Fatigue & Fracture of Engineering Materials & Structures 1

7/28/2019 Fatigue & Fracture of Engineering Materials & Structures 1

1/13

doi: 10.1111/j.1460-2695.2008.01234.x

Correspondence: J. R. Tarpani. E-mail: [email protected]

Fatigue behaviour of friction stir welded AA2024-T3 alloy: longitudinaland transverse crack growth

M . T . M I L A N , W . W . B O S E F I L H O , C . O . F . T . R U C K E R T a n d J . R . T A R P A N I

Department of Materials, Aeronautics and Automotive Engineering, Engineering School of Sao Carlos, University of Sao Paulo, Av. Trabalhador

Sao-Carlense, 400, Centro, CEP. 13.566-590, Sao Carlos-SP, Brazil

Received in final form 20 April 2007

A B S T R A C T The fatigue crack growth properties of friction stir welded joints of 2024-T3 aluminiumalloy have been studied under constant load amplitude (increasing-K), with specialemphasis on the residual stress (inverse weight function) effects on longitudinal andtransverse crack growth rate predictions (Glinkas method). In general, welded joints

were more resistant to longitudinally growing fatigue cracks than the parent material atthreshold K values, when beneficial thermal residual stresses decelerated crack growthrate, while the opposite behaviour was observed next to KC instability, basically due tomonotonic fracture modes intercepting fatigue crack growth in weld microstructures.

As a result, fatigue crack growth rate (FCGR) predictions were conservative at lowerpropagation rates and non-conservative for faster cracks. Regarding transverse cracks,intense compressive residual stresses rendered welded plates more fatigue resistant thanneat parent plate. However, once the crack tip entered the more brittle weld regionsubstantial acceleration of FCGR occurred due to operative monotonic tensile modes offracture, leading to non-conservative crack growth rate predictions next to KC instability.

At threshold K values non-conservative predictions values resulted from residual stressrelaxation. Improvements on predicted FCGR values were strongly dependent on howthe progressive plastic relaxation of the residual stress field was considered.

Keywords aluminium alloy; crack growth rate prediction; fatigue; friction stir welding;residual stress.

N O M E N C L A T U R E S a = crack lengthAA2024-T3 = high-strength aluminium alloy grade

d= slot apertureda/dN= crack growth rate

E= plane-stress Youngs modulusE = plane-strain Youngs modulus

EL = elongation at fractureFCGR

=fatigue crack growth rate(s)

FSW= friction stir weldingh(x,a) =weight functionHAZ = heat-affected zone

KC = critical stress intensity factorKMAX = maximum applied stress intensity in fatigue

KIr = residual stress intensity factor in mode I of crack opening

526 c 2008 The Authors. Journal Compilation c 2008 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct. 31, 526538

Fatigue & Fracture of

Engineering Materials & Structures


2/13

F A T IG U E B E H AV I O UR O F F R I CT I O N S T I R W E L DE D AA 2 0 24 - T 3 A L L OY 527

Krx, Kry= residual stress intensity factor distribution on x and y directionsL = length of generic test piece

L0 = original gage lengthM= location of strain-gageR = stress ratio

R = effective stress ratioRA = reduction in area at fracture

S= engineering, nominal, remote or gross stressS, L(x), T(y) = three main orthogonal metallographic axes or directions

TP= test pieceTMAZ = thermo-mechanically affected zone

UTS = ultimate tensile strengthW=width of generic test piece

WEDM =wire electro-discharge machineYS =yield strength

Z(a) = influence functionK(th) = range of stress intensity factor in fatigue (threshold value)

= strainM

=strain measured at M position

= Poissons coefficientrx, ry= residual stress distribution on x and y directions

I N T R O D U C T I O N

In recent years, friction stir welding (FSW), a solid-statejoining technique, has been considered a potential tech-nique to replace conventional riveting operations and fu-sion welding methods (e.g. laser and electron beam) inaircraft manufacture. However, the weld still results ina continuous medium for crack propagation and hence,

knowledge of the fatigue and fracture properties of suchclasses of materials is vital if a damage-tolerant design isadopted.Advantages of FSW process include: design simpli-

fication (easy periodical inspection, less macro stress-concentrators), low distortion, poreless welding process(less micro-stress concentrators), static strength as highas 80100% of parent material, improved fatigue perfor-mance and better load distribution. FSW disadvantagescomprise: lack of extensive data on mechanical proper-ties, continuous crack propagation medium, mismatchingbetween plastic properties of weld and parent metals andresidual stress effects.

