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Experimental and numerical investigation on the fatigue behaviour of friction stirred channel plates
Engineering Failure Analysis 103 (2019) 57–69
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Experimental and numerical investigation on the fatigue behaviour of friction stirred channel plates
T
⁎
Catarina Vidala,b, , Ricardo Baptistac,d, Virgínia Infanted a
Instituto Superior Técnico, Universidade de Lisboa, Av. Rovisco Pais, 1049-001 Lisboa, Portugal UNIDEMI; Departamento de Engenharia Mecânica e Industrial, Faculdade de Ciências e Tecnologia, Universidade NOVA de Lisboa, Caparica 2829516, Portugal c CDP2T, Department of Mechanical Engineering, ESTSetúbal, Instituto Politécnico de Setúbal, Campus do IPS, Estefanilha, 2914-761 Setúbal, Portugal d LAETA, IDMEC, Instituto Superior Técnico, Universidade de Lisboa, Av. Rovisco Pais, 1049-001 Lisboa, Portugal b
A R T IC LE I N F O
ABS TRA CT
Keywords: Friction stir channelling Channel geometry Bending strength Fatigue life Fatigue crack growth
The friction stir channelling (FSC) process is a technological innovation based on the friction stir fundamentals. To produce friction stirred channel components, a non-consumable tool, similar to that used for friction stir welding (FSW), is used. Friction stir channelling is a single step manufacturing process in which a continuous channel with a specific path is produced in a monolithic metal component. This work discusses the influence of the channel geometry on the bending strength of friction stirred channel aluminium alloy specimens, as well as their fatigue behaviour. Moreover, fatigue analyses using the finite element method were also carried out, considering the channels' geometrical features in order to assess and compare the fatigue lives. The stress intensity factor for different crack lengths was determined using ABAQUS. The fatigue crack growth curve was established according to the Paris Law, and the crack was only allowed to propagate along the materials interface. It was observed that the critical zones are located in the vicinity of the channel corners and found that the fatigue crack propagation period on FSC specimens is very short compared to the crack initiation period. Base and FSW material properties play a major role on fatigue crack propagation simulation. The accuracy of the predicted results increased when considering lower stress amplitudes and by using the FSW material parameters.
1. Introduction Developments in the friction stir channelling (FSC) technology have made possible the material removal by the tool as the channel is being produced, leaving the metal component with the same level and geometry it had before being processed [1,2]. As for all friction stir based manufacturing technologies, a visco-plasticised solid-state region is generated and processed into a new shape and properties. This region whilst remaining solid presents a three-dimensional material flow pattern almost as a liquid, enabling easy mixing and blending between different materials. This phenomenon is generally referred to as the third-body region concept [3]. The material flow during the FSC process leads to the formation of a channel surrounded by four different material regions: the advancing side (AS), the retreating side (RS), the top (nugget) and the bottom of the channel. In the nugget zone, the grains are dynamically recrystalized. Fig. 1 shows a typical friction stirred channel cross-section macrograph. ⁎
Corresponding author at: Instituto Superior Técnico, Universidade de Lisboa, Av. Rovisco Pais, 1049-001 Lisboa, Portugal. E-mail address: [email protected] (C. Vidal).
https://doi.org/10.1016/j.engfailanal.2019.04.068 Received 23 January 2019; Received in revised form 15 April 2019; Accepted 26 April 2019 Available online 29 April 2019 1350-6307/ © 2019 Elsevier Ltd. All rights reserved.
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Fig. 1. Cross-section macrograph of a friction stirred channel produced in an aluminium alloy component.
The interface between the nugget and the thermo-mechanically affected zone (TMAZ) is the main structural concern of the friction stirred channel components, particularly from the fatigue point of view, as the corner defined by the AS and the nugget is a region of stress concentration between very fine and coarse grains, amplified by its geometry. Generally, the fatigue life can be divided into two major steps [4]: 1) Crack initiation, due to local accumulation of dislocations, stress hot points, material defects or plastic deformation due to inhomogeneous inclusions; 2) Crack propagation, due to crack opening and closing and introducing permanent damage to the material. Crack initiation mode and local depends on material microstructure, applied loads, and specimen geometry [5]. The duration of the fatigue crack initiation period can be influenced by different factors and can occur either on the specimen surface or under the specimen surface. It can be observed near local stress concentration points, where the local resistant area is smaller. Therefore, the total number of fatigue cycles, Nf, required to achieve the final fracture can be obtained from the number of cycles required for crack initiation, Ni, and the number of cycles required for crack propagation to occur, Np, from the initial to the critical crack length (Nf = Ni + Np). When considering the interface between two different materials, one must consider that this zone shows lower fatigue resistance than both, the base and the added materials. The interface between materials can be used to represent different situations: 1) a crack between a base material and a coating [6], 2) a crack between two dissimilar materials joined by friction stir welding [7] or 3) a crack between two non-metallic materials [8]. Regardless the case, the crack propagation rate on the interface will always be higher than on the base material. Accordingly, to Shah et al. [8], this can be attributed to the mismatch in the Young's modulus between the different materials. Different fracture parameters, as the energy release-rate, must be used when considering fatigue crack propagation along materials' interfaces, such as the interface between different elastic layers [9] or friction stir welds [7]. The different behaviour can also be attributed to the fact that a crack between two different material interfaces, inherently propagates in mixed mode, and not in pure mode I or pure mode II [8]. Due to the mismatch between the Young's modulus, even a pure mode I problem will not be symmetric, and the crack tip will be subjected to normal and shear stresses. Therefore, crack propagation will always depend on the crack length and the relationships between the two materials properties. The stress field near the crack tip will then show an oscillatory behaviour, eq. (1), characterized by an oscillatory index ϵ introduced by Dundurs [10], eq. (2). Unlike the linear homogeneous conditions, where the stress filed shows an inverse square root singularity. The nature of the stress field and stress singularity leads to oscillatory crack propagation, with the crack inter-penetrating both materials. Therefore, crack opening mode I and mode II cannot be decoupled, rendering analytical solutions inadmissible [8]. 1
σ ~r − 2 + ϵ ϵ=
(1)
1 − β⎞ 1 ln ⎛⎜ ⎟ 2π ⎝ 1 + β ⎠
(2)
Parameters β and α are the material interface parameters defined by Dundurs [10]. Parameter β is the elastic mismatch parameter, a function of the Poisson's ratio, υ, and the material shear modulus, μ, (eq. (3)). Parameter α measures the relative materials stiffness (eq. (4)), where E denotes the equivalent modulus of elasticity for plane stress or plane strain and is a function of the materials' mismatch. 1 and 2 indices denote material 1 and material 2, respectively. Both parameters will be equal to zero in the case of an interface between equal properties materials.
β=
μ1 (1 − 2υ2) − μ 2 (1 − 2υ1 ) 2(μ1 (1 − υ2) + μ 2 (1 − υ1 ))
(3)
α=
E1 − E2 E1 + E2
(4)
Rice [11] defined the behaviour of the stress field along the crack interface. Eq. (5) establishes the oscillatory behaviour of the stress singularity and, as mentioned, the stress intensity factors for mode I and mode II, cannot be decoupled.
σyy + iσyy =
K1 + iK2 iϵ r 2πr
(5)
It is then required to determine the solution for the coupled stress intensity factors, in order to analyse the crack propagation along the interface between different materials. One solution for this problem is to use the energy release rate, G, or J-integral to study the interface crack propagation, as reported by Kakiuchi et al. [7] and Xuan et al. [12]. The finite element method (FEM) 58
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provides a simple solution to numerically estimate the energy release rate or J-integral. The J-integral can then be related to the stress intensity factor using eq. (6), even for a mixed material problem.
