Assessing the Effects of Riveting Induced Residual Stresses on Fatigue Crack Behaviour in Lap Joints by means of Fractography

Transcription

1 Assessing the Effects of Riveting Induced Residual Stresses on Fatigue Crack Behaviour in Lap Joints by means of Fractography C.D. Rans a, R.C. Alderliesten b, P.V. Straznicky a a Department of Mechanical and Aerospace Engineering, Carleton University, 25 Colonel By Drive, KS 5B6, Ottawa, Canada b Faculty of Aerospace Engineering, Delft University of Technology, Kluyverweg, 2629 HS, Delft, the Netherlands Abstract The interference fit provided by solid rivets induces a residual stress field beneficial to the fatigue performance of riveted lap joints. Attributing benefits in crack growth nucleation and growth behaviour within this residual stress field is non-trivial. Once fatigue cracks become visible on the surface of a joint they have already grown beyond the region of beneficial residual stress. In order to circumvent this problem, fractographic techniques were employed to evaluate postmortem the influence of the rivet squeeze force and resulting residual stress field on crack behaviour. Results demonstrated that within the range of rivet squeeze forces studied, a 3-fold reduction in crack growth rate is achievable at high rivet squeeze forces, representing a marked improvement in damage tolerance of the joint.

2 2 Introduction The hole filling properties of solid rivets are a key contributor to the fatigue performance of riveted lap joints. As a rivet is installed, it first expands to fill the hole then further expands, cold working the surrounding sheet material and generating a beneficial residual compressive stress field. The nature of this residual stress field plays an important role in the nucleation and growth of cracks in the vicinity of the rivet hole. Studying the effects of riveting induced residual stresses on the fatigue performance of lap joints requires consideration of several factors. The region of beneficial residual stresses is localized around the rivet hole and is obstructed by the rivet head, reducing the effectiveness of conventional crack detection and measurement techniques. The secondary bending stresses resulting from rivet rotation produce highly elliptical crack fronts, causing the crack to remain below the visible sheet surface through most of the region of interest. As a result, overall fatigue lives, crack time-todetection lives, and growth rates for crack tips beyond the region of residual stress are relied on as measures; however, such secondary effects can only provide a limited understanding of the residual stress field. Ideally, knowledge of the impact on nucleation and crack growth within the region of residual stress is required. Fractography provides an alternative means for studying the effects of residual stress on crack nucleation and growth. Postmortem examination of fracture surfaces from failed lap joints, typically using a scanning electron microscope (SEM), can provide direct evidence of crack growth behaviour in a region that is otherwise not observable during the fatigue process. Two methods are often employed in extracting crack growth information in fractographic examinations. The first method examines the spacing between individual fatigue striations on a fatigue surface [2, 6, 9]. By measuring the spacing at a specific point, the crack growth rate at that point can be determined directly. This method provides an accurate and flexible means to determine crack growth rates at specific points on the fracture surface, but does not provide an efficient means for visualizing crack front shapes. The second method involves identifying striation patterns resulting from load variations within the fatigue spectrum [3,5,7,]. These variations may naturally occur in the fatigue spectrum or can be introduced with periodic under- or over-loads known as marker loads. These patterns outline the shape of the crack front giving an indication of shape and location of the crack at a given moment in the spectrum. If these patterns can be associated with a period of time within the spectrum, then the crack growth history can be reconstructed. The present investigation aims to study the crack growth behaviour within the region of residual stress in riveted lap joints. To this end, a series of fatigue tests using a marker load spectrum developed by Piasick and Willard [] were undertaken. Fractography methods were employed to reconstruct crack growth behaviour for these tests allowing the effects of rivet squeeze force and the resulting residual stress state on fatigue performance to be evaluated. The work presented in this paper represents the second part of a larger research effort to better understand the role of rivet installation on fatigue performance in riveted lap joints. The first part of this research effort focused on quantifying the effects of rivet installation on the stress state in riveted lap joints. Using finite element methods, the riveting process was simulated in order to quantify the effect of rivet installation force (squeeze force or F Sq ) on the resulting residual stress field [ 3]. Follow-up finite element studies examined the effects of rivet installation on the stress state in loaded lap joints [,3]. Particular attention was placed on secondary bending stresses due