Residual stresses are invariably present in welded struc-tures after fabrication.They are likely to affect mechanicaland corrosion properties of the materials and thereforeinfluence the in-service performance of structural com-ponents. The effects of residual stresses on fatigue crackpropagation have been reported by several authors suchas Itoh et al.,1 Bussu and Irving2 and Milan and Bowen.3,4

Based on Parkers superposition principle,5 they con-cluded that tensile residual stresses increase the crackgrowth rate due to increasing effective stress ratio (R).

On the other hand, compressive residual stresses reducethe fatigue crack growth rate (FCGR) by decreasing theeffective stress ratio. Additionally, residual stresses werefound to affect initiation fracture toughness values (KC)of aluminium alloys.6,7

Transverse cracks present an even more complex situa-tion, inasmuch as the defect propagates from the parentmaterial towards the weld region. The crack tip inter-

sects different microstructures and distinct intrinsic fa-tigue cracking behaviours can be expected.This paper presents data obtained from studying the fa-

tigue crack resistance of FSW joints of aeronautical gradeAA2024-T3 high-strength alloy containing either lon-gitudinal or transverse cracks. The effective stress ratiomethod (Glinkas R)8 was employed to predict FCGR inthe welded alloy taking into account the residual stress in-tensity factor calculated by the slitting or cut compliancemethod9 and parent material fatigue properties. Predictedcrack growth rates are then compared to experimental

values.The authors expect to contribute to the still limited

body of knowledge regarding FSW materials by provid-ing useful fatigue crack growth data for both safe-life anddamage-tolerant designs.

I N C R E M E N T A L S L I T T I N G A N D W E I G H T

F U N C T I O N M E T H O D S

The cut compliance or incremental slitting method is ahelpful technique to determine both the near surface andthrough thickness residual stress profiles. It is based on

c 2008 The Authors. Journal Compilation c 2008 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct. 31, 526538


3/13

528 M . T . M I L A N et al.

the fact that when a cut, simulating a crack, is incremen-tally introduced into a part, the residual stresses are re-lieved, causing the part to deform. Such deformation canbe sensed by strain gauges attached at specific positionsof the part (Fig. 1) and the residual stress intensity factorprofile can be derived.9,10Assuming a sufficiently narrow

slot (d a), linear elastic fracture mechanics (LEFM) canbe employed to establish a relationship (Eq. 1) betweenthe measured strains, , and the corresponding residualstress intensity factor in opening mode, KIr:

9

KIr(a) =E

Z(a)

dMd a

, (1)

where M is the measured strain at the back face posi-tion M during the cutting procedure, a is the slot length,

E, the generalized form of the Youngs modulus (E = Efor plane-stress, and E = E/1 2 for plane-strain con-ditions), and Z(a) the influence function that dependson the test-piece geometry, cut plane location and strain

measurement position. Here Z(a) is considered as beingindependent of the residual stress profile.

For a rectangular plate, whereL > 2W, and taking strainmeasurements at the back face (position M), Z(a) is givenas follows:11

- for a/W< 0.2 (shallow crack):

Z(a) = 2,532(Wa)1.5

1 25

aW 0.2

2

5.926 0.2 aW

2

0.288 0.2 aW+ 1

(2)

- for 0.2


4/13


Table 1 Mechanical properties of the Al-alloy tested at ambient temperature

Orientation/Property E (GPa) UTS (MPa) YS (MPa) ELa (%) RA (%)

AA2024-T3 longitudinal 79 477 350 21.1 20.8

Standard deviation 4.6 4.2 6.0 0.36 2.46

AA2024-T3 transverse 81 463 308 21.4 23.0

Standard deviation 2.0 3.8 4.5 0.31 1.45

Average of three test pieces for each specimen orientation.aL0 = 25 mm.

Fig. 3 Typical microstructures of the rolled A1-2024-T3 alloy: (a) Kellers etching; (b) Simply polished. Note different image

magnifications (approximately 2:1).