J=
1 − β2 2 1 2 (K1 + K22) + K3 E∗ 2G∗ 1 E∗
1 (E1 2
1 G∗
(6)
1 (G1 2
= + E2) and = + G2.) where and G denotes the materials' shear modulus. FEM is therefore a suitable method to solve this type of problem. Finally, the Paris Law can be used to calculate the fatigue crack propagation life, using eq. (7). da = B ∆G n dN
(7)
Where B and n are material parameters. In this paper FEM analysis was used to estimate the energy release rate for different crack lengths along the interface between the nugget and the TMAZ. Using eq. (6), the number of elapsed cycles was calculated for each crack increment and added to the previous obtained value, to calculate the total fatigue life. 2. Material and experimental details 2.1. Material The strain hardened aluminium alloy AA5083-H111 was used in this work. Its nominal chemical composition is Al–5.26 Mg–1.02Mn–0.19Fe–0.15Cr (wt%) [13]. The tensile properties, at room temperature are: σy = 210 MPa, E = 68.9 GPa, σUTS = 375 MPa and A = 18.3% [13], where, σy is the yield stress for 0.2% offset parallel line to the elastic line, with Young's modulus, E; σUTS is the ultimate tensile stress and A is the elongation. 2.2. Friction stirred channel production An ESAB LEGIO™ 3UL numeric control equipment was used to produced friction stirred channels on 15 mm-thick rolled plates along the rolling direction. Plunge and dwell periods (zero tool travel speed) were performed under vertical position control and processing period (non-zero tool travel speed) was carried out under vertical force control. In order to obtain channels with different dimensions, four sets of FSC parameters were implemented as shown in Table 1. Channels U# were produced with a probe length of 6 mm and channels L# with a probe length of 8 mm. The channels were produced using a modular H13 steel tool comprised by a body, a shoulder and a probe. All tool's arrangement had a threaded cylindrical probe with 8 mm diameter. Table 2 shows the three different shoulders used. The tool tilt angle was 0° for all the runs. 2.3. Four-point bend testing Four-point bending tests were performed to obtain the mechanical strength of the channels produced. This test was selected because the purpose is to assess the mechanical strength of a specific and large area, the processed zone, rather than a specific and small zone as, for example, a weld root [14]. Specimens were manufactured according to the standard ASTM E290-97a requirements as shown in Fig. 2. From each FSC condition (Table 1), two bending specimens were taken and tested with the processed surface under tensile stress, as depicted in Fig. 3. An electromechanical Zwick Roell testing machine with a load cell of 200 kN was used to carry out the four-point bending tests. The rollers have a diameter of 15 mm. The span distances, the values of which are given in Fig. 3, were set according to the ASTM E 290-97a standard. Four-point bend testing was performed at a loading speed of 1 mm/min in laboratory air at room temperature. 2.4. Fatigue testing The fatigue specimens were manufactured with the same dimensions of the four-point bending specimens (Fig. 2). The fatigue tests were carried out at room temperature in a servo-hydraulic test machine (Instron® 8502), with a load capacity of 100 kN. The Table 1 FSC parameters. FSC set
Tool rotation speed (rpm)
Tool travel speed (mm/min.)
Vertical force (kN)
U1 U2 L1 L2
1100
50 80 50 100
4 2 1.5 2.2
400
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Table 2 Tool shoulders used to produce the channels. FSC set
Tool shoulder
U1, L2
U2
L1
Fig. 2. Geometry and dimensions of the four-point bending specimens [15].
15 P/2
30 90
P/2
30
Fig. 3. Schematic representation of a FSC specimen under four-point bending.
fatigue cycle was constant amplitude loading, with a sinusoidal load wave of 5–10 Hz, with stress ratio of R = 0.1 and a four-point bending set up. The fatigue tests were performed with the maximum stress levels ranging from 30 to 150 MPa and it was defined the run-out limit at 3 × 106 cycles.
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Fig. 4. Channel nugget configurations. Crack propagation location along interfaces.
The applied load, P, was calculated by eq. (8) where w is the smallest span distance (30 mm), h is the specimen's thickness, b is the specimen's width and σ is the maximum stress.
P=
σbh2 3w
(8)
The load was applied out of the processed zone and the specimens were tested with the processed surface under tensile stress (Fig. 3). 3. Numerical simulation 3.1. Finite element model Four different channel geometries (U1, U2, L1 and L2) were modelled (Fig. 4) in ABAQUS. The goal was to model crack propagation along the materials' interface between the nugget and the TMAZ on the advancing side (AS) of the channel. Using the spline function to reconstruct the different channel and the nugget geometries, four two-dimensional solid models, with a constant 30 mm depth, were built. Each specimen with a 180 × 30 mm section was subjected to four-point bending loads and boundary conditions, considering a 90 mm span. 3.2. Fatigue failure simulation Crack propagation simulations were performed with different loading amplitudes, in order to simulate the experimental conditions. Initially four rollers were fully modelled, including contact between the rollers and specimen, in order to simulate four-point bending. The simulation of the rollers-specimen contact resulted in a complex and numerically demanding simulation. It was then verified that by applying the loads and boundary conditions to four nodes directly, the results were not affected, and the resulting linear simulation could be solved faster. Base and channel nugget materials were considered as linear elastic, with properties described in section 2.1 and Table 3, respectively. The effect of FSC residual stresses was not modelled in our work. It is believed that FSW and FSC produce generally low residual stresses, due to low temperature solid-state process [16]. Compressive residual stresses lead to lower crack propagation rates, but this effect is neglected when using higher R [16] values. Tension residual stresses accelerate crack propagation. By not simulating this effect, our results will be less conservative. Each of the four models used quadratic plane strain elements. In order to define the stress singularity around the crack tip, a special meshing technique was used (Fig. 5 c)). Using wedge elements with collapsed nodes, five contour integrals were used to estimate the energy release rate (J-integral). The different models used up to 50,000 nodes. Considering a constant crack increment (Δa) and eq. (9), the fatigue crack propagation life for each specimen and loading amplitude was calculated.
N=
∆a B ∆G n
(9) Table 3 Linear elastic properties of the processed material used in FEM analyses [18].
Nugget (L1 and L2) Nugget (U1 and U2)
Young's modulus (GPa)
Poisson rate
69.8 67.3
0.3 0.3
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Fig. 5. Mesh and crack details. a) L2 channel and nugget, b) crack and mesh along the crack, c) spider-web mesh for crack tip contour integral calculation.
From the work of Ilman et al. [17], the parameters for the Paris law were estimated. These authors considered the crack propagation on similar base material and nugget material resulting from FSW (Table 4), although not along the materials interface. 4. Results 4.1. Four-point bend testing Four-point bending tests were performed to evaluate the mechanical performance of the friction stir channels. A typical loaddisplacement curve was plotted for each test, as presented in Fig. 6. The displacement was obtained from the machine cross-head movement. To assess the channels' strength, their ultimate bending strength was additionally calculated by using eq. (8). The maximum load, P, was extracted from the curves. A summary of the mechanical performance of the channels is presented in Table 5. To better understand the influence of the channel shape and dimensions on the bending strength, as well as to establish feasible comparisons among the channels analysed, their geometric parameters were measured. The results are schematic illustrated in Fig. 7. 4.2. Fatigue testing Table 6 presents the fatigue test results for each one of the four FSC conditions analysed. Table 4 Fatigue crack propagation parameters for the base material and FSW material, adapted from [17]. B (m/cycle) Base material FSW
−1
1.365 × 10 1.116 × 10−4
62
n 2.222 1.078
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Fig. 6. Load-displacement curves.
