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Fatigue behaviour and damage evolution of single lap bonded joints in composite material
COMPOSITES SCIENCE AND TECHNOLOGY Composites Science and Technology 66 (2006) 176–187 www.elsevier.com/locate/compscitech
Fatigue behaviour and damage evolution of single lap bonded joints in composite material M. Quaresimin *, M. Ricotta Department of Management and Engineering, University of Padova, Stradella San Nicola 3, 36100 Vicenza, Italy Received 7 March 2005; accepted 12 April 2005 Available online 27 June 2005
Abstract The paper presents the results of an experimental investigation on the fatigue behaviour of single lap bonded joints. Carbon/ epoxy laminates bonded with epoxy adhesive were tested under tension–tension fatigue loading and the effect of overlap length and corner geometry was discussed by using the classical stress-life approach. Significant improvements in fatigue strength can be obtained by adopting long overlap length and spew fillet corner geometry. A careful analysis of the evolution of the fatigue damage was also carried out and it was observed that a significant fraction of the fatigue life of the joint is spent in the nucleation of one or more cracks which then propagate up to failure of the joint. The duration of nucleation process, which can last from 20% up to 70% of the joint life, suggests the need of incorporating this phase in the development of future predictive models. The fatigue crack propagation process is also discussed and the rates of crack propagation evaluated. 2005 Elsevier Ltd. All rights reserved. Keywords: Fatigue; Damage evolution; Composite material
1. Introduction This paper is the first of a series of three which present an overview on the activity done at University of Padova in the framework of a project for the development of fatigue design methodologies for bonded joints in composite materials. In this first paper a synthesis of the experimental activity on single lap bonded joints in composite material is presented and discussed. The second paper [1] deals with the finite element analyses carried out to investigate the influence of the main design parameters on the stress distributions in the joints and to provide suitable tools for the definition of the life prediction methodology, which is presented in the third paper [2]. *
Corresponding author. Tel.: +39 0444 998723; fax: +39 0444 998888. E-mail address: [email protected] (M. Quaresimin). 0266-3538/$ - see front matter 2005 Elsevier Ltd. All rights reserved. doi:10.1016/j.compscitech.2005.04.026
As it will be better discussed and explained later, the fatigue life of the bonded joints tested here consists of a first phase spent in the nucleation of a macroscopic crack and the subsequent evolution of this crack which propagates at the adhesive–adherend interface up to a critical length corresponding to the joint failure. The fraction of the fatigue life spent in the first phase was found to be always greater than 20%. The life prediction models available in the literature usually neglect this phase and this could lead to strong underestimations of the joint fatigue life. Even though on the conservative side, this way of design is not very efficient from an engineering point of view. To overcome this problem, the model developed during the project and presented in [2] estimates the fatigue life of a joint taking into account the actual evolution of the fatigue damage: the first phase of the fatigue life, required to nucleate a crack, is calculated by using a non-conventional extension of the Stress Intensity Factor approach, as sug-
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gested in [3,4], whereas the subsequent propagation phase is estimated through the integration of a Paris-like curve relating the crack growth rate da/dN to the range of the Strain Energy Release Rate (SERR), as commonly done in the literature. In this paper, after a brief survey of some of the experimental analyses available in the literature, the results obtained on single lap joints made by bonding carbon/epoxy laminates with an epoxy adhesive are presented, with particular reference to the description of the joint fatigue behaviour and the accurate analysis of the damage evolution under fatigue loading. The investigations published in the literature on the fatigue behaviour of bonded joints can be divided into two main groups: those oriented to evaluate the loadcarrying capability of the joints, seen as load-bearing parts of a structure, and those more interested in analysing the evolution of the damage and measuring the rate of crack propagation. In the former case the results are presented, according to the stress-life approach, by using the S-N curves from nominal stress and number of cycles to complete failure. In the latter case, the crack propagation rate is usually related, in a Paris-like curve, to an energetic or fracture mechanics parameter like the Strain Energy Release Rate (G), the J-integral (J) or the Stress Intensity Factor (K or SIF). Since both approaches can provide useful and synergical design indications, papers and results referred to both approaches are briefly discussed, with particular attention to those illustrating the influence of design parameters and describing the mechanics of the fatigue damage evolution. The first experimental investigations date back to the early Õ70s. Renton and Vinson [5] found an increase in the static tensile strength of single lap joints with the overlap length, while joints with shorter overlap exhibited higher normalised fatigue strength values, independently of the ply orientation at the interface. However, the absolute values of fatigue strength were higher for joints with longer overlap. Ishii and co-authors [6] noticed the same trend, investigating the fatigue behaviour of CFRP/aluminium single and double lap joints. The length of the overlap was also found to have influence on the damage evolution: for short overlap (10 mm) the fatigue life was spent almost entirely for nucleating a fatigue crack, whereas for long overlap two distinct phases, crack onset and crack propagation, were observed. The beneficial contribution of longer overlap length is also reported in [7]. The fatigue curves are presented here in terms of shear stress on the adhesive and therefore need to be rearranged in terms of nominal tensile stress on the adhered to allow a comparison to be made with the previous data. Zeng and Sun [8] proved that the static properties of a single lap bonded joint can be almost doubled by an im-
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proved shape of the overlap zone. The wavy design adopted for the joints leads also to significant improvements in the fatigue strength. Mall et al. [9], investigating the influence of the load ratio R on the crack propagation rate in double cantilever beam (DCB) and cracked lap shear (CLS) joints, found an apparent improvement in the fatigue strength at the higher values of the load ratio, with lower propagation rate for a given values of the maximum Strain Energy Release Rate, Gmax. The improvement disappeared and the data collapsed in one single curve, when the crack propagation data obtained under different positive stress ratio were plotted against the SERR range DG rather than against Gmax. The same situation was reported in [10], where the rate of crack growth was found to be almost insensitive to the load ratio, at least in the range R = 0.1–0.5, when using DG. On the other hand, a drastic reduction in the crack growth resistance under fully reversed loading (R = 1) with respect to the case of no reversal (R = 0.1) was reported for Mode II fatigue loading of composite ENF joints bonded with a toughened epoxy adhesive [11]. The study also indicated in the total Strain Energy Release Rate the driving parameter for the crack propagation. In another work on the same adhesive [12], Mall and Ramamurthy investigated the effect of the adhesive layer thickness on the Mode I propagation: DCB joints with different adhesive thickness exhibited about the same threshold at lower crack growth rates, whereas a thicker adhesive layer resulted in an improved resistance to the crack growth in the high propagation rate region. A careful description of the evolution of the fatigue damage in bonded joint was presented by Dessureault and Spelt [13]. Although on aluminium-epoxy bonded joints, the paper is worth to be discussed as a bright example of illustration of the fatigue damage mechanics. Studying the behaviour of DCB and CLS under different mode loading (I, II and mixed-mode) and constant load ratio (R = 0.1), the authors correlated the life to crack initiation and the crack propagation rate to the Strain Energy Release Rate. The analysis of the damage evolution indicated a high variability in the sites of crack nucleation as well as a significant part of the fatigue life spent in the crack nucleation. This fraction of life was found to increase as the maximum applied SERR decreased; the fatigue threshold Gth, derived as the asymptotic values of SERR at which no cracks initiated before 107 cycles, resulted almost unaffected by the mode-ratio. When plotted against the Gmax, the crack propagation rates were similar under Mode I and mixed-mode loading whereas under pure Mode II conditions the propagation rates were significantly lower. However, the fatigue threshold Gth estimated from the Paris-like curves at very low propagation rates was in good agreement with that obtained from the crack initiation analysis.