Fig. 4 T-L testpiece configurations for measuring transverse residual stress intensity factor on y direction, Kry, and for fatigue crack

propagation tests along the rolling or longitudinal (x) direction: (a) Slot at the centre of the weld line (0 mm), and (b) 5 mm displaced from

zero position. Tool shoulder width is 16 mm. FSW tools travelling from the bottom to the top of the page and rotating in counter-clockwise

direction.

were used to connect the strain gauges to the strain dataacquisition apparatus in order to minimize electromag-netic interferences. Plate cutting was performed by wireelectro-discharge machining (WEDM) either longitudi-

nally or transversally to the weldline. For the former case,as depicted in Fig. 4, the slot was introduced in two dif-ferent positions: 0 mm and 5 mm distant from the weldnugget centre line (on the FSW tool advancing side), to



5/13

530 M . T . M I L A N et al.

Fig. 5 L-T testpiece configuration for measuring longitudinal residual stress intensity factor on x direction, Krx, and for fatigue crack

propagation tests along transverse (y) direction. Strain measurments are taken at the point M. Tool shoulder width is 16 mm.

simulate a crack growing in the weld and in the heat af-fected zone, respectively. Recent work by Milan et al.13

has found WEDM more practical, precise and less likely

to introduce additional stresses compared to abrasive sawmethod. In all experiments, cutting increments of 0.5 mm

were chosen. Readings were taken 5 min after each singleslotting procedure had finished.After the strain data were obtained as a function of the

slot length, the secant method was used to derive d/dadata, which were then employed in Eq. (1) to determinecorresponding Kry and Krx values, for longitudinal andtransverse cracks, respectively.

Fatigue crack growth tests

A 5-mm-long notch was introduced at the edge of thewelded test pieces, in the positions indicated in Figs. 4 and5. Once a 5-mm-long fatigue pre-crack emanated fromthe notch tip, mode I fatigue loading tests were conductedat ambient temperature under constant amplitude loadingcondition (increasing-K testing). Sinusoidal waveformunder a frequency of 60 Hz was applied to all the testspecimens. The parent metal wastestedunder stress ratios(R) of 0, 0.3 and 0.5, respectively, in order to provide datafor the Glincas R method predictions, while the weldmetals where tested atR= 0.5 only. A detailed descriptionof the R method can be found in Milan and Bowen.4

R E S U L T S A N D D I S C U S S I O N

Residual stress intensity factor profile

Longitudinal cracking

Figure 6a presents transverse residual stress intensity fac-tor profiles, Kry, obtained for longitudinal slot test pieces.For both cases, that is, for the slot introduced at 0 mm and5 mm from the weld centre line, Kry values remain low

and negative in the entire range of acquired data. Rigor-ously speaking, a negative Kr value does not exist for aclosed crack, but it means that when an external load is ap-

plied, the thermal residual stress profile shields the crack,that is, effectively reduces the locally applied K factor,and hence decelerates FCGR. Clearly, this effect is a wel-come safety margin against failure in damage-tolerant ap-proaches. It is observed thatKry profiles vary only slightlyfrom one test piece to another, so that no systematic trendcan be established regarding the slot position in relationto the weldline, which means that nugget and thermo-mechanically affected zone (TMAZ) microstructures aresimilar to some extent with regard to fatigue properties.

Using the inverse weight function method, as detailedby Schindler,9 it was possible to obtain the initial residualstress profile present in the material before the slot wasintroduced, as depicted in Fig. 6b. Tensile residual stressin the centre of the test piece and balancing compres-sive residual stresses near the edges can be noticed. FromFig. 6, it canbe observed that compressive residual stressesmay delay crack nucleation at the edge of a pristine com-ponent. On the other hand, tensile residual stresses mayfavour nucleation at the centre of the piece. These find-ings have important implications in safe-life approaches.

Transverse cracking

Figure 7a shows longitudinal residual stress intensity fac-

tor profile, Krx, for test pieces with the slot axis perpen-dicular to the weldline. A negative peak value developsat approximately 16 mm from the edge of the test piece,so that Krx remains negative up to the weld centre line(30 mm position from the specimen edge). Therefore,the FCGR of a growing crack approaching perpendic-ularly to the weldline is expected to be slowed down.However, after the crack crosses the weld region (i.e. pos-itive Krx values in Fig. 7a) the FCGR will accelerate ascompared to the crack growth rate in the parent material



6/13


Fig. 6 (a) Transverse residual stress intensity factor profiles, and (b) Initial residual stress profiles derived for longitudinal slot testpieces.

under identically applied stress intensity factor range andR ratio.

