Fatigue crack growth rate behaviour of friction-stir aluminium alloy AA2024-T3 welds under transient thermal tensioning

Materials and Design 50 (2013) 235–243

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Fatigue crack growth rate behaviour of friction-stir aluminium alloy AA2024-T3 welds under transient thermal tensioning M.N. Ilman ⇑, Kusmono, P.T. Iswanto Department of Mechanical and Industrial Engineering, Gadjah Mada University, Indonesia

a r t i c l e

i n f o

Article history: Received 22 December 2012 Accepted 27 February 2013 Available online 14 March 2013 Keywords: Friction stir welding Aluminium alloys Transient thermal tensioning Fatigue crack growth rate

a b s t r a c t Friction stir welding (FSW) has become a serious candidate technology to join metallic fuselage panels for the next generation of civil aircrafts. However, residual stress introduced during welding which subsequently affects fatigue performance is still a major problem that needs to be paid attention. The present investigation aims to improve fatigue crack growth resistance of friction stir aluminium alloy AA2024-T3 welds using transient thermal tensioning (TTT) treatment. In this investigation, aluminium alloy AA2024T3 plates were joined using FSW process with and without TTT. The welding parameters used including tool rotation speed (Rt) and the plate travelling speed (v) were 1450 rpm and 30 mm/min respectively. The TTT treatments were carried out by heating both sides of friction stir weld line using moving electric heaters ahead of, beside and behind the tool at a heating temperature of 200 °C. Subsequently, a sequence of tests was carried out including microstructural examination, hardness measurement, tensile test and fatigue crack growth rate (FCGR) test in combination with fractography using scanning electron microscopy (SEM). The FCGR test was carried out using a constant amplitude fatigue experiment with stress ratio (R) of 0.1 and frequency (f) of 11 Hz whereas specimens used were centre-crack tension (CCT) type with the initial crack located at the weld nugget. Results of this investigation showed that at low DK, typically below 9 MPa m0.5, the friction stir welds under TTT treatments lowered fatigue crack growth rate (da/dN) and the lowest (da/dN) was achieved as the heaters were located ahead of the tool. This improved weld fatigue performance was associated with stretching effect generated by movingly localised secondary heating which might alter the magnitude and distribution of residual stress in weld region and in a such condition, TTT seemed to act as local preheating. Ó 2013 Elsevier Ltd. All rights reserved.

1. Introduction Aluminium alloy AA2024-T3 has long been used for aircraft structure for example fuselage or lower wing surface due to its high strength-to-weight ratio, extremely good damage tolerance and high resistance to fatigue crack growth. Aluminium alloy AA 2024-T3 is Al–Cu alloy with the copper content in the range of 3.8–4.9% whereas T3 represents solution heat treated and naturally aged to achieve significant hardening [1]. However, this particular alloy is considered to be unweldable metal and therefore conventional arc welding such as metal inert gas (MIG) or tungsten inert gas (TIG) is not recommended due to solidification cracking. A significant advance has been made in welding technology marked by the invention of friction stir welding (FSW) at The Welding Institute in 1991 [2]. The FSW is a solid state joining process in which the weld joint is formed by plunging a rotating shouldered tool equipped with a pin into the adjoining edges of the plates to be welded until the tool shoulder is in contact with ⇑ Corresponding author. Tel./fax: +62 274 521673. E-mail address: [email protected] (M.N. Ilman). 0261-3069/$ - see front matter Ó 2013 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.matdes.2013.02.081

the plate surfaces [3–6]. Such welding process offers several benefits, i.e. FSW is a solid state joining process which enables unweldable aluminium alloys such as AA 2024-T3 to be joined without melting hence avoiding hot cracking [7]. As a result, FSW finds its potential applications in the fabrication of light weight materials suitable for transportation industries such as automotive, rail, marine and aerospace industries [8]. In mechanical design, strength is an important consideration for components or structures under static loads. The strength of FSW joints is usually lower than that of the base metal due recrystallisation in the weld nugget. Recent investigations have shown that the strength of FSW joints can be improved by optimising process parameters such as pin profile, tool rotating speed and tool travelling speed [9–11] or post weld heat treatment for age hardenable aluminium alloys [12,13]. In aircraft manufacture, the riveting technique is commonly used for joining aircraft structure rather than arc welding since welding of aluminium alloys often produce problems such as porosity, hot cracking, residual stress and distortion [14–16]. To date, riveting technique is being challenged by FSW process because of benefits offered by FSW, i.e. it does not need