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An interesting method to control the fatigue damage evolution and the crack paths, improving therefore the prediction capability, was presented by Potter et al. [14]. By introducing polymeric films in a suitable position within the adhesive layer, the authors induced the cracks to propagate within the adhesive layer rather than into the composite adherend, improving also the fatigue performances of the joints. An extensive analysis of the damage mechanisms is also presented. In composite bonded joints, temperature, humidity and water absorption can have a significant influence on both adhesive and adherend properties and therefore the possible effects of the environment on fatigue behaviour and damage evolution were extensively investigated. Just to mention some significant examples, Krause et al. [15] found a greater sensitivity to the immersion in water at 40 C for single lap joints bonded with epoxy adhesives compared to the same joints bonded with urethane adhesives. They also noted that only the fatigue strength was affected by the water absorption whereas no appreciable loss in the static strength was observed. Gilmore and Shaw [16] studied the influence of temperature and humidity on the fatigue damage evolution. Both hot (90 C) and cold (50 C) environment accelerated the nucleation of a fatigue crack, whereas the humidity had significant influence only at higher temperature, inducing the adhesive to fail cohesively. Ashcroft et al. [17] tested lap strap joints bonded with modified epoxy adhesive under tension–tension fatigue loading at different temperatures (50, 22, 90 C) and humidity degrees (ambient humidity and 95% R.H.). Similar values of the threshold load were observed for the ‘‘ dry’’ joints tested at the different temperature. On the other hand moisture-conditioned joints tested at 90 C or ‘‘dry’’ joints tested at 90 C in wet environment exhibited a significant reduction in the threshold load for damage initiation. The authors explained this behaviour indicating a reduction in the fatigue resistance for test conditions near the glass transition temperature Tg of the adhesive and noted also a significant decrease in the adhesive Tg due to the moisture absorption. Sites and mode of fatigue crack initiation as well as damage patterns, duration of the propagation phase and failure modes were found also to be heavily influenced by the environmental conditions. Ashcroft and Shaw [18] studied the influence of the temperature on the Mode I fatigue crack propagation of DCB bonded joints produced with the same modified epoxy adhesive. They found the threshold value for SERR, Gth, to increase from 80 J/m2 at 50 C to 130 J/m2 at 90 C. At low temperature the propagation rate was very high (slope of the Paris-like curve m = 8.8), whereas lower and similar values (m = 3.4– 3.7) were observed at 22 and 90 C. The failure occurred
in the composite adherends at low temperature, in the adhesive layer, cohesively, at higher temperature. The possible influence of variable amplitude loading on the fatigue properties of bonded joints is another topic of recent interest for the scientific community. In two recent papers [10,19], Erpolat and co-authors observed the presence of a severe acceleration of the crack growth due to the load interactions under variable amplitude loading conditions with respect to the constant amplitude fatigue. This led numerical crack growth integration [10] and application of linear rules for the fatigue damage summation [19] based on constant amplitude data to provide unreliable life estimations under variable amplitude loading. To conclude this brief discussion, it is also worth noting the detailed analysis presented by De Goeij and coauthors [20], which provided a quite comprehensive overview of the literature on the influence of the main design parameters on static and fatigue behaviour of bonded joints in composite.
2. Joint configuration, materials and test procedures The experimental activity was carried out with the aim to evaluate the static and fatigue behaviour under tension loadings of single lap bonded joints in composite materials and to investigate the mechanics of the damage evolution. Due to the out-of-plane bending induced by the adherend misalignment, the single lap joints do not have indeed a very efficient geometry in terms of load-bearing capability. However, they are frequently present in the practical applications and relatively easy to manufacture; in many cases, due to the restrictions imposed by the structure geometry, they represent also the only possible design solution. The geometry of the joints under investigation is shown in Fig. 1. The joints were manufactured from autoclave-moulded laminates (Seal Texipreg CC206, T300 twill 2 · 2 carbon fibre fabric/ET442 toughened epoxy matrix) and bonded with a two-part epoxy adhesive 9323 B/A by 3M [22]. The laminate lay-up was [0]6
Fig. 1. Geometry of single lap bonded joint (dimensions in millimetres), SE = square edge joints, F = spew fillet joints.
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As mentioned before, one of the aims of the work is the analysis of the damage evolution during the fatigue life of the joints, for a better understanding of the damage mechanics but also to put an experimental basis for the development of a physically based life prediction model. In particular, the analysis was oriented to the assessment of the life spent for the initiation of a fatigue crack and the evaluation of the rates of crack propagation. Some joints were therefore subjected to repeated blocks of fatigue loading at constant amplitude, up to the failure of the joint. At the end of each block, the damage patterns were analysed by means of microscopic observation of the polished edges of the joints to identify the onset of a fatigue crack or to measure the length of the cracks, when present. This procedure was used according to the results presented by Dessureault and Spelt [13] proving that the intermittent cycling between observations has negligible influence on the crack behaviour. Moreover, it was decided to use optical microscopy for the crack detection because, although very time consuming, it is far more accurate than the compliancebased methods which are more sensitive, instead, to the overall response of the joint.
Table 1 Properties of carbon/epoxy laminates [22] and Scotch Weld 9323 B/A adhesive [21] Adherends
EL (MPa) 58,050
ET (MPa) 58,050
GLT (MPa) 3300
mLT 0.06
Adhesive
E (MPa) 2870
G (MPa) 1050
m 0.37
sr (MPa) 39.1
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and the properties of laminates and adhesive are given in Table 1. The joint width (24 mm) and overall length (260 mm) were kept constant as well as the laminate and adhesive layer thickness, 1.65 and 0.15 mm, respectively. Three overlap lengths (20, 30 and 40 mm) and two geometry corners (square edge and spew fillet) were investigated in the test programme. In the preliminary phase of the programme, the influence of the surface preparation was also analysed by comparing the static tensile behaviour of joints made from panels of laminates laid up with or without an external layer of peelply. The joints were produced by bonding the laminate panels and curing them in oven at 65 ± 2 C under a pressure of 0.1 MPa for two hours. The required pressure was obtained by clamping the panels with suitable spring devices. To obtain the appropriate adhesive thickness, about 1% in weight of 150 lm diameter glass spheres were mixed with the adhesive. Prior to bonding, the panels without peel-ply layer were grit blasted and degreased with methyl-ethyl-ketone, the panels with peel-ply only degreased. All the joints were cut from the bonded panels and tested at room temperature on a servo-hydraulic MTS 809 machine with 10/100 kN load cells. The static behaviour of the joints was investigated by means of tensile tests under displacement control with a crosshead speed equal to 2 mm/min. The fatigue tests were carried out under load control, with a sinusoidal wave, nominal load ratio R = 0.05 and a test frequency variable in the range of 10–15 Hz depending on the applied stress level. On the basis of the results obtained during the static tests, only joints produced from laminates with peelply were considered for fatigue testing.
3. Static test results The results of the static tests are presented in terms of both nominal tensile stress on the adherends and shear stress on the adhesive, with the aim to provide information on load-carrying capability of the joints and adhesive properties. Tensile and shear stresses are calculated simply dividing the applied tensile load by the transversal section of one adherend and by the overlap area, respectively. The non-uniformity of the stress fields and, in particular, the stress singularity at the overlap end [1,4], inducing very high stress concentrations, were deliberately neglected at this stage. Tables 2 and 3 summarise the strength data of the joints manufactured from panels without and with peel-ply. The results are, a part from one case, averaged on at least three repeated tests for each condition. During static tests, the final failure was observed to occur more frequently at the end of the overlap as the
Table 2 Tensile static strength of joints without peel-ply Corner geometry
Overlap (mm)
rUTS (MPa)
c.o.v. (%)
sUTS (MPa)
c.o.v. (%)
Failure modea
Square edge
20 40
336.1 392.2
5.3 3.5
28.1 16.4
5.3 4.2
LF LF
Spew fillet
20 30 40
375.2 426.2 491.3
4.1 7.7 1.1
31.0 23.6 20.3
3.7 8.2 0.6
LF LF LF
a
LF, laminate failure (see Fig. 2(a)); A, failure at adherend–adhesive interface (see Fig. 2(b)).
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Table 3 Tensile static strength of joints with peel-ply Corner geometry
Overlap (mm)
rUTS (MPa)
c.o.v. (%)
sUTS (MPa)
c.o.v. (%)
Failure modea
Square edge
20 30 40
347.9 407.9 440.1
4.0 10.8 6.9
29.2 22.6 18.5
4.9 11.5 7.4
A LF LF
Spew fillet
20 30 40
408.9 501.7 551.5
– 2.6 4.2
33.8 27.6 22.7
– 2.9 10.9
A LF LF
a
LF, laminate failure (see Fig. 2(a)); A, failure at adherend–adhesive interface (see Fig. 2(b)).
Fig. 2. Side view of static failures: (a) typical laminate failure, (b) failure in the adhesive layer.
propagation of crack through the laminate, due to the very high stress concentrations. An example of this failure mode, in agreement to the results presented by Cheuck and Tong [23], is given in the side view of the failure zone reported in Fig. 2(a). In few cases, for short overlap joints, the failure was caused by the propagation of a crack at the adhesive–laminate interface, as shown in Fig. 2(b). The typical failure patterns observed for the different joint configurations are listed in Tables 2 and 3. The average results of the static tests are also shown in Fig. 3, for an easier comparison and evaluation of the influence of overlap length, corner geometry and surface preparation on the tensile strength of the joint. According to the results commonly reported in the literature, a general increase in the joint strength can be observed in Fig. 3 as the overlap length increases, to-
Tensile strength [MPa]
600
F PP F GB-D SE PP SE GB-D
500
400
gether with a reduction of the nominal shear stress calculated on the adhesive layer. The increase in the load-carrying capability of the joints with the overlap length can be justified, obviously, on the basis of the greater bonded surface but also considering the results of the stress analysis reported in [1]. A reduced value of the normalised stress intensity factor associated to the local stress field at the critical location in the joint was found, in fact, for joints with longer overlap, at least in the range of dimensions investigated. The reduced values of both intensity and degree of singularity calculated for spew fillet with respect to square edge corners [1,4], can also be of help in explaining the significant influence of the corner geometry and the strong improvement in the strength of the spew fillet joints. As expected and frequently reported in the literature, the surface preparation procedure is another key issue for obtaining good performances from adhesive joints: the results indicate an average increase of more than 10% in the tensile strength of the joints manufactured from panels with an external layer of peel-ply. This can be due to the improved adhesion at the adhesive– composite interface, as indicated by the higher values of the shear strength measured, at failure, for peel-ply joints. The effect seems to be even more evident for the spew fillet joints.