Table 5 Four-point bend testing results. Condition ID
Specimen ID
Maximum load (kN)
Average value (kN)
Deviation (kN)
Ultimate bending strength (MPa)
U1
1 2 1 2 1 2 1 2
30.0 32.0 22.1 21.2 20.0 20.6 10.4 9.9
31.00
± 1.00 (3.2%) ± 0.45 (2%) ± 0.30 (1.5%) ± 0.25 (2.5%)
400 427 295 283 267 275 139 132
U2 L1 L2
21.65 20.30 10.15
The obtained fatigue strength values for conditions U1, U2 and L1 were significantly higher than those for condition L2. Therefore, these results indicate that, for the L2 condition, a reduced fatigue strength is obtained due to the channel geometry. It was also found that when the maximum applied stress increases the number of cycles to failure decreases, as expected, with a maximum stress applied of 50 MPa for a run-out of 3 × 106 cycles. 4.3. Numerical simulations On each of the four types of fatigue specimen numerical models, a crack on the materials' interface was allowed to propagate. An initial crack size of 0.1 mm was considered for all specimen models. These cracks propagated from the closest corner, of the modelled channel, to the free surface of the specimen. The crack propagation simulation ended when loss of tightness of the channel occurred, i.e., the crack reaches the specimens free surface. Fig. 8 shows the energy release rate as a function of crack length. The number of elapsed fatigue cycles between increments, was calculated using eq. (9). Considering the geometry along the materials interface of specimen U1, the maximum crack length is 1.7 mm, while on specimen U2 allowed for a maximum crack length of 1.95 mm. On these two specimens when a load that produces a normal stress of 50 MPa is applied, the maximum energy release rates were 5 J/mm2 and 11 J/mm2, respectively. For specimen L2 the maximum crack length is 1.2 mm, as the distance between the channel and the free surface is much smaller. In fact, Fig. 5 shows that the crack propagates in a constant direction toward the free surface, and therefore the energy release rate increases exponentially with the crack length. The crack propagation path is vertical, unlike on specimens U1 and U2 (Fig. 4), and the crack opening mode is almost pure mode I, therefore the maximum energy release rate is higher, reaching 21 J/mm2. Specimen L1 allowed a 3.3 mm maximum crack length. As one can see in Fig. 4, the materials interface changes direction, and therefore the crack propagation path is longer. The crack opening mode changes from pure mode I to a mixed mode and back to almost pure mode I again. When changing from mode I to mixed mode the energy release rate slightly decreases, slowing down the crack propagation rate. As the crack length reaches the critical value, and the opening mode returns to mode I, the crack propagation rate increases as the energy release rate increases to 21 J/mm2. Fig. 9 shows the corresponding crack propagation curves, for a fatigue load with a maximum normal stress of 50 MPa. The fatigue life of the specimen L2 resulted in 23,069 cycles. This is the shortest fatigue life, as the maximum crack length is low, and the energy release rate is high. The fatigue life for specimens U1 and U2, was 117,260 and 138,150 cycles, respectively. The behaviour of these cracks is very similar, as well as the evolution of the energy release rate, that governs the crack propagation rate. Finally, the fatigue 63
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details (a) to (d)
a)
3
1.5
6.5
b)
3.7
1.6
8
c)
2.7
4.6
8
d)
1.3
5.8
2.8
8
Fig. 7. Geometry and dimensions of the specimens. Details showing the schematic representation of the channels produced with FSC set a) U1, b) U2, c) L1 and d) L2 [18].
life of specimen L1 was 274,780 cycles. As previously mentioned, this is the longest crack and the one that shows the lowest energy release rates, as the crack changes directions along the interface. One can see in Fig. 9, that the crack propagation rate decreases around the 150,000 cycles mark, and increases again, around the 250,000 cycles mark. Several fatigue crack propagation simulation models were created, considering different fatigue loads. The models included different crack propagation properties for the base and FSW materials. The obtained results as summarized in Table 7. Considering the base material crack propagation properties, the obtained fatigue life is always higher than the one obtained with the processed material properties. Therefore, the base material has a higher fatigue crack propagation resistance, as concluded by Ilman et al. [17].
5. Discussion From Fig. 6 and Table 5, it is possible to conclude that the channels produced with FSC conditions U1 are the most resistant to bending loads; channels produced with FSC set of parameters U2 and L1 have similar performances and channels performed at conditions L2 exhibit almost elastic behaviour before final failure. Channels produced with FSC conditions U1 also have the largest elongation. The height of the closing layer thickness of channels produced with parameters U1 and U2 is almost equal (Fig. 7), however, channels U1 have a smaller area and were performed at a higher vertical force, which seems to indicate that improves the channels' mechanical strength and ductility. Regarding the influence of channel depth on the mechanical behaviour of the channels, no feasible conclusion can be drawn from the results of conditions U2 and L1. One can observe that a thicker closing layer did not lead to a better mechanical performance which can be explained by the presence of small internal voids on the channel nugget of condition L1 (Fig. 10). From the global analysis of fatigue test results, it is possible to conclude that FSC specimens exhibit a very low fatigue strength comparing to the unprocessed base material, even though they have no defects. As can be extrapolated from Table 6 the scattering of the results is quite low which demonstrates that, on one hand it is possible to predict the fatigue behaviour of FSC components and, on the other hand, the production process is stable, and the channels are reproducible. Regarding the fatigue strength, the results indicate that specimens produced with condition L1 revealed a slightly better fatigue performance than those performed with parameters U1 and U2, mainly at lower stress values, which seems to suggest that, for the 64
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Table 6 Fatigue test results [18]. Fatigue specimen ID
σmax (MPa)
Freq. (Hz)
Nf, Cycles to failure
U1_1 U1_2 U1_3 U1_4 U1_5 U1_6 U1_7 U2_1 U2_2 U2_3 U2_4 U2_5 U2_6 U2_7 U2_8 U2_9 L1_1 L1_2 L1_3 L1_4 L1_5 L1_6 L1_7 L1_8 L1_9 L2_1 L2_2 L2_3 L2_4 L2_5 L2_6
125 125 100 100 65 65 50 100 100 80 80 65 65 60 60 50 150 125 100 100 80 80 65 65 50 65 65 50 50 40 40
10 10 10 10 10 10 10 5 5 5 5 7.5 7.5 7.5 7.5 7.5 3 4 4 4 5 5 8 8 8 7.5 7.5 7.5 7.5 10 10
13,702 15,794 49,038 54,881 385,977 291,320 3,000,000 58,595 51,856 142,555 149,081 330,259 329,134 353,401 374,398 3,000,000 15,259 17,094 46,165 67,293 165,572 120,796 773,309 808,165 3,000,000 14,953 19,890 54,954 45,628 88,099 196,253
Obs.
Run-out
Run-out
Run-out
Fig. 8. Energy release rate vs crack length, for a fatigue load with a maximum normal stress of 50 MPa.
conditions analysed, a thicker closing layer tends to improve the fatigue strength of FSC specimens, regardless the channel area. In fact, a thicker closing layer contributes to longer fatigue crack propagation periods, although it has been observed that the crack propagation period for FSC specimens is very small when compared with the number of cycles required for crack initiation. The similarity among the SeN curves of defect free FSC specimens can be explained by the similar crack initiation periods observed for conditions U1, U2 and L1. The worst fatigue performance presented by the specimens produced with FSC conditions L2 is related with the defect observed in the nugget/TMAZ interface. The defect works as a notch, leading to higher local stresses, which contributes to failure at a lower number of cycles. When considering the base material and FSW material models offered by Ilman et al. [17], one can see (Figs. 11 to 14) that the fatigue lives predicted by the crack propagation simulations presented herein, using the base material properties, are always higher than the experimental ones. This shows that the current model is not suitable for the fatigue crack propagation at the interface
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Fig. 9. Fatigue crack propagation curves, for a fatigue load with a maximum normal stress of 50 MPa. Table 7 Numerical fatigue simulation results. Fatigue specimen
σmax [MPa]
FSW - N [Cycles]
BM - N [Cycles]
L1
150 125 100 80 65 50 65 50 40 125 100 65 50 100 80 65 50
25,728 38,115 61,662 99,756 156,080 274,780 13,105 23,069 37,320 16,265 26,314 66,605 117,260 31,003 50,156 78,474 138,150
40,676 91,454 246,520 664,510 1,672,000 5,365,000 30,544 98,008 264,190 38,287 103,210 699,970 2,246,000 165,820 446,970 1,124,600 3,608,700
L2
U1
U2
Fig. 10. Optical micrograph under polarized light from an electrolytic etched L1 sample showing small internal voids on the channel nugget.
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Fig. 11. Fatigue life simulations for the L1 specimens.
Fig. 12. Fatigue life simulations for the L2 specimens.