In terms of non-destructive inspection programs ofdamage-tolerant structures, the above provided results

point out the need for a much more complex surveillanceschedule for welded structures as compared to monolithic(non-welded) ones in order to guarantee structural in-tegrity for both short- and long-fatigue growing cracks.

Figure 7b depicts theoriginal residual stressprofile in theflawless material, that is, before the slot introduction, asderivedthrough the inverse weight function method.9 Itispossible to observe the tensile residual stress approaching80% of the yield strength of the parent plate in the weldregion (in the advancing side of the tool), while near theedges the residual stresses remain compressive. A likely

crack nucleation at the TMAZ/HAZ zone can then bepredicted.

FCGR curves (experimental vs. predicted)

Longitudinal cracking

Fatigue crack propagation rate curves of longitudinallyslotted test pieces are presented in Fig. 8. For the sameapplied K level, a crack positioned on the centre ofthe weld nugget (0 mm position) exhibits fatigue growthrates slightly lower than a crack located along the TMAZat about 4 mm from the weld centre line. Compared tothe parent material however, all the welded test pieces



7/13

532 M . T . M I L A N et al.

Fig. 7 (a) Residual stress intensity factor, and (b) Corresponding initial residual stress profiles for transverse slot testpieces.

presented higher fatigue crack growth resistance at lowapplied K values. This is consistent with the presenceof a negative peak of Kry at a slot length of order of

5 mm (Fig. 6a). Nonetheless for higher applied K val-ues (above 10 MPa.m), that is, longer slots (cracks) andfaster crack growth rate, although Kry remains negativeduringthe entire range of testing, thewelded material pre-sented higher values of FCGR. This leads to the assump-tion that whereas microstructure and residual stresses actsimultaneously in the threshold region, atK values ap-proaching KC, monotonic fracture modes (dimple and/orcleavage micromechanisms) intercept fatigue crack prop-agation, that is, they accelerate FCGR as compared to

the parent plate behaviour. It is worth mentioning thatat these crack tip positions only microstructural effectscan play a role, because Kry values approach zero (refer to

Fig. 6a).Figure 9 shows the fatigue crack propagation rates esti-

mated via the so-called R method, for a test piece con-taining a crack in the weld centre line. For K valuesbelow 8 MPa.m, predicted values are conservative com-pared to the experimental ones, whereas for K valuesabove 10 MPa.m, predicted values are otherwise non-conservative. Because the prediction model takes intoaccount only the parent metal fatigue properties and the

Kry profile of the welded material, it is possible that



8/13


Fig. 8 Fatigue crack propagation curves of parent and welded

materials, with cracks propagating along the weld centre line (0

mm) and TMAZ (5 mm), respectively.

Fig. 9 Experimental and predicted FCGR of the welded material,

with the crack propagating along the weld centre line (0 mm).

the microstructure of the weld nugget is responsible forthe differences observed between predicted and experi-mentally measured figures. The nugget region is formedby fine equiaxed dynamically recrystallized grains,2 whilethe parent metal presents elongated grains due to therolling process, as can be seen in Fig. 10. Therefore, it is

likely that the intrinsic fatigue crack growth resistance ofthe nugget microstructure differs significantly from thatof the parent plate. Such a hypothesis is supported bythe fact that the discrepancies between predicted and ex-perimental values are higher at lowK values and nearthe final fracture region, where FCGR is much more de-

pendent on microstructural features. Conversely, in theParis region, where microstructure plays a minor role inFCGR, there is good correlation between experimentaland predicted values.

Briefly, recalling that at the threshold region (i.e. nextto the Kth value) conservative FCGR predictions werederived for the weld microstructure on the basis of theintrinsic fatigue behaviour of the parent plate, the resultsstrongly suggest that at low applied K values the formermicrostructure is more damage tolerant than the parentmetal. On the other hand, near the final catastrophicfracture region (i.e. approaching KC value), the non-conservatism of FCGR predictions for the weld nugget

indicates that monotonic fracture modes (i.e. KMAX-controlled mechanisms) are not precisely accounted forby Glinkas model, and evidences the parent metal as thetoughest microstructure at high applied K values.