300 15
20
25 30 35 Overlap length [mm]
40
45
Fig. 3. Influence of overlap length, corner geometry and surface preparation on the static strength (PP = peel-ply, GB-D = grit blasted and degreased, error bar = one standard deviation).
4. S–N curves and stress-life approach All the fatigue data were preliminarily analysed by using the classical stress-life approach and drawing the
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Table 4 Results of the statistical analysis for the different series of fatigue data (fatigue strength values at 2 · 106 cycles, r.o. = run out specimens) Corner geometry
Overlap (mm)
Square edge
20 40
Spew fillet
20 30 40
rMAX,50% (MPa)
rMAX,90% (MPa)
K
Tr
sMAX,50% (MPa)
sMAX,90% (MPa)
No. of data
53.9 85.7
46.0 70.0
6.13 5.84
1.372 1.498
4.4 3.6
3.6 2.8
4 12
77.5 106.2 108.5
69.2 98.3 87.6
6.50 7.08 6.07
1.253 1.166 1.534
6.3 5.9 4.4
5.6 5.3 3.5
14 (2 r.o.) 8 13 (1 r.o.)
fatigue curves based on nominal stress and number of cycles to failure, identified as the complete separation of the joints. Again, tensile stresses on the laminates and shear stresses on the adhesive were considered and the fatigue data were expressed in terms of the maximum component of the applied loading cycle. The data were then analysed by assuming log-normal distribution of the number of cycles to failure and the fatigue curve for each series was calculated in the form of N rkmax ¼ cost
or
N skmax ¼ cost.
ð1Þ
Table 4 lists the results of the analysis: the reference stress values at 2 · 106 cycles for a probability of survival of 50% and 90%, the values of the inverse slope k and scatter index Tr (Tr = rMAX,10%/rMAX,90%) are reported. The fatigue data, the fatigue curves and the associated scatter bands are also presented in Figs. 4–6. The stress-life approach allows the influence of the design parameters under investigation to be clearly identified. In fact, an increase in the fatigue strength of the joints and therefore in their load-bearing capability can be observed when the overlap length increases for both the corner geometries. Even the influence of the corner geometry itself can be easily identified in Fig. 6. The improvement in the average fatigue strength is very significant indeed, about 27% for joints with overlap length equal to 40 mm and 44% for joints with 20 mm
Fig. 5. Influence of overlap length on the fatigue strength of spew fillet joints (average fatigue curves and 10–90% probability of survival scatter bands).
overlap, and could be justified, again, with the reduced values of both intensity and degree of singularity associated to the spew fillet corner geometry when compared to the square edge [1]. It is also interesting to note, a part from the series spew fillet with 30 mm overlap, the limited variation of the inverse slopes for the different series. This seems to indicate a shift of the fatigue curves as a function mainly of overlap length and corner geometry. However, with the adoption the stress-life approach difficulties may arise when attempting to develop a prediction tool of general validity, suitable to be applied also to joints of different type. In fact, even calculating the fatigue strength in terms of shear stress on the adhesive layer, the nominal stress approach does not allow the fatigue data to be summarised in a single scatter band. Moreover, no information about the damage evolution can be obtained and therefore it is impossible to develop a life prediction model suitable to incorporate, at least in part, the mechanics of the fatigue damage evolution.
5. Crack nucleation under fatigue loading Fig. 4. Influence of the overlap length on the fatigue strength of square edge joints (average fatigue curves and 10–90% probability of survival scatter bands).
The visual and microscopic observation of the joints during the block loading fatigue tests indicated in the central zone of the bondline at the end of the overlap
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Fig. 6. Influence of corner geometry on the fatigue strength of joints with 40 mm overlap.
edge the more frequent site of crack initiation. An example is presented in Fig. 7(a), where the just-nucleated cracks are evidenced by the whitening and crazing of the bondline. After the nucleation, the cracks propagated very quickly toward the edges of the joints. In other cases, the cracks nucleated in one of the four corners of the overlap. It was therefore decided to identify the life to crack initiation as the number of cycles in correspondence of which a crack could be practically observed on the polished edges of the joint. A crack length of 0.3 mm was arbitrarily fixed as reference threshold for the crack initiation. It is worth noting, however, that small variations in the initial crack length were found to have negligible influence on the estimated fatigue life, when using a life prediction model suitable to simulate the crack propagation [2]. In spite of the different geometry of corner and bondline at the end of the overlap, the crack nucleation behaviour for square edge and spew fillet joints was found to be quite similar, with crack onset in the centre of the bondline and fast propagation toward the edges.
Figs. 7(b) and 8(a) show the crack onset as it appears from the side view of the edges of both types of joints. Fig. 7(b) shows also the small radius present sometimes in the bondline of square edge joints due to the bonding procedure adopted. In few cases only, the cracks were found to nucleate also in the upper zone of the fillet, as illustrated in Fig. 8(b). The life to crack initiation (Ni) measured during the block loading tests are listed in Tables 5 and 6 together with the fatigue life to complete failure (Nf) and the cycles spent for propagation (Np). As can be seen from the tables, the fraction of life for crack initiation represents a significant part of the joint fatigue life, ranging from a minimum of about 20% up to more than 70%. Similar results are available in the literature for bonded joints made from metallic or composite adherends [6,13,14,17,24,25]. The relative fraction of life spent in the crack initiation is also reported to depend on several parameters like joint materials, overall and corner geometry, load level and presence of defects in the bondline. These results clearly indicate that neglecting the nucleation phase and considering the fatigue life of the joint entirely spent in crack propagation, as done by some of the methodologies available for the life prediction of bonded joints [26–29], could result in too conservative assessment of the fatigue life of the joints. A clear influence of the overlap length was observed: in joints with shorter overlap a greater fraction of life was spent for nucleating a crack, according to the results presented by Ishii et al. [6]. The better fatigue performances described before for joints with longer overlap could be therefore justified by the longer duration of the propagation phase. The corner geometry was found to have a small influence, with the spew fillet joints characterised by values of the relative life to crack initiation only slightly higher with respect to those of the square edge joints. However, the behaviour at the different stress levels is completely different. For the square edge
Fig. 7. (a) Crack nucleation site in the bondline of a spew fillet joint. (b) Side view of the crack nucleation in a square edge joint.
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Fig. 8. Crack nucleation sites and crack paths in spew fillet joints.
Table 5 Results of block loading fatigue tests for square edge joints Corner geometry
Overlap (mm)
rmax (MPa)
Ni cycles
Nf cycles
Np cycles
Square edge
20 20 20 20
110 110 60 60
9000 9000 540,000 700,000
19,013 33,611 1,053,991 1,020,556
10,013 24,611 513,991 320,556
0.47 0.27 0.51 0.69
40 40 40 40 40 40 40
180 160 160 160 80 80 80
3000 22,500 6000 8600 1,730,000 520,000 560,000
13,317 48,100 32,146 30,805 3,921,000 2,572,021 1,352,934
10,317 25,600 26,146 22,205 2,191,000 2,052,021 792,934
0.23 0.47 0.19 0.28 0.44 0.20 0.41
Ni/Nf
Ni, cycles to crack initiation; Nf, cycles to failure; Np, cycles for crack propagation.
Table 6 Results of block loading fatigue tests for spew fillet joints Corner geometry
Overlap (mm)
rmax (MPa)
Ni cycles
Nf cycles
Np cycles
Ni/Nf
Spew fillet
20 20 20 20 20 20 20
188 188 168 126 113 113 84
2700 4500 9000 42,000 77,000 60,000 325,000
3906 6580 19,695 106,603 152,000 200,100 1,341,464
1206 2080 10,695 64,603 75,000 140,100 1,016,464
0.69 0.68 0.46 0.39 0.51 0.30 0.24
30 30 30
211 168 126
6800 33,700 205,000
15,228 89,802 828,018
8428 52,802 623,018
0.45 0.41 0.25
40 40 40 40 40 40
253 200 200 168 126 120
4800 48,000 25,500 26,000 475,000 200,000
21,428 64,235 62,730 130,180 2,200,105 612,628
16,628 16,235 37,230 104,180 1,726,105 412,628
0.23 0.75 0.41 0.20 0.21 0.33
Ni, cycles to crack initiation; Nf, cycles to failure; Np, cycles for crack propagation.