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Table 1 Chemical composition of AA 2024-T3 (wt.%). Si

Fe

Cu

Mn

Mg

Zn

Cr

Ti

0.50

0.50

3.9

0.60

1.5

0.25

0.10

0.15

consumables and shielding gas hence reducing manufacturing cost and weight. However, stringent design criterion on superior resistance to the growth of fatigue crack under cyclic loading conditions remains a critical issue that needs to be paid attention when considering FSW process for aircraft construction. As it is commonly observed in conventional arc welding processes, residual stress that forms during welding process is still a major problem in FSW since it reduces fatigue crack growth resistance of friction stir aluminium weld joints [17–21]. One of the promising methods for control of residual stress and distortion is transient thermal tensioning (TTT) method. This method consists of heat sources located at both sides of weld line, in front of, beside or behind the moving weld torch or FSW tool. The presence of movingly localised heat sources produces temperature gradient and gives thermal tensioning or stretching effect which opposes weld residual stress [22,23]. The advantage of TTT is based on the fact that TTT is an in-process method, therefore from the view point of manufacturing efficiency and cost, the TTT is more desirable than other post weld treatments. In recent years, the mechanisms in which a TTT process to control distortion and residual stress have been, in general, explained based on thermal elastic–plastic theory using computational [24– 26] and experimental [27,28] approaches. However, there is little published data on how the TTT process to affect fatigue crack growth behaviour in friction stir hardenable aluminium alloy welds such as AA2024-T3 and therefore, is the subject of the present investigation. 2. Materials and experimental methods 2.1. Materials Aluminium alloy AA2024-T3 plates were used as the base metal in the present investigation. The chemical composition of the plates is given in Table 1. 2.2. Welding process and TTT treatments Two AA 2024-T3 plates with the dimensions of 400 mm long, 100 mm wide and 3 mm thick were butt welded along 400 mm long using FSW process with tool rotating speed and tool travelling

speed of 1450 rpm and 30 mm/min (0.5 mm/s) respectively. The shoulder diameter of the tool was 16 mm whereas diameter and length of the pin were 4 mm and 2.9 mm respectively. TTT treatments were carried out by locating electric heaters as secondary heating on both sides of weld line ahead of, beside or behind the moving tool at a temperature typically of 200 °C as shown Fig. 1. This temperature was likely to be close to the optimum value for TTT [24]. For each specimen, a K-type thermocouple was attached to the region near the weld, typically of 10 mm from the weld line to monitor the thermal cycle during welding process using a data acquisition system. The transient temperature field at the plates subjected to moving electric heaters as shown in Fig. 2 can be analysed using Goldak’s double ellipsoidal heat source [29] as follows:

! pffiffiffi 6 3Q w f ð3X b þ v tÞ2 3Y 2b 3Z 2b pffiffiffiffi exp  Qðx; y; zÞ ¼  2  2 c2 a abcp p b

ð1Þ

where Q(x,y,z) is the double ellipsoid volume source heat input, Qw is heat input, f is scaling factor, v is heat source travelling speed, t is time; x, y and z are the local coordinates of the double ellipsoidal model; and a, b and c are the transverse, through the depth and longitudinal ellipsoidal axes. By using Goldak’s analytical solution, the heat flux, q(y), resulted from thermal tensioning heat sources applied on the top surface of the plates (Fig. 2) can be calculated using the following equation [30]:

qðyÞ ¼ 4

3Q s g

p

exp 

3X 2s 2

d2



3Y 2s

!

2

d1

ð2Þ

where Qs is heat flux per unit area, g is the heating efficiency whereas Xs and Ys are the local coordinates of the surface ellipsoidal heat source:

X s ¼ v t þ d3  0:5d2

ð3Þ

Y s ¼ jyj  d4  0:5d1

ð4Þ

The dimensions of d1, d2, d3 and d4 as shown in Fig. 2 represent width of the heaters, length of the heaters, offset from the tool and offset from the weld line respectively. For electric heaters, heat flux per unit area (Qs) is defined as:

Q s ¼ Q w =Aw ¼ VIeff =ðd1 d2 Þ

ð5Þ

where Aw is heater area, V is voltage and Ieff is effective current (in the case of alternating current). During TTT process, the distance between heaters and the tool along x direction, i.e. d3 was varied whereas other parameters were maintained constant. The voltage (V) and current (Ieff) were care-

Fig. 1. Welding process: (a) as welded, and TTT with heater locations: (b) beside (c) ahead of, (d) behind the tool.