3 to their role in crack nucleation. Details on these studies can be found in publications by the present authors [ 3]. 3 2 Fatigue Crack Growth Tests Two sets of fatigue tests were carried out in order to investigate the effects of rivet installation on the fatigue performance of riveted lap joints. The first set was undertaken to quantify the effects of rivet squeeze force and rivet type on overall fatigue life. The second set was designated for fractographic crack growth reconstruction. Due to the time consuming nature of such reconstructions, these tests were not run until final fracture. Rather, once cracks of a given size were present, the tests were stopped and the specimen disassembled for fractographic analysis. 2. Test Specimen Geometry Fatigue tests were conducted on 2-row riveted lap joint specimens as illustrated in Figure. Specimens were manufactured with 224-T3 Alclad sheet and 27-T4 rivets. Alclad refers to a thin layer of pure aluminum placed on the sheets for corrosion resistance. Specimens containing countersunk (NAS97AD4-4) and universal (MS2426AD4-4) rivets were made. Specimens manufactured using rivet squeeze forces of 5, 2, 25, and 3 lbf, corresponding to D/D o ratios of.4,.5,.6 and.7 respectively, were studied. These D/D o ratios fall within a range of generally accepted driven rivet head geometries for aerospace applications [8]. The orientation of the rivets in the two rivet rows was reversed giving a true anti-symmetric geometry. This modification, although not practical in a commercial application, is useful for experimental investigations as it ensures the same conditions are present in the two fatigue critical locations. This anti-symmetric configuration results in two specimen types, designated type and type 2 (Figure ). For the type configuration, crack nucleation is assured to occur next to a manufactured rivet head (countersunk or universal rivet head), simulating failure within the outer or exposed sheet in a practical joint. For specimen type 2, failure at the driven rivet head is ensured as would be the case for failure of the inner joint sheet. Use of these two specimen types permits the effects of rivet installation on inner and outer sheet failure to be studied independently. 2.2 Fatigue Test Spectrum In order to facilitate the fractographic reconstructions of crack growth, a specialized test spectrum which would introduce distinctive features on the fracture surface was needed for the current tests. For this purpose, a constant amplitude spectrum that included periodic underload cycles to create band features on the fracture surface was used. This spectrum, illustrated in Figure 2 and used in previous fractography studies [3, 5, ], introduced patterns of, 4, and 6 bands which could easily be identified on the fracture surface. A maximum stress level, S max, of MPa was specified. To prevent damage of the marker bands on the fracture surface, a stress ratio R = S max /S min =. was specified to minimize contact between the opposing fracture surfaces during unloading. All tests were carried out at a frequency of Hz on an MTS 38. load frame equipped with a kn load cell and a TestStar IIs TM controller. Although the marker band spectrum was only required for the set of tests designated for crack growth reconstructions, the spectrum was used in all tests. To account for the difference in crack

4 growth at the 75% S max cycles, a correction factor derived using Elber s plasticity induced crack closure concept [4] was used. This approach was developed by Fawaz [5] and applied in addition studies completed by de Rijck [3]. For the loading case used in this investigation, de Rijck [3] determined that the ratio of the crack growth rate for the 75% S max and S max cycles is 9.22% in 224-T3 and verified this for open-hole fatigue specimens. For the derivation of this factor, the reader is encouraged to refer to the work of de Rijck [3]. Given the 9.22% crack growth correction factor, one program of the marker band fatigue spectrum, which consists of 28 cycles, reduces to 3538 baseline cycles. This correction factor is based on differences in crack growth and does not necessarily account for differences in crack nucleation between the baseline and marker load cycles. All fatigue lives and crack growth rates within this paper will be presented with the applied correction factor unless otherwise noted. 2.3 Test Matrix Two series of fatigue tests were undertaken in this investigation: a series to determine overall fatigue life, and a series for fractographic crack growth reconstruction (designated F and SEM respectively). Furthermore, two specimen configurations (Type and 2 in Figure ) and specimens containing two rivet types were tested. In order to distinguish between the various specimen configurations and test types undertaken, the following specimen naming convention was used: CSK3.-T-F- Specimen number Specimen series (F or SEM) Specimen type (T or T2) F Sq in s lbf Rivet type (U or CSK) Where the specimen series designators F and SEM refer to the overall fatigue life and fractographic reconstruction series respectively. The specimen type designators refer to the type numbers illustrated in Figure and the rivet type designators of U and CSK refer to the rivet types MS2426AD4-4 and NAS97AD4-4 respectively. All specimens were made with mm thick 224-T3 Alclad sheet. All of the tests completed are summarized in the test matrix given in Table. There are a lesser number of tests conducted on Type 2 specimens as these tests were performed only to verify that the outer sheet failure which occurs in the Type specimens is more critical. Additionally, as difficulties were encountered with fractographic reconstructions for specimens containing universal type rivets (MS2426AD4-4), a larger number of tests were conducted on countersunk rivet (NAS97AD4-4) specimens. 2.4 Crack Growth Reconstruction Crack growth reconstructions were performed using a JEOL JSM-64 scanning electron microscope. The reconstruction procedure was as follows. A photographic survey of the fracture surface was taken in the SEM facility at 45x magnification, with a probe voltage of 2 kv and a working distance of 2 mm. The resulting images from the survey were manually assembled to generate a single high resolution image of the fracture surface. From this high resolution composite image, the locations of the advancing crack front could be identified from the pattern of -, 4-, and 6-band 4