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joints the fraction of life spent in initiation increases at the lower stress levels, as reported also by Crocombe and Richardson for steel bonded joints [25]. In the case of spew fillet joints, instead, this fraction seems to increase as the applied stress increases.
6. Crack propagation under fatigue loading and growth rates After the identification of the crack onset, the crack evolution during the propagation phase was monitored by measuring the length of the cracks running from each of the four corners of the overlap, for the subsequent evaluation of the rate of crack growth. The propagation of the cracks occurred mainly at the adhesive–adherend interface, as indicated in Fig. 8(a). In some cases, in correspondence of resin rich areas or fibre bundles perpendicular to the crack front, deviations toward the inner layers were observed (Fig. 8(b)). The woven architecture of the reinforcement, however, forced the crack propagation to remain and proceed mainly at the adhesive/adherend interface. Usually, two quite symmetric crack fronts were observed to grow uniformly from both sides of the overlap, as indicated by crack lengths plotted in Figs. 9 and 10 against the normalised fatigue life. The corner geometry, the stress level and the overlap length were found to have only negligible influence on the crack front evolution. However, the spew fillet at the end of the overlap induced in some cases, at low stress level, an irregular propagation with non-symmetric and/or non-uniform crack front as shown in Fig. 11. A possible reason for this behaviour, a part from the natural experimental variation, can be the significant change in the severity of the local stress field in the fillet once the crack is nucleated, far greater with respect to that occurring in the square edge joints. This fact leads
Fig. 9. Symmetric crack growth in a square edge joint.
Fig. 10. Symmetric crack growth in a spew fillet joint.
Fig. 11. Non-uniform, non-symmetric crack growth in a spew fillet joint.
the propagation to concentrate more in one location rather than to proceed uniformly. The Crack Growth Rate (CGR) was calculated by using the incremental polynomial method, suggested by the ASTM E 647-00 standard [30], under the assumption of a uniform crack front propagating symmetrically from both sides of the overlap, which is representative of the majority of the experimental observations. Moreover, as discussed in [1], symmetric cracks running from both edges of a single lap joints represent the more severe condition in terms of Strain Energy Release Rate trends. In view of the development of a prediction model this choice should therefore produce predictions on the conservative side. For the length of the crack to be used in the calculations, two alternative possibilities were considered: the crack averaged from the measures taken on four corners of the overlap (average crack) or the crack with the greatest length at failure (dominant crack). The former seems to be more appropriate to describe the general conditions of propagation while the latter define a more conservative scenario. Fig. 12 com-
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pares the propagation rates calculated with the two alternatives for tests on square edge joints with 20 and 40 mm overlap. It can be easily observed that the two alternatives provide, from a practical point of view, the same results and therefore they can be used alternatively. Furthermore, it is demonstrated in [2] that the choice of average or dominant crack length for the calculation of the crack growth rates has a negligible influence in the Paris-like curve describing the propagation behaviour of the joints investigated here and, more important, on life predictions made by using these curves. For an easier and direct comparison, the data in Figs. 12–14 are plotted against the crack length a normalised with respect to the overlap length w. Moreover, the stress levels were selected, for each configuration, such as to produce an estimated life of about 50,000 cycles (referred to as high stress level) and about one million cycles (referred to as low stress level). The trends of the CGR measured on square edge and spew fillet joints of different overlap length during their fatigue life, i.e., for increasing crack length are presented in Figs. 13 and 14. For the square edge joints the data are roughly divided into two groups, as a function of the stress level only. The length of the joint overlap seems therefore not to have any influence. The significant influence of the stress level can be explained considering the increase in the SERR associated to an increase of applied stress level [1]. However, another factor which can further contribute to accelerate the propagation process is the predominant contribution of the Mode I component of the SERR at the higher stress levels [1]. The situation seems to be different for the spew fillet joints. In fact the crack growth rates at high stress level are quite different for joints with different overlap length. This fact, however, is not caused by the influence of the overlap length but to the too high stress level
185
Fig. 13. Crack growth rate for square edge joints (from average crack length).
Fig. 14. Crack growth rate for spew fillet joints (from average crack length).
inadvertently chosen for the joints with 20 mm overlap, which results in very short fatigue lives when compared to those of the joints with longer overlap (see Table 6). An increase in the delamination growth rate with the increase of applied stress levels was observed also during tension–tension fatigue testing of co-cured stepped lap joints [31].
7. Conclusions
Fig. 12. Comparison of the crack growth rates calculated from the average or dominant crack length.
The behaviour of single lap bonded joints in composite material under static and fatigue loading was investigated. The influence of design parameters like surface preparation, overlap length and geometry corner was analysed and the evolution of fatigue damage was carefully monitored by visual and microscopic inspection.
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The main results can be summarised as follows: • The static and fatigue strength, in terms of nominal tensile stress on the adherends, were found to increase with the overlap length. • A beneficial contribution was provided also by the presence of a spew fillet at the end of the overlap length. The improvement in fatigue strength with respect to square edge joints was greater than 25%. • The analysis of the fatigue damage evolution indicated the presence of a nucleation phase and the subsequent propagation, at the adhesive–adherend interface, of a crack front up to the joint failure. • The fraction of fatigue life spent in the crack initiation phase was in the range from 20% to more than 70% depending, mainly, on overlap length and stress level. This fact confirms the need of developing life prediction models suitable to incorporate and describe the crack nucleation phase. • The crack growth rates increased considerably with the applied stress level, whereas the length of the overlap seems not having any significant influence. The fatigue data presented here (S–N curves, life to crack initiation and crack growth rates), together with the stress intensity factors and the strain energy release rates trends calculated in [1], represent the basis for the development and validation of a life prediction methodology for bonded joints in composite materials presented in [2], incorporating the actual mechanics of the damage evolution.
Acknowledgement The work was carried out in the frame of the project ‘‘Fatigue design methodologies of joints in composite material’’ (2001092449) financially supported by the Italian Ministry of Research and University.
References [1] Quaresimin M, Ricotta M. Stress intensity factors and strain energy release rates in single lap bonded joints in composite materials. Compos Sci Technol [accepted for publication]. [2] Quaresimin M, Ricotta M. Life prediction of bonded joints in composite materials. Int J Fatigue [submitted]. [3] Lefebvre DR, Dillard DA. A stress singularity approach for the prediction of fatigue crack initiation in adhesive bonds. Part 1: theory. J Adhesion 1999;70:119–38. [4] Lazzarin P, Quaresimin M, Ferro P. A two-term stress function approach to evaluate stress distributions in bonded joints of different geometries. J Strain Anal Eng Des 2002;37:385–98. [5] Renton WJ, Vinson JR. Fatigue behavior of bonded joints in composite material structures. J Aircraft 1975;12:442–7.
[6] Ishii K, Imanaka M, Nakayama H, Kodama H. Evaluation of the fatigue strength of adhesively bonded CFRP/metal single and single-step double-lap joints. Compos Sci Technol 1999;59:1675–83. [7] Ferreira JAM, Reis PN, Costa JDM, Richardson MOW. Fatigue behaviour of composite adhesive lap joints. Compos Sci Technol 2002;62:1373–9. [8] Zeng Q, Sun CT. Fatigue performance of a bonded wavy composite lap joint. Fatigue Fract Eng Mater Struct 2004;27:413–22. [9] Mall S, Ramamurthy G, Rezaizdeh MA. Stress ratio effect on cyclic debonding in adhesively bonded composite joints. Compos Struct 1987;8:31–45. [10] Erpolat S, Ashcroft IA, Crocombe AD, Abdel Wahab MM. Fatigue crack growth acceleration due to intermittent overstressing in adhesively bonded CFRP joints. Composites A 2004;35:1175–83. [11] Mall S, Kochhar NK. Characterization of debond growth mechanism in adhesively bonded composites under mode II static and fatigue loadings. Eng Fract Mech 1988;31:747–58. [12] Mall S, Ramamurthy G. Effect of bond thickness on fracture and fatigue strength of adhesively bonded composite joints. Int J Adhes Adhes 1989;9:33–7. [13] Dessureault M, Spelt JK. Observations of fatigue crack initiation and propagation in an epoxy adhesive. Int J Adhes Adhes 1997;17:183–95. [14] Potter KD, Guild FJ, Harvey HJ, Wisnom MR, Adams RD. Understanding and control of adhesive crack propagation in bonded joints between carbon fibre composite adherends I. Experimental. Int J Adhes Adhes 2001;21:435–43. [15] Krause AR, Holubka JW, Chun W. The effect of adhesive chemistry on the fatigue resistance of adhesive bonds in service environment. In: Proceedings of the 2nd conference on advanced composites: the latest developments, 18–20 November, Dearborn (MI); 1986. p. 193–201. [16] Gilmore RB, Shaw SJ. The effect of temperature and humidity on the fatigue behaviour of composite bonded joints. ASTM STP 1227 – composite bonding, vol. 82–95. Philadelphia: American Society for Testing and Materials; 1994. [17] Ashcroft IA, Abdel Wahab MM, Crocombe AD, Hughes DJ, Shaw SJ. The effect of environment on the fatigue of bonded composite joints. Part 1: testing and fractography. Composites A 2001;32:45–58. [18] Ashcroft IA, Shaw SJ. Mode I fracture of epoxy bonded composite joints 2. Fatigue loading. Int J Adhes Adhes 2002;22:151–67. [19] Erpolat S, Ashcroft IA, Crocombe AD, Abdel-Wahab MM. A study of adhesively bonded joints subjected to constant and variable amplitude fatigue. Int J Fatigue 2004;26:1189–96. [20] De Goeij WC, van Tooren MJL, Beukers A. Composite adhesive joints under cyclic loading. Mater Des 1999;20:213–21. [21] Quaresimin M. Fatigue of woven composite laminates under tensile and compressive loading. 10th European conference on composite materials, Brugge. ESCM; 2002. [22] 3M. Aerospace Technical Data Sheet Nb13. Scotch-Weld 9323 B/ A; 1998. [23] Cheuk PT, Tong L. Failure of adhesive bonded composite lap shear joints with embedded precrack. Compos Sci Technol 2002;62:1079–95. [24] Zhang Z, Shang JK, Lawrence FV. Backface strain technique for detecting fatigue crack initiation in adhesive joints. J Adhesion 1995;49:23–36. [25] Crocombe AD, Richardson G. Assessing stress state and mean load effects on the fatigue response of adhesively bonded joints. Int J Adhes Adhes 1999;19:19–27. [26] Kinloch AJ, Osiyemi SO. Predicting the fatigue life of adhesivelybonded joints. J Adhesion 1993;43:79–90.