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Fig. 4. Centre-cracked tension (CCT) specimen.

2.5. Fatigue tests

Fig. 2. Transient thermal tensioning analysis.

fully adjusted to reach a preset temperature of 200 °C at certain value of Qs.

2.3. Microstructural examination Microscopical examination was carried out using an optical microscopy. The specimens were prepared using standard metallograpic procedure consisting of grinding, polishing and etching using Keller’s reagent made of 5 ml HNO3 (95% concentration), 2 ml HF, 3 ml HCl, 190 ml H2O. The examination was focused on the cross section perpendicular to weld centre. Scanning electron microscopy (SEM) equipped with EDX-analysis was also employed for microanalysis.

Specimens for fatigue crack growth rate test were prepared according to ASTM: E647-13 standard as shown in Fig. 4. Centrecracked tension (CCT) or middle tension (MT) specimens were selected with the initial crack (ao) of 18 mm long were located at the weld metal region. Fatigue experiment was carried out using a servo-hydraulic universal testing machine and a sinusoidal load was selected with the stress ratio (R) of 0.1 and a frequency (f) of 11 Hz. A stress level used was around 20% of yield stress. The fatigue crack growth rate (da/dN) of the Paris power law was analysed using Secant method as follows:



da dN

 ¼ a

aiþ1  ai Niþ1  Ni

ð6Þ

a ¼ ðaiþ1 þ ai Þ=2

ð7Þ

where a is average crack length and subscripts i and (i + 1) represent ith and (i + 1)th cycle. The stress intensity factor range, DK for centre crack tension (CCT) geometry was calculated using the following equation:

DK ¼

DP B

"rffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi#

pa

2W

sec

pa 2

ð8Þ

2.4. Mechanical properties tests Mechanical properties of friction sitr welds were assessed using hardness measurement and tensile test. Vickers microhardness with the load of 100 grf was carried out to allow hardness distribution along weld nugget zone (WNZ), thermomechanically affected zone (TMAZ), heat affected zone (HAZ) and the base metal. Subsequently, the strength of friction stir welds was assessed using tensile test with the specimens in the form of transverse weld specimen as shown in Fig. 3.

Fig. 3. Transverse weld specimen for tensile test.

DP ¼ Pmax  Pmin

ð9Þ

where B is the specimen thickness, W is the specimen width and a = 2a/W. Subsequently, the fractured surfaces of the specimens were examined using SEM. 3. Results and discussion 3.1. Thermal cycles The origin of residual stress is the highly localised transient heat input and the mechanism in which transient thermal tensioning treatment to control weld residual stress can be analysed based on welding temperature field. Fig. 5 shows thermal cycles of regions close to the weld metal at the distance of 10 mm from the weld centre line (Fig. 1) with and without TTT treatments. The thermal cycle of as welded FSW with no TTT treatment was characterised by rapid heating to a peak temperature (Tp) followed by a slightly lower cooling rate to ambient temperature. The effect of TTT seemed to increase the peak temperature and to lower cooling rate. The highest peak temperature was observed in FSW with the heaters located beside the tool. At this condition, offset from the

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Fig. 5. Thermal cycles measured at a region close to weld zone for FSW joints with and without TTT.

tool (d3) was minimum and according to Eqs. (2)–(4), a decrease in d3 reduces Xs resulting in high heat flux q(y). A classical Rosenthal analytical solution [31,32] seemed to apply to FSW where the temperature profile in weld zone and its adjacent area for thin plates is given by:

T  To ¼

Q =v dð4pkqctÞ1=2

  r2 exp  4at

ð10Þ

where To is initial temperature (or preheat), k is thermal conductivity, (qc) is specific heat per unit volume, a is thermal diffusivity defined as k/(qc), d is plate thickness, v is travelling speed, t is time and r is radial/lateral distance from the weld. In the case of FSW, the total heat Q generated by the tool is given by [33,34]:

qs ¼

3Qr 2pR3

ð11Þ

or

Q¼

2pR3 qs 3r

ð12Þ

where qs is heat flux under the tool shoulder, r is the radius originating at the centre of the tool and R is the outside radius of the tool shoulder. Results from the present investigation seem to suggest that increasing heat flux generated by heaters q(y) or qs increases the plate initial temperature (To) and according to Eq. (10), the peak temperature Tp is expected to increase. 3.2. Microstructure Fig. 6 shows a typical friction stir weld profile marked by unsymmetrically inverted trapezoidal form which elongated towards advancing side. The microstructures present in FSW can be classified into various zones, i.e. weld nugget zone (WNZ), ther-

momechanically affected zone (TMAZ), heat affected zone (HAZ) and unaffected base metal (BM) as reported by Fratini et al. [35]. Referring to Fig. 7, it can be seen that the base metal microstructure was characterised by elongated grains which were oriented parallel to rolling direction. At the region close to the weld, the grains became coarser due to heat effect known as HAZ region. Microstructure of TMAZ consisted of deformed grains as a result of combined effect of heat and mechanical forces from the tool. The weld nugget zone showed fine equiaxed grained structure with the grain size was significantly smaller than the base metal grains. This grain refinement was caused by recrystallization under high temperature and large deformation in the weld centre due to stirring process. In addition, the presences of oxides and/or precipitates (dark etched) were also observed in WNZ with no evidence of defects. Fig. 8 shows weld nugget microstructure under various TTT treatments. A significantly microstructural change was observed especially in the weld treated by a TTT process with heat sources beside the tool. This particular microstructure was marked by coarser grain structure. This grain coarsening was caused by additional heating from secondary heat sources, i.e. heaters which had the closest distance with the the tool and such condition increased peak temperature Tp and slowed down cooling rate. 3.3. Hardness distribution and strength The hardness distributions of all welded specimens are shown in Fig. 9. The hardness value of unaffected AA2024-T3 base metal was around 160.3 VHN. For all specimens under study, the hardness started to decrease at the region approaching HAZ, and further decrease was observed at HAZ region close to TMAZ until the lowest hardness was achieved at the transition between HAZ-TMAZ and WNZ. The averaged hardness of the weld nugget regions for all the FSW joints was around 80 VHN except the weld nugget resulted from TTT treatment with heaters behind the tool. This weld nugget showed a sharp increase in hardness at the weld centre forming ‘W’-shaped hardness profile. This type of W-shaped hardness profile is commonly observed in hardenable aluminium alloys [21]. It seemed that the ‘W’-shaped hardness profile in the weld with heating position behind the tool was associated with reprecipitation. This argument was supported by SEM microanalysis results as shown in Fig. 10. It can be seen that the weld nugget was marked by the presence of uniformly distributed fine particles (light etched) and some of them were aligned forming arrays of precipitates. EDXspectra taken from these particles as seen in Fig. 11 showed a strong peak of Al with the trace of Cu which could be in the form of Al2Cu precipitates. Since the heaters were placed behind the tool, the fully developed weld nugget region behind the tool was reheated by the heaters at a temperature within the temperature range of precipitation reaction. As a result, reprecipitation took place and under such condition, TTT treatment acted as postweld heat treatment. Transverse strength properties of FSW joints are presented in Fig. 12. During tensile test, all welded specimens exhibited fracture at weld nugget zone with the their yield stress around 0.6–0.7 of the ultimate tensile stress (UTS). The test error was relatively minor and it was statistically accepted. The joint efficiency (defined as the ratio of the tensile strength of the friction stir weld to that of the base metal) was in the range of 40–60%. It can be seen that the minimum joint efficiency occurred as the heaters were placed beside the tool. This finding was consistent with Hall–Petch relationship [36] where the low strength of this particular weld nugget was associated with its coarse grained structure as given by:

ry ¼ ro þ ky d1=2 Fig. 6. A typical FSW profile consisting of various regions: base metal (BM), heat affected zone (HAZ), thermomechanically affected zone (TMAZ) and weld nugget zone (WNZ).

ð13Þ

where ry is yield stress, d is grain size, ro is friction stress and ky is a positive yield constant. However, the Hall–Petch equation may not

M.N. Ilman et al. / Materials and Design 50 (2013) 235–243

Fig. 7. Typical FSW microstructures: (a) BM, (b) HAZ, (c) TMAZ and (d) WNZ.