For cracks positioned on the TMAZ and subjected tolow K values, predicted and experimental FCGR arein good agreement (Fig. 11), indicating that the TMAZmicrostructure exhibits the same intrinsic fatigue crackgrowth resistance as the parent material. It should bementioned that as the crack grew during the increasing-K tests, it deflected towards the interface between the

TMAZ and the heat affected zone (HAZ), as seen in

Fig. 12. As a result, FCGR predictions at higher K val-ues were underestimated (i.e.non-conservative approach).Assuming that residual stress profiles for both TMAZand HAZ regions are similar, the underestimated predic-tion suggests that the microstructure of the TMAZ/HAZinterface offers lower resistance to fatigue crack propa-gation than the parent plate material probably due to alower fracture toughness value, which would contributeto stronger KMAX effect on fatigue crack growth resis-tance. This point will be revisited in the next section fortransverse cracks. On the basis of Fig. 6a data, one couldargue that a less compressive stress field would have alsocontributed to higher crack growth rates. Regardless of

fact that the accelerated FCGR was generated by eithera less tough microstructure or a less compressive stressfield, or even by both effects simultaneously, the truth isthat the experiment did fail to test Glinkas model.

Transverse cracking

Fatigue crack growth curves for cracks running perpen-dicularly to the weldline are presented in Fig. 13. Upto a position corresponding to the border of the tool



9/13

534 M . T . M I L A N et al.

Fig. 10 Typical microstructures developed in FSW aluminium alloy. FSW tool advances in the right hand side region. The nugget consists

of recrystallized equiaxial grains, TMAZ of stretched grains due to rotational tool movement, and HAZ possesses similar appearance to the

parent metal.

Fig. 11 Experimental and predicted FCGR of welded material,

with the longitudinal crack propagating 5 mm from the weld centre

line.

shoulder, results clearly show that the welded jointpresents higher fatigue crack growth resistance than theparent material, for the same applied K value, or KMAX

value since the nominalRvalue is the same. This is mainly

Fig. 12 Longitudinal fatigue crack growth specimen showing crack

deflection towards the TMAZ/HAZ interface distant

approximately 8 mm from weld centre line. Penetrant-dye testing

was performed in order to facilitate crack visualization.

attributed to the strong negative Krx values shown inFig. 7a, in a position still located in the parent metal. Anegative Krx value reduces both the stress ratio, R, andthe effective K values, so decreasing the net crack tipdriving force. However, when the crack tip indeed en-ters the weld region, FCGR are found to be higher forthe welded joint than for the parent material, althoughthe Krx values still remain considerably negative (Fig. 7a).



10/13


Fig. 13 Fatigue crack propagation curves of welded joint and

parent material, with crack growing transverse to the weld line.

This fact suggests that the weld microstructure is brit-tle than the parent metal because at this position residualstresses approach zero.

Figure 14 provides Vickers microhardness profile trans-verse to the weldline at the mid-thickness position of thewelded joint. Although it agrees quite well with literaturedata for 2024-T3 Al-alloys,14 the relatively soft nuggetcould lead to the conclusion of a tougher microstructureif compared to the parent plate. However, recent studyby Kamp et al.15 has shown that alloys containing precip-itate distributions that are unstable at elevated tempera-ture, such as aerospace aluminium alloys, are quite proneto profound microstructural changes when friction stir

Fig. 14 Vickers hardness profile transversal

to the friction stir weld line.

welded. These include significant intermetallic particlescoarsening at grain boundaries in the HAZ and nuggetregions, hence leading to poor fracture properties even-tually associated to only modest bulk hardness values.

Furthermore, non-conservative Glinkas predictionsgiven in Fig. 15 endorse the assumption of brittle nugget

microstructure. Inasmuch as the R method takes intoaccount the fatigue properties of the parent material,it is possible, in the final fracture region (i.e. KMAX

KC), that the low-initiation fracture toughness (KC value)of the nugget microstructure enhances the FCGR bystrengthening monotonic tensile modes of fracture, re-sulting in somewhat underestimated FCGR predictions.

This was verified as well during longitudinal crackinganalysis, and it seems again that Glinkas model is notable to cope with this particular condition. At low appliedK values, however, a possible microstructural effect isdiscarded because the crack tip is still in the parent mate-rial. Thus, in this case, it is more likely that residual stress

relaxation is taking place due to the cyclic load applica-tion, which would result in non-conservative predictions.

The R method is only valid if linear elastic conditions aremaintained during the test, that is, if crack tip plasticity issmall. If this condition is broken, the obtained Krx profileis no longer the best choice to carry out the prediction.