M. Quaresimin, M. Ricotta / Composites Science and Technology 66 (2006) 176–187 [27] Curley AJ, Hadavinia H, Kinloch AJ, Taylor AC. Predicting the service-life of adhesively-bonded joints. Int J Fracture 2000;103:41–69. [28] Abdel Wahab MM, Ashcroft IA, Crocombe AD, Smith PA. Numerical prediction of fatigue crack propagation lifetime in adhesively bonded structures. Int J Fatigue 2002;24:705–9. [29] Hadavinia H, Kinloch AJ, Little MSG, Taylor AC. The prediction of crackgrowth in bonded joints under cyclic-fatigue loading
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II. Analytical and finite element studies. Int J Adhes Adhes 2003;23:463–71. [30] ASTM E 647-00. Standard test method for measurement of fatigue crack growth rates. American Society for Testing and Materials; 2000. [31] Quaresimin M, Ricotta M. Delamination growth analysis of cocured joints in composite materials. Advances in experimental mechanics – XII ICEM Proceedings (Bari, Italy). New York: McGraw-Hill; 2004.
Fatigue behaviour and damage evolution of single lap bonded joints in composite material M. Quaresimin *, M. Ricotta Department of Management and Engineering, University of Padova, Stradella San Nicola 3, 36100 Vicenza, Italy Received 7 March 2005; accepted 12 April 2005 Available online 27 June 2005
Abstract The paper presents the results of an experimental investigation on the fatigue behaviour of single lap bonded joints. Carbon/ epoxy laminates bonded with epoxy adhesive were tested under tension–tension fatigue loading and the effect of overlap length and corner geometry was discussed by using the classical stress-life approach. Significant improvements in fatigue strength can be obtained by adopting long overlap length and spew fillet corner geometry. A careful analysis of the evolution of the fatigue damage was also carried out and it was observed that a significant fraction of the fatigue life of the joint is spent in the nucleation of one or more cracks which then propagate up to failure of the joint. The duration of nucleation process, which can last from 20% up to 70% of the joint life, suggests the need of incorporating this phase in the development of future predictive models. The fatigue crack propagation process is also discussed and the rates of crack propagation evaluated. 2005 Elsevier Ltd. All rights reserved. Keywords: Fatigue; Damage evolution; Composite material
1. Introduction This paper is the first of a series of three which present an overview on the activity done at University of Padova in the framework of a project for the development of fatigue design methodologies for bonded joints in composite materials. In this first paper a synthesis of the experimental activity on single lap bonded joints in composite material is presented and discussed. The second paper [1] deals with the finite element analyses carried out to investigate the influence of the main design parameters on the stress distributions in the joints and to provide suitable tools for the definition of the life prediction methodology, which is presented in the third paper [2]. *
Corresponding author. Tel.: +39 0444 998723; fax: +39 0444 998888. E-mail address: [email protected] (M. Quaresimin). 0266-3538/$ - see front matter 2005 Elsevier Ltd. All rights reserved. doi:10.1016/j.compscitech.2005.04.026
As it will be better discussed and explained later, the fatigue life of the bonded joints tested here consists of a first phase spent in the nucleation of a macroscopic crack and the subsequent evolution of this crack which propagates at the adhesive–adherend interface up to a critical length corresponding to the joint failure. The fraction of the fatigue life spent in the first phase was found to be always greater than 20%. The life prediction models available in the literature usually neglect this phase and this could lead to strong underestimations of the joint fatigue life. Even though on the conservative side, this way of design is not very efficient from an engineering point of view. To overcome this problem, the model developed during the project and presented in [2] estimates the fatigue life of a joint taking into account the actual evolution of the fatigue damage: the first phase of the fatigue life, required to nucleate a crack, is calculated by using a non-conventional extension of the Stress Intensity Factor approach, as sug-
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gested in [3,4], whereas the subsequent propagation phase is estimated through the integration of a Paris-like curve relating the crack growth rate da/dN to the range of the Strain Energy Release Rate (SERR), as commonly done in the literature. In this paper, after a brief survey of some of the experimental analyses available in the literature, the results obtained on single lap joints made by bonding carbon/epoxy laminates with an epoxy adhesive are presented, with particular reference to the description of the joint fatigue behaviour and the accurate analysis of the damage evolution under fatigue loading. The investigations published in the literature on the fatigue behaviour of bonded joints can be divided into two main groups: those oriented to evaluate the loadcarrying capability of the joints, seen as load-bearing parts of a structure, and those more interested in analysing the evolution of the damage and measuring the rate of crack propagation. In the former case the results are presented, according to the stress-life approach, by using the S-N curves from nominal stress and number of cycles to complete failure. In the latter case, the crack propagation rate is usually related, in a Paris-like curve, to an energetic or fracture mechanics parameter like the Strain Energy Release Rate (G), the J-integral (J) or the Stress Intensity Factor (K or SIF). Since both approaches can provide useful and synergical design indications, papers and results referred to both approaches are briefly discussed, with particular attention to those illustrating the influence of design parameters and describing the mechanics of the fatigue damage evolution. The first experimental investigations date back to the early Õ70s. Renton and Vinson [5] found an increase in the static tensile strength of single lap joints with the overlap length, while joints with shorter overlap exhibited higher normalised fatigue strength values, independently of the ply orientation at the interface. However, the absolute values of fatigue strength were higher for joints with longer overlap. Ishii and co-authors [6] noticed the same trend, investigating the fatigue behaviour of CFRP/aluminium single and double lap joints. The length of the overlap was also found to have influence on the damage evolution: for short overlap (10 mm) the fatigue life was spent almost entirely for nucleating a fatigue crack, whereas for long overlap two distinct phases, crack onset and crack propagation, were observed. The beneficial contribution of longer overlap length is also reported in [7]. The fatigue curves are presented here in terms of shear stress on the adhesive and therefore need to be rearranged in terms of nominal tensile stress on the adhered to allow a comparison to be made with the previous data. Zeng and Sun [8] proved that the static properties of a single lap bonded joint can be almost doubled by an im-
177
proved shape of the overlap zone. The wavy design adopted for the joints leads also to significant improvements in the fatigue strength. Mall et al. [9], investigating the influence of the load ratio R on the crack propagation rate in double cantilever beam (DCB) and cracked lap shear (CLS) joints, found an apparent improvement in the fatigue strength at the higher values of the load ratio, with lower propagation rate for a given values of the maximum Strain Energy Release Rate, Gmax. The improvement disappeared and the data collapsed in one single curve, when the crack propagation data obtained under different positive stress ratio were plotted against the SERR range DG rather than against Gmax. The same situation was reported in [10], where the rate of crack growth was found to be almost insensitive to the load ratio, at least in the range R = 0.1–0.5, when using DG. On the other hand, a drastic reduction in the crack growth resistance under fully reversed loading (R = 1) with respect to the case of no reversal (R = 0.1) was reported for Mode II fatigue loading of composite ENF joints bonded with a toughened epoxy adhesive [11]. The study also indicated in the total Strain Energy Release Rate the driving parameter for the crack propagation. In another work on the same adhesive [12], Mall and Ramamurthy investigated the effect of the adhesive layer thickness on the Mode I propagation: DCB joints with different adhesive thickness exhibited about the same threshold at lower crack growth rates, whereas a thicker adhesive layer resulted in an improved resistance to the crack growth in the high propagation rate region. A careful description of the evolution of the fatigue damage in bonded joint was presented by Dessureault and Spelt [13]. Although on aluminium-epoxy bonded joints, the paper is worth to be discussed as a bright example of illustration of the fatigue damage mechanics. Studying the behaviour of DCB and CLS under different mode loading (I, II and mixed-mode) and constant load ratio (R = 0.1), the authors correlated the life to crack initiation and the crack propagation rate to the Strain Energy Release Rate. The analysis of the damage evolution indicated a high variability in the sites of crack nucleation as well as a significant part of the fatigue life spent in the crack nucleation. This fraction of life was found to increase as the maximum applied SERR decreased; the fatigue threshold Gth, derived as the asymptotic values of SERR at which no cracks initiated before 107 cycles, resulted almost unaffected by the mode-ratio. When plotted against the Gmax, the crack propagation rates were similar under Mode I and mixed-mode loading whereas under pure Mode II conditions the propagation rates were significantly lower. However, the fatigue threshold Gth estimated from the Paris-like curves at very low propagation rates was in good agreement with that obtained from the crack initiation analysis.