5 5 Table : Test matrix for -4-6 marker band fatigue tests F Sq (lb f ) Rivet Type Specimen Type No. F-Specimens No. SEM-Specimens 5 NAS97-AD NAS97-AD NAS97-AD NAS97-AD MS2426-AD MS2426-AD MS2426-AD MS2426-AD NAS97-AD NAS97-AD NAS97-AD NAS97-AD MS2426-AD MS2426-AD MS2426-AD MS2426-AD markers visible on the fracture surface (Figure 3). Due to the smaller striation spacing, blocks containing the lower load cycles (75% S max ) show up as dark lines or featureless bands at the given magnification. Once all the crack fronts were identified, the life of the specimen for each front could be determined by identifying the interface between fatigue and static failure (associated with the final life of the specimen, see Figure 3) and counting backwards in blocks of cycles for each crack front. In some instances, it was difficult to detect particular crack fronts; however, it was easy to determine and account for such missing fronts due to the alternating -4-6 marker band pattern expected. Final crack growth measurements were made relative to the hole edge as observed on the fracture surface. For cases where cracks did not intersect the rivet hole, crack front positions were measured relative to the rivet centre line. Several factors limited the extent to which the crack growth reconstructions could be performed. First, as crack length increases, the morphology of the fracture surface changes. At large crack lengths, the fracture surface becomes dominated by micro voids which impede the detection of the marker bands. Second, the reconstruction process requires many hours to complete. For both these reasons, testing of fatigue specimens destined for crack growth reconstruction were stopped once a crack of approximately 2-3 mm was visible. As each specimen contained 6 rivet locations where cracking could occur, the rivet containing the largest cracks was always selected for crack growth reconstruction. 3 Overall Fatigue Life Results The overall fatigue life results for both specimen configurations (Type I and II) manufactured with the universal and countersunk rivet types are shown in Figure 4. Type I and II specimen configurations correspond to outer sheet and inner sheet failure respectively. As different population sizes