Moreover, it deserves to be emphasized that FCGR pre-diction is based on initial KIr profile, which presents highnegative values as shown in Fig. 7a. In order to verify

whether there was plastic relaxation or not, both Krx andcorresponding rx profiles were again derived from thebroken halves of a fatigued test piece.

Results exhibited in Fig. 16 indicate that the material in-deed suffered significant stress relaxation by plastic flow.This is in line with recent findings from Liljedahl et al.,16

in which good prediction of residual stress redistribu-tion was achieved for growing cracks at constant stress



11/13

536 M . T . M I L A N et al.

Fig. 15 FCGR predictions for the welded joint compared to

experimental values, with crack growing transverse to the weld line.

intensity range, a scenario farther from situations in prac-tice than that studied in this work (i.e. increasing-Ktesting).The maximum nominal stress initially applied to the test

piece was approximately 80 MPa. By adding this valueto the peak tensile residual stress observed on a positioncorresponding to the advancing side of the rotating tool

(Fig. 7b), an effective tensile stress approaching 300 MPais obtained (Fig. 16b). Such stress level is certainly highenough to cause plastic deformation as the elastic limitof the material is around 250 MPa. Plastic deforma-tion reduces both the peak tensile residual stress andthe compressive residual stress near the edges of the testpiece because a zero-stress global balance must be main-tained. Thus, the initial Krx values calculated consider-ing the non-relaxed residual stress profile are likely to beoverestimated, resulting in non-conservative predictionofFCGR.A new FCGR prediction of the welded joint based on the

relaxedKrx profile is given in Fig. 17.Results indicate that

even considering the residual stress relaxation, forecastvalues for the final fracture region within the weld regionare still non-conservative, so confirming that monotonicfracture modes do prevail at higher K values. However,for K values below 17 MPa.

m it is now observed thatpredicted values are invariably higher than experimentalones, evidencing that the relaxedKrx profile is now under-estimated. Therefore, the real Krx profile during fatiguetesting must range from that obtained with the intacttest pieces to that derived from the test specimens halves,

inasmuch as the relaxation takes place progressively asloading cycles are applied to the test samples. This can beconfirmed in Fig. 16b, when a single load cycle is enoughto cause a detectable stress relaxation and hence a rear-ranged residual stress profile.

Last but not least, it should be noted that the trans-

verse crack growth case illustrates very well the limi-tation of the presently adopted approach in identifyingthe main controlling factor of FCGR and drawing con-clusive facts about it. Aimed at separating the effects ofthe magnitude of the residual stresses and the weld mi-crostructure on fatigue behaviour of the FSW plates,FCGR experiments employing pre-strained test piecesare being carefully planned. FCGR testing of these stress-relieved specimens certainly will clarify current uncertain-ties and allow the role of the secondary stresses to be fullyunderstood.

C O N C L U D I N G R E M A R K S

Longitudinal and transverse fatigue crack growth prop-erties of friction stir welded joints of thin rolled plates of2024-T3 aluminium alloy have been studied under con-stant load amplitude conditions (increasing-K).

Corresponding residual stress intensity factors in mode I(KIr) profiles were determined using a fracture mechanicsapproach, and the inverse weight function method wasutilized to obtain the residual stress profiles.

Fatigue properties of the parent metal were employedalongside KIr profiles and Glinkas R method in order topredict FCGR of the welded joints.

Main conclusions that may apply to the weld region, butnot to necessarily a welded component as a whole, are asfollows:

1. Welded joints are more resistant to longitudinally grow-

ing fatigue cracks than the parent metal at threshold

K values probably due to beneficial thermal resid-

ual stresses arresting crack propagation, but less resis-

tant next to the final instability point probably owing

to monotonic tensile fracture modes intercepting sub-

critical crack growth in weld microstructures. Conse-

quently, FCGR predictions trend to be conservative at

lower propagation rates and non-conservative for fastercracks;

2. Intense compressive residual stresses render welded

plates more fatigue resistant to transverse cracks than

unmixed parent plate. Nonetheless, once the crack

tip enters the brittle weld region FCGR accelerates

due to operative monotonic tensile modes of fracture,

leading to non-conservative crack growth rate predic-

tions next to KC instability. At threshold K values

non-conservative predictions values result from residual



12/13


13/13

538 M . T . M I L A N et al.

Fig. 17 FCGR predictions for the welded joint (crack growth

transverse to the weld line) compared to the experimental values,

considering the Krx profiles obtained, respectively, from intact

testpieces and broken halves of fatigue specimens.