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An interesting method to control the fatigue damage evolution and the crack paths, improving therefore the prediction capability, was presented by Potter et al. [14]. By introducing polymeric films in a suitable position within the adhesive layer, the authors induced the cracks to propagate within the adhesive layer rather than into the composite adherend, improving also the fatigue performances of the joints. An extensive analysis of the damage mechanisms is also presented. In composite bonded joints, temperature, humidity and water absorption can have a significant influence on both adhesive and adherend properties and therefore the possible effects of the environment on fatigue behaviour and damage evolution were extensively investigated. Just to mention some significant examples, Krause et al. [15] found a greater sensitivity to the immersion in water at 40 C for single lap joints bonded with epoxy adhesives compared to the same joints bonded with urethane adhesives. They also noted that only the fatigue strength was affected by the water absorption whereas no appreciable loss in the static strength was observed. Gilmore and Shaw [16] studied the influence of temperature and humidity on the fatigue damage evolution. Both hot (90 C) and cold (50 C) environment accelerated the nucleation of a fatigue crack, whereas the humidity had significant influence only at higher temperature, inducing the adhesive to fail cohesively. Ashcroft et al. [17] tested lap strap joints bonded with modified epoxy adhesive under tension–tension fatigue loading at different temperatures (50, 22, 90 C) and humidity degrees (ambient humidity and 95% R.H.). Similar values of the threshold load were observed for the ‘‘ dry’’ joints tested at the different temperature. On the other hand moisture-conditioned joints tested at 90 C or ‘‘dry’’ joints tested at 90 C in wet environment exhibited a significant reduction in the threshold load for damage initiation. The authors explained this behaviour indicating a reduction in the fatigue resistance for test conditions near the glass transition temperature Tg of the adhesive and noted also a significant decrease in the adhesive Tg due to the moisture absorption. Sites and mode of fatigue crack initiation as well as damage patterns, duration of the propagation phase and failure modes were found also to be heavily influenced by the environmental conditions. Ashcroft and Shaw [18] studied the influence of the temperature on the Mode I fatigue crack propagation of DCB bonded joints produced with the same modified epoxy adhesive. They found the threshold value for SERR, Gth, to increase from 80 J/m2 at 50 C to 130 J/m2 at 90 C. At low temperature the propagation rate was very high (slope of the Paris-like curve m = 8.8), whereas lower and similar values (m = 3.4– 3.7) were observed at 22 and 90 C. The failure occurred
in the composite adherends at low temperature, in the adhesive layer, cohesively, at higher temperature. The possible influence of variable amplitude loading on the fatigue properties of bonded joints is another topic of recent interest for the scientific community. In two recent papers [10,19], Erpolat and co-authors observed the presence of a severe acceleration of the crack growth due to the load interactions under variable amplitude loading conditions with respect to the constant amplitude fatigue. This led numerical crack growth integration [10] and application of linear rules for the fatigue damage summation [19] based on constant amplitude data to provide unreliable life estimations under variable amplitude loading. To conclude this brief discussion, it is also worth noting the detailed analysis presented by De Goeij and coauthors [20], which provided a quite comprehensive overview of the literature on the influence of the main design parameters on static and fatigue behaviour of bonded joints in composite.
2. Joint configuration, materials and test procedures The experimental activity was carried out with the aim to evaluate the static and fatigue behaviour under tension loadings of single lap bonded joints in composite materials and to investigate the mechanics of the damage evolution. Due to the out-of-plane bending induced by the adherend misalignment, the single lap joints do not have indeed a very efficient geometry in terms of load-bearing capability. However, they are frequently present in the practical applications and relatively easy to manufacture; in many cases, due to the restrictions imposed by the structure geometry, they represent also the only possible design solution. The geometry of the joints under investigation is shown in Fig. 1. The joints were manufactured from autoclave-moulded laminates (Seal Texipreg CC206, T300 twill 2 · 2 carbon fibre fabric/ET442 toughened epoxy matrix) and bonded with a two-part epoxy adhesive 9323 B/A by 3M [22]. The laminate lay-up was [0]6
Fig. 1. Geometry of single lap bonded joint (dimensions in millimetres), SE = square edge joints, F = spew fillet joints.
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As mentioned before, one of the aims of the work is the analysis of the damage evolution during the fatigue life of the joints, for a better understanding of the damage mechanics but also to put an experimental basis for the development of a physically based life prediction model. In particular, the analysis was oriented to the assessment of the life spent for the initiation of a fatigue crack and the evaluation of the rates of crack propagation. Some joints were therefore subjected to repeated blocks of fatigue loading at constant amplitude, up to the failure of the joint. At the end of each block, the damage patterns were analysed by means of microscopic observation of the polished edges of the joints to identify the onset of a fatigue crack or to measure the length of the cracks, when present. This procedure was used according to the results presented by Dessureault and Spelt [13] proving that the intermittent cycling between observations has negligible influence on the crack behaviour. Moreover, it was decided to use optical microscopy for the crack detection because, although very time consuming, it is far more accurate than the compliancebased methods which are more sensitive, instead, to the overall response of the joint.
Table 1 Properties of carbon/epoxy laminates [22] and Scotch Weld 9323 B/A adhesive [21] Adherends
EL (MPa) 58,050
ET (MPa) 58,050
GLT (MPa) 3300
mLT 0.06
Adhesive
E (MPa) 2870
G (MPa) 1050
m 0.37
sr (MPa) 39.1
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and the properties of laminates and adhesive are given in Table 1. The joint width (24 mm) and overall length (260 mm) were kept constant as well as the laminate and adhesive layer thickness, 1.65 and 0.15 mm, respectively. Three overlap lengths (20, 30 and 40 mm) and two geometry corners (square edge and spew fillet) were investigated in the test programme. In the preliminary phase of the programme, the influence of the surface preparation was also analysed by comparing the static tensile behaviour of joints made from panels of laminates laid up with or without an external layer of peelply. The joints were produced by bonding the laminate panels and curing them in oven at 65 ± 2 C under a pressure of 0.1 MPa for two hours. The required pressure was obtained by clamping the panels with suitable spring devices. To obtain the appropriate adhesive thickness, about 1% in weight of 150 lm diameter glass spheres were mixed with the adhesive. Prior to bonding, the panels without peel-ply layer were grit blasted and degreased with methyl-ethyl-ketone, the panels with peel-ply only degreased. All the joints were cut from the bonded panels and tested at room temperature on a servo-hydraulic MTS 809 machine with 10/100 kN load cells. The static behaviour of the joints was investigated by means of tensile tests under displacement control with a crosshead speed equal to 2 mm/min. The fatigue tests were carried out under load control, with a sinusoidal wave, nominal load ratio R = 0.05 and a test frequency variable in the range of 10–15 Hz depending on the applied stress level. On the basis of the results obtained during the static tests, only joints produced from laminates with peelply were considered for fatigue testing.
