Fatigue crack propagation into the residual stress ﬁeld along and.pdf

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doi: 10.1111/j.1460-2695.2011.01631.x

Fatigue crack propagation into the residual stress field along andperpendicular to laser beam butt-weld in aluminium alloy AA6056

W . V . V A I D Y A , P . S T A R O N a n d M . H O R S T M A N N

Helmholtz-Zentrum Geesthacht, Institute of Materials Research, Max-Planck-Strasse 1, D-21502 Geesthacht, Germany

Received in final form 31 August 2011

A B S T R A C T Laser beam butt welds in Al-alloys are very narrow and are accompanied by steep residualstress gradients. In such a case, how the initial crack orientation and the distance of thenotch tip relative to the weld affect fatigue crack propagation has not been investigated.

Therefore, this investigation was undertaken with two different crack orientations: alongthe mid-weld and perpendicular to the weld. Fatigue crack propagation ‘along the mid-

weld’ was found to be faster in middle crack tension specimens than in compact tensionspecimens. For the crack orientation ‘perpendicular to the weld’, the relative distancebetween the notch tip and the weld was varied using compact tension specimens to

generate either tensile or compressive residual stresses near the notch tip. When tensileresidual stresses were generated near the notch tip, fatigue crack propagation was found tobe faster than that in the base material, irrespective of the difference in the initial residualstress level and whether the crack propagated along the mid-weld or perpendicular tothe weld. In contrast, when compressive weld residual stresses were generated near thenotch tip, fatigue crack arrest, slow crack propagation, multiple crack branching and out of plane deviation occurred. The results are discussed by considering the superpositionprinciple and possible practical implications are mentioned.

Keywords AA6056; aluminium alloy; fatigue crack propagation; laser beam weld; resid-ual stresses; weld orientation.

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

Construction by welding is very common in industries.In civil aircrafts, however, welding of Al-alloys had not been so common in the past, because when fusion weldedby conventional technique, airframe Al-alloys, particu-larly AA2024 and AA7075, tend to become susceptibleto hot cracking.1 With suitable alloys and the appro-priate welding technique, Neye and Heider2 could show that laser beam welding can produce narrow fusion weldsof airworthiness quality. This has opened new perspec-

tives of savings in weight and cost through welded in-tegral structures,3 and laser beam welded structures arenow becoming a regular feature of some civil aircrafts.4,5

Additionally, application of friction stir welding6 (solid-state process) and other joining techniques7 are also beingconsidered. A side effect accompanying the weld is the generation

of residual stresses. The latter are in general known to

Correspondence: W. V. Vaidya. E-mail: waman.vaidya@ hzg.de

affect fatigue crack propagation.8 Some reviews (e.g. Refs[9–11]) show how fatigue life can be affected by residualstresses. When the crack initiation phase is likely to beshort or absent due to weld defects, it is safer to assumea limited life restricted only to fatigue crack propagation.

The role of residual stresses thereby is to modify the localstress by changing the mean stress. Using the principleof superposition,12 a residual stress intensity factor, K rs,may be defined at a given crack length and then added tothe applied (cyclic) stress intensity factor. This is shownschematically in Fig. 1 fortwo cases assuming hypothetical

tensile and compressive weld residual stresses at a fatiguecrack tip. Although then theK level remains unchanged,tensile residual stresses increase the local stress ratio from

Rappl (= P min / P max = S min / S max = K min / K max)toarealvalue Rlocal [ =( K min + K rs)/( K max + K rs)]. When a certain stressintensity level K op is exceeded, the crack remains open,the shielding effect of crack closure is lost and crack prop-agation rate increases. Hence, although welds are actually tested at low Rappl, a high local stress ratio is likely to bereached.13 On the other hand, near crack tip compressive

c 2011 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 35, 399–411 399



400 W . V . V A I D Y A et al .

Fig. 1 Modification of the applied loading pattern through the

superposition due to tensile and compressive weld residual stresses

(schematics based on Parker,12 Ohta et al .13 and Withers and

Bhadeshia.10 Residual stress patterns shown by hatching on the

specimens are hypothetical ).

residual stresses (i.e. K rs → − ve) have an opposite effect and are beneficial, at least as long as plastic deformationat the crack tip does not occur under the applied tensileloading or when compressive residual stresses are not re-lieved or under both conditions.11 Effects such as slow propagation or retardation are very often observed dueto compressive residual stresses induced by local treat-ments such as laser glazing,14 laser surface annealing,15,16

laser shock processing,17 laser heating18 and spot heat-ing.19 However, it has been also observed in a partially laser shock-peened structural steel that through-thicknesstensile residual stresses, sandwiched between compressiveresidual stresses,keep the crack open anddominate fatigue

propagation, whereby relaxation of residual stresses wasfound to be negligible and crack retardation did not oc-cur.20 In a forged AA7XXX alloy, tested with and without stress relief, residual stresses are found to affect fatiguecrack propagation, particularly in the low K-regime.8

Therefore, the nature (in terms of the sign) and the dis-tribution of residual stresses deserve consideration. Ten-sile residual stresses are generally detrimental (Fig. 1).Such stresses are released, redistributed or relaxed when(weld) coupons are cut 8,13 and specimens are tested un-

der fatigue loading.11,21 Hence, Rlocal is a variable (andnot a constant). Cold working,22 stress relieving8 or ap-propriate post-weld heat treatment as for steels13 can re-move residual stresses. In a laser beam welded Al-alloy,however, post-weld ageing did not reduce weld resid-ual stresses significantly due to low ageing temperature

(190

◦

C).

23

Thus, whether weld residual stresses are re-moved should be dependent on temperature of post-weldtreatment,13 extent of cycling11,21 and deformation11,22

and the K level because the threshold regime is affectedmore than the Paris regime.8,13 The latter regime, be-ing of interest for the damage tolerance analysis, is theone mostly investigated.22,24–34 Although residual stresspattern differs basically between middle crack tension,

M(T), and compact tension, C(T), weld-specimens13 andmay lead to specimen-geometry-dependent fatigue crack propagation,33,34 the C(T) geometry is the one mostly used22,24–28,30–34 due to material saving and experimentaladaptability to crack orientation relative to the weld.

The crack orientation, whether along the mid-weld orperpendicular to theweld,and thedistanceof thenotch tiprelative to the weld (which determines the residual stressgradient) are known to affect fatigue crack propagationin fusion welds.24,25 The crack orientation effect 24 hasbeen confirmed by recent results on friction stir welded

Al-alloys, in that compared to a crack propagating alongthe mid-weld, a crack propagating perpendicular to the

weld can slow down, retard or stop,22,26–30 also when thedistance between the notch tip and the weld centre lineis varied.22,30 This behaviour, which is similar to that ob-served in locally treated materials,14–19 is attributed to the

existence of compressive residual stresses22,24–30

and ex-plained by considering the superposition principle.24,26,27

The crack orientation effect 24 is confirmed also in laserbeam welded and laser scanned 304 austenitic steel, andis found to be lost, when the distance to notch tip wasdecreased or after stress relief treatment due to modi-fication of residual stresses.31 In contrast, tungsten inert gas welds containing phase-transformed martensite in the

weld zone of Ti-6Al-4V are found to exhibit a different behaviour, namely, in the low K regime and comparedto the base material, the resistance to crack propagation

was better along the mid-weld and within the heat affectedzone, but inferior when perpendicular to the weld.32 Such

difference was lost in the Paris regime and all crack ori-entations exhibited behaviour similar to the base mate-rial. Thus, microstructure may dominate residual stressesresulting from welding32 or quenching35,36 when defor-mation is restricted (due to martensite or Widmanst ¨ attenmicrostructure) or affected. This possibility is, of course,unlikely in Al-alloys, because phase transformations aremostly restricted to precipitation or to dissolution of phases in welds37 and the resulting microstructure con-tinues to be deformable.

c 2011 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 35, 399–411



F A T IG U E C RA C K P RO P A GA T I ON I N TO T H E R ES I D UA L S T RE S S F IE L D O F L A S E R B E A M WE L D A A6 0 5 6 401

Compared to Al-alloy welds by friction stir22,26–30 andconventional fusion techniques,2,24,33,34 laser beam Al-alloy welds aredue to high power density used very narrow in width.2,37 In turn, the gradient in the residual stresspattern nearer to the laser beam weld is steeper23 thanthat in broad welds.24–30,33,34 Whereas the fatigue (S-N)

behaviour of laser beam welded Al-alloys has received at-tention,38,39 data on fatigue crack propagation are scarce40

and are lacking completely for a crack perpendicular tothe weld. Therefore, this work was undertaken to study the effect of crack orientation22,24–34 and the distance be-tween the notch tip and the weld22,25,30,31 on fatigue crack propagation in a laser beam welded butt joint of the air-frame Al-alloy AA6056.41–43

E X P E R I M E N T A L P R O C E D U R E

The alloy AA6056 is an Al-Mg-Si-Cu precipitation hard-enable alloy and is considered for replacement of AA2024

in fuselage.3,41–43 Whereas AA2024 is likely to be proneto hot cracking by conventional fusion welding,1 AA6056is fusion weldable and is suitable for welded integral con-structions.3,41–43 The sheets investigated were bare andhad a nominal thickness of 6 mm (=B). The reason forusing thicker sheets for this investigation was the sharp

weld residual stress profile and the high relative peak stress value (190 MPa), which was about 90% weld material yield strength.23 The presence of such high tensile resid-ual stresses is expected to facilitate fatigue crack propa-gation, because tensile residual stresses are detrimental(Fig. 1).

Sheets were peak-aged to T6 condition (190◦

C–4h)be-fore welding. Blanks (250 mm width and 500 mm length) were cut in the rolling direction, milled, cleaned and de-greased. The blanks were vacuum clamped and welded tothe butt joint configuration with two 3.5 kW CO2 lasers(Rofin Sinar DC035) using the dual beam technique andthe filler wire AlSi12 (AA4047). Welding power used was

2 × 3.25 kW at a speed of 2.0 m/min with Ar inert gasshielding. The welded coupons were 500 mm wide and500 mm long with the weld in the centre and along therolling direction. When welded in the T6 condition, themicro-hardness profile, hence, properties are found tostabilise after a post-weld natural ageing of 7 weeks.44

Therefore, all testing was carried out after this timeinterval.Specimens were prepared by water jet cutting and ma-

chining. To facilitate surface observations by microscope,the excess weld material was skimmed off by milling andthe specimen surface polished. Due to the possibility of tailoring the residual stress gradient, tests were carriedout mostly on C(T) specimens, except for one test on

M(T) specimens. C(T) specimens were extracted in twonotch tip orientations; along the mid-weld and perpen-dicular to the weld. In the latter case, the distance be-tween the notch tip and the weld was varied (as to bereported below). Because residual stress pattern changes

basically between the weld coupon and the C(T) specimenextracted,13 measurements were undertaken directly onrepresentative C(T) specimensof a given variant using ourGeesthacht Neutron Facility as reported previously.23,37

The principal axis ‘ x’ was placed in the direction of crack propagation, ‘ y’ in plane and perpendicular to ‘ x’ and ‘ z’out of specimen plane as shown in Fig. 2. The slit width

was 3 mm and the nominal size of the gauge volume mea-sured was about 27 mm3. This is depicted on the weldcross-section in Fig. 3. Due to divergence of the neutronbeam, the true gauge volume was slightly larger than that intended. Moreover, adjacent to the weld interface some

volume from the base material was also scanned by theneutron beam due to the difference in the shape of the weld (Fig. 3). Measurements were performed in allthree directions, x, y and z and residual stresses were de-termined under the assumption of a plane stress state.Because elastic anisotropy in aluminium is small, values

E = 70 GPa and ν = 0.33 were used for (311)Al reflection.

Fig. 2 Specimens used along with the

coordinate notation for residual stress

measurements by neutron diffraction.

Whereas ‘ z’ is out plane of the paper, ‘ x’ and

‘ y’ are coplanar and ‘ x’ is placed always in

the direction of fatigue crack propagation,

irrespective of the weld position. The cut at

the specimen top was used for fixing the

on-top displacement gauge for indirect

crack length measurement (dimensions in

millimetre).




402 W . V . V A I D Y A et al .

Fig. 3 Shape of the laser beam weld investigated. The square

indicates the approximate size and shape of the gauge volume for

measurements of strains in the two directions perpendicular to the

weld.

Scatter in the given residual stress measurement data is

about ±10 MPa.C(T) specimens had width W = 100 mm and the initial

notch length an was 20 or 25 mm. To study specimengeometry effect additionally, M(T) specimens with W =

200 mm [equivalent to C(T)100 specimen] were tested. The weld was located at the centre of the M(T) specimensuch that the loading was perpendicular to the weld andcrack propagation occurred along the mid-weld. The ini-tial notch length 2an for these specimens consisted of acentral bore, Ø1 mm, extended by slits of each 1.5 mm oneither side. Specimens were tested using servo-hydraulicmachines of appropriate capacities, 25 kN for C(T), and

400 kN for M(T) specimens, and by operating in the load-controlled sine-wave mode in close compliance with the ASTM Standard E647.45 Testing was carried out underconstant amplitude conditions (K increasing tests) at load ratio Rappl = 0.1. The test frequency, f , was 10 Hz.Due to limited availability of the welded material, we didnot investigate redistribution of residual stresses and de-termined K rs through techniques such as sawcuts8 (cut-compliance technique). This is an open point and needs tobe covered in a future investigation. Presently, the crack propagation data are considered in terms of K and thereference to K rs is qualitative.

Crack length was monitored continuously and indirectly

using an on-top (clip-on) displacement gauge for C(T)specimens (attached to the cut at the specimen top; Fig. 2)and a load-line displacement gauge for M(T) specimens,and calculated from the maximum crack opening dis-placement (CODmax) as described in detail elsewhere.46

In brief, the CODmax value was recorded continuously at a signal interval of 20 mV. Occasionally, crack length wasalso measured directly on opposite specimen surfaces us-ing optical microscope, and the averaged optical value wasused for calibration of the corresponding CODmax value.

Because the shape of CODmax versus direct crack lengthdiffers from specimen to specimen (due to characteris-tics specific to a given specimen), the calibration curvehas to be obtained for each and every specimen. On theother hand, the specimen-specific information (say resid-ual stresses) and its effect on crack length are then already

accounted for (in the shape of CODmax curve). Moreover,the calibration curve obtained has nearly 1:1 correlationbetween the direct and the indirect crack length measure-ment and can be used directly for the crack length con-

version without any correction as required for the com-pliance technique. In contrast to other techniques suchas potential drop or compliance, the CODmax values, be-ing obtained at the maximum load, are less affected by fracture surface contact or crack closure. The importanceof this point would be clear when it is considered that at a low Rappl the level of crack closure would increase inthe presence of compressive residual stresses (Fig. 1; K rs

→ − ve). In that case, the applicability and the accuracy

of potential drop or compliance technique would be ad- versely affected. In this context the CODmax technique46

was found to be less susceptible.Base material specimens for comparison were extracted

from the weld coupons. All fatigue crack propagation data were considered over a normalised crack length a/W ≤

0.65 (i.e. over crack length, a, shorter than 65 mm), al-though data were achieved over a longer crack length.

After testing, specimens were selectively investigated by radiography to understand the interaction between thefatigue crack and isolated macro-pores.

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

The weld was symmetrical about the butt joint line,but was broad at the top (laser incidence) and narrow at the bottom surface as shown in Fig. 3. Nonetheless,the micro-hardness profile was nearly comparable on thetop and the bottom surface, and over the mid-thickness,except for lateral displacement along the weld distance.Hence, hardness in the vertical direction varied within the

weld (and along the later crack front). As inferred fromthe mid-thickness profile, the fusion zone extended over±2.5 mm from the weld centre and had the least hard-

ness (76 VHN). The hardness trough resulting from thefusion zone and the heat affected zone extended on eitherside up to ±8 mm from the weld centre before reachingthe average base material hardness level (128 VHN). Be-cause the excess weld material was skimmed off by millingfor fatigue crack observations, the only weld defect re-maining was porosity, which is common in Al-welds dueto entrapped gaseous impurities, particularly hydrogen.47

Large isolated pores (Ø ≤ 1 mm) were occasionally ob-served, however, compared to micro-pores (Ø ≤ 100 μm)





and small pores (Ø ≤ 500 μm) had no adverse effect onfatigue crack propagation. On the contrary, such pores in-duced crack retardation.48 The mid-weld residual stressesin the coupon (before cutting and machining) were tensilein nature (60 MPa ≤ σ x ≤ 190 MPa and 5 MPa ≤ σ y ≤

25 MPa).23

Notch orientations and residual stress patterns

To recall the experimental background, the initial crack orientation24 and the relative distance between the notchtip and the weld22,25,30,31 can affect residual stress patternand fatigue crack propagation. To study these aspects, twonotch tip orientations were selected:

• along the mid-weld and

• perpendicular to the weld

as shown in Fig. 4. Whereas the former position ‘alongthe mid-weld’ is unique (Fig. 4a), crack location in thelatter case canbe anywhere, say within the fusion zone, theheat affected zone or the base material containing residualstresses. Our interest was to let the crack propagate under

different residual stress gradients by changing the notchtip position relative to the beginning of the weld (denotedby X ). Therefore, following positions were selected forcrack ‘perpendicular to the weld’:

1 Notch tip at the beginning of the weld ( X = 0 mm):

The crack propagates across the fusion zone, across the

heat affected zone and into the base material containing

residual stresses, Fig. 4b.

2 Notch tip at the end of the weld ( X = 5 mm): This

position is comparable to one above, however, the crack

starts at the end of the fusion zone and propagates across

the heat affected zone into the base material containing

residual stresses, Fig. 4c.

3 Weld nearer to the notch tip ( X = 15 mm): The crack

propagates from the base material containing residual

stresses towards the heat affected zone before entering

the fusion zone. After crossing the fusion zone, propaga-

tion would continue across the heat affected zone on theother side and into the base material containing residual

stresses, Fig. 4d.

4 Weld in the mid-ligament ( X = 40 mm): This position

is comparable to that above, except that the weld is now

Fig. 4 Location of the weld relative to the notch tip, (a) along the mid-weld and (b)–(f) perpendicular to the weld with the given distance X

between the notch tip and the weld as shown ( X is in millimetre).




404 W . V . V A I D Y A et al .

in the middle of the ligament and away from the notch

tip, Fig. 4e.

5 Weld at the end of the specimen ( X = 75 mm): The

crack propagates fromthe basematerial containing resid-

ual stresses and across the heat affected zone towards the

fusion zone. Because results were obtained for limited

crack length (a/W ≤ 0.65; a ≤ 65 mm), this was basically

crack propagation only in the base material (L-T ori-

entation), Fig. 4f, with the difference that this ligament

portion contained residual stresses.

Positions in Figs 4b and c simulate cracks propagatingfrom the weld, and in Figs 4d–f cracks propagating fromstress raisers in the vicinity of the weld. The positions inFigs 4b, c and f are also called ‘terminal positions’ andthose in Figs 4d and e ‘non-terminal positions’.

Residual stress profiles for the cases described above areshown in Fig. 5. The σ x was placed always along the direc-

tion of crack propagation (Fig.2). The nature of the resid-ual stress component acting perpendicular to the crack contributes to opening (when tensile in nature; K rs →

+ ve) or closing (when compressive in nature; K rs → − ve)of the crack.24 Hence, the profile of σ y (shown by open

symbols in Fig. 5) is decisive for crack propagation. Here,the base material specimens from the welded coupon werealso investigated for comparison. Because residual stressesare not restricted only to the weld, but extend over the

whole coupon, the base material specimens were not com-pletely free from residual stresses. However, the averaged

stress level was very low (σ

x≤ −

25 MPa) andσ

y be-ing tensile in nature (≤ 20 MPa) should not affect crack propagation of the base material significantly. This wasconfirmed within the experimental tolerance limit.49 It isas well likely that these residual stresses are resulting frommicro-stresses, because comparable values were found inheat treated (peak-aged), but non-welded base material.Such specimens, in the present context and in comparison

with the weld containing specimens, are considered hereas nearly stress-free.

Over the ligament investigated (a/W ≤ 0.65; a ≤ 65 mm),the residual stress level within the weld was also very low (5 MPa ≤ σ x ≤ 50 MPa and −15 MPa ≤ σ y ≤ 25 MPa)

and was reduced compared to that before cutting (60 MPa≤ σ x ≤ 190 MPa and 5 MPa ≤ σ y ≤ 25 MPa),23 as to beexpected8,13 due to stress relaxation on the extraction of specimens from the welded plates. The stress gradients‘along the mid-weld’ were very weak, Fig. 5a. In contrast,

Fig. 5 Residual stress gradients measured by neutron diffraction in the specimen batches investigated, (a) along the mid-weld and (b)–(f)

perpendicular to the weld. Gray shading shows the relative weld position. In (c), the notch tip is positioned at the end of the weld.





steep gradients were generated in the case of crack ‘per-pendicular to the weld’, Figs 5b–f. For the case ‘notchtip near the weld’, the residual stress profile was nearly identical, irrespective of whether the notch tip was placedat the beginning or at the end of the weld, Figs 5b andc. Whereas the nature of residual stresses (σ y ) adjacent to

the weld was always tensile, the nature of near-notch tipresidual stress was different, ranging from tensile (Fig. 5b)to compressive (Figs 5d & e) and the stress level differedsignificantly (−100 MPa ≤ σ y ≤ 150 MPa). Compressiveresidual stresses were present along the gradient at notchtip (Figs 5d & e), and away from the notch tip (Figs 5b, c& f).

Fatigue crack propagation

For a number of testing parameters, fatigue crack prop-agation data in the mid-regime of the base material arefound to be confined to an experimental tolerance limit of about ±35%.49 This tolerance limit is also assumed in thisinvestigation. In the following, typical trends observed in

welds are reported on selected specimens. As mentioned,the base material fatigue crack propagation data were ob-tained on nearly stress-free specimens extracted from the

welded coupon, and data are shown here for compari-son with the welds. The base material had partially re-crystallised grain structure, was pan-caked with long axisin the rolling direction and compared to the L-T orien-tation, exhibited transition in fatigue crack propagationafter about 20 MPam0.5 in the T-L orientation. More-over, M(T) specimens exhibited faster fatigue crack prop-

agation and yielded more conservative data than C(T)specimens.49 This is in agreement with the literature andthe difference due to specimen geometries is to be at-tributed to the asymmetrical loading, increased crack clo-sure, smaller process zone and higher crack tip constraint in C(T) specimens. In welds, crack closure is expectedto be dominated by the superposition of tensile residualstresses with the applied stresses (Fig. 1).

Fatigue crack propagation along the mid-weld

This configuration was unique (Fig. 4a), except for thedifference in the specimen geometry, C(T) and M(T).

The difference in fatigue crack propagation due to spec-imen geometries was evident more in welds than in thebase material specimens, Fig. 6. Weld containing M(T)specimens exhibited faster crack propagation than weldcontaining C(T) specimens. Such difference in the basematerial specimens was much lower and limited to thetolerance limit.

In C(T) specimens, the residual stress gradient was weak and the near-notch tip stress level was nearly zero(Fig. 5a). Hence, a pronounced effect on fatigue crack

Fig. 6 Fatigue crack propagation along the mid-weld and itsdependency on the specimen geometry. Hatching patterns mark

the base material data for M(T) and C(T) specimens in the same

orientation. Arrows indicate local crack retardation due to

macro-pores.

propagation was not to be expected and was also not ob-served, as for ‘weld-C(T)1’ specimen (shown by invertedtriangles in Fig. 6). If,however, isolated macro-pores wereencountered, initial crack propagation was delayed and

was even slower than that in the base material as for ‘weld-C(T)2’ specimen (upright triangles in Fig. 6). The crack

retardation effect due to isolated macro-pores48

is indi-cated in Fig. 6 by downward arrows and was observed overthe entire regime, whereby weld containing C(T) speci-mens exhibited inferior fatigue crack propagation than thebase material after about 18 MPam0.5. Because the resid-ual stress level was low (Fig. 5a), effect due to superpo-sition would have been negligible. Hence, the differencein fatigue crack propagation should be attributed to thedifference in the microstructure. In other words, here thecast microstructure (weld) is found to be slightly inferiorto the wrought microstructure (basematerial). In contrast,such difference is not found in metal arc inert gas welds of

AA7005 and the base material.24 Insofar the difference in

the welding process, the faster cooling rate in laser beam welds and the difference in material composition in thisstudy appear to have caused this difference. A given residual stress pattern keeps on changing

through different processes, say on specimen extraction ortesting. A good discussion of the changes in residual stresspatterns in test specimens that occur upon extraction fromparent plates or forgings, as well as the influence that thishas on fatigue crack propagation rate properties was givenby Bucci.8 Ohta et al .13 have pointed out that the nature




406 W . V . V A I D Y A et al .

and the extent of residual stress pattern can change onextraction of specimens, and that such difference is lessbetween the welded coupon and M(T) specimens, but more between the welded coupon and C(T) specimens.

The weld in our M(T) specimens is estimated to havehigher residual stress level than C(T) specimens, compa-

rable to that in the coupon (60 MPa≤ σ

x≤

190 MPaand 5 MPa ≤ σ y ≤ 25 MPa).23 In turn, as to be expectedfrom the superposition (Fig. 1), residual stresses and alsothe cast microstructure have contributed to faster propa-gation. The crack retardation was very weakly developed,because of few pores encountered and also due to low constraint in M(T) specimens. Thus, as for the base ma-terial,49 M(T) specimens provide more conservative dataalso for welds. The data on C(T) specimens indicate that tensile residual stresses, although small in level, acceleratecrack propagation and, hence, deserve precaution.

Fatigue crack propagation perpendicular to the weld Assuming σ y to be decisive for crack propagation,24 twotrends were observed in terms of the nature of residualstresses: the near-notch tip tensile stresses for the terminal

weld positions (Figs 4b & c and Figs 5b & c; K rs → + ve)and the near-notch tip compressive stresses for the near-notch and the mid-ligament weld positions (Figs 4d & eand Figs 5d & e; K rs → − ve). For the weld at the end of specimen (Fig. 4f), the near-notch tip residual stress wasnearer to zero (Fig. 5f).

Even though in the terminal weld positions the near-notch tip residual stress levels differed slightly along the

crack plane (−25 MPa ≤ σ x ≤ 5 MPa) and substantially inthe perpendicular direction (−15 MPa ≤ σ y ≤ 150 MPa),and compressive stresses were present at different loca-tions along the specimen centre line, fatigue crack prop-agation in the weld containing base material was nearly identical in these positions, Fig. 7. Except for near thenotch tip propagation across the weld, fatigue crack prop-agation occurred mainlyin thebase material in these cases.Nonetheless, the stress-free base material and the weldcontaining base material did not exhibit the identical fa-tigue crack propagation behaviour (Fig. 7). Such obser-

vation has not been reported in the literature, probably because specimens with small ligaments [such as C(T)50]

are usually tested. Smaller the specimen, less would bethe extent of residual stresses initially retained and flatter

would be the residual stress gradient along the specimencentre line. In that case, residual stresses may be removedefficiently and the difference, as we have observed (Fig. 7),may not be detectable.

Returning to the data, although fatigue crack propaga-tion occurs in the base material, the ligament in the weld-containing specimens did contain residual stresses (Fig. 5)in contrast to the base material (which was free from weld

Fig. 7 Fatigue crack propagation perpendicular to the weld in theterminal positions. The weld position is shown in the inset along

with the symbols. Hatching pattern marks the base material data

for C(T) specimens in the same orientation.

residual stresses). The difference in the data in Fig. 7indicates that the base material containing weld residualstresses is prone to fatigue crack propagation, even thoughresidual stresses should have been redistributed, reducedor relieved by the crack extension.11 Moreover, becausecrack propagation was (slightly) faster and not identical tothat in the base material over a wide range of K, it seems

that crack propagation was dominated by tensile residualstresses; those already present and those redistributed dueto the crack advance. It seems that the initial compressivestresses, being small (Fig. 5f), have been removed andredistributed during crack propagation. Hence, the dif-ference in the initial level of σ y for the terminal positions‘notch tip at the beginning of the weld’ and ‘weld at theend of the specimen’ seems to loose its effectiveness. The crack appearance in these specimens is shown in

Fig. 8. Initially (a/W ≥ 0.4; a ≥ 40 mm), the crack prop-agated along the mid-plane. Thereafter, deviation wasobserved and the tests were stopped when the local devi-ation became more excessive than the validity criterion.45

Nonetheless, the deviation has not caused data variability (Fig. 7). Interestingly, we did not observe such an ex-cessive deviation in the base material for different crack orientations (L-T, T–L) or for the crack along the mid-

weld, either previously,48 or presently. Thus, this effect is attributed to the residual stress gradient and could ap-proximately be correlated to the change in the natureof residual stresses (Fig. 5). From the crack appearance,it is tempting to speculate that during crack propaga-tion a redistribution of residual stresses has occurred such





Fig. 8 Fatigue crack appearance in the

terminal weld positions, (a) notch tip at the

beginning of the weld, X = 0 mm, (b)

notch tip at the end of the weld, X = 5 mm

and (c) weld at the end of the specimen, X

= 75 mm. (Specimens slightly tilted toimprove crack visibility. The near-notch tip

weld positions, being polished out, are

indicated by black dotted lines and the local

crack deviation from the mid-plane is

indicated in each case).

that compressive residual stresses have piled up along themid-plane and have deviated the crack, which propagatesfurther under the applied load and redistributed tensileresidual stresses. This aspect needs to be investigated fur-ther, because similar observations have not been reportedexplicitly in the literature. We are planning to look into

redistribution of residual stresses by using progressivesawcuts8 (cut-compliance technique) and modelling us-ing finite element analysis. As regards safety aspects, cracks propagating from the

terminal weld positions shall need monitoring in the pres-ence of tensile residual stresses. Although apparently smallin quantity, tensile residual stresses appear to acceleratecrack propagation through the superposition (Fig. 1).

On the other hand, the near-notch tip and the mid-ligament weld positions may not be that critical whencompressive residual stresses exist near the notch tip(Fig. 5d & e; K rs → − ve). In such cases, the crack isfound to slow down as the weld is approached.24–31,34

This effect occurs in Al-alloys, either fusion welded24,34

or friction stir welded26–30 and in fusionwelded steels.25,31

Hence, it is independent of material welded and weldingprocess, but dependent on the nature of weld residualstresses. Crack arrest is, however, observed seldom andis reported for fusion welded steel25 and fusion welded

AA2024-T351 by variable polarity plasma arc (VPPA)technique.34 Whereas friction stir welded AA5083-H32exhibited crack arrest, AA6061-T651 did not.27 Also, in

locally treated materials crackretardationis reported,14–19

but not crack arrest. Thus, in addition to the extent of compressive residual stresses, the K level applied andthe material investigated seem also to be decisive.

Our purpose to investigate the near-notch tip and themid-ligament weld positions (Figs 4d & e and Figs 5d &

e) was to let the crack propagate against the compressiveresidual stresses. If crack propagation is reduced or ob-structed, a residual stress level may then be found whichmay be non-damaging. The major difference betweenthe two weld positions was the extent of the near-notchtip compressive residual stresses induced, which differednearly by a factor of 2 (Fig. 5d & e). The results are shownin Figs 9 and 10. Indeed, fatigue crack propagation wasstopped or delayed in both cases. However, because the

ASTM criterion45 for pre-cracked length or the maxi-mum out of plane crack deviation or both could not befulfilled, the data became invalid as reported below.

Even though a fatigue crack could be generated at thenotch tip, it proved to be very difficult to make it grow in a self-similar fashion and its further propagation dif-fered significantly depending on the extent of the near-notch tip compressive residual stresses. When the near-notch tip residual stress was high (σ y ≈ −100 MPa;Fig. 5d), pre-cracking became a problem. The crack couldbe generated, however, was arrested after a short exten-sion (0.2 mm ≤ a ≤ 1 mm), and did not propagate onextended cycling, even after more than 5 million cycles




408 W . V . V A I D Y A et al .

Fig. 9 Fatigue crack propagation perpendicular to the weld in thenon-terminal positions. The weld position is shown in the inset

along with the symbols. Hatching pattern marks the base material

data for C(T) specimens in the same orientation.

(data shown by circle in Fig. 9). Therefore, the test wasstopped. In the next test, load levels were increased fol-lowing non-propagation ( N ≥ 0.2 million cycles; datashown by squares in Fig. 9). But except for a small incre-ment (0.1 mm ≤ a ≤ 1 mm) following the load increase,non-propagation was observed often. Finally, the crack deviated out of the mid-plane completely (5◦ ≤ θ ≤ 30◦)

and unexpected cracking occurred on extended cycling at blunt stress raisers such as notch corner or loading boresor both (Fig. 10a). Hence, this test became invalid. Tofacilitate pre-cracking, a test was started at a higher K level (14.7 MPam0.5; data shown by upright triangles inFig. 9). After some initial extension (a ≤ 5 mm) out of plane deviation and cracking at blunt stress raisers wasobserved, making the test invalid again. When the near-notch tip residual stress level was lower

(σ y ≈ −50 MPa; Fig. 5e), pre-cracking was no more aproblem. Interestingly, instead of a steady increase inpropagation rate, thecrack propagated at a nearly constant rate initially and this effect was reproducible (data shown

by inverted triangles and diamonds in Fig. 9). However,during further propagation (5 mm ≤ a ≤ 10 mm) crack branching was very often observed, propagation was de-layed and was followed by excessive out of plane deviation(Fig. 10b). Thus, the tests became invalid for these casesas well. Although valid data could not be obtained for the non-

terminal positions, these tests show that a near-notch tipcompressive residual stress level of more than −50 MPashould be non-damaging. The crack cannot be made to

Fig. 10 Cracks leading to data invalidity of fatigue crack

propagation perpendicular to the weld in the non-terminal

positions, (a) weld nearer to the notch tip, X = 15 mm and (b)

weld in the mid-ligament, X = 40 mm. (Specimens slightly tilted

to improve crack visibility).

propagate along the mid-plane, instead a deviation oc-curs and crack propagates on a new plane. This suggeststhat the initial compressive stresses (which resist propa-gation), when released (due to crack propagation) are re-distributed and equilibrating tensile residual stresses aregenerated. Fatigue crack extension is found to decreasethe residual stress level.11 It follows that if then the lo-cal loading conditions ( Rlocal and K rs) progressively shift from the compressive towards the tensile region (Fig. 1),





propagation would be facilitated on extended cycling. Thus, due to the superposition the whole loading pat-tern would have been depressed (partly or fully) below

K op in the mid-plane (Fig. 1). Hence, the driving forcefor crack propagation decreases and can lead to crack ar-rest. We have observed non-propagation similar to that in

Fig. 9 (shown by square symbols) in a laser beam weldeddissimilar joint.50 Such crack arrest seems to be specificto the (narrow) size and the steep gradient induced by the laser beam weld. Similar observations are not knownon broad welds, except for that by Ignat’eva et al .25 ona fusion welded steel. They also observed global crack-ing following crack arrest, however, with the near-notchtip compressive residual stress of about −300 MPa inC(T) specimens. Thus, although due to crack arrest the

weld becomes self-protecting against fatigue crack prop-agation, extended cycling may cause unexpected damageat other stress raisers (Fig. 10). Therefore, also the levelof compressive residual stresses deserves consideration.

Broad fusion welds of AA2024-T351 by VPPA techniqueare reported to exhibit crack arrest, which could, however,be overcome when the K level was slightly increased.34

Thus, the steep residual stress gradient causing multiplecrack arrests and non-propagation seems to be typical forthe present narrow fusion welds produced by laser beamtechnique.

Concluding remarks

This work shows that whether a crack can be stopped isdependent on the K level, the relative distance between

the crack and the weld, the residual stress gradient and thelevel and the nature of residual stresses. We had already pointed outthat as regards safety aspects a crack propagat-ing from the weld should need monitoring in the presenceof tensile residual stresses. Because tensile residual stressesare present within, along and adjacent to the weld (Fig. 5),the case that a crack is initiated within the weld (Fig. 6) orfrom the weld (Fig. 7) and propagates along the mid-weld(Fig. 6) or into the residual stress field containing basematerial (Fig. 7) is rather critical and unfavourable. Onthe other hand, the weld can be useful when compressiveresidual stresses (of sufficiently high level) are present at a site where a crack is likely to emerge outside the weld.

Because such crack cannot grow ( K rs → − ve), the weldmay then prove to be a crack stopper (Fig. 9).

Further research in this direction may open new per-spectives for crack stoppers through local treatments suchas laser shock peening. Because specimen geometry canbe a deciding factor for fatigue crack propagation due tothe residual stress gradient 13 (see Fig. 6 also), laboratory data should preferably be obtained on M(T) specimens.

A crack propagating perpendicular to stiffeners in M(T)panel is reported to have no effect for extruded integral

stringers in AA7075-T6,51 but to slow down for weldedstringers in a steel.52 Thus, weld residual stresses may infact be useful, if they can be manipulated in their naturethrough design. We are looking into such aspects, includ-ing modelling on the lines laid down in the literature (e.g.Ref. [34]).

C O N C L U S I O N S

The effect of weld residual stresses on fatigue crack prop-agation was investigated using different crack orientationsin a laser beam welded butt joint of the Al-alloy AA6056.Residual stress gradients were tailored by changing thedistance of the notch tip relative to the weld in C(T)specimens. The conclusions are as follows:

1 Whereas the nature of residual stresses adjacent to the

weld was always tensile, the nature of the near-notch tip

residual stresses could be varied systematically, rangingfrom tensile to compressive, and the stress level per-

pendicular to the direction of fatigue crack propagation

differed significantly initially (−100 MPa ≤ σ y ≤ 150

MPa).

2 The crack orientation ‘along the mid-weld’ exhibited

faster crack propagation than that in the base material.

Because the initial near-notch tip residual stress level was

nearly zero in C(T) specimens, the difference in fatigue

crack propagation is attributed to the microstructural

difference, cast against wrought. The delay caused by the

interaction with macro-pores was a phenomenon specific

to this (6-mm thick sheet) weld.3 Near-notch tip tensile residual stresses contributed to

faster crack propagation for the crack orientation ‘per-

pendicular to the weld’, irrespective of the weld position

relative to the notch tip. This trend was not affected by

compressive stresses initially present along the residual

stress gradient away from the notch tip. Insofar tensile

residual stresses, although small in level, can acceler-

ate crack propagation and, hence, deserve precaution.

Change in the nature of residual stresses along the gra-

dient is presumed to cause crack deviation.

4 When compressive residual stresses were induced near

the notch tip, the fatigue crack, propagating perpendic-

ular to the weld, was either arrested or when propaga-tionoccurred, exhibited multiple crack branching, partial

crack arrest and out of plane deviation. In such a case,

valid data along the mid-plane could not be obtained.

Even cracking was observed on extended cycling at blunt

stress raisers such as bores and notch corners depending

on the level of the initial compressive residual stresses.

Thus, compressive residual stresses may turn-up to be

harmful, even when fatigue crack propagation does not

occur.




410 W . V . V A I D Y A et al .

5 M(T) specimens provided conservative data for welds

as well as the base material. Hence, when conservative

laboratory data are required, M(T) specimen geometry

should be preferred to C(T).

Acknowledgements

The material investigated originates from an AerospaceProject (LuFoII: Luftfahrt-Forschung II) initiated by

AIRBUS. We thank Dr. J. Hackius, AIRBUS Bremen,Germany, for supplying us the material and gratefully acknowledge his interest. This work was supported by in-ternal funding at HZG (previously GKSS). Interest of Dr.

M. Kocak, now at GEDIK Holding, Istanbul, Turkey, isalso acknowledged.

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