameter

natural logarithm of applied stress as a function of the Larson-Miller Parameter (LMP). This

relationship has been successfully demonstrated to offer a reliable means of constructing

creep curves performed within and outside the test matrix; i.e. interpolation and extrapolation

[4].

Figure 14: Relationship between natural logarithm of applied stress and the Larson-Miller

parameter.

2.2.2 Sample B: Strain-Hardened (Hot-Rolled) Aluminum Alloy

Creep curves collected for sample B-Thick at a constant temperature of 350°C under varying

stress levels in the range 70 to 125 MPa are shown in Figure 15.

Figure 15: Creep curves for the Type B-Thick alloy at a constant temperature of 350°C under

different stress level loadings.

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4.2 4.3 4.4 4.5 4.6 4.7 4.8 4.9-4.2

-4.1

-4.0

-3.9

-3.8

-3.7

-3.6

-3.5

-3.4

ln (

str

ain

at ze

ro)

ln (σ)

From the curves shown in Figure 15, the steady creep strain rate ( )IIεɺ was determined. The

natural logarithm of the ( )IIεɺ was then plotted against the logarithm of stress at 350°C, Figure

16. There is a linear dependency of ln ( )IIεɺ on the natural logarithm of the applied stress, ln(σ)

from which a value of n = –3.96 as contained in ( ) )/exp( RTQA n

II −= σεɺ is obtained from the

slope.

Figure 16: Plot of ln ( )IIεɺ versus ln(σ ) at 350°C for the Type B -Thick alloy.

In addition, the natural logarithm of the strain at time t = 0, ( )0ε , was also plotted against the

natural logarithm of the applied stress, Figure 17. The data is scattered and no mathematical

expression could be used to describe this data set. The same observation was made in our

previous studies wherein a different alloy 5083-H116 was investigated [5].

Figure 17: Plot of ln ( )0ε versus ln(σ) for the extruded Type B-Thick alloy.

4.2 4.3 4.4 4.5 4.6 4.7 4.8 4.9-14

-13

-12

-11

-10

-9

-8

ln (strain rate) = 3.96*ln (σ) - 28.73

ln (

str

ain

rate

)

ln (σ)

11

0 5000 10000 15000 20000 250000.0

0.1

0.2

0.3

0.4

0.5

40

0oC

36

5oC

41

0oC

40

5oC

35

0oC

Cre

ep

str

ain

Time (s)

From another set of experiments, creep curves obtained at a constant stress of 70 MPa at

different temperatures in the range of 350 to 420°C are shown in Figure 18. Due to the limited

sample size and the variability in the behaviour of the test specimens, only a limited number

of creep test data sets were usable.

Figure 18: Creep curves for the Type B-Thick alloy at a constant stress of 70 MPa at different

temperatures.

From the curves shown in Figure 18, the steady creep strain rate ( )IIεɺ was determined. The

natural logarithm of the ( )IIεɺ was then plotted against the inverse of the absolute test

temperatures, (1/T), Figure 19. There is a linear dependency of ln ( )IIεɺ on the inverse of

absolute test temperature (1/T). The activation energy, Q, calculated from the slope of the

curve is 163 kJ/mol. By taking the natural logarithm of ( ) )/exp( RTQA n

II −= σεɺ , the intercept of

the plot of ln(strain rate) versus ln(σ) can be used to calculated the frequency factor, A. A

value of 16.76 s-1

was determined for the parameter A.

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0.00148 0.00152 0.00156 0.00160-12.5

-12.0

-11.5

-11.0

-10.5

-10.0

-9.5

-9.0

-8.5

ln(strain rate) = 19.5377 - 19655.3347*(1/T)

ln(s

tra

in r

ate

)

1/T(K-1)

0.00148 0.00152 0.00156 0.00160

-6.5

-6.0

-5.5

-5.0

-4.5

-4.0

-3.5

-3.0

ln(s

tra

in a

t zero

)

1/T (K-1)

Figure 19: Plot of ln ( )IIεɺ versus 1/T for the Type B-Thick alloy.

The natural logarithm of the strain at time t = 0, ( )0ε , was also plotted against 1/T, Figure 20.

Figure 20: Plot of ln ( )0ε versus1/T for the Type B-Thick alloy.

The data is scattered and no mathematical expression could be used to describe this data set.

The same observation was made in our previous studied wherein a different ally 5083-H116

was investigated [5].

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References

1. EN 1999:1-1, 2009, ‘Eurocode 9 – Design of Aluminum Structures – Part 1-1: General

Rules,’ British Standard.

2. EN 1999-1-2:2007, 2009, ‘Eurocode 9 – Design of Aluminum Structures – Part 1-2:

Structural Fire Design,’ British Standard.

3. Fogle, E.J., Lattimer, B.Y., Feih, S., Kandare, E., Mouritz, A.P., Case, S.W. ‘Compression

load failure of aluminum plates due to fire’, Engineering Structures Journal, Submitted

2011.

4. Feih, S., Kandare, E., Lattimer, B.Y., Mouritz, A.P. ‘Structural analysis of compression

deformation and failure of aluminium in fire’, Journal of Structural Engineering.

http://dx.doi.org/10.1061/(ASCE)ST.1943-541X.0000313.

5. Feih, S., Kandare, E., Mouritz, A.P. ‘Aluminium creep data and modelling approaches’,

CRC-ACS TM 10025, June 2010.