3. Static test results The results of the static tests are presented in terms of both nominal tensile stress on the adherends and shear stress on the adhesive, with the aim to provide information on load-carrying capability of the joints and adhesive properties. Tensile and shear stresses are calculated simply dividing the applied tensile load by the transversal section of one adherend and by the overlap area, respectively. The non-uniformity of the stress fields and, in particular, the stress singularity at the overlap end [1,4], inducing very high stress concentrations, were deliberately neglected at this stage. Tables 2 and 3 summarise the strength data of the joints manufactured from panels without and with peel-ply. The results are, a part from one case, averaged on at least three repeated tests for each condition. During static tests, the final failure was observed to occur more frequently at the end of the overlap as the
Table 2 Tensile static strength of joints without peel-ply Corner geometry
Overlap (mm)
rUTS (MPa)
c.o.v. (%)
sUTS (MPa)
c.o.v. (%)
Failure modea
Square edge
20 40
336.1 392.2
5.3 3.5
28.1 16.4
5.3 4.2
LF LF
Spew fillet
20 30 40
375.2 426.2 491.3
4.1 7.7 1.1
31.0 23.6 20.3
3.7 8.2 0.6
LF LF LF
a
LF, laminate failure (see Fig. 2(a)); A, failure at adherend–adhesive interface (see Fig. 2(b)).
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Table 3 Tensile static strength of joints with peel-ply Corner geometry
Overlap (mm)
rUTS (MPa)
c.o.v. (%)
sUTS (MPa)
c.o.v. (%)
Failure modea
Square edge
20 30 40
347.9 407.9 440.1
4.0 10.8 6.9
29.2 22.6 18.5
4.9 11.5 7.4
A LF LF
Spew fillet
20 30 40
408.9 501.7 551.5
– 2.6 4.2
33.8 27.6 22.7
– 2.9 10.9
A LF LF
a
LF, laminate failure (see Fig. 2(a)); A, failure at adherend–adhesive interface (see Fig. 2(b)).
Fig. 2. Side view of static failures: (a) typical laminate failure, (b) failure in the adhesive layer.
propagation of crack through the laminate, due to the very high stress concentrations. An example of this failure mode, in agreement to the results presented by Cheuck and Tong [23], is given in the side view of the failure zone reported in Fig. 2(a). In few cases, for short overlap joints, the failure was caused by the propagation of a crack at the adhesive–laminate interface, as shown in Fig. 2(b). The typical failure patterns observed for the different joint configurations are listed in Tables 2 and 3. The average results of the static tests are also shown in Fig. 3, for an easier comparison and evaluation of the influence of overlap length, corner geometry and surface preparation on the tensile strength of the joint. According to the results commonly reported in the literature, a general increase in the joint strength can be observed in Fig. 3 as the overlap length increases, to-
Tensile strength [MPa]
600
F PP F GB-D SE PP SE GB-D
500
400
gether with a reduction of the nominal shear stress calculated on the adhesive layer. The increase in the load-carrying capability of the joints with the overlap length can be justified, obviously, on the basis of the greater bonded surface but also considering the results of the stress analysis reported in [1]. A reduced value of the normalised stress intensity factor associated to the local stress field at the critical location in the joint was found, in fact, for joints with longer overlap, at least in the range of dimensions investigated. The reduced values of both intensity and degree of singularity calculated for spew fillet with respect to square edge corners [1,4], can also be of help in explaining the significant influence of the corner geometry and the strong improvement in the strength of the spew fillet joints. As expected and frequently reported in the literature, the surface preparation procedure is another key issue for obtaining good performances from adhesive joints: the results indicate an average increase of more than 10% in the tensile strength of the joints manufactured from panels with an external layer of peel-ply. This can be due to the improved adhesion at the adhesive– composite interface, as indicated by the higher values of the shear strength measured, at failure, for peel-ply joints. The effect seems to be even more evident for the spew fillet joints.
300 15
20
25 30 35 Overlap length [mm]
40
45
Fig. 3. Influence of overlap length, corner geometry and surface preparation on the static strength (PP = peel-ply, GB-D = grit blasted and degreased, error bar = one standard deviation).
4. S–N curves and stress-life approach All the fatigue data were preliminarily analysed by using the classical stress-life approach and drawing the
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Table 4 Results of the statistical analysis for the different series of fatigue data (fatigue strength values at 2 · 106 cycles, r.o. = run out specimens) Corner geometry
Overlap (mm)
Square edge
20 40
Spew fillet
20 30 40
rMAX,50% (MPa)
rMAX,90% (MPa)
K
Tr
sMAX,50% (MPa)
sMAX,90% (MPa)
No. of data
53.9 85.7
46.0 70.0
6.13 5.84
1.372 1.498
4.4 3.6
3.6 2.8
4 12
77.5 106.2 108.5
69.2 98.3 87.6
6.50 7.08 6.07
1.253 1.166 1.534
6.3 5.9 4.4
5.6 5.3 3.5
14 (2 r.o.) 8 13 (1 r.o.)
fatigue curves based on nominal stress and number of cycles to failure, identified as the complete separation of the joints. Again, tensile stresses on the laminates and shear stresses on the adhesive were considered and the fatigue data were expressed in terms of the maximum component of the applied loading cycle. The data were then analysed by assuming log-normal distribution of the number of cycles to failure and the fatigue curve for each series was calculated in the form of N rkmax ¼ cost
or
N skmax ¼ cost.
ð1Þ
Table 4 lists the results of the analysis: the reference stress values at 2 · 106 cycles for a probability of survival of 50% and 90%, the values of the inverse slope k and scatter index Tr (Tr = rMAX,10%/rMAX,90%) are reported. The fatigue data, the fatigue curves and the associated scatter bands are also presented in Figs. 4–6. The stress-life approach allows the influence of the design parameters under investigation to be clearly identified. In fact, an increase in the fatigue strength of the joints and therefore in their load-bearing capability can be observed when the overlap length increases for both the corner geometries. Even the influence of the corner geometry itself can be easily identified in Fig. 6. The improvement in the average fatigue strength is very significant indeed, about 27% for joints with overlap length equal to 40 mm and 44% for joints with 20 mm
Fig. 5. Influence of overlap length on the fatigue strength of spew fillet joints (average fatigue curves and 10–90% probability of survival scatter bands).
overlap, and could be justified, again, with the reduced values of both intensity and degree of singularity associated to the spew fillet corner geometry when compared to the square edge [1]. It is also interesting to note, a part from the series spew fillet with 30 mm overlap, the limited variation of the inverse slopes for the different series. This seems to indicate a shift of the fatigue curves as a function mainly of overlap length and corner geometry. However, with the adoption the stress-life approach difficulties may arise when attempting to develop a prediction tool of general validity, suitable to be applied also to joints of different type. In fact, even calculating the fatigue strength in terms of shear stress on the adhesive layer, the nominal stress approach does not allow the fatigue data to be summarised in a single scatter band. Moreover, no information about the damage evolution can be obtained and therefore it is impossible to develop a life prediction model suitable to incorporate, at least in part, the mechanics of the fatigue damage evolution.
5. Crack nucleation under fatigue loading Fig. 4. Influence of the overlap length on the fatigue strength of square edge joints (average fatigue curves and 10–90% probability of survival scatter bands).
The visual and microscopic observation of the joints during the block loading fatigue tests indicated in the central zone of the bondline at the end of the overlap
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Fig. 6. Influence of corner geometry on the fatigue strength of joints with 40 mm overlap.
edge the more frequent site of crack initiation. An example is presented in Fig. 7(a), where the just-nucleated cracks are evidenced by the whitening and crazing of the bondline. After the nucleation, the cracks propagated very quickly toward the edges of the joints. In other cases, the cracks nucleated in one of the four corners of the overlap. It was therefore decided to identify the life to crack initiation as the number of cycles in correspondence of which a crack could be practically observed on the polished edges of the joint. A crack length of 0.3 mm was arbitrarily fixed as reference threshold for the crack initiation. It is worth noting, however, that small variations in the initial crack length were found to have negligible influence on the estimated fatigue life, when using a life prediction model suitable to simulate the crack propagation [2]. In spite of the different geometry of corner and bondline at the end of the overlap, the crack nucleation behaviour for square edge and spew fillet joints was found to be quite similar, with crack onset in the centre of the bondline and fast propagation toward the edges.
Figs. 7(b) and 8(a) show the crack onset as it appears from the side view of the edges of both types of joints. Fig. 7(b) shows also the small radius present sometimes in the bondline of square edge joints due to the bonding procedure adopted. In few cases only, the cracks were found to nucleate also in the upper zone of the fillet, as illustrated in Fig. 8(b). The life to crack initiation (Ni) measured during the block loading tests are listed in Tables 5 and 6 together with the fatigue life to complete failure (Nf) and the cycles spent for propagation (Np). As can be seen from the tables, the fraction of life for crack initiation represents a significant part of the joint fatigue life, ranging from a minimum of about 20% up to more than 70%. Similar results are available in the literature for bonded joints made from metallic or composite adherends [6,13,14,17,24,25]. The relative fraction of life spent in the crack initiation is also reported to depend on several parameters like joint materials, overall and corner geometry, load level and presence of defects in the bondline. These results clearly indicate that neglecting the nucleation phase and considering the fatigue life of the joint entirely spent in crack propagation, as done by some of the methodologies available for the life prediction of bonded joints [26–29], could result in too conservative assessment of the fatigue life of the joints. A clear influence of the overlap length was observed: in joints with shorter overlap a greater fraction of life was spent for nucleating a crack, according to the results presented by Ishii et al. [6]. The better fatigue performances described before for joints with longer overlap could be therefore justified by the longer duration of the propagation phase. The corner geometry was found to have a small influence, with the spew fillet joints characterised by values of the relative life to crack initiation only slightly higher with respect to those of the square edge joints. However, the behaviour at the different stress levels is completely different. For the square edge
Fig. 7. (a) Crack nucleation site in the bondline of a spew fillet joint. (b) Side view of the crack nucleation in a square edge joint.