Fig. 13. Fatigue life simulations for the U1 specimens.
between base material and FSW material. As mentioned, this interface is considered to have lower fatigue crack propagation resistance. When considering the FSW material model for the L1 specimen and a load amplitude of 50 MPa, as seen in Fig. 11, the predicted fatigue life represents 9% of the total fatigue life. Fig. 13 shows a similar behavior for the U1 specimen. For a fatigue load of 50 MPa, the simulated fatigue life represents 4% of the total life. The remaining 96% would represent the initiation fatigue life that is not considered by our model. But if one considers a fatigue load of 125 MPa, the simulated crack propagation life represents 110% of the experimental one. The model is not suitable for fatigue predictions when considering higher load amplitudes. Figs. 12 and 14 67
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Fig. 14. Fatigue life simulations for the U2 specimens.
show similar behaviors for the L2 and U2 specimens, respectively. As explained before, the fatigue crack propagation was obtained using the Paris' law where the predicted life was obtained under Linear Elastic Fracture Mechanical (LEFM). Predictions of fatigue life in crack propagation are compared with experimental data of fatigue life and the determination of fatigue crack initiation (Ni) is, in this study, defined by the difference between the experimental numbers of cycles (Nf) and the propagation number of cycles obtained in the computational simulations (Np). Table 8 presents the experimental and numeric fatigue failure results and the prediction of the fatigue crack initiation cycles. As expected, the number of cycles Ni required for crack initiation increase with the decreasing of maximum stress applied. It was found that the total number of cycles (obtained from experimental tests) is similar to the propagation number of cycles (obtained in numerical simulations) for the highest load amplitudes. Increasing the load amplitude means that the initiation phase is practically non-existent. On the other hand, for the smallest applied stresses (σmax = 50 MPa for the specimens U1, U2 and L1) where the experimental tests showed infinite lives it is possible to conclude that the stress intensity factor range is less than the threshold value for the propagation of the crack suggesting that for this condition the defects would never initiate. 6. Conclusions The channels produced with FSC conditions U1 have shown to be the most resistant to bending loads and present the largest elongation; channels produced with FSC set of parameters U2 and L1 have similar performances and channels performed at conditions L2 exhibit almost elastic behaviour before final failure. Regarding the influence of channel depth on the mechanical behaviour of the channels, no feasible conclusion can be drawn from the results of conditions U2 and L1. Even considering that FSC specimens showed no defects, the analysis of their fatigue strength has revealed to be lower than the unprocessed base material. A thicker closing layer, increasing the distance between the channel and the specimen free surface, contributes to longer fatigue crack propagation periods. It was also verified that on the FSC specimens this period is very short when compared with the fatigue crack initiation period. Table 8 Comparison of the experimental and numerical results. Specimen
σmax (MPa)
Experimental results, Nf (number of cycles)
Computational results, Np (number of cycles)
Ni = Nf - Np (number of cycles)
U1
125 100 65 50 100 80 65 50 150 125 100 80 65 50 65 50 40
14,748 51,960 338,649 Run-out 55,226 145,818 329,697 Run-out 15,259 17,094 56,729 143,184 790,737 Run-out 17,422 50,291 142,176
16,265 26,314 66,605 117,260 31,003 50,156 78,474 138,150 25,728 38,115 61,662 99,756 156,080 274,780 13,105 23,069 37,320
– 25,645 272,043 – 24,222 95,662 251,222 – – – – 43,428 634,657 – 4316 27,222 104,856
U2
L1
L2
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Using the base material crack propagation properties resulted in a longer fatigue life, when compared to the experimental results or the numerical ones obtained with the FSW parameters. Therefore, the base material model used is not compatible with fatigue crack propagation simulations on the interface between materials. When considering the FSW material parameters and a load amplitude of 80 MPa, the estimated fatigue crack propagation life represents a maximum of 70% of the total fatigue life. Lower load amplitudes predicted lower crack propagation life percentages. Therefore, the simulations are more accurate for lower load amplitudes. Considering crack initiation was not modelled, it is recommended to use the FSW material parameters, when the interface data is not available. Future work should be applied to the numerical model, and to experimentally quantify the relation between the crack initiation and propagation periods, allowing to choose a more suitable material model for fatigue crack propagation analyses. Including the effect of tension residual stresses will accelerate the crack propagation rate, decreasing the final fatigue propagation life, also making the calculated values more in-line with the experimental results. Acknowledgements The authors would like to acknowledge the Portuguese Foundation for the Science and Technology (FCT) for its financial support through the PhD scholarship FCT SFRH/BD/62963/2009, IDMEC, under LAETA project UID/EMS/50022/2019 and project UID/ EMS/00667/2019. References [1] C. Vidal, P. Vilaça, Processo de Abertura de Canais Internos Contínuos em Componentes Maciços Sem Alteração da Cota de Superfície Processada e Respectiva Ferramenta Modular Ajustável, Patent number: PT 105628 (B), April 15 (2011). [2] C. Vidal, V. Infante, P. Vilaça, Fatigue assessment of friction stir channels, Int. J. Fatigue 62 (2014) 77–84, https://doi.org/10.1016/j.ijfatigue.2013.10.009 May. [3] W.M. Thomas, An Investigation and Study into Friction Stir Welding of Ferrous-Based Material, PhD Thesis University of Bolton, 2009. [4] G. Fajdiga, M. Sraml, Fatigue crack initiation and propagation under cyclic contact loading, Eng. Fract. Mech. 76 (2009) 1320–1335. [5] W. Cheng, H.S. Cheng, T. Mura, L.M. Keer, Micromechanics modelling of crack initiation under contact fatigue, ASME J. Tribol. 116 (1994) 2–8. [6] H. Zhang, J. Yang, J. Hu, X. Li, M. Li, C. Wang, An experimental and simulation study of interface crack on zinc coating/304 stainless steel, Constr. Build. Mater. 161 (2018) 112–123, https://doi.org/10.1016/j.conbuildmat.2017.11.022. [7] T. Kakiuchi, Y. Uematsu, K. Suzuki, Evaluation of fatigue crack propagation in dissimilar Al/steel friction stir welds, Procedia Struct. Integr. 2 (2016) 1007–1014, https://doi.org/10.1016/j.prostr.2016.06.129. [8] S.G. Shah, S. Ray, J.M. Chandra Kishen, Fatigue crack propagation at concrete-concrete bi-material interfaces, Int. J. Fatigue 63 (2014) 118–126, https://doi. org/10.1016/j.ijfatigue.2014.01.015. [9] Z.G. Suo, J.W. Hutchinson, Interface crack between 2 elastic layers, Int. J. Fract. 43 (5) (1990) 1–18, https://doi.org/10.1007/Bf00018123. [10] J. Dundurs, Edge bonded dissimilar orthogonal elastic wedges under normal and shear loading, J. Appl. Mech. 36 (1969) 650. [11] J.R. Rice, G.C. Sih, Plane problems of cracks in dissimilar media, ASME. J. Appl. Mech. 32 (2) (1965) 418–423, https://doi.org/10.1115/1.3625816. [12] F.Z. Xuan, S.T. Tu, Z. Wang, A modification of ASTM E 1457 C* estimation equation for compact tension specimen with a mismatched cross-weld, Eng. Fract. Mech. 72 (17) (2005) 2602–2614, https://doi.org/10.1016/j.engfracmech.2005.05.002. [13] C. Vidal, V. Infante, P. Vilaça, Fatigue behaviour at elevated temperature of friction stir channelling solid plates of AA5083-H111 aluminium alloy, Int. J. Fatigue 62 (2014) 85–92, https://doi.org/10.1016/j.ijfatigue.2013.10.012 May. [14] C. Vidal, V. Infante, P. Vilaça, Assessment of performance parameters for friction stir channelling, Proceedings of the IIW 2011 International Conference on Global Trends in Joining, Cutting and Surfacing Technology, Chennai, India, (2011) 21-22 July. (ISBN 978-81-8487-152-4, paper IC_99). [15] C. Vidal, V. Infante, P. Vilaça, Characterisation of fatigue fracture surfaces of friction stir channelling specimens tested at different temperatures, Eng. Fail. Anal. 56 (2015) 204–215, https://doi.org/10.1016/j.engfailanal.2015.02.009. [16] J.B. Jordon, H. Rao, R. Amaro, P. Allison, Fatigue modeling of friction stir welding, Mater. Sci. Eng. R 50 (2005) 1–78. [17] M.N. Ilman, Fatigue crack growth behaviour of shot peened 5083 aluminium alloy friction stir welds, International Conference on Joining Materials, (May 2013), 2013. [18] C. Vidal, V. Infante, P. Vilaça, Effect of microstructure on the fatigue behavior of a friction Stirred Channel aluminium alloy, Procedia Eng. 66 (2013) 264–273, https://doi.org/10.1016/j.proeng.2013.12.081.