6 were used to determine the mean fatigue life, statistical confidence intervals based on an ANalysis Of VAriance (ANOVA) [4] study are included. Inner sheet failures are undesirable in airframe skin joints as access to the inner surface of the fuselage is necessary during inspections to detect such cracks. For this reason, only a limited number of tests were performed on the Type II specimen configuration with the primary objective of establishing if inner or outer sheet failure would likely occur first for a given rivet squeeze force. Outer sheet failure was favoured for rivet squeeze forces equal to or greater than 2 lbf for both rivet types. At high squeeze forces, the disparity in compressive residual stresses between the inner and outer sheet promote failure in the outer sheet. At lower squeeze forces, this disparity is reduced and the probability of failure in the inner sheet increases. This was observed in the universal rivet specimens where inner sheet failure became critical for the lowest squeeze force examined. At this squeeze force, outer sheet failure remained critical for the countersunk rivet specimens as a result of the added stress concentration provided by the machined countersink. The remaining discussion on fatigue life will refer to outer sheet failure only. For the universal rivet specimens, a small increase in mean fatigue life was observed with increasing squeeze force. Taking into account the scatter observed in the tests (indicated by the 9% confidence interval bars in Figure 4), however, the mean fatigue lives observed for squeeze forces below 3 lbf are not statistically distinct according to the ANOVA analysis. This relative insensitivity of overall fatigue life to rivet squeeze force results from the location of fatigue damage observed in the universal rivet specimens as shown in Figure 5. Fatigue cracks nucleated above the rivet hole and grew around it. This location corresponds to the region of maximum secondary bending stress for universal rivet joints observed in a finite element study by the present authors [, 3] and marked in the Figure 5. This location for crack nucleation and growth is also beyond the region of beneficial residual stress observed in finite element simulations [2, 3] for the majority of rivet squeeze forces examined (indicated by the plastic zone radii in Figure 5). This explains why a statistically distinct increase in fatigue life was only observed once the squeeze force became large enough to induce a beneficial residual stress state at the location of cracking. Conversely, the results for the countersunk rivet specimens show a strong dependency between squeeze force and overall fatigue life. In contrast to the universal rivet specimens, fatigue damage occurs close to the rivet row centre line as shown in Figure 6. Cracks nucleate at the hole edge and continue to grow along the rivet row through the region of residual compressive stress. For higher rivet squeeze forces, the region of beneficial residual stress through which the crack has to grow increases, increasing the overall fatigue life. 6 4 Crack Growth Reconstruction Results Fractographic crack growth reconstructions were completed in order to quantify the effect of rivet installation on the behaviour of fatigue cracks. Based on the overall fatigue life results presented in the previous section and the required hours for each reconstruction, the scope of the crack growth reconstructions was reduced to two squeeze forces: 2 lbf and 3 lbf. Discussion of the results will be divided by rivet type.

7 7 4. Countersunk Rivet Joint Reconstructions The complete set of crack front plots generated for countersunk riveted specimens is shown in Figures 7 and 8 for rivet squeeze forces of 2 and 3 lbf respectively. For the 3 lbf results, only one third of the crack fronts (corresponding to the -band markers) are plotted for clarity. Thus, the number of cycles between plotted crack fronts must be noted when comparing the plots for the two squeeze forces. Several key differences in crack growth behaviour are evident when comparing the crack front plots for the two examined squeeze forces. First, crack nucleation location appears to be influenced by the rivet squeeze force. For the lower squeeze force, crack nucleation occurs at or very near the rivet hole location. At the higher squeeze force, the nucleation location is shifted approximately.5 mm away from the rivet hole. The reason for this shift is a change in the dominate nucleation mechanism. At the lower squeeze force, crack nucleation is dominated by the presence of a notch (rivet hole) and its associated stress concentration. At the higher squeeze force crack nucleation shifts to a region of heavy fretting damage. Within this region, cold-welding followed by disbonding of the asperities of the contacting joint surfaces occurs under cyclic loading, resulting in the formation of surface micro-cracks. Given sufficient bulk stress, these micro-cracks can propagate into the substrate, forming a fatigue crack. Several factors contribute to the shift in crack nucleation location and mechanism. First, at the higher rivet squeeze force, larger residual compressive stresses are generated at the hole edge [2, 3]. These compressive stresses reduce the peak tensile stress at the hole edge, thus reducing its fatigue criticality. Second, the rivet clamping force increases with rivet squeeze force, exacerbating the fretting process. Although fretting and fretting fatigue are beyond the scope of the present investigation, a complementary investigation is being completed by Brown [] on fretting fatigue in riveted lap joints. Preliminary results from Brown have confirmed the crack nucleation location observed at the higher squeeze force correlates with the edge-of-contact between the contacting sheets, where fretting damage and micro-crack densities are expected to be highest. The crack reconstructions also provide an indication of when crack nucleation occurs. Although the exact time of nucleation cannot be established, crack front locations can be associated with specific times in the fatigue life, allowing the period over which nucleation occurred to be narrowed down. Results are shown in Table 2. The difficulty encountered with this method is that it is based on the earliest detectable crack front that can be accounted for. In some instances, additional crack fronts may have been present but too small or obscured to detect, or discontinuities in the fracture surface thwarted efforts to associate a fatigue life with a particular crack front. As a result, although the results in Table 2 seem to indicate that an increased squeeze force delays crack nucleation, this conclusion can not be made. What can be said, however, is that comparing the detection lives in Table 2 to the overall average fatigue lives in Section 3, crack nucleation was found to occur during the first 35% of the average overall fatigue life. An additional difference in crack growth behaviour evident when comparing the crack front plots in Figures 7 and 8 is crack shape. At the higher rivet squeeze force, significant crack growth occurs along the faying joint surface while the crack is still a part-through crack. This results in elongated crack fronts that deviate from the classical quarter elliptical front shapes at the lower squeeze force. The more favourable residual compressive stress field at the higher squeeze force combined with secondary bending in the joint greatly reduces the rate of propagation of the crack through the thickness of the sheet. Crack growth data from the reconstructions is summarized in terms of crack length and crack

