- Email: [email protected]
Fatigue characterization and crack propagation mechanism of self-piercing riveted joints in titanium plates
International Journal of Fatigue 134 (2020) 105465
Contents lists available at ScienceDirect
International Journal of Fatigue journal homepage: www.elsevier.com/locate/ijfatigue
Fatigue characterization and crack propagation mechanism of self-piercing riveted joints in titanium plates
T
⁎
Xianlian Zhang, Xiaocong He , Wenjie Wei, Jiawei Lu, Kai Zeng Innovative Manufacturing Research Centre, Kunming University of Science and Technology, Kunming 650500, PR China
A R T I C LE I N FO
A B S T R A C T
Keywords: Self-piercing riveting Titanium Fretting wear Fatigue failure Crack propagation
The fatigue characterization and failure modes of self-piercing riveted joints in titanium plates were studied experimentally. The self-piercing riveting (SPR) process was optimised to improve the forming quality, mechanical property and failure behaviour of titanium joints through the comparative trials. The crack propagation mechanisms and fretting behaviours are quite different for different failure modes and analysed in detail from macro- and micro-perspective. The static and fatigue properties are both enhanced and the failure mode of rivet fracture is avoided for the optimised joints. The failures of upper sheet fracture and lower sheet fracture occur in the fatigue tests. Heavy fretting wear exists precisely at the critical faying interfaces and results in the fatigue crack initiations. The fatigue crack initiates at the faying interface between both sheets for the upper sheet fracture and propagates to both sides of the upper sheet simultaneously. Concerning the lower sheet fracture, the fatigue crack initiates at the faying interface between the rivet and the lower sheet, and then propagates to the close side of the lower sheet in the first stage, extends to the far side in the second stage.
1. Introduction
joining similar or dissimilar metallic and non-metallic sheet materials. Mori et al. [5] investigated the SPR process for joining the ultra high strength steel and aluminium alloy sheets, and the die shape was optimized based on numerical simulations to attain the joining of the hard ultra sheets. Rao et al. [6] presented the fatigue properties of SPR joints in aluminium alloy 6111 and carbon fibre reinforced polymers (CFRP) using lap-shear and cross-tension samples, and the joints with different configurations were produced with two distinctive rivet head heights. Calabrese et al. [7] carried out a galvanic corrosion test on dissimilar SPR joints in steel/aluminium sheet materials, and the relationship among failure mechanism, joint configuration and corrosion damage in salt spray environment was discussed in detail through the long term aging tests. Di Franco et al. [8] investigated the possibility of joining carbon fibre panels with aluminium alloy sheets using SPR and an adhesive layer. The effect of varying distances between the rivets was considered through the static and fatigue tests. Zhang et al. [9] made a discussion on the SPR joints in aluminium-lithium alloy sheets, and the mechanical properties and fretting wear of the joints were obtained and detected experimentally. Porcaro et al. [10] studied the static behaviours of SPR joints in aluminium alloy AA6060 under three different loading directions using an experimental study. Mucha [11] presented an intensive research on the joining of aluminium alloy sheets and discussed the effect of various riveting process parameters on the rivet
With the emphases on the energy consumption in all walks of life, more and more attention has been paid to the application of the lightweight structures. Therefore, the lightweight structures, particularly in lightweight materials, have gradually become a hot topic in current researches. In this work, the lightweight materials of titanium plates have been taken into consideration. Titanium and its alloys have been successfully widely used in aerospace, military, power and chemical industries [1], due to their remarkable characteristics: high specific strength and stiffness, low density, effective deformability and excellent corrosion resistance [2]. However, the welding of titanium is complicated and the inert gases or vacuum shielding is required, which often leads to inferior mechanical properties or hot cracking [3]. Wider applications of the lightweight materials are limited by hard or even impossible to be joined using the conventional joining techniques. Selfpiercing riveting (SPR) is a mechanical cold-forming joining technique, which has been considered as an economical and effective process to join lightweight sheet materials in many industrial fields [4]. It mainly depends on a firm mechanical interlock to fasten the sheet materials with two or more layers. The schematics of the joining process are illustrated in Fig. 1. Generally, it has been proved that SPR possesses the capability of
⁎
Corresponding author. E-mail address: [email protected] (X. He).
https://doi.org/10.1016/j.ijfatigue.2019.105465 Received 8 July 2019; Received in revised form 31 December 2019; Accepted 31 December 2019 Available online 02 January 2020 0142-1123/ © 2020 Elsevier Ltd. All rights reserved.
International Journal of Fatigue 134 (2020) 105465
X. Zhang, et al.
Fig. 1. Schematic illustration of SPR process.
forming characteristics. The mechanical properties and failure modes of the optimised joints in titanium plates were discussed through a contrastive analysis. The fatigue crack propagation and fretting behaviour were intensively analysed from macro- and micro-perspective.
deformation of SPR joints. Xie et al. [12] investigated the cold-formed steel (CFS) shear walls with self-piercing riveted (SPR) connections under monotonic and reversed cyclic loading through a series of experiments. Moraes et al. [13] analysed the influence of the process deformation history on the mechanical performance of dissimilar SPR joints between magnesium and aluminium alloys. Particularly, the joining of similar titanium plates using SPR process was originally implemented with a pre-heating operation on the titanium plates by He et al. [14]. The influence of relief annealing on the static and fatigue performances of the pre-heated joints was carried out experimentally by Zhang et al. [15]. The regular SPR joints in titanium plates without pre-heating were realised later and the effects of sheet thickness and fatigue load on the mechanical property and failure were discussed by Zhao et al. [16]. However, it was found that the mechanical properties of the regular joints were notably lower than those with pre-heating. A majority of the mentioned titanium joints failed due to the rivet fracture in both quasi-static and fatigue tests. The fracture type of the rivet is the brittle fracture, which is quite dangerous without any plastic deformation for the mechanical structure design. Therefore, the corresponding efforts should be made to further improve the mechanical properties and avoid the sudden failure of the SPR joints in titanium plates. In addition, the available literature about the fatigue crack propagation in SPR joints is quite limited. Li et al. [17] discussed the fatigue crack initiation and development of the double rivets lap shear joints in aluminium alloy AA5754, which was based on the local images and SEM microstructures of the fracture interfaces. They thought that the fatigue cracks initiated at the bottom surface of the upper sheet at high load amplitudes, and then developed in transverse direction and along the sheet thickness direction to the top surface of the upper sheet; while at low load amplitudes, the fatigue cracks initiated at the top surface of the lower sheet. Kang and Kim [18] did a brief analysis using the partial enlarged images of the fractures in AA5052 joints and found a similar situation as Li et al. [17] at high-loading range, while the fatigue cracks initiated on the lower sheet near the rivet tail at low-loading range and then propagated along the sheet thickness direction. Moraes et al. [19] and Huang et al. [20] investigated the fatigue crack initiation mainly based on a detailed analysis on the fretting wear in dissimilar aluminium/steel and similar aluminium joints. They discovered that the crack initiation points were quite different for different failure modes. Particularly, it could be found from the previous researches on the mechanical fasteners, the fretting fatigue behaviours potentially took a pivotal role in the fatigue crack initiation that the fatigue cracks typically originated in the severe fretting region at the faying interface of the fastener [20–22]. The fatigue performance was obviously impaired in certain cases due to the fretting wear [23,24]. This work aims to deal with the above circumstances of SPR process for the titanium plates and investigate the fatigue characterization and crack propagation mechanism of corresponding joints in detail. The SPR process was optimised for a better forming structure through the comparative trials. A micro-hardness analysis was performed for the