Fig. 8. Weld nugget microstructures: (a) as welded and TTT-treated welds with the heaters: (b) ahead of, (c) beside, (d) behind the tool.

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Fig. 12. Tensile stresses of FSW joints. Fig. 9. Hardness distribution of FSW joints.

Fig. 10. SEM micrograph of the TTT-treated weld nugget with heater position behind the tool.

can be seen that TTT treatments delayed the crack growth of friction stir welds resulting in longer fatigue life compared to the as welded friction stir welds. The most effective TTT treatment in improving fatigue crack growth resistance was observed as the heater position was ahead of the tool followed by the positions beside and behind the tool. It is surprising that TTT process with heater position ahead of the tool effectively improved fatigue performance close to that of the base metal. Previous reports [37,38] have shown that the crack growth retardation could be related to microstructure, crack closure and residual stress. In the case of TTT treatments under study, residual stress could be the main factor controlling fatigue crack growth rate behaviour. If this mechanism is important, then the crack growth inhibition in the TTT-treated welds would be associated with compressive residual stress in the plastic zone a head of the crack tip. Fatigue crack growth rate (FCGR) of the friction stir welds under study can be analysed using double logarithmic plots of da/dN–DK as shown in Fig. 14. In the stage II of the da/dN–DK curves, the fatigue crack growth rates (da/dN) were linearly proportional to stress intensity factor range (DK) and the trendlines taken from the region II as shown in Fig. 15 seemed to follow a power-law relation, i.e. da/dN = C(DK)n where C is constant and n is a power law exponent. Of note is that the n value represents the slope of the fatigue growth rate curve in the log–log plot whereas changes in C alters the position of the data. The values of n and C at

Fig. 11. EDX-spectra taken from fine particles (light etched) in Fig. 10 showing Al with the trace of Cu.

directly apply to the FSW joints with appreciable precipitation hardening such as the TTT-treated weld with heaters behind the tool. 3.4. Fatigue crack growth rate behaviour Results of fatigue crack growth test for all friction stir welds with the base metal as reference fatigue are shown in Fig. 13. It

Fig. 13. Schematic diagram of a–N curves for FSW joints.

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M.N. Ilman et al. / Materials and Design 50 (2013) 235–243 Table 2 Paris constants.

Fig. 14. Plots of (da/dN)–DK for FSW joints.

Fig. 15. Trendlines taken from region II of (da/dN)–DK curves.

p DK = 1 MPa m for the friction stir welds and the base metal under study are given in Table 2. It can be seen that the n values were in the range of 2–7 in agreement with previous reports [39,40]. Referring to Fig. 15, it can be seen that at lower DK, typically bep low 9 MPa m, the TTT-treated friction stir weld with heater position ahead of the tool had the lowest da/dN, followed by the second lowest da/dN, i.e. the weld with the heater position beside the tool. Finally, a small decrease in da/dN was observed as the TTT treatment was applied with the heaters behind the tool. Base metal, on the other hand, was characterised by relatively low n and C suggesting that da/dN was low. The conclusion obtained from this study is that the retardation of fatigue crack growth due to TTT treatments occurs in the early stage of stable crack growth. One possible explanation could be that TTT treatments cause redistribution of residual stresses and probably result in compressive residual stress in the weld zone. As the crack propagates and gets longer, the compressive residual stress ahead of crack tip decreases eventually disappears at higher DK. In a such condition, fatigue crack growth in the stable crack growth period is no longer affected by the material surface conditions and the crack growth becomes a bulk material phenomenon. Fig. 16 shows SEM fractographs of fractured surfaces of base metal and friction stir weld nuggets with and without TTT treatments taken from fatigue stable crack growth region (region II).