8 8 Table 2: Nucleation period obtained from SEM reconstructions Specimen Crack () Specimen Crack () CSK2.-T-SEM- A-left 26.2 B3-left 3.7 CSK3.-T-SEM- A-right 26.2 B3-right 32. CSK2.-T-SEM-2 B3-left 4.6 B2-left 8.2 CSK3.-T-SEM-2 B3-right 4.6 B2-right 8.2 CSK2.-T-SEM-3 B-left 32. A-left 54.2 CSK3.-T-SEM-3 B-right 4.6 A-right 67.7 average growth rate in Figures 9 and. The highly elongated crack front shapes observed at the higher rivet squeeze force made determining the crack length at the free joint surface difficult. As a result, only the faying surface crack length behaviour is shown. These results clearly show the effects of riveting induced residual stresses on fatigue crack growth. Excellent repeatability was observed in the crack growth behaviour amongst specimens manufactured with the same rivet squeeze force. Furthermore, a distinct difference in crack growth rate was observed for the two rivet squeeze forces. A 3-fold reduction in crack growth rate was observed for the higher squeeze force as a result of the more favourable residual stress field induced during rivet installation [ 3]. 4.2 Universal Rivet Joint Reconstructions The universal rivet joint reconstructions were less successful. The reconstruction process relied on being able to identify crack front locations and associate a fatigue life and crack length with these locations. Identifying the crack front locations was not problematic. Associating a fatigue life with each crack front location, however, was hampered by the presence of multiple fatigue cracks and fracture plane intersections. Defining a robust measure for crack length in the presence of multiple fatigue cracks was also difficult. For these reasons, only a limited number of reconstructions were performed for universal rivet joints. The above described difficulties are illustrated in the reconstruction results shown in Figures and 2. Figure best illustrates the difficulty of defining a robust measure for crack length. The main fracture surface was formed by two distinct fatigue cracks (centred at approximately y = - and.7 mm) that eventually linked up to form a larger crack. During this link-up process, the shape and propagation rate of the crack front is dependant on the proximity and number of cracks involved. Simply tracking the location of the faying surface crack front location does not account for these factors, making it difficult to compare such results between different reconstructions. Reconstructions were further hampered by the presence of crack plane intersections. Such discontinuities in the fracture surface prevent the necessary accounting in determining the fatigue life associated with each identified crack front. As an example, in Figure 2, although crack fronts could be identified between y = mm, associated fatigue lives could not be determined for the fronts due to the discontinuities in the fracture surface at y = -3 and 2.5 mm. Despite the limited quantitative data that could be obtained from the universal rivet joint reconstruction, some qualitative observations could be made. The presence of multiple fatigue cracks and fracture plane intersections indicate that fretting fatigue and the formation of multiple microcracks is a dominant crack nucleation mechanism for the universal joints. Although fretting was