R E F E R E N C E S

1 Itoh, Y. Z., Suruga, S. and Kashiwaya, H. (1989) Prediction of

fatigue crack-growth rate in welding residual-stress field.

Engng. Fract. Mech. 33, 397407.

2 Bussu, G. and Irving, P. E. (2003) The role of residual stress

and affected zone properties on fatigue crack propagation infriction stir welded 2024-T351 aluminium joints. Intl. Jnl.

Fatigue 25, 7788.

3 Milan, M. T. and Bowen, P. (2002) Effects of particle size,

particle volume fraction and matrix composition on the

fatigue crack growth resistance of Al alloy / Al alloy+ SiCbimaterials. Proceedings of the Institute of Mechanical

Engineering L. Jnl Mater. Des. Appl. 216, 245255.

4 Milan, M. T. and Bowen, P. (2003) Experimental and

predicted fatigue crack growth resistance in Al2124/Al2124+35%SiC biomaterial. Intl. Jnl. Fatigue 25, 649659.

5 Parker, A. P. (1984) An overview of the mechanics of fracture

and fatigue in the presence of residual stress. J. Mech. Work.

Tech. 10, 165174.

6 Milan, M. T. and Bowen, P. (2004a) Fracture toughness of

selectively reinforced Al2124 alloy: precrack tip in the

composite side. Metal. Mater. Trans. A 35A, 13931401.

7 Milan, M. T. and Bowen, P. (2004b) Fracture toughness of

selectively reinforced Al2124 alloy: precrack tip in thealuminium alloy side. Mater. Sci. Tech. 20, 783789.

8 Glinka, G. (1979) Effect of Residual Stresses on Fatigue Crack

Growth in Steel Weldments under Constant and Variable

Amplitude Loads, ASTM Special Technical Publication (STP)

677, American Society for Testing and Materials,

Philadelphia, PA, USA,2 pp. 198214.

9 Schindler, H. J. (1996) Determination of residual stress

distributions from measured stress intensity factors. Intl. Jnl.

Fract. 74, R23R30.

10 Prime, M. (1999) Residual stress measurement by successive

extension of a slot: the crack compliance method. Appl. Mech.

Rev. 52, 7596.

11 Schindler, H. J. and Bertschinger, P. (1997) Some step

towards automation of the crack compliance method to

measure residual stress distributions. In: Proceedings of the 5th

International Conference on Residual Stresses ICRS-5. (Edited by

T. Ericsson, M. Oden and A. Andersson), Linkoping,

Sweden, Vol. 1.

12 Fett, T. and Munz, D. (1997) Stress Intensity Factors and

Weight Functions, Computational Mechanics Publications,

Southampton, UK.

13 Milan, M. T., Bose Filho, W. W., Malafaia, A. M. S., Silva, C.

P. O. and Pellizer, B. C. (2006) Slot machining effects on

residual stress measurements using the crack compliance

method. Jnl. Test. Eval. 34, 149152.

14 Khaled, T. (2005) An Outsider Looks at Friction Stir Welding.

Report #: ANM-112N-05-06. Federal Aviation

Administration, Lakewood, USA.

15 Kamp, N., Sullivan, A., Tomasi, R. and Robson, J. D. (2006)Modelling of heterogeneous precipitate evolution during

friction stir welding process. Acta Mater. 54, 20032014.

16 Liljedahl C. D. M., Brouard, J., Zanellato, O., Lin, J., Tan, J.

F., Ganguly, S., Irving, P. E., Fitzpatrick, M. E., Zhang, X.

and Edwards, L. (2007) Weld residual stress effects on fatigue

crack growth behaviour of aluminium alloy 2024-T3. In:

Proceedings of the First International Conference on Damage

Tolerance of Aircraft Structures. (Edited by R. Benedictus, J.

Schijve, R. C. Alderliesten and J. J. Homan), Delft, The

Netherlands.

c 2008 Th A th J l C il ti c 2008 Bl k ll P bli hi Ltd F ti F t E M t St t 31 526 538

Documents

Fatigue & Fracture of Engineering Materials & Structures 1