Weldments

C

n

As welded TTT-behind TTT-beside TTT-ahead of Base metal

1.5294E10 3.2974E10 3.1296E13 3.5490E14 4.2621E11

3.0410 2.6086 6.2419 6.8144 3.2565

Of note is that the crack direction started from the left to the right. It can be seen that all weld nuggets revealed transgranular cleavage fracture indicative of brittle fracture feature with the striations were not clearly seen. In contrast, the fracture surface of base metal showed the presence of very fine striations. This finding seemed to suggest that microstructural characteristics had an important contribution in term of fracture surface type but not in the case of fatigue crack growth rate. Results obtained from the present investigation seem to suggest that improved fatigue performance of TTT-treated FSW joints is related to changes in the magnitude and distribution of residual stress. The mechanism of thermal tensioning for control of residual residual stress may be explained based on a model proposed by Guan [41] as shown in Fig. 17. It can be seen that localised heat sources produce temperature curves T1 and T2 with the corresponding thermal stresses curves rx1 and rx2. A highly localised heating with a peak temperature of T1max results in local expansion but the presence of restraint of dimensional changes by the surrounding unheated base material induces maximum compressive thermal stress rx1max. Similarly, a peak temperature of T2max produces rx2max. Since thermal stresses are in static equilibrium so that the formation of compressive stress in the heated plate regions could induce tensile stress in other regions, i.e. weld region and its adjacent area. In such a case as this, thermal tensioning effect is defined as the value of rx at Y = 0 as shown in Fig. 17. The magnitude of thermal stress is proportional to temperature gradient ð@T=@YÞ. Accordingly, the T1 curve which has steep temperature gradient, induces higher value of compressive stress, i.e. r1-max compared to T2 curve for a given rx . Based on analysis of thermal tensioning as discussed above, Burak et al. [42,43] conducted an experiment to control residual stress by creating temperature gradient using stationary heating bands located on both sides along the weld line. This method produced steady state temperature differential known as static thermal tensioning (STT). It seems that the basic principle of this thermal tensioning effect may also apply to TTT treatment which uses movingly localised heat sources. Results of the present investigation have shown that in general, TTT treatments improved weld fatigue crack growth resistance. Based on previous studies [23,27,41], it may be argued that movingly local heating resulted from the heaters produces thermal gradient which subsequently gives a stretching effect in a similar manner to that observed in mechanical stretching method and alters the magnitude and distribution of normal weld residual stress developed during welding. The effectiveness of TTT treatments was dependent on the position of the heaters with respect to the heat source (tool). The TTT with heaters ahead of the tool was the most effective TTT treatments for improving fatigue crack growth resistance followed by the heater positions beside and behind the tool. In the case of the TTT-treated weld with heaters behind the tool, stretching effect induced by thermal gradient was not effective since the position of heaters was lagging and residual stress had already developed in the weld zone and its adjacent area as the stretching was taking place. In contrast, the heaters ahead of the tool were operative prior to or during the formation of weld residual stress so that the stretching effect induced by temperature gradient seemed to

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Fig. 16. SEM fractographs of: (a) as welded weld nugget, (b) wed nugget with TTT ahead of the tool, (c) weld nugget with TTT behind the tool and (d) base metal.

Fig. 17. Mechanism of thermal tensioning [41].

be effective. In such as case as this, TTT treatment acted as local preheating [44]. The effect of TTT with heaters beside the tool on the weld fatigue crack growth resistance fell within the effectiveness range of the two previous heater positions. Results of this investigation lead to the conclusions that apart from thermal gradient, the position of heaters with respect to the tool plays an important role in reducing weld residual stress and hence improving fatigue crack growth resistance. The present investigation has endeavoured to elucidate effect of TTT on fatigue crack growth rate (FCGR) by analising transient temperature field. However, residual stress was not measured in the present investigation. Therefore, additional work will be carried out to assess the magnitude and distribution of residual stress in FSW joint under TTT treatments. 4. Conclusions Conclusions that can be drawn from this investigation are as follows:

(1) In general, TTT treatments improved fatigue crack growth resistance of friction stir welds. This improved fatigue resistance seemed to be associated with stretching effect generated by heaters which might alter the magnitude and distribution of residual stress in weld region. (2) The TTT treatment was found to be effective for stress relieving as the heaters were located ahead of the heat source (tool). In a such position, the stretching effect took place prior to and during the formation of residual stress in a similar manner to that observed in local preheating. (3) Apart from stretching effect, the TTT treatments had a potency to affect microstructure developed in the weld nugget. These microstructural changes, i.e. grain size and reprecipitation in heat-treatable aluminium alloys such as AA2024-T3 were associated with additional heat generated by heaters which increased a peak temperature and lowered a cooling rate during welding. However, the effect of microstructure on fatigue crack growth rate was not significant compared to residual stress.

Acknowledgments This work was carried out under research grant (Contract No. 085/Dir.Keu/KN/DIPA-UGM/2010:4 May 2010). The authors acknowledge funding of LPPM UGM and DIKTI.

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