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Fig. 8. Crack nucleation sites and crack paths in spew fillet joints.
Table 5 Results of block loading fatigue tests for square edge joints Corner geometry
Overlap (mm)
rmax (MPa)
Ni cycles
Nf cycles
Np cycles
Square edge
20 20 20 20
110 110 60 60
9000 9000 540,000 700,000
19,013 33,611 1,053,991 1,020,556
10,013 24,611 513,991 320,556
0.47 0.27 0.51 0.69
40 40 40 40 40 40 40
180 160 160 160 80 80 80
3000 22,500 6000 8600 1,730,000 520,000 560,000
13,317 48,100 32,146 30,805 3,921,000 2,572,021 1,352,934
10,317 25,600 26,146 22,205 2,191,000 2,052,021 792,934
0.23 0.47 0.19 0.28 0.44 0.20 0.41
Ni/Nf
Ni, cycles to crack initiation; Nf, cycles to failure; Np, cycles for crack propagation.
Table 6 Results of block loading fatigue tests for spew fillet joints Corner geometry
Overlap (mm)
rmax (MPa)
Ni cycles
Nf cycles
Np cycles
Ni/Nf
Spew fillet
20 20 20 20 20 20 20
188 188 168 126 113 113 84
2700 4500 9000 42,000 77,000 60,000 325,000
3906 6580 19,695 106,603 152,000 200,100 1,341,464
1206 2080 10,695 64,603 75,000 140,100 1,016,464
0.69 0.68 0.46 0.39 0.51 0.30 0.24
30 30 30
211 168 126
6800 33,700 205,000
15,228 89,802 828,018
8428 52,802 623,018
0.45 0.41 0.25
40 40 40 40 40 40
253 200 200 168 126 120
4800 48,000 25,500 26,000 475,000 200,000
21,428 64,235 62,730 130,180 2,200,105 612,628
16,628 16,235 37,230 104,180 1,726,105 412,628
0.23 0.75 0.41 0.20 0.21 0.33
Ni, cycles to crack initiation; Nf, cycles to failure; Np, cycles for crack propagation.
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joints the fraction of life spent in initiation increases at the lower stress levels, as reported also by Crocombe and Richardson for steel bonded joints [25]. In the case of spew fillet joints, instead, this fraction seems to increase as the applied stress increases.
6. Crack propagation under fatigue loading and growth rates After the identification of the crack onset, the crack evolution during the propagation phase was monitored by measuring the length of the cracks running from each of the four corners of the overlap, for the subsequent evaluation of the rate of crack growth. The propagation of the cracks occurred mainly at the adhesive–adherend interface, as indicated in Fig. 8(a). In some cases, in correspondence of resin rich areas or fibre bundles perpendicular to the crack front, deviations toward the inner layers were observed (Fig. 8(b)). The woven architecture of the reinforcement, however, forced the crack propagation to remain and proceed mainly at the adhesive/adherend interface. Usually, two quite symmetric crack fronts were observed to grow uniformly from both sides of the overlap, as indicated by crack lengths plotted in Figs. 9 and 10 against the normalised fatigue life. The corner geometry, the stress level and the overlap length were found to have only negligible influence on the crack front evolution. However, the spew fillet at the end of the overlap induced in some cases, at low stress level, an irregular propagation with non-symmetric and/or non-uniform crack front as shown in Fig. 11. A possible reason for this behaviour, a part from the natural experimental variation, can be the significant change in the severity of the local stress field in the fillet once the crack is nucleated, far greater with respect to that occurring in the square edge joints. This fact leads
Fig. 9. Symmetric crack growth in a square edge joint.
Fig. 10. Symmetric crack growth in a spew fillet joint.
Fig. 11. Non-uniform, non-symmetric crack growth in a spew fillet joint.
the propagation to concentrate more in one location rather than to proceed uniformly. The Crack Growth Rate (CGR) was calculated by using the incremental polynomial method, suggested by the ASTM E 647-00 standard [30], under the assumption of a uniform crack front propagating symmetrically from both sides of the overlap, which is representative of the majority of the experimental observations. Moreover, as discussed in [1], symmetric cracks running from both edges of a single lap joints represent the more severe condition in terms of Strain Energy Release Rate trends. In view of the development of a prediction model this choice should therefore produce predictions on the conservative side. For the length of the crack to be used in the calculations, two alternative possibilities were considered: the crack averaged from the measures taken on four corners of the overlap (average crack) or the crack with the greatest length at failure (dominant crack). The former seems to be more appropriate to describe the general conditions of propagation while the latter define a more conservative scenario. Fig. 12 com-
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pares the propagation rates calculated with the two alternatives for tests on square edge joints with 20 and 40 mm overlap. It can be easily observed that the two alternatives provide, from a practical point of view, the same results and therefore they can be used alternatively. Furthermore, it is demonstrated in [2] that the choice of average or dominant crack length for the calculation of the crack growth rates has a negligible influence in the Paris-like curve describing the propagation behaviour of the joints investigated here and, more important, on life predictions made by using these curves. For an easier and direct comparison, the data in Figs. 12–14 are plotted against the crack length a normalised with respect to the overlap length w. Moreover, the stress levels were selected, for each configuration, such as to produce an estimated life of about 50,000 cycles (referred to as high stress level) and about one million cycles (referred to as low stress level). The trends of the CGR measured on square edge and spew fillet joints of different overlap length during their fatigue life, i.e., for increasing crack length are presented in Figs. 13 and 14. For the square edge joints the data are roughly divided into two groups, as a function of the stress level only. The length of the joint overlap seems therefore not to have any influence. The significant influence of the stress level can be explained considering the increase in the SERR associated to an increase of applied stress level [1]. However, another factor which can further contribute to accelerate the propagation process is the predominant contribution of the Mode I component of the SERR at the higher stress levels [1]. The situation seems to be different for the spew fillet joints. In fact the crack growth rates at high stress level are quite different for joints with different overlap length. This fact, however, is not caused by the influence of the overlap length but to the too high stress level
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Fig. 13. Crack growth rate for square edge joints (from average crack length).
Fig. 14. Crack growth rate for spew fillet joints (from average crack length).
inadvertently chosen for the joints with 20 mm overlap, which results in very short fatigue lives when compared to those of the joints with longer overlap (see Table 6). An increase in the delamination growth rate with the increase of applied stress levels was observed also during tension–tension fatigue testing of co-cured stepped lap joints [31].
7. Conclusions
Fig. 12. Comparison of the crack growth rates calculated from the average or dominant crack length.
The behaviour of single lap bonded joints in composite material under static and fatigue loading was investigated. The influence of design parameters like surface preparation, overlap length and geometry corner was analysed and the evolution of fatigue damage was carefully monitored by visual and microscopic inspection.
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The main results can be summarised as follows: • The static and fatigue strength, in terms of nominal tensile stress on the adherends, were found to increase with the overlap length. • A beneficial contribution was provided also by the presence of a spew fillet at the end of the overlap length. The improvement in fatigue strength with respect to square edge joints was greater than 25%. • The analysis of the fatigue damage evolution indicated the presence of a nucleation phase and the subsequent propagation, at the adhesive–adherend interface, of a crack front up to the joint failure. • The fraction of fatigue life spent in the crack initiation phase was in the range from 20% to more than 70% depending, mainly, on overlap length and stress level. This fact confirms the need of developing life prediction models suitable to incorporate and describe the crack nucleation phase. • The crack growth rates increased considerably with the applied stress level, whereas the length of the overlap seems not having any significant influence. The fatigue data presented here (S–N curves, life to crack initiation and crack growth rates), together with the stress intensity factors and the strain energy release rates trends calculated in [1], represent the basis for the development and validation of a life prediction methodology for bonded joints in composite materials presented in [2], incorporating the actual mechanics of the damage evolution.
Acknowledgement The work was carried out in the frame of the project ‘‘Fatigue design methodologies of joints in composite material’’ (2001092449) financially supported by the Italian Ministry of Research and University.
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