9 also observed as a nucleation mechanism for the countersunk rivet joints, the presence of the rivet hole along the fracture plane for these joints prevents a large number of micro-cracks from growing. Typically, a micro-crack would link-up with the rivet hole and become the dominant crack. For the universal rivet joints, the rivet hole does not lie along the fracture plane. As a result, multiple micro-cracks tend to grow and subsequently coalesce into larger fatigue crack. 9 5 Conclusions Two series of fatigue tests have been carried out in order to quantify the effects of riveting induced residual stresses on the fatigue performance of riveted lap joints. This work was carried out as a follow-up to finite element studies carried out by the present authors on quantifying the effects of rivet installation on the stress field in loaded and unloaded lap joints [ 3]. The first test series focused on quantifying the effect of rivet installation on the overall fatigue performance. Based on the results from this series, the following conclusions can be made: In general, fatigue life increased with increasing rivet squeeze forces for both countersunk and universal rivet joints for the sheet thickness and joint geometry examined. The influence of rivet squeeze force on fatigue life is dependant on the location of cracking. For the universal rivet case, little improvement in fatigue performance was observed for the lower range of squeeze forces due to crack nucleation occurring away from the rivet hole outside the region of beneficial residual stresses. Conversely, cracking in the case of countersunk joints occurred at or near the joint net section where it is highly influenced by residual stresses. Thus, the countersunk joints displayed a greater sensitivity between rivet squeeze force and fatigue performance. The second series of fatigue tests were used to study the effects of riveting induced residual stresses on crack growth. Based on the results of this test series, the following conclusions can be made: The SEM reconstruction techniques employed in this study provided little quantifiable data in terms of the fatigue performance of universal riveted joints. The presence of multiple fatigue cracks and multiple fracture planes impeded the reconstruction process. For countersunk rivet joint specimens, the reconstructions demonstrated that use of force controlled riveting techniques resulted in repeatable fatigue crack growth behaviour for the same rivet squeeze force. Reductions in crack growth rate as great as 3-fold were observed when the rivet squeeze force was increased from 2 to 3 lbf. For the rivet/joint geometry studied, these squeeze forces produce upset rivet head geometries which fall within typical riveting practice standards [8]. Crack nucleation was observed to occur within the first 35% of the average overall fatigue life. Although exact nucleation times could not be established, a time period within which it occurred was established from the earliest detectable fracture surface marker.

10 A3 A2 A B3 B2 B 2.7. TK 224-T3 Alclad SHEET *ALL UNITS IN MM type type 2 Figure : Fatigue specimen geometry (adapted from coupon design presented in [8]) S 264 cycles 76 cycles 44 cycles 4 cycles 4 cycles S max 75% S max 4 cycles 4 cycles 4 cycles Program = 28 cycles Figure 2: Schematic of modified -6-4 marker-band spectrum

11 Crack propagation direction -band 6 4 Fatigue failure Static failure (a) visible marker bands (b) fatigue-static failure interface Figure 3: Examples of SEM reconstruction features 35 > 6 35 corrected fatigue life () error bars show 9% confidence levels inner sheet failure outer sheet failure 2-row 224-T3 splice with MS247AD4-4 s = 25.4mm, p = 38.mm, t =.mm D o = 3.2mm, Marker-band spectrum corrected fatigue life () error bars show 9% confidence levels inner sheet failure outer sheet failure 2-row 224-T3 splice with NAS97AD4-4 s = 25.4mm, p = 38.mm, t =.mm D o = 3.2mm, Marker-band spectrum F Sq (lbf) (a) universal joint F Sq (lbf) (b) countersunk joint Figure 4: Overall fatigue lives for inner and outer sheet failure

12 2 Figure 5: Typical location of fatigue damage observed in universal rivet joint fatigue specimens. FE results [2, 3] showing outer sheet faying surface plastic zone (circular regions) and region of maximum bending included for comparison. Figure 6: Typical location of fatigue damage observed in countersunk rivet joint fatigue specimens. FE results [2, 3] showing outer sheet faying surface plastic zone (circular regions) and region of maximum bending included for comparison.

13 3.5 CSK2.-T-SEM- Crack: A2 left elliptical curve fit kcycle band marker CSK2.-T-SEM- Crack: A2 right elliptical curve fit band marker CSK2.-T-SEM-2 Crack: B3 left elliptical curve fit band marker CSK2.-T-SEM-2 Crack: B3 right elliptical curve fit band marker CSK2.-T-SEM-3 Crack: B left polynomial curve fit band marker CSK2.-T-SEM-3 Crack: B right polynomial curve fit band marker Figure 7: Crack front reconstruction for a countersunk rivet joint with FSq = 2 lbf

14 4.5 CSK3.-T-SEM- Crack: B3 left polynomial curve fit kcycle (uncorrected) band marker CSK3.-T-SEM- Crack: B3 right polynomial curve fit band marker CSK3.-T-SEM-2 Crack: B2 left polynomial curve fit band marker CSK3.-T-SEM-2 Crack: B2 right polynomial curve fit band marker.5 CSK3.-T-SEM-3 Crack: A left polynomial curve fit band marker CSK3.-T-SEM-3 Crack: A right polynomial curve fit band marker Figure 8: Crack front reconstruction for a countersunk rivet joint with FSq = 3 lbf (6- and 4- band markers removed for clarity)