2. Experimental details 2.1. Materials and process The materials used are 1.5 mm thick plates of commercially pure titanium TA1 and the chemical composition (wt%) is: 0.06 Fe, 0.01 C, 0.01 N, 0.001 H, 0.04 O and Ti balance. All the titanium plates in this work were sheared along the rolling direction to 110 mm × 20 mm × 1.5 mm. A RIVSET VARIO-FC (VTF) hydraulic driven riveting machine was adopted in the riveting experiments, which was fabricated and supplied by Böllhoff GmbH & Co. For joining the stiffer sheet materials, the rivet hardness needs to be enhanced using modified metal heat treatment process, which may increase the brittleness and reduce the shear resistance of the rivet. The enhanced semi-tubular rivets were used to join the 1.5 mm thick titanium plates without pre-heating in the previous research [16]. Particularly, the enhanced semi-tubular rivets were used in this work and possessed the mechanical properties: young’s modulus 206 GPa, poisson’s ratio 0.3, tensile yield strength 1719.7 MPa and compressive strength 1595.6 MPa. The rivets are made of boron steel 36MnB4 and coated with zinc-tin alloy, and the chemical composition (wt%) is: 0.33 C, 0.80 Mn, 0.30 Si, 0.025 S, 0.025 P, 0.30 Cr and Fe balance. Considering these backgrounds, the comparative trials of the joints in titanium plates were carried out. The riveting qualities are evaluated using the cross section visual inspection method based on three structural parameters [4]: rivet head height (h), rivet spread (r) and remaining bottom thickness (t). It is known that the joint strength mainly depends on the mechanical interlock in the SPR structure and the value of rivet spread (r) is the critical indicator. The joints manufactured with varying riveting parameters are evaluated based on the three structural parameters and joint strengths. Finally, the optimised parameters are confirmed and listed as below: riveting pressure 270 bar, rivet hardness 50 ± 2 HRC, rivet length 5.5 mm and die cavity φ9.2 × 1.8 mm. Comparing with the previous research, the riveting pressure was increased by 60 bar; the values of rivet hardness were the same; the rivet length was decreased from 6 mm to 5.5 mm; and the die cavity was changed from φ9.0 × 1.25 mm to φ9.2 × 1.8 mm. The geometrics of the riveting tools are shown in Fig. 2. More than fifty optimised joints were fabricated for the following research. Ten previous joints were made for the static tests, which has not been discussed in the previous research. The dimensions of the specimens were illustrated in Fig. 3. 2.2. Test set-up A bit deeper analysis about the forming structures of the corresponding joints was implemented using a digital micro-hardness tester 2
International Journal of Fatigue 134 (2020) 105465
X. Zhang, et al.
Fig. 2. The riveting tools for the optimised joints (dimensions in mm): (a) punch, (b) rivet, and (c) die.
Fig. 4(a). Concerning the three structural parameters, the rivet head height (h = –0.1 mm) was mainly improved by a higher riveting pressure (270 bar) and a decrease of the rivet length (5.5 mm), which could make the optimised joints have a better sealing performance compared to the previous joints (h = 0.2 mm). The values of rivet spread (r = 1.0 mm) and remaining bottom thickness (t = 0.7 mm) for the optimised joints are superior to those (r = 0.8 mm, t = 0.3 mm) of the previous joints. Due to the increase of die cavity and the decrease of the rivet length, the extrusion of the lower sheet could be assuaged effectively, which could be observed from the extrusion lines marked on the cross sections in Fig. 4. The hardness (HV0.05) distribution on the corresponding cross sections is summarised in Fig. 5. The initial hardness of the plates could be considered as 140 HV, which was measured from the edge of cross section. It can be found that the hardness profiles on the left and right sides of the rivet are different for both joints, which was mainly affected by the inhomogeneity of the plate itself and the mechanical polishing process. The plates of the previous joint are more hardened in the slight deformation region (Nos. 1–3 and 11–13). It is noteworthy that the values of hardness points near the rivet (Nos. 4–10) are higher than those in the slight deformation region. The optimised joint basically presents the same feature. The hardness distribution for the optimised joint presents a shape “M”, and two peak values occur at points No. 4 and 10 on both sheet that are close to the outer surface of the rivet leg. Generally, the hardness values of the previous joint are higher than those of the optimised joint. Especially for the region inside the rivet (Nos. 6–8), the previous joint (203.8 HV for the upper sheet, 218.7 HV for the lower sheet) is obviously superior to the optimised one (169.3 HV for the upper sheet, 193.9 HV for the lower sheet). Consequently, both the upper and lower sheets of the previous joint possess a better work hardening effect comparing with the optimised joint, which means that the plastic deformation of the plates is more serious and the lower plate binds the rivet more tightly in the previous joint.
(HVST-1000Z, China). A loading force of 0.49 N and a holding time of 15 s were adopted for the cross section samples, which were cut along the rivet diameter and then mechanical polished. All the hardness values were measured three times and averaged. The quasi-static tensile-shear tests and fatigue tests were performed on a MTS Landmark100 servo-hydraulic testing machine. Two spacers with equal thickness of 1.5 mm were glued on both ends of the specimens to avoid additional bending, as shown in Fig. 3. A tensile speed of 5 mm/min was adopted and ten specimens were repeated in the quasistatic tests for both previous and optimised joints. A sinusoidal wave at the frequency of f = 10 Hz and a load ratio of r = 0.1 was performed in the fatigue tests using the tension-tension loading mode. The fatigue loads were chosen based on the static strength for reflecting the entire fatigue lives. Five specimens were tested at each different loads. Attaining over 2 million cycles or the specimen failure was regarded as the termination criterion. All the tests were carried out at a room temperature. Particularly, the fatigue data of the previous joints by Zhao et al. [16], which were obtained under the same testing condition, were cited in this work. The fracture features of the typical failed specimens and the micro morphologies of the typical fracture surfaces were observed and analysed using a scanning electron microscopy (SEM, TESCAN Inc., Czech Republic). The effort was focused on trying to illuminate the fatigue crack initiations and propagation stages in more detail using SEM technique, and the SEM images of the whole macro fracture were stitched from more than a dozen pictures so as to illustrate the rough trend and direction of the fatigue cracks. Furthermore, the fretting debris in the fretting region was detected using an energy dispersive X-ray spectroscopy (EDS, EDAX Inc., USA).