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Engineering Failure Analysis journal homepage: www.elsevier.com/locate/engfailanal
Experimental and numerical investigation on the fatigue behaviour of friction stirred channel plates
T
⁎
Catarina Vidala,b, , Ricardo Baptistac,d, Virgínia Infanted a
Instituto Superior Técnico, Universidade de Lisboa, Av. Rovisco Pais, 1049-001 Lisboa, Portugal UNIDEMI; Departamento de Engenharia Mecânica e Industrial, Faculdade de Ciências e Tecnologia, Universidade NOVA de Lisboa, Caparica 2829516, Portugal c CDP2T, Department of Mechanical Engineering, ESTSetúbal, Instituto Politécnico de Setúbal, Campus do IPS, Estefanilha, 2914-761 Setúbal, Portugal d LAETA, IDMEC, Instituto Superior Técnico, Universidade de Lisboa, Av. Rovisco Pais, 1049-001 Lisboa, Portugal b
A R T IC LE I N F O
ABS TRA CT
Keywords: Friction stir channelling Channel geometry Bending strength Fatigue life Fatigue crack growth
The friction stir channelling (FSC) process is a technological innovation based on the friction stir fundamentals. To produce friction stirred channel components, a non-consumable tool, similar to that used for friction stir welding (FSW), is used. Friction stir channelling is a single step manufacturing process in which a continuous channel with a specific path is produced in a monolithic metal component. This work discusses the influence of the channel geometry on the bending strength of friction stirred channel aluminium alloy specimens, as well as their fatigue behaviour. Moreover, fatigue analyses using the finite element method were also carried out, considering the channels' geometrical features in order to assess and compare the fatigue lives. The stress intensity factor for different crack lengths was determined using ABAQUS. The fatigue crack growth curve was established according to the Paris Law, and the crack was only allowed to propagate along the materials interface. It was observed that the critical zones are located in the vicinity of the channel corners and found that the fatigue crack propagation period on FSC specimens is very short compared to the crack initiation period. Base and FSW material properties play a major role on fatigue crack propagation simulation. The accuracy of the predicted results increased when considering lower stress amplitudes and by using the FSW material parameters.
1. Introduction Developments in the friction stir channelling (FSC) technology have made possible the material removal by the tool as the channel is being produced, leaving the metal component with the same level and geometry it had before being processed [1,2]. As for all friction stir based manufacturing technologies, a visco-plasticised solid-state region is generated and processed into a new shape and properties. This region whilst remaining solid presents a three-dimensional material flow pattern almost as a liquid, enabling easy mixing and blending between different materials. This phenomenon is generally referred to as the third-body region concept [3]. The material flow during the FSC process leads to the formation of a channel surrounded by four different material regions: the advancing side (AS), the retreating side (RS), the top (nugget) and the bottom of the channel. In the nugget zone, the grains are dynamically recrystalized. Fig. 1 shows a typical friction stirred channel cross-section macrograph. ⁎
Corresponding author at: Instituto Superior Técnico, Universidade de Lisboa, Av. Rovisco Pais, 1049-001 Lisboa, Portugal. E-mail address: [email protected] (C. Vidal).
https://doi.org/10.1016/j.engfailanal.2019.04.068 Received 23 January 2019; Received in revised form 15 April 2019; Accepted 26 April 2019 Available online 29 April 2019 1350-6307/ © 2019 Elsevier Ltd. All rights reserved.
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Fig. 1. Cross-section macrograph of a friction stirred channel produced in an aluminium alloy component.
The interface between the nugget and the thermo-mechanically affected zone (TMAZ) is the main structural concern of the friction stirred channel components, particularly from the fatigue point of view, as the corner defined by the AS and the nugget is a region of stress concentration between very fine and coarse grains, amplified by its geometry. Generally, the fatigue life can be divided into two major steps [4]: 1) Crack initiation, due to local accumulation of dislocations, stress hot points, material defects or plastic deformation due to inhomogeneous inclusions; 2) Crack propagation, due to crack opening and closing and introducing permanent damage to the material. Crack initiation mode and local depends on material microstructure, applied loads, and specimen geometry [5]. The duration of the fatigue crack initiation period can be influenced by different factors and can occur either on the specimen surface or under the specimen surface. It can be observed near local stress concentration points, where the local resistant area is smaller. Therefore, the total number of fatigue cycles, Nf, required to achieve the final fracture can be obtained from the number of cycles required for crack initiation, Ni, and the number of cycles required for crack propagation to occur, Np, from the initial to the critical crack length (Nf = Ni + Np). When considering the interface between two different materials, one must consider that this zone shows lower fatigue resistance than both, the base and the added materials. The interface between materials can be used to represent different situations: 1) a crack between a base material and a coating [6], 2) a crack between two dissimilar materials joined by friction stir welding [7] or 3) a crack between two non-metallic materials [8]. Regardless the case, the crack propagation rate on the interface will always be higher than on the base material. Accordingly, to Shah et al. [8], this can be attributed to the mismatch in the Young's modulus between the different materials. Different fracture parameters, as the energy release-rate, must be used when considering fatigue crack propagation along materials' interfaces, such as the interface between different elastic layers [9] or friction stir welds [7]. The different behaviour can also be attributed to the fact that a crack between two different material interfaces, inherently propagates in mixed mode, and not in pure mode I or pure mode II [8]. Due to the mismatch between the Young's modulus, even a pure mode I problem will not be symmetric, and the crack tip will be subjected to normal and shear stresses. Therefore, crack propagation will always depend on the crack length and the relationships between the two materials properties. The stress field near the crack tip will then show an oscillatory behaviour, eq. (1), characterized by an oscillatory index ϵ introduced by Dundurs [10], eq. (2). Unlike the linear homogeneous conditions, where the stress filed shows an inverse square root singularity. The nature of the stress field and stress singularity leads to oscillatory crack propagation, with the crack inter-penetrating both materials. Therefore, crack opening mode I and mode II cannot be decoupled, rendering analytical solutions inadmissible [8]. 1
σ ~r − 2 + ϵ ϵ=
(1)
1 − β⎞ 1 ln ⎛⎜ ⎟ 2π ⎝ 1 + β ⎠
(2)
Parameters β and α are the material interface parameters defined by Dundurs [10]. Parameter β is the elastic mismatch parameter, a function of the Poisson's ratio, υ, and the material shear modulus, μ, (eq. (3)). Parameter α measures the relative materials stiffness (eq. (4)), where E denotes the equivalent modulus of elasticity for plane stress or plane strain and is a function of the materials' mismatch. 1 and 2 indices denote material 1 and material 2, respectively. Both parameters will be equal to zero in the case of an interface between equal properties materials.
β=
μ1 (1 − 2υ2) − μ 2 (1 − 2υ1 ) 2(μ1 (1 − υ2) + μ 2 (1 − υ1 ))
(3)
α=
E1 − E2 E1 + E2
(4)
Rice [11] defined the behaviour of the stress field along the crack interface. Eq. (5) establishes the oscillatory behaviour of the stress singularity and, as mentioned, the stress intensity factors for mode I and mode II, cannot be decoupled.
σyy + iσyy =
K1 + iK2 iϵ r 2πr
(5)
It is then required to determine the solution for the coupled stress intensity factors, in order to analyse the crack propagation along the interface between different materials. One solution for this problem is to use the energy release rate, G, or J-integral to study the interface crack propagation, as reported by Kakiuchi et al. [7] and Xuan et al. [12]. The finite element method (FEM) 58
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provides a simple solution to numerically estimate the energy release rate or J-integral. The J-integral can then be related to the stress intensity factor using eq. (6), even for a mixed material problem.
J=
1 − β2 2 1 2 (K1 + K22) + K3 E∗ 2G∗ 1 E∗
1 (E1 2
1 G∗
(6)
1 (G1 2
= + E2) and = + G2.) where and G denotes the materials' shear modulus. FEM is therefore a suitable method to solve this type of problem. Finally, the Paris Law can be used to calculate the fatigue crack propagation life, using eq. (7). da = B ∆G n dN
(7)
Where B and n are material parameters. In this paper FEM analysis was used to estimate the energy release rate for different crack lengths along the interface between the nugget and the TMAZ. Using eq. (6), the number of elapsed cycles was calculated for each crack increment and added to the previous obtained value, to calculate the total fatigue life. 2. Material and experimental details 2.1. Material The strain hardened aluminium alloy AA5083-H111 was used in this work. Its nominal chemical composition is Al–5.26 Mg–1.02Mn–0.19Fe–0.15Cr (wt%) [13]. The tensile properties, at room temperature are: σy = 210 MPa, E = 68.9 GPa, σUTS = 375 MPa and A = 18.3% [13], where, σy is the yield stress for 0.2% offset parallel line to the elastic line, with Young's modulus, E; σUTS is the ultimate tensile stress and A is the elongation. 2.2. Friction stirred channel production An ESAB LEGIO™ 3UL numeric control equipment was used to produced friction stirred channels on 15 mm-thick rolled plates along the rolling direction. Plunge and dwell periods (zero tool travel speed) were performed under vertical position control and processing period (non-zero tool travel speed) was carried out under vertical force control. In order to obtain channels with different dimensions, four sets of FSC parameters were implemented as shown in Table 1. Channels U# were produced with a probe length of 6 mm and channels L# with a probe length of 8 mm. The channels were produced using a modular H13 steel tool comprised by a body, a shoulder and a probe. All tool's arrangement had a threaded cylindrical probe with 8 mm diameter. Table 2 shows the three different shoulders used. The tool tilt angle was 0° for all the runs. 2.3. Four-point bend testing Four-point bending tests were performed to obtain the mechanical strength of the channels produced. This test was selected because the purpose is to assess the mechanical strength of a specific and large area, the processed zone, rather than a specific and small zone as, for example, a weld root [14]. Specimens were manufactured according to the standard ASTM E290-97a requirements as shown in Fig. 2. From each FSC condition (Table 1), two bending specimens were taken and tested with the processed surface under tensile stress, as depicted in Fig. 3. An electromechanical Zwick Roell testing machine with a load cell of 200 kN was used to carry out the four-point bending tests. The rollers have a diameter of 15 mm. The span distances, the values of which are given in Fig. 3, were set according to the ASTM E 290-97a standard. Four-point bend testing was performed at a loading speed of 1 mm/min in laboratory air at room temperature. 2.4. Fatigue testing The fatigue specimens were manufactured with the same dimensions of the four-point bending specimens (Fig. 2). The fatigue tests were carried out at room temperature in a servo-hydraulic test machine (Instron® 8502), with a load capacity of 100 kN. The Table 1 FSC parameters. FSC set
Tool rotation speed (rpm)
Tool travel speed (mm/min.)
Vertical force (kN)
U1 U2 L1 L2
1100
50 80 50 100
4 2 1.5 2.2
400
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Table 2 Tool shoulders used to produce the channels. FSC set
Tool shoulder
U1, L2
U2
L1
Fig. 2. Geometry and dimensions of the four-point bending specimens [15].
15 P/2
30 90
P/2
30
Fig. 3. Schematic representation of a FSC specimen under four-point bending.
fatigue cycle was constant amplitude loading, with a sinusoidal load wave of 5–10 Hz, with stress ratio of R = 0.1 and a four-point bending set up. The fatigue tests were performed with the maximum stress levels ranging from 30 to 150 MPa and it was defined the run-out limit at 3 × 106 cycles.
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Fig. 4. Channel nugget configurations. Crack propagation location along interfaces.
The applied load, P, was calculated by eq. (8) where w is the smallest span distance (30 mm), h is the specimen's thickness, b is the specimen's width and σ is the maximum stress.
P=
σbh2 3w
(8)
The load was applied out of the processed zone and the specimens were tested with the processed surface under tensile stress (Fig. 3). 3. Numerical simulation 3.1. Finite element model Four different channel geometries (U1, U2, L1 and L2) were modelled (Fig. 4) in ABAQUS. The goal was to model crack propagation along the materials' interface between the nugget and the TMAZ on the advancing side (AS) of the channel. Using the spline function to reconstruct the different channel and the nugget geometries, four two-dimensional solid models, with a constant 30 mm depth, were built. Each specimen with a 180 × 30 mm section was subjected to four-point bending loads and boundary conditions, considering a 90 mm span. 3.2. Fatigue failure simulation Crack propagation simulations were performed with different loading amplitudes, in order to simulate the experimental conditions. Initially four rollers were fully modelled, including contact between the rollers and specimen, in order to simulate four-point bending. The simulation of the rollers-specimen contact resulted in a complex and numerically demanding simulation. It was then verified that by applying the loads and boundary conditions to four nodes directly, the results were not affected, and the resulting linear simulation could be solved faster. Base and channel nugget materials were considered as linear elastic, with properties described in section 2.1 and Table 3, respectively. The effect of FSC residual stresses was not modelled in our work. It is believed that FSW and FSC produce generally low residual stresses, due to low temperature solid-state process [16]. Compressive residual stresses lead to lower crack propagation rates, but this effect is neglected when using higher R [16] values. Tension residual stresses accelerate crack propagation. By not simulating this effect, our results will be less conservative. Each of the four models used quadratic plane strain elements. In order to define the stress singularity around the crack tip, a special meshing technique was used (Fig. 5 c)). Using wedge elements with collapsed nodes, five contour integrals were used to estimate the energy release rate (J-integral). The different models used up to 50,000 nodes. Considering a constant crack increment (Δa) and eq. (9), the fatigue crack propagation life for each specimen and loading amplitude was calculated.
N=
∆a B ∆G n
(9) Table 3 Linear elastic properties of the processed material used in FEM analyses [18].
Nugget (L1 and L2) Nugget (U1 and U2)
Young's modulus (GPa)
Poisson rate
69.8 67.3
0.3 0.3
61
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Fig. 5. Mesh and crack details. a) L2 channel and nugget, b) crack and mesh along the crack, c) spider-web mesh for crack tip contour integral calculation.
From the work of Ilman et al. [17], the parameters for the Paris law were estimated. These authors considered the crack propagation on similar base material and nugget material resulting from FSW (Table 4), although not along the materials interface. 4. Results 4.1. Four-point bend testing Four-point bending tests were performed to evaluate the mechanical performance of the friction stir channels. A typical loaddisplacement curve was plotted for each test, as presented in Fig. 6. The displacement was obtained from the machine cross-head movement. To assess the channels' strength, their ultimate bending strength was additionally calculated by using eq. (8). The maximum load, P, was extracted from the curves. A summary of the mechanical performance of the channels is presented in Table 5. To better understand the influence of the channel shape and dimensions on the bending strength, as well as to establish feasible comparisons among the channels analysed, their geometric parameters were measured. The results are schematic illustrated in Fig. 7. 4.2. Fatigue testing Table 6 presents the fatigue test results for each one of the four FSC conditions analysed. Table 4 Fatigue crack propagation parameters for the base material and FSW material, adapted from [17]. B (m/cycle) Base material FSW
−1
1.365 × 10 1.116 × 10−4
62
n 2.222 1.078
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Fig. 6. Load-displacement curves.
Table 5 Four-point bend testing results. Condition ID
Specimen ID
Maximum load (kN)
Average value (kN)
Deviation (kN)
Ultimate bending strength (MPa)
U1
1 2 1 2 1 2 1 2
30.0 32.0 22.1 21.2 20.0 20.6 10.4 9.9
31.00
± 1.00 (3.2%) ± 0.45 (2%) ± 0.30 (1.5%) ± 0.25 (2.5%)
400 427 295 283 267 275 139 132
U2 L1 L2
21.65 20.30 10.15
The obtained fatigue strength values for conditions U1, U2 and L1 were significantly higher than those for condition L2. Therefore, these results indicate that, for the L2 condition, a reduced fatigue strength is obtained due to the channel geometry. It was also found that when the maximum applied stress increases the number of cycles to failure decreases, as expected, with a maximum stress applied of 50 MPa for a run-out of 3 × 106 cycles. 4.3. Numerical simulations On each of the four types of fatigue specimen numerical models, a crack on the materials' interface was allowed to propagate. An initial crack size of 0.1 mm was considered for all specimen models. These cracks propagated from the closest corner, of the modelled channel, to the free surface of the specimen. The crack propagation simulation ended when loss of tightness of the channel occurred, i.e., the crack reaches the specimens free surface. Fig. 8 shows the energy release rate as a function of crack length. The number of elapsed fatigue cycles between increments, was calculated using eq. (9). Considering the geometry along the materials interface of specimen U1, the maximum crack length is 1.7 mm, while on specimen U2 allowed for a maximum crack length of 1.95 mm. On these two specimens when a load that produces a normal stress of 50 MPa is applied, the maximum energy release rates were 5 J/mm2 and 11 J/mm2, respectively. For specimen L2 the maximum crack length is 1.2 mm, as the distance between the channel and the free surface is much smaller. In fact, Fig. 5 shows that the crack propagates in a constant direction toward the free surface, and therefore the energy release rate increases exponentially with the crack length. The crack propagation path is vertical, unlike on specimens U1 and U2 (Fig. 4), and the crack opening mode is almost pure mode I, therefore the maximum energy release rate is higher, reaching 21 J/mm2. Specimen L1 allowed a 3.3 mm maximum crack length. As one can see in Fig. 4, the materials interface changes direction, and therefore the crack propagation path is longer. The crack opening mode changes from pure mode I to a mixed mode and back to almost pure mode I again. When changing from mode I to mixed mode the energy release rate slightly decreases, slowing down the crack propagation rate. As the crack length reaches the critical value, and the opening mode returns to mode I, the crack propagation rate increases as the energy release rate increases to 21 J/mm2. Fig. 9 shows the corresponding crack propagation curves, for a fatigue load with a maximum normal stress of 50 MPa. The fatigue life of the specimen L2 resulted in 23,069 cycles. This is the shortest fatigue life, as the maximum crack length is low, and the energy release rate is high. The fatigue life for specimens U1 and U2, was 117,260 and 138,150 cycles, respectively. The behaviour of these cracks is very similar, as well as the evolution of the energy release rate, that governs the crack propagation rate. Finally, the fatigue 63
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C. Vidal, et al.
details (a) to (d)
a)
3
1.5
6.5
b)
3.7
1.6
8
c)
2.7
4.6
8
d)
1.3
5.8
2.8
8
Fig. 7. Geometry and dimensions of the specimens. Details showing the schematic representation of the channels produced with FSC set a) U1, b) U2, c) L1 and d) L2 [18].
life of specimen L1 was 274,780 cycles. As previously mentioned, this is the longest crack and the one that shows the lowest energy release rates, as the crack changes directions along the interface. One can see in Fig. 9, that the crack propagation rate decreases around the 150,000 cycles mark, and increases again, around the 250,000 cycles mark. Several fatigue crack propagation simulation models were created, considering different fatigue loads. The models included different crack propagation properties for the base and FSW materials. The obtained results as summarized in Table 7. Considering the base material crack propagation properties, the obtained fatigue life is always higher than the one obtained with the processed material properties. Therefore, the base material has a higher fatigue crack propagation resistance, as concluded by Ilman et al. [17].
5. Discussion From Fig. 6 and Table 5, it is possible to conclude that the channels produced with FSC conditions U1 are the most resistant to bending loads; channels produced with FSC set of parameters U2 and L1 have similar performances and channels performed at conditions L2 exhibit almost elastic behaviour before final failure. Channels produced with FSC conditions U1 also have the largest elongation. The height of the closing layer thickness of channels produced with parameters U1 and U2 is almost equal (Fig. 7), however, channels U1 have a smaller area and were performed at a higher vertical force, which seems to indicate that improves the channels' mechanical strength and ductility. Regarding the influence of channel depth on the mechanical behaviour of the channels, no feasible conclusion can be drawn from the results of conditions U2 and L1. One can observe that a thicker closing layer did not lead to a better mechanical performance which can be explained by the presence of small internal voids on the channel nugget of condition L1 (Fig. 10). From the global analysis of fatigue test results, it is possible to conclude that FSC specimens exhibit a very low fatigue strength comparing to the unprocessed base material, even though they have no defects. As can be extrapolated from Table 6 the scattering of the results is quite low which demonstrates that, on one hand it is possible to predict the fatigue behaviour of FSC components and, on the other hand, the production process is stable, and the channels are reproducible. Regarding the fatigue strength, the results indicate that specimens produced with condition L1 revealed a slightly better fatigue performance than those performed with parameters U1 and U2, mainly at lower stress values, which seems to suggest that, for the 64
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Table 6 Fatigue test results [18]. Fatigue specimen ID
σmax (MPa)
Freq. (Hz)
Nf, Cycles to failure
U1_1 U1_2 U1_3 U1_4 U1_5 U1_6 U1_7 U2_1 U2_2 U2_3 U2_4 U2_5 U2_6 U2_7 U2_8 U2_9 L1_1 L1_2 L1_3 L1_4 L1_5 L1_6 L1_7 L1_8 L1_9 L2_1 L2_2 L2_3 L2_4 L2_5 L2_6
125 125 100 100 65 65 50 100 100 80 80 65 65 60 60 50 150 125 100 100 80 80 65 65 50 65 65 50 50 40 40
10 10 10 10 10 10 10 5 5 5 5 7.5 7.5 7.5 7.5 7.5 3 4 4 4 5 5 8 8 8 7.5 7.5 7.5 7.5 10 10
13,702 15,794 49,038 54,881 385,977 291,320 3,000,000 58,595 51,856 142,555 149,081 330,259 329,134 353,401 374,398 3,000,000 15,259 17,094 46,165 67,293 165,572 120,796 773,309 808,165 3,000,000 14,953 19,890 54,954 45,628 88,099 196,253
Obs.
Run-out
Run-out
Run-out
Fig. 8. Energy release rate vs crack length, for a fatigue load with a maximum normal stress of 50 MPa.
conditions analysed, a thicker closing layer tends to improve the fatigue strength of FSC specimens, regardless the channel area. In fact, a thicker closing layer contributes to longer fatigue crack propagation periods, although it has been observed that the crack propagation period for FSC specimens is very small when compared with the number of cycles required for crack initiation. The similarity among the SeN curves of defect free FSC specimens can be explained by the similar crack initiation periods observed for conditions U1, U2 and L1. The worst fatigue performance presented by the specimens produced with FSC conditions L2 is related with the defect observed in the nugget/TMAZ interface. The defect works as a notch, leading to higher local stresses, which contributes to failure at a lower number of cycles. When considering the base material and FSW material models offered by Ilman et al. [17], one can see (Figs. 11 to 14) that the fatigue lives predicted by the crack propagation simulations presented herein, using the base material properties, are always higher than the experimental ones. This shows that the current model is not suitable for the fatigue crack propagation at the interface
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Fig. 9. Fatigue crack propagation curves, for a fatigue load with a maximum normal stress of 50 MPa. Table 7 Numerical fatigue simulation results. Fatigue specimen
σmax [MPa]
FSW - N [Cycles]
BM - N [Cycles]
L1
150 125 100 80 65 50 65 50 40 125 100 65 50 100 80 65 50
25,728 38,115 61,662 99,756 156,080 274,780 13,105 23,069 37,320 16,265 26,314 66,605 117,260 31,003 50,156 78,474 138,150
40,676 91,454 246,520 664,510 1,672,000 5,365,000 30,544 98,008 264,190 38,287 103,210 699,970 2,246,000 165,820 446,970 1,124,600 3,608,700
L2
U1
U2
Fig. 10. Optical micrograph under polarized light from an electrolytic etched L1 sample showing small internal voids on the channel nugget.
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Fig. 11. Fatigue life simulations for the L1 specimens.
Fig. 12. Fatigue life simulations for the L2 specimens.
Fig. 13. Fatigue life simulations for the U1 specimens.
between base material and FSW material. As mentioned, this interface is considered to have lower fatigue crack propagation resistance. When considering the FSW material model for the L1 specimen and a load amplitude of 50 MPa, as seen in Fig. 11, the predicted fatigue life represents 9% of the total fatigue life. Fig. 13 shows a similar behavior for the U1 specimen. For a fatigue load of 50 MPa, the simulated fatigue life represents 4% of the total life. The remaining 96% would represent the initiation fatigue life that is not considered by our model. But if one considers a fatigue load of 125 MPa, the simulated crack propagation life represents 110% of the experimental one. The model is not suitable for fatigue predictions when considering higher load amplitudes. Figs. 12 and 14 67
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Fig. 14. Fatigue life simulations for the U2 specimens.
show similar behaviors for the L2 and U2 specimens, respectively. As explained before, the fatigue crack propagation was obtained using the Paris' law where the predicted life was obtained under Linear Elastic Fracture Mechanical (LEFM). Predictions of fatigue life in crack propagation are compared with experimental data of fatigue life and the determination of fatigue crack initiation (Ni) is, in this study, defined by the difference between the experimental numbers of cycles (Nf) and the propagation number of cycles obtained in the computational simulations (Np). Table 8 presents the experimental and numeric fatigue failure results and the prediction of the fatigue crack initiation cycles. As expected, the number of cycles Ni required for crack initiation increase with the decreasing of maximum stress applied. It was found that the total number of cycles (obtained from experimental tests) is similar to the propagation number of cycles (obtained in numerical simulations) for the highest load amplitudes. Increasing the load amplitude means that the initiation phase is practically non-existent. On the other hand, for the smallest applied stresses (σmax = 50 MPa for the specimens U1, U2 and L1) where the experimental tests showed infinite lives it is possible to conclude that the stress intensity factor range is less than the threshold value for the propagation of the crack suggesting that for this condition the defects would never initiate. 6. Conclusions The channels produced with FSC conditions U1 have shown to be the most resistant to bending loads and present the largest elongation; channels produced with FSC set of parameters U2 and L1 have similar performances and channels performed at conditions L2 exhibit almost elastic behaviour before final failure. Regarding the influence of channel depth on the mechanical behaviour of the channels, no feasible conclusion can be drawn from the results of conditions U2 and L1. Even considering that FSC specimens showed no defects, the analysis of their fatigue strength has revealed to be lower than the unprocessed base material. A thicker closing layer, increasing the distance between the channel and the specimen free surface, contributes to longer fatigue crack propagation periods. It was also verified that on the FSC specimens this period is very short when compared with the fatigue crack initiation period. Table 8 Comparison of the experimental and numerical results. Specimen
σmax (MPa)
Experimental results, Nf (number of cycles)
Computational results, Np (number of cycles)
Ni = Nf - Np (number of cycles)
U1
125 100 65 50 100 80 65 50 150 125 100 80 65 50 65 50 40
14,748 51,960 338,649 Run-out 55,226 145,818 329,697 Run-out 15,259 17,094 56,729 143,184 790,737 Run-out 17,422 50,291 142,176
16,265 26,314 66,605 117,260 31,003 50,156 78,474 138,150 25,728 38,115 61,662 99,756 156,080 274,780 13,105 23,069 37,320
– 25,645 272,043 – 24,222 95,662 251,222 – – – – 43,428 634,657 – 4316 27,222 104,856
U2
L1
L2
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Using the base material crack propagation properties resulted in a longer fatigue life, when compared to the experimental results or the numerical ones obtained with the FSW parameters. Therefore, the base material model used is not compatible with fatigue crack propagation simulations on the interface between materials. When considering the FSW material parameters and a load amplitude of 80 MPa, the estimated fatigue crack propagation life represents a maximum of 70% of the total fatigue life. Lower load amplitudes predicted lower crack propagation life percentages. Therefore, the simulations are more accurate for lower load amplitudes. Considering crack initiation was not modelled, it is recommended to use the FSW material parameters, when the interface data is not available. Future work should be applied to the numerical model, and to experimentally quantify the relation between the crack initiation and propagation periods, allowing to choose a more suitable material model for fatigue crack propagation analyses. Including the effect of tension residual stresses will accelerate the crack propagation rate, decreasing the final fatigue propagation life, also making the calculated values more in-line with the experimental results. Acknowledgements The authors would like to acknowledge the Portuguese Foundation for the Science and Technology (FCT) for its financial support through the PhD scholarship FCT SFRH/BD/62963/2009, IDMEC, under LAETA project UID/EMS/50022/2019 and project UID/ EMS/00667/2019. References [1] C. Vidal, P. Vilaça, Processo de Abertura de Canais Internos Contínuos em Componentes Maciços Sem Alteração da Cota de Superfície Processada e Respectiva Ferramenta Modular Ajustável, Patent number: PT 105628 (B), April 15 (2011). [2] C. Vidal, V. Infante, P. Vilaça, Fatigue assessment of friction stir channels, Int. J. Fatigue 62 (2014) 77–84, https://doi.org/10.1016/j.ijfatigue.2013.10.009 May. [3] W.M. Thomas, An Investigation and Study into Friction Stir Welding of Ferrous-Based Material, PhD Thesis University of Bolton, 2009. [4] G. Fajdiga, M. Sraml, Fatigue crack initiation and propagation under cyclic contact loading, Eng. Fract. Mech. 76 (2009) 1320–1335. [5] W. Cheng, H.S. Cheng, T. Mura, L.M. Keer, Micromechanics modelling of crack initiation under contact fatigue, ASME J. Tribol. 116 (1994) 2–8. [6] H. Zhang, J. Yang, J. Hu, X. Li, M. Li, C. Wang, An experimental and simulation study of interface crack on zinc coating/304 stainless steel, Constr. Build. Mater. 161 (2018) 112–123, https://doi.org/10.1016/j.conbuildmat.2017.11.022. [7] T. Kakiuchi, Y. Uematsu, K. Suzuki, Evaluation of fatigue crack propagation in dissimilar Al/steel friction stir welds, Procedia Struct. Integr. 2 (2016) 1007–1014, https://doi.org/10.1016/j.prostr.2016.06.129. [8] S.G. Shah, S. Ray, J.M. Chandra Kishen, Fatigue crack propagation at concrete-concrete bi-material interfaces, Int. J. Fatigue 63 (2014) 118–126, https://doi. org/10.1016/j.ijfatigue.2014.01.015. [9] Z.G. Suo, J.W. Hutchinson, Interface crack between 2 elastic layers, Int. J. Fract. 43 (5) (1990) 1–18, https://doi.org/10.1007/Bf00018123. [10] J. Dundurs, Edge bonded dissimilar orthogonal elastic wedges under normal and shear loading, J. Appl. Mech. 36 (1969) 650. [11] J.R. Rice, G.C. Sih, Plane problems of cracks in dissimilar media, ASME. J. Appl. Mech. 32 (2) (1965) 418–423, https://doi.org/10.1115/1.3625816. [12] F.Z. Xuan, S.T. Tu, Z. Wang, A modification of ASTM E 1457 C* estimation equation for compact tension specimen with a mismatched cross-weld, Eng. Fract. Mech. 72 (17) (2005) 2602–2614, https://doi.org/10.1016/j.engfracmech.2005.05.002. [13] C. Vidal, V. Infante, P. Vilaça, Fatigue behaviour at elevated temperature of friction stir channelling solid plates of AA5083-H111 aluminium alloy, Int. J. Fatigue 62 (2014) 85–92, https://doi.org/10.1016/j.ijfatigue.2013.10.012 May. [14] C. Vidal, V. Infante, P. Vilaça, Assessment of performance parameters for friction stir channelling, Proceedings of the IIW 2011 International Conference on Global Trends in Joining, Cutting and Surfacing Technology, Chennai, India, (2011) 21-22 July. (ISBN 978-81-8487-152-4, paper IC_99). [15] C. Vidal, V. Infante, P. Vilaça, Characterisation of fatigue fracture surfaces of friction stir channelling specimens tested at different temperatures, Eng. Fail. Anal. 56 (2015) 204–215, https://doi.org/10.1016/j.engfailanal.2015.02.009. [16] J.B. Jordon, H. Rao, R. Amaro, P. Allison, Fatigue modeling of friction stir welding, Mater. Sci. Eng. R 50 (2005) 1–78. [17] M.N. Ilman, Fatigue crack growth behaviour of shot peened 5083 aluminium alloy friction stir welds, International Conference on Joining Materials, (May 2013), 2013. [18] C. Vidal, V. Infante, P. Vilaça, Effect of microstructure on the fatigue behavior of a friction Stirred Channel aluminium alloy, Procedia Eng. 66 (2013) 264–273, https://doi.org/10.1016/j.proeng.2013.12.081.
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