15 5 7 6 F Sq = 3lbf (3.3kN) 5 c (mm) 4 3 F Sq = 2lbf (8.9kN) 2 All SEM Specimens with NAS97AD4-4 rivets Figure 9: Crack growth curves from SEM reconstructions of countersunk rivet joints da/dn (µm/cycle) - CSK2.-T-SEM- A2-left CSK2.-T-SEM- A2-right CSK2.-T-SEM-2 B3-left CSK2.-T-SEM-2 B3-right CSK2.-T-SEM-3 B-left CSK2.-T-SEM-3 B-right CSK3.-T-SEM- B3-left CSK3.-T-SEM- B3-right CSK3.-T-SEM-2 B2-left CSK3.-T-SEM-2 B2-right CSK3.-T-SEM-3 A-left CSK3.-T-SEM-3 A-right F Sq = 2lbf (8.9kN) F Sq = 3lbf (3.3kN) c (mm) Figure : Crack growth rates from SEM reconstructions of countersunk rivet joints

16 6 y -band marker distance from rivet centre, y (mm) Figure : Crack front reconstruction for U2.-T-SEM- (arrows indicate intersecting fracture planes, shaded area indicates rivet hole location) -band marker -band marker y(mm) y(mm) Figure 2: Crack front reconstruction for U3.-T-SEM- (arrows indicate intersecting fracture planes)

17 7 References [] A. M. Brown and P. V. Straznicky. Modelling fretting contact stresses in a single lap splice. In Proceedings of The Fith International Symposium on Fretting Fatigue, ISFF5, Montreal, Canada, April, 27. [2] P. F. P. de Matos, P. M. G. P. Moreira, I. Nedbal, and P. M. S. T. de Castro. Reconstruction of fatigue crack growth in Al-alloy 224-T3 open-hole specimens using microfractographic techniques. Engineering Fracture Mechanics, pages , 25. [3] R. de Rijck. Stress Analysis of Fatigue Cracks in Mechanically Fastened Joints. PhD dissertation, Delft University of Technology, 25. [4] W. Elber. The significance of fatigue closure. In Damage Tolerance in Aircraft Structures, number ASTM STP 486, pages , 97. [5] S. A. Fawaz. Fatigue Crack Growth in Riveted Joints. PhD Dissertation, Delft University of Technology, 997. [6] N.M. Grinberg. Spacing of fatigue striations and crack growth rate. Soviet Materials Science, 2(2):53 6, 985. [7] P. M. G. P. Moreira, P. F. P. de Matos, and P. M. S. T. de Castro. Fatigue striation spacing and equivalent initial flaw size in al 224-t3 riveted specimens. Theoretical and Applied Fracture Mechanics, 43:89 99, 25. [8] R. P. G. Müller. An Experimental and Analytical Investigation on the Fatigue Behaviour of Fuselage Riveted Lap Joints: the Significance of the Rivet Squeeze Force and a Comparison of 224-T3 and Glare 3. PhD dissertation, Delft University of Technology, 995. [9] I. Nedbal, J. Siegl, and J. Kunz. Relation between striation spacing and fatigue crack growth rate in al alloy sheets. Advances in Fracture Research. Proceedings of the 7th International Conference on Fracture (ICF7), pages , 989. [] R. S. Piascik and S. A. Willard. The characterization of wide spread fatigue damage in the fuselage riveted lap splice. Technical Report NASA-TP , NASA, 997. [] C. D. Rans. The Role of Rivet Installation on the Fatigue Performance of Riveted Lap Joints. PhD dissertation, Carleton University, 27. [2] C. D. Rans, R. C. Alderliesten, and P. V. Straznicky. Riveting process induced residual stresses around solid rivets in mechanical joints. Journal of Aircraft, 44(): , 27. [3] C. D. Rans, R. C. Alderliesten, and P. V. Straznicky. Effects of rivet installation on residual stresses and secondary bending in a riveted lap joint. In Proceedings of The 48 th AIAA Structures, Structural Dynamics, and Materials Conference, Waikiki, USA, Apr. 27. [4] D. B. Schiff and R. B. d Agostino. Practical Engineering Statistics. Wiley-Interscience, 996.