3. Results and discussion 3.1. Forming characteristics The cross sections of the previous and optimised joints in titanium plates are presented in Fig. 4. It is notable that there is an obvious gap between the rivet head and the upper sheet for the optimised joints in Fig. 4(b), while the gap is much smaller due to a rivet length of 6 mm in
3.2. Static property and failure Fig. 6 shows the results of the joints in titanium plates in quasi-static tests. For evaluating the validity of the data, the Grubbs criterion was
Fig. 3. Schematic illustration of the investigated specimens (dimensions in mm). 3
International Journal of Fatigue 134 (2020) 105465
X. Zhang, et al.
Fig. 4. The cross sections and marked hardness point Nos. of the joints (dimensions in mm): (a) for the previous joints, and (b) for the optimised joints.
rivet tail and the lower sheet. It is worthy to note that the mechanical interlock of the optimised one was totally damaged under tensile-shear loading. The failure mode was changed from the rivet fracture to the rivet tail pulled-out using the optimised parameters, which could help prevent sudden failure from a practical standpoint. It is thought that the previous joints are not prone to plastic deformation under tensile-shear loading because of a better work hardening effect on the plates, which indicates that the interlock structure is firmer compared to the optimised joints. As a consequence, the rivet of the previous joints bore more shear loads during quasi-static tests, which allowed the previous joints to fail in the rivet breakage.
used to eliminate the outliers, and it was found that there was no outlier for both joints. Intuitively, the uniformity of load-displacement curves for the optimised joints is quite well, as shown in Fig. 6(b). After calculation, the average lap shear strength and energy absorption of the optimised joints are 7.19 kN and 30.84 J (Standard deviation: 0.18 kN and 1.95 J) respectively, and 5.41 kN and 4.09 J (Standard deviation: 0.53 kN and 1.62 J) are for the previous joints. The optimised joints possessed a better static property, which was even superior to those with pre-heating or relief annealing by He et al. [14] and Zhang et al. [15]. In addition, the load-displacement curve of the optimised joints (Fig. 6(b)) could be divided into three stages. In the 1st stage, the joint bore the increasing shear load and the peak load occurred at the displacement of 2 mm. The rivet started to be pulled out from the lower sheet and the upper sheet was gradually bended under tensile-shear loading in the 2nd stage. In the 3rd stage, it was the final failure of the joint with the load dropping rapidly. For the previous ones, only the 1st and 3rd stage can be observed in Fig. 6(a) due to the failure mode of rivet fracture. Fig. 7 presents the failure modes in quasi-static tests and the schematic illustrations. For the previous joints, the failure mode of rivet fracture occurred for all specimens under tensile-shear loading. As shown in Fig. 7(a), the upper sheet was slightly deformed during the quasi-static test, while the lower sheet including the joint bottom was barely deformed. It is noticed that the separated part of the upper sheet was left in the interlock structure due to the action of the broken rivet tail, as marked in the schematic illustration. The brittle rupture features was detected on the rivet fracture surface using SEM technique, which led to the sudden failure of the previous joints in quasi-static tests. For the optimised joints in Fig. 7(b), all specimens failed due to the rivet tail pulled-out from the lower sheet, which is consistent with the previous researches by Sun et al. [25] and Porcaro et al. [10]. The separated part of the upper sheet was also pulled out for the action of the rivet leg. Both sheets presented obvious deformation during the separation of the
3.3. Fatigue property and failure The fatigue test results are summarised in Table 1. Six fatigue load amplitudes were adopted to evaluate the fatigue life of the optimised joints. The dispersion of fatigue cycles for different joints increases with the decrease of fatigue load amplitudes. Besides, two failure modes occurred in fatigue tests: failure mode A: upper sheet fracture and failure mode B: lower sheet fracture, which were quite different from the previous joints. For a comparison analysis, the fatigue data and the fatigue load-fatigue life (F-N) curves were plotted in Fig. 8 for both joints. It is noticed that three interesting regions are presented for the optimised joints: Region Ⅰ: mainly failure mode A; Region Ⅱ: mainly failure mode B; Region III: no failure over 2 million cycles. The region distribution could help to predict the fatigue lives and failure modes of the optimised joints at different fatigue loads. However, it is found that the previous joints mainly failed in the lower sheet when the fatigue cycles were between 105 and 106, while failed in the rivet when the fatigue cycles were below 105 or over 106. The distribution feature of the fatigue data is obviously different for the previous and optimised joints. In addition, it is observed from Fig. 8, while attaining the same fatigue cycles, the fatigue load amplitudes of the optimised joints are
Fig. 5. The Vickers hardness for the cross sections. 4
International Journal of Fatigue 134 (2020) 105465
X. Zhang, et al.
Fig. 6. Quasi-static lap shear test results for the joints in titanium plates: (a) load-displacement curves for the previous joints, and (b) load-displacement curves for the optimised joints.
recorded by a digital camera (Sony ILCE-A7RM2). As shown in Fig. 10(a), the fatigue crack firstly appeared at the ring of the rivet head and extended from the pierced hole to both sides of the upper sheet. In Fig. 10(b), the fatigue crack was observed in the lower sheet and near the joint bottom. It is thought that the fatigue crack was initiated near the riveted part and extended along the sheet width direction. Meanwhile, the cross sections of the failed specimens were fabricated to explore the failure paths along the sheet thickness direction in both plates, and the partial enlarged drawings of cross sections D1 and D2 are presented in Fig. 10(c) and (d), respectively. The upper sheet fracture path occurred underneath the edge of the rivet head for failure mode A, which is consistent to the conclusion by Khanna et al. [27]. The lower sheet fracture path existed under the rivet tail, which was the region of the remaining bottom.
higher than that of the previous joints. The F-N curve of the optimised joints is completely above that of the previous joints at all fatigue phases. Consequently, it means that the fatigue property of the titanium joints is improved significantly with the optimised parameters. Based on the above analysis, the previous joints had a more serious working harden effect on the lower sheet and an inferior remaining bottom thickness of 0.3 mm comparing with the optimised joints. This indicates that the lower sheet and rivet of the previous joint are hard to deform effectively and easily destroyed under cyclic fatigue loading, which made the optimised joints failed in the lower sheet exhibit the higher fatigue lives than the previous joints failed in the same type of lower sheet. Particularly, due to the modified forming structure obtained using the optimised parameters, the failure modes of rivet fracture were successfully avoided for the optimised joints, and the failure modes were changed from the rivet fracture to the upper sheet fracture. Fig. 9 presents the failures at different fatigue load amplitudes of the optimised joints in fatigue tests, and the failures are similar to the results from Iyer et al. [26]. It is worthy to note that the appearances of the specimens in failure mode A seem quite similar and all the specimens are fractured close to the rivet in Fig. 9(a). The visible plastic deformation could be observed on both sides (near the sheet edges) of the fractured plates. For failure mode B, the appearances are a bit different and the visible cracks (marked in red arrow in Fig. 9(b)) present in the upper sheet at high loads of 3.60 kN and 2.88 kN, which may be caused by the differential stress distribution on the faying interface. Unlike the failure mode A, the visible plastic deformation only occurred on one side of the fractured plates. Further, the locations of fatigue cracks during the fatigue tests were
3.4. Fatigue crack propagation Considering the failure mode A, Fig. 11 presents the characteristics of the typical upper sheet fracture surface. The corresponding enlarged views of the marked parts in Fig. 11a are shown in Fig. 11(b)–(e), respectively. In terms of the morphological features, the fracture surface could be divided into three zones: fatigue source (zone i), fatigue crack propagation (zone ii) and final fracture (zone iii), which have been marked in the enlarged views. The river patterns on the fracture surface indicate the fatigue crack initiation location in zone i. The fatigue crack initiated on the faying surface of the upper sheet and propagated outward as the yellow arrow marked in Fig. 11(d). Fig. 12(c) presents the micro morphology of the crack initiation location. The cleavage step
Rivet tail
Joint bottom
Rivet
Separated part
(a)
Separated part
(b)
Fig. 7. Quasi-static failure modes: (a) rivet fracture for the previous joints, and (b) rivet tail pulled-out for the optimised joints. 5
International Journal of Fatigue 134 (2020) 105465
X. Zhang, et al.
Table 1 Fatigue test results of the optimised joints. Specimen Nos.
Fatigue load amplitude/kN
Fatigue cycles
Failure mode
Specimen Nos.
Fatigue load amplitude/kN
Fatigue cycles
Failure mode
01 02 05 20 28 03 04 11 12 26 07 09 19 23 24
3.60 3.60 3.60 3.60 3.60 2.88 2.88 2.88 2.88 2.88 2.52 2.52 2.52 2.52 2.52
54,923 71,447 75,724 80,352 73,096 198,723 215,729 154,960 174,313 145,010 368,039 317,213 383,087 205,165 304,332
A A A A B B A A B A B A A A A
08 10 13 22 25 16 17 18 21 27 14 15
2.16 2.16 2.16 2.16 2.16 1.94 1.94 1.94 1.94 1.94 1.80 1.80
591,330 707,532 601,146 896,506 1,082,870 1,029,673 1,381,648 1,297,992 924,917 1,051,983 2,172,564 2,269,167
A B B B B B B B B B No failure No failure
and both lines lean towards the crack initiation location. This indicates that the fatigue cracks propagate faster in the sheet width direction than in the sheet thickness direction, which is consistent with the standpoint by Li et al. [28]. Considering the failure mode B, the features of the lower sheet fracture surface are presented in Fig. 13, and Fig. 13(b)–(e) are the enlarged views of the marked parts in the macro fracture in Fig. 13(a), respectively. The three zones as the upper sheet fracture also present on the lower sheet fracture. In the fatigue source region (zone i), a ridge could be observed at the crack initiation location, marked with a blue arrow in Fig. 13(d). It means that the fatigue crack initiates at the faying interface between the rivet and the lower sheet and near the bending point of the lower sheet, and then propagates along the direction marked with a yellow arrow. The micro morphology of the marked area in a red rectangle is presented in Fig. 14(c). The tiny fatigue striations and step patterns present the transgranular fracture in the fatigue source region. The crack propagation direction is indicated in zone ii, marked with the yellow arrows. The fatigue striations and the river marks can be seen in a more obvious perspective than in the fatigue source region. However, the appearances of zone ii look quite different for the far side (Fig. 13(b) and (c)) and close side (Fig. 13(e)) of zone i, and their micro morphology is shown in Fig. 14(b) and (d), respectively. The fatigue striations and ridges in Fig. 14(d), which are the features of the cleavage fracture, indicate that the first stage of the fatigue crack propagation is to extend to the close side of zone i, as the visible initial crack indicated in Fig. 10(b). The river marks and secondary cracks in Fig. 14(b) proved that the joint bottom failed due to the transgranular fracture. The far side of zone i is the second stage of the fatigue crack propagation. Finally, the completed fracture occurred when the residual structure of the joint could no longer bear the continuous cyclic fatigue loading. The micropores and dimples in the microstructure (Fig. 14(a)) indicated that the residual structure of the joint failed due to the ductile fracture. Besides, the crack propagation
Fig. 8. Fatigue load-fatigue life (F-N) curves for the joints in titanium plates.
patterns indicated the fatigue crack initiated through the grains. In zone ii, the fatigue striations and river patterns point out the crack propagation direction, marked with the yellow arrows. The micro morphology of zone ii is presented in Fig. 12(b) and (d) which are located at the left and right side of the crack initiation location, respectively. The morphological features seem to be quite similar for the both sides. This indicated that the fatigue crack propagated to zone ii simultaneously, as the fatigue crack indicated in Fig. 10(a). The typical fatigue striations and secondary cracks can be seen obviously in zone ii, which is the transgranular fracture. Fig. 11(a) and (d) present the micro fracture features of the final fracture in zone iii caused by the loading amplitude exceeded the fracture limit of the residual structure of the joint. From the micro morphology in Fig. 12(a), many dimples and tearing ridges could be found, which shows the typical ductile fracture feature in zone iii. Particularly, the crack propagation front lines can be observed clearly on both sides of the upper sheet fracture in Fig. 11(b) and (e),
Deformation
Reverse side 3.60 kN
2.88 kN
Deformation
Crack
2.52 kN
2.16 kN
(a)
Crack 3.60 kN 2.88 kN 2.52 kN
2.16 kN 1.88 kN
(b)
Fig. 9. Fatigue failures for the optimised joints: (a) the fatigue failure mode A, and (b) the fatigue failure mode B. 6
International Journal of Fatigue 134 (2020) 105465
X. Zhang, et al.
D1
Crack
Joint bottom Upper sheet
Rivet Rivet
Crack
(a)
Rivet
Upper sheet Lower sheet Crack Lower sheet
D2
Crack
(b)
(c)
(d)
Fig. 10. Fatigue failure cracks for the optimised joints: (a) crack in the upper sheet, (b) crack in the lower sheet, (c) fracture path for failure mode A, and (d) fracture path for failure mode B.
length at high load amplitude of 3.60 kN is shorter than those at the other amplitudes, which could be caused by the fatigue crack grown faster at high load amplitude. However, on the lower sheet fracture (failure mode B), the length at low load amplitude of 1.94 kN is shorter than those at the other amplitudes, which is considered to be affected by a lower fatigue crack growth rate at low load amplitude. This indicates that the fatigue crack growth rate plays an important role on the crack growth region length. Due to the obvious plastic deformation on the lower sheet and different propagation stages on the lower sheet fracture, the crack growth region length decreases with the fatigue load increasing, which is totally in contrast with the upper sheet fracture. In addition, the crack growth region lengths are generally higher on the upper sheet fracture than on the lower sheet fracture, which could be also caused by the differences in the fatigue crack propagation stages and the plastic deformation.
front lines also prove that the fatigue cracks propagate faster in the sheet width direction than in the sheet thickness direction. As the discussion above, the fatigue crack initiated at the faying interface between both sheets, and simultaneously propagated outward to both sides of the upper sheet for the failure mode A; while the fatigue crack initiated at the faying interface between the rivet and the lower sheet, and firstly propagated to the close side of the lower sheet, then extended to the far side of the lower sheet for the failure mode B. Until the residual structure of the corresponding plate could not withstand the fatigue load amplitude, the final fracture happened. The schematic diagrams of the fatigue crack propagation are illustrated in Fig. 15, and the propagation directions are indicated using the yellow arrows. Further, it shows during the fatigue tests that the fatigue source zone i and fatigue crack propagation zone ii sustain most of the fatigue cycles. The length of the crack growth region, included zone i and zone ii, is measured using an industrial microscope (ZW-H3000, China) with a measuring software. The detailed values are listed in Table 2, and Fig. 16 presents the average crack growth region length at different load amplitudes. On the upper sheet fracture (failure mode A), the
3.5. Fretting fatigue behaviour The fretting fatigue behaviour existed precisely and was highly
(a)
Crack propagation front line
(b)
(c)
Crack initiation location
Crack propagation front line
(d)
(e)
Fig. 11. Fatigue failure analysis for failure mode A: (a) macro view of the upper sheet fracture (b) enlarged view of Part 1, (c) enlarged view of Part 2, (d) enlarged view of Part 3, and (e) enlarged view of Part 4. 7
International Journal of Fatigue 134 (2020) 105465
X. Zhang, et al.
Striations Striations Dimples
Step pattern Secondary crack
Ridges Secondary crack
(a)
(b)
(c)
(d)
Fig. 12. Micro morphology at different zones marked with red rectangles for failure mode A: (a) at zone iii marked in Fig. 11b, (b) at zone ii marked in Fig. 11c, (c) at zone ii marked in Fig. 11d, and (d) at zone ii marked in Fig. 11e. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
(marked with the dot-dashed line), the constituent of the fretting debris could be inferred to be one of the titanium oxides. Therefore, it is considered that due to the fretting wear in the faying interface between both sheets, the elements on the sheet surface are scratched out and oxidized. The fatigue crack initiate in the fretting region S at the bottom surface of the upper sheet fracture because of the continuous cyclic loads and micro-vibrations. According to the previous analysis for failure mode B, the typical rivet of the failed joints is detected using micro test methods, and the upward view of the rivet is presented in Fig. 18(b). There is no any signs of wear on the bottom surface of rivet tail, and the separated part of the upper sheet is unharmed. It means that for the forming structure of the optimised joints, no fretting occur in the contact interface between the bottom of rivet tail and the lower sheet. The front view of the rivet is shown in Fig. 18(c). The fretting debris, especially area β, could be observed on the fretting region R. Probably, this is the reason of the fatigue crack initiation, which is consistent with the conclusions by Sun et al. [25]. Besides, Fig. 18(d) shows the enlarged view of area γ on the
correlated with the fatigue failures of the optimised joints. Due to the optimised riveting process, the fretting behaviour of the optimised joints is quite different from that of the previous joints which has been discussed in Zhao et al. [16]. The fracture paths and fretting regions in different fatigue failure modes for the optimised joints are illustrated and marked in Fig. 17 for a better understanding. Particularly, the fretting regions R (rivet) and S (sheets) indicate the faying interfaces between the rivet tail and lower sheet and between both sheets, respectively. It was found that the heavy fretting and fracture path always occurred at the side along the direction of cyclic loading for all failed optimised joints. Fig. 18(a) presents the upward view of the upper sheet fracture (failure mode A). The fretting phenomenon near the fatigue crack initiation point and the fatigue crack in fretting region S can be observed clearly. Further, the area α at the edge of the fretting region was analysed using EDS, as shown in Fig. 18(e). The energy spectrum of area α contains some Ti, O, N and other elements in an infinitesimal quantity. Concerning the distributions of the fretting debris and oxygen element
(a)
Crack propagation front line
(b)
(c)
Crack initiation location
Crack propagation front line
(d)
(e)
Fig. 13. Fatigue failure analysis for failure mode B: (a) macro view of the lower sheet fracture, (b) enlarged view of Part 1, (c) enlarged view of Part 2, (d) enlarged view of Part 3, and (e) enlarged view of Part 4. 8
International Journal of Fatigue 134 (2020) 105465
X. Zhang, et al.
Ridges
Striations Striations Secondary crack Step pattern
Micropores
(a)
(b)
(c)
(d)
Fig. 14. Micro morphology at different zones marked with red rectangles for failure mode B: (a) at zone iii marked in Fig. 13b, (b) at zone ii marked in Fig. 13c, (c) at zone i marked in Fig. 13d, and (d) at zone ii marked in Fig. 13e. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
Final fracture
Crack initiation
Final fracture
Crack growth region length Crack initiation
Final fracture
Crack growth region length
(a)
Final fracture
(b)
Fig. 15. Schematic diagrams for the macro fractures in different failure modes: (a) the upper sheet fracture surface, and (b) the lower sheet fracture surface. Fracture path for failure mode A
Table 2 Statistics for the fatigue crack growth region length at different load amplitudes. The upper sheet fracture (Failure mode A) Specimen Nos.
Cyclic load
The lower sheet fracture (Failure mode B)
Fatigue load amplitude/ kN
Crack growth region length/ mm
Specimen Nos.
3.60 3.60 3.60 3.60 2.88 2.88 2.88 2.52 2.52 2.52 2.52 2.16
15.04 15.06 15.47 15.11 16.41 16.20 16.39 16.11 16.68 16.66 16.80 16.64
28 03 12 07 10 13 22 25 16 17 18 21 27
Fatigue load amplitude/ kN
Crack growth region length/ mm
3.60 2.88 2.88 2.52 2.16 2.16 2.16 2.16 1.94 1.94 1.94 1.94 1.94
15.41 15.35 15.72 14.54 15.94 15.68 15.10 15.44 14.46 14.43 14.77 14.41 14.22
Cyclic load Fretting region S Fretting region R Fracture path for failure mode B
Fig. 17. Schematic illustration for the fatigue failures of the optimised joints. 01 02 05 20 04 11 26 09 19 23 24 08
lower sheet fracture which belongs to the fatigue crack propagation zone (Fig. 13(e)). The obvious fretting layer, thickness over 50 μm, exists on the sheet surface. It is thought that due to the continuous cyclic loads and micro-vibrations between both sheets, the fretting wear on the lower sheet fracture is generated in the course of the fatigue crack propagation, and it accelerates the propagation of the fatigue cracks. 4. Conclusions The paper is highlighted to the mechanical improvement, fatigue failure, crack propagation and fretting behaviour of self-piercing riveted joints in titanium plates. The main conclusions are summarised as following: 1. The forming feature of self-piercing riveting in titanium plates can be optimised in this work using a riveting pressure of 270 bar, rivet length of 5.5 mm and rivet hardness of 50 ± 2 HRC; 2. The static and fatigue properties are both improved effectively, and the rivet fracture in the previous joints is avoided in the optimised joints. The failure in the quasi-static tests presents the pull-out of the rivet tail from the lower sheet, and two failure modes occurred in the fatigue tests: upper sheet fracture and lower sheet fracture. 3. For the upper sheet fracture, the fatigue crack initiates at the faying interface between both sheets, and then propagates to both sides of the upper sheet simultaneously. For the lower sheet fracture, the fatigue crack initiates at the faying interface between the rivet and the lower sheet, and propagates to the close side of the lower sheet at the first stage, extends to the far side at the second stage. The final
Fig. 16. Average crack growth region length of the fractures at different fatigue load amplitudes.
9
International Journal of Fatigue 134 (2020) 105465
X. Zhang, et al.
Fig. 18. Fretting in the optimised joints: (a) upward view of the upper sheet fracture for failure mode A, (b) upward view of the rivet for failure mode B, (c) front view of the rivet for failure mode B, (d) enlarged view of area γ in Fig. 13(e) for failure mode B, and (e) energy spectrum of area α in Fig. 18(a) for failure mode A.
fracture occurs until the residual structure could no longer bear the continuous fatigue loading. 4. The fretting fatigue behaviours occurred precisely at the faying interfaces between the rivet tail and lower sheet and between both sheets where the fatigue crack initiated. The fretting was highly correlated with the fatigue failures for self-piercing riveted joints.
[8] Di Franco G, Fratini L, Pasta A. Influence of the distance between rivets in self-piercing riveting bonded joints made of carbon fiber panels and AA2024 blanks. Mater Des 2012;35:342–9. [9] Zhang X, He X, Gu F, Ball A. Self-piercing riveting of aluminium–lithium alloy sheet materials. J Mater Process Technol 2019;268:192–200. [10] Porcaro R, Hanssen A, Langseth M, Aalberg A. An experimental investigation on the behaviour of self-piercing riveted connections in aluminium alloy AA6060. Int J Crashworth 2006;11(5):397–417. [11] Mucha J. A study of quality parameters and behaviour of self-piercing riveted aluminium sheets with different joining conditions. Stroj Vestn-J Mech Eng 2011;57(4):323–33. [12] Xie Z, Yan W, Yu C, Mu T, Song L. Experimental investigation of cold-formed steel shear walls with self-piercing riveted connections. Thin Wall Struct 2018;131:1–15. [13] Moraes JFC, Jordon JB, Su X, Barkey ME, Jiang C, Ilieva E. Effect of process deformation history on mechanical performance of AM60B to AA6082 self-pierce riveted joints. Eng Fract Mech 2019;209:92–104. [14] He X, Wang Y, Lu Y, Zeng K, Gu F, Ball A. Self-piercing riveting of similar and dissimilar titanium sheet materials. Int J Adv Manuf Technol 2015;80(9–12):2105–15. [15] Zhang X, He X, Xing B, Zhao L, Lu Y, Gu F, et al. Influence of heat treatment on fatigue performances for self-piercing riveting similar and dissimilar titanium, aluminium and copper alloys. Mater Des 2016;97:108–17. [16] Zhao L, He X, Xing B, Zhang X, Cheng Q, Gu F, et al. Fretting behavior of self-piercing riveted joints in titanium sheet materials. J Mater Process Technol 2017;249:246–54. [17] Li D, Han L, Thornton M, Shergold M. Influence of rivet to sheet edge distance on fatigue strength of self-piercing riveted aluminium joints. Mater Sci Eng A 2012;558:242–52. [18] Kang SH, Kim HK. Fatigue strength evaluation of self-piercing riveted Al-5052 joints under different specimen configurations. Int J Fatigue 2015;80:58–68. [19] Moraes JFC, Rao HM, Jordon JB, Barkey ME. High cycle fatigue mechanisms of aluminum self-piercing riveted joints. Fatigue Fract Eng Mater Struct 2018;41(1):57–70. [20] Huang L, Bonnen J, Lasecki J, Guo H, Su X. Fatigue and fretting of mixed metal selfpiercing riveted joint. Int J Fatigue 2016;83:230–9. [21] Chen Y, Han L, Chrysanthou A, O’Sullivan J. Fretting wear in self-piercing riveted aluminium alloy sheet. Wear 2003;255:1463–70. [22] Han L, Chrysanthou A, O’Sullivan J. Fretting behaviour of self-piercing riveted aluminium alloy joints under different interfacial conditions. Mater Des 2006;27:200–8. [23] Szolwinski M, Farris T. Linking riveting process parameters to the fatigue performance of riveted aircraft structures. J Aircraft 2000;37:130–7. [24] Wagle S, Kato H. Ultrasonic detection of fretting fatigue damage at bolt joints of aluminum alloy plates. Int J Fatigue 2009;31:1378–85. [25] Sun X, Khaleel MA. Strength estimation of self-piercing rivets using lower bound limit load analysis. Sci Technol Weld Joi 2005;10(5):624–35. [26] Iyer K, Hu SJ, Brittman FL, Wang PC, Hayden DB, Marin SP. Fatigue of single- and doublerivet self-piercing riveted lap joints. Fatigue Fract. Eng. Mater Struct 2005;28(11):997–1007. [27] Khanna SK, Long X, Krishnamoorthy S, Agrawal HN. Fatigue properties and failure characterisation of self-piercing riveted 6111 aluminium sheet joints. Sci Technol Weld Joi 2006;11(5):544–9. [28] Li D, Han L, Thornton M, Shergold M, Williams G. The influence of fatigue on the stiffness and remaining static strength of self-piercing riveted aluminium joints. Mater Des 2014;54:301–14.
Declaration of Competing Interest The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper. Acknowledgements This work was supported by the National Natural Science Foundation of China (Grant No. 51565023). The authors would like to thank Mr. Yun Lei from Faculty of Materials Science and Engineering of Kunming University of Science and Technology for his support during micro-hardness testing. References [1] Bishoyi B, Sabat R, Sahu J, Sahoo S. Effect of temperature on microstructure and texture evolution during uniaxial tension of commercially pure titanium. Mater Sci Eng A 2017;703:399–412. [2] Gurao N, Kapoor R, Suwas S. Deformation behaviour of commercially pure titanium at extreme strain rates. Acta Mater 2001;59:3431–46. [3] Sridharan N, Wolcott P, Dapino M, Babu S. Microstructure and texture evolution in aluminum and commercially pure titanium dissimilar welds fabricated using ultrasonic additive manufacturing. Scripta Mater 2016;117:1–5. [4] Haque R. Quality of self-piercing riveting (SPR) joints from cross-sectional perspective: a review. Arch Civ Mech Eng 2018;8(1):83–93. [5] Mori K, Kato T, Abe Y, Ravshanbek Y. Plastic joining of ultra high strength steel and aluminium alloy sheets by self piercing rivet. CIRP Ann-Manuf Technol 2006;55(1):283–6. [6] Rao H, Kang J, Huff G, Avery K. Impact of specimen configuration on fatigue properties of self-piercing riveted aluminum to carbon fiber reinforced polymer composite. Int J Fatigue 2018;113:11–22. [7] Calabrese L, Proverbio E, Pollicino E, Galtieri G, Borsellino C. Effect of galvanic corrosion on durability of aluminium/steel self-piercing rivet joints. Corros Eng Sci Technol 2015;50(1):10–7.
10
Contents lists available at ScienceDirect
International Journal of Fatigue journal homepage: www.elsevier.com/locate/ijfatigue
Fatigue characterization and crack propagation mechanism of self-piercing riveted joints in titanium plates
T
⁎
Xianlian Zhang, Xiaocong He , Wenjie Wei, Jiawei Lu, Kai Zeng Innovative Manufacturing Research Centre, Kunming University of Science and Technology, Kunming 650500, PR China
A R T I C LE I N FO
A B S T R A C T
Keywords: Self-piercing riveting Titanium Fretting wear Fatigue failure Crack propagation
The fatigue characterization and failure modes of self-piercing riveted joints in titanium plates were studied experimentally. The self-piercing riveting (SPR) process was optimised to improve the forming quality, mechanical property and failure behaviour of titanium joints through the comparative trials. The crack propagation mechanisms and fretting behaviours are quite different for different failure modes and analysed in detail from macro- and micro-perspective. The static and fatigue properties are both enhanced and the failure mode of rivet fracture is avoided for the optimised joints. The failures of upper sheet fracture and lower sheet fracture occur in the fatigue tests. Heavy fretting wear exists precisely at the critical faying interfaces and results in the fatigue crack initiations. The fatigue crack initiates at the faying interface between both sheets for the upper sheet fracture and propagates to both sides of the upper sheet simultaneously. Concerning the lower sheet fracture, the fatigue crack initiates at the faying interface between the rivet and the lower sheet, and then propagates to the close side of the lower sheet in the first stage, extends to the far side in the second stage.
1. Introduction
joining similar or dissimilar metallic and non-metallic sheet materials. Mori et al. [5] investigated the SPR process for joining the ultra high strength steel and aluminium alloy sheets, and the die shape was optimized based on numerical simulations to attain the joining of the hard ultra sheets. Rao et al. [6] presented the fatigue properties of SPR joints in aluminium alloy 6111 and carbon fibre reinforced polymers (CFRP) using lap-shear and cross-tension samples, and the joints with different configurations were produced with two distinctive rivet head heights. Calabrese et al. [7] carried out a galvanic corrosion test on dissimilar SPR joints in steel/aluminium sheet materials, and the relationship among failure mechanism, joint configuration and corrosion damage in salt spray environment was discussed in detail through the long term aging tests. Di Franco et al. [8] investigated the possibility of joining carbon fibre panels with aluminium alloy sheets using SPR and an adhesive layer. The effect of varying distances between the rivets was considered through the static and fatigue tests. Zhang et al. [9] made a discussion on the SPR joints in aluminium-lithium alloy sheets, and the mechanical properties and fretting wear of the joints were obtained and detected experimentally. Porcaro et al. [10] studied the static behaviours of SPR joints in aluminium alloy AA6060 under three different loading directions using an experimental study. Mucha [11] presented an intensive research on the joining of aluminium alloy sheets and discussed the effect of various riveting process parameters on the rivet
With the emphases on the energy consumption in all walks of life, more and more attention has been paid to the application of the lightweight structures. Therefore, the lightweight structures, particularly in lightweight materials, have gradually become a hot topic in current researches. In this work, the lightweight materials of titanium plates have been taken into consideration. Titanium and its alloys have been successfully widely used in aerospace, military, power and chemical industries [1], due to their remarkable characteristics: high specific strength and stiffness, low density, effective deformability and excellent corrosion resistance [2]. However, the welding of titanium is complicated and the inert gases or vacuum shielding is required, which often leads to inferior mechanical properties or hot cracking [3]. Wider applications of the lightweight materials are limited by hard or even impossible to be joined using the conventional joining techniques. Selfpiercing riveting (SPR) is a mechanical cold-forming joining technique, which has been considered as an economical and effective process to join lightweight sheet materials in many industrial fields [4]. It mainly depends on a firm mechanical interlock to fasten the sheet materials with two or more layers. The schematics of the joining process are illustrated in Fig. 1. Generally, it has been proved that SPR possesses the capability of
⁎
Corresponding author. E-mail address: [email protected] (X. He).
https://doi.org/10.1016/j.ijfatigue.2019.105465 Received 8 July 2019; Received in revised form 31 December 2019; Accepted 31 December 2019 Available online 02 January 2020 0142-1123/ © 2020 Elsevier Ltd. All rights reserved.
International Journal of Fatigue 134 (2020) 105465
X. Zhang, et al.
Fig. 1. Schematic illustration of SPR process.
forming characteristics. The mechanical properties and failure modes of the optimised joints in titanium plates were discussed through a contrastive analysis. The fatigue crack propagation and fretting behaviour were intensively analysed from macro- and micro-perspective.
deformation of SPR joints. Xie et al. [12] investigated the cold-formed steel (CFS) shear walls with self-piercing riveted (SPR) connections under monotonic and reversed cyclic loading through a series of experiments. Moraes et al. [13] analysed the influence of the process deformation history on the mechanical performance of dissimilar SPR joints between magnesium and aluminium alloys. Particularly, the joining of similar titanium plates using SPR process was originally implemented with a pre-heating operation on the titanium plates by He et al. [14]. The influence of relief annealing on the static and fatigue performances of the pre-heated joints was carried out experimentally by Zhang et al. [15]. The regular SPR joints in titanium plates without pre-heating were realised later and the effects of sheet thickness and fatigue load on the mechanical property and failure were discussed by Zhao et al. [16]. However, it was found that the mechanical properties of the regular joints were notably lower than those with pre-heating. A majority of the mentioned titanium joints failed due to the rivet fracture in both quasi-static and fatigue tests. The fracture type of the rivet is the brittle fracture, which is quite dangerous without any plastic deformation for the mechanical structure design. Therefore, the corresponding efforts should be made to further improve the mechanical properties and avoid the sudden failure of the SPR joints in titanium plates. In addition, the available literature about the fatigue crack propagation in SPR joints is quite limited. Li et al. [17] discussed the fatigue crack initiation and development of the double rivets lap shear joints in aluminium alloy AA5754, which was based on the local images and SEM microstructures of the fracture interfaces. They thought that the fatigue cracks initiated at the bottom surface of the upper sheet at high load amplitudes, and then developed in transverse direction and along the sheet thickness direction to the top surface of the upper sheet; while at low load amplitudes, the fatigue cracks initiated at the top surface of the lower sheet. Kang and Kim [18] did a brief analysis using the partial enlarged images of the fractures in AA5052 joints and found a similar situation as Li et al. [17] at high-loading range, while the fatigue cracks initiated on the lower sheet near the rivet tail at low-loading range and then propagated along the sheet thickness direction. Moraes et al. [19] and Huang et al. [20] investigated the fatigue crack initiation mainly based on a detailed analysis on the fretting wear in dissimilar aluminium/steel and similar aluminium joints. They discovered that the crack initiation points were quite different for different failure modes. Particularly, it could be found from the previous researches on the mechanical fasteners, the fretting fatigue behaviours potentially took a pivotal role in the fatigue crack initiation that the fatigue cracks typically originated in the severe fretting region at the faying interface of the fastener [20–22]. The fatigue performance was obviously impaired in certain cases due to the fretting wear [23,24]. This work aims to deal with the above circumstances of SPR process for the titanium plates and investigate the fatigue characterization and crack propagation mechanism of corresponding joints in detail. The SPR process was optimised for a better forming structure through the comparative trials. A micro-hardness analysis was performed for the
2. Experimental details 2.1. Materials and process The materials used are 1.5 mm thick plates of commercially pure titanium TA1 and the chemical composition (wt%) is: 0.06 Fe, 0.01 C, 0.01 N, 0.001 H, 0.04 O and Ti balance. All the titanium plates in this work were sheared along the rolling direction to 110 mm × 20 mm × 1.5 mm. A RIVSET VARIO-FC (VTF) hydraulic driven riveting machine was adopted in the riveting experiments, which was fabricated and supplied by Böllhoff GmbH & Co. For joining the stiffer sheet materials, the rivet hardness needs to be enhanced using modified metal heat treatment process, which may increase the brittleness and reduce the shear resistance of the rivet. The enhanced semi-tubular rivets were used to join the 1.5 mm thick titanium plates without pre-heating in the previous research [16]. Particularly, the enhanced semi-tubular rivets were used in this work and possessed the mechanical properties: young’s modulus 206 GPa, poisson’s ratio 0.3, tensile yield strength 1719.7 MPa and compressive strength 1595.6 MPa. The rivets are made of boron steel 36MnB4 and coated with zinc-tin alloy, and the chemical composition (wt%) is: 0.33 C, 0.80 Mn, 0.30 Si, 0.025 S, 0.025 P, 0.30 Cr and Fe balance. Considering these backgrounds, the comparative trials of the joints in titanium plates were carried out. The riveting qualities are evaluated using the cross section visual inspection method based on three structural parameters [4]: rivet head height (h), rivet spread (r) and remaining bottom thickness (t). It is known that the joint strength mainly depends on the mechanical interlock in the SPR structure and the value of rivet spread (r) is the critical indicator. The joints manufactured with varying riveting parameters are evaluated based on the three structural parameters and joint strengths. Finally, the optimised parameters are confirmed and listed as below: riveting pressure 270 bar, rivet hardness 50 ± 2 HRC, rivet length 5.5 mm and die cavity φ9.2 × 1.8 mm. Comparing with the previous research, the riveting pressure was increased by 60 bar; the values of rivet hardness were the same; the rivet length was decreased from 6 mm to 5.5 mm; and the die cavity was changed from φ9.0 × 1.25 mm to φ9.2 × 1.8 mm. The geometrics of the riveting tools are shown in Fig. 2. More than fifty optimised joints were fabricated for the following research. Ten previous joints were made for the static tests, which has not been discussed in the previous research. The dimensions of the specimens were illustrated in Fig. 3. 2.2. Test set-up A bit deeper analysis about the forming structures of the corresponding joints was implemented using a digital micro-hardness tester 2
International Journal of Fatigue 134 (2020) 105465
X. Zhang, et al.
Fig. 2. The riveting tools for the optimised joints (dimensions in mm): (a) punch, (b) rivet, and (c) die.
Fig. 4(a). Concerning the three structural parameters, the rivet head height (h = –0.1 mm) was mainly improved by a higher riveting pressure (270 bar) and a decrease of the rivet length (5.5 mm), which could make the optimised joints have a better sealing performance compared to the previous joints (h = 0.2 mm). The values of rivet spread (r = 1.0 mm) and remaining bottom thickness (t = 0.7 mm) for the optimised joints are superior to those (r = 0.8 mm, t = 0.3 mm) of the previous joints. Due to the increase of die cavity and the decrease of the rivet length, the extrusion of the lower sheet could be assuaged effectively, which could be observed from the extrusion lines marked on the cross sections in Fig. 4. The hardness (HV0.05) distribution on the corresponding cross sections is summarised in Fig. 5. The initial hardness of the plates could be considered as 140 HV, which was measured from the edge of cross section. It can be found that the hardness profiles on the left and right sides of the rivet are different for both joints, which was mainly affected by the inhomogeneity of the plate itself and the mechanical polishing process. The plates of the previous joint are more hardened in the slight deformation region (Nos. 1–3 and 11–13). It is noteworthy that the values of hardness points near the rivet (Nos. 4–10) are higher than those in the slight deformation region. The optimised joint basically presents the same feature. The hardness distribution for the optimised joint presents a shape “M”, and two peak values occur at points No. 4 and 10 on both sheet that are close to the outer surface of the rivet leg. Generally, the hardness values of the previous joint are higher than those of the optimised joint. Especially for the region inside the rivet (Nos. 6–8), the previous joint (203.8 HV for the upper sheet, 218.7 HV for the lower sheet) is obviously superior to the optimised one (169.3 HV for the upper sheet, 193.9 HV for the lower sheet). Consequently, both the upper and lower sheets of the previous joint possess a better work hardening effect comparing with the optimised joint, which means that the plastic deformation of the plates is more serious and the lower plate binds the rivet more tightly in the previous joint.
(HVST-1000Z, China). A loading force of 0.49 N and a holding time of 15 s were adopted for the cross section samples, which were cut along the rivet diameter and then mechanical polished. All the hardness values were measured three times and averaged. The quasi-static tensile-shear tests and fatigue tests were performed on a MTS Landmark100 servo-hydraulic testing machine. Two spacers with equal thickness of 1.5 mm were glued on both ends of the specimens to avoid additional bending, as shown in Fig. 3. A tensile speed of 5 mm/min was adopted and ten specimens were repeated in the quasistatic tests for both previous and optimised joints. A sinusoidal wave at the frequency of f = 10 Hz and a load ratio of r = 0.1 was performed in the fatigue tests using the tension-tension loading mode. The fatigue loads were chosen based on the static strength for reflecting the entire fatigue lives. Five specimens were tested at each different loads. Attaining over 2 million cycles or the specimen failure was regarded as the termination criterion. All the tests were carried out at a room temperature. Particularly, the fatigue data of the previous joints by Zhao et al. [16], which were obtained under the same testing condition, were cited in this work. The fracture features of the typical failed specimens and the micro morphologies of the typical fracture surfaces were observed and analysed using a scanning electron microscopy (SEM, TESCAN Inc., Czech Republic). The effort was focused on trying to illuminate the fatigue crack initiations and propagation stages in more detail using SEM technique, and the SEM images of the whole macro fracture were stitched from more than a dozen pictures so as to illustrate the rough trend and direction of the fatigue cracks. Furthermore, the fretting debris in the fretting region was detected using an energy dispersive X-ray spectroscopy (EDS, EDAX Inc., USA).
3. Results and discussion 3.1. Forming characteristics The cross sections of the previous and optimised joints in titanium plates are presented in Fig. 4. It is notable that there is an obvious gap between the rivet head and the upper sheet for the optimised joints in Fig. 4(b), while the gap is much smaller due to a rivet length of 6 mm in
3.2. Static property and failure Fig. 6 shows the results of the joints in titanium plates in quasi-static tests. For evaluating the validity of the data, the Grubbs criterion was
Fig. 3. Schematic illustration of the investigated specimens (dimensions in mm). 3
International Journal of Fatigue 134 (2020) 105465
X. Zhang, et al.
Fig. 4. The cross sections and marked hardness point Nos. of the joints (dimensions in mm): (a) for the previous joints, and (b) for the optimised joints.
rivet tail and the lower sheet. It is worthy to note that the mechanical interlock of the optimised one was totally damaged under tensile-shear loading. The failure mode was changed from the rivet fracture to the rivet tail pulled-out using the optimised parameters, which could help prevent sudden failure from a practical standpoint. It is thought that the previous joints are not prone to plastic deformation under tensile-shear loading because of a better work hardening effect on the plates, which indicates that the interlock structure is firmer compared to the optimised joints. As a consequence, the rivet of the previous joints bore more shear loads during quasi-static tests, which allowed the previous joints to fail in the rivet breakage.
used to eliminate the outliers, and it was found that there was no outlier for both joints. Intuitively, the uniformity of load-displacement curves for the optimised joints is quite well, as shown in Fig. 6(b). After calculation, the average lap shear strength and energy absorption of the optimised joints are 7.19 kN and 30.84 J (Standard deviation: 0.18 kN and 1.95 J) respectively, and 5.41 kN and 4.09 J (Standard deviation: 0.53 kN and 1.62 J) are for the previous joints. The optimised joints possessed a better static property, which was even superior to those with pre-heating or relief annealing by He et al. [14] and Zhang et al. [15]. In addition, the load-displacement curve of the optimised joints (Fig. 6(b)) could be divided into three stages. In the 1st stage, the joint bore the increasing shear load and the peak load occurred at the displacement of 2 mm. The rivet started to be pulled out from the lower sheet and the upper sheet was gradually bended under tensile-shear loading in the 2nd stage. In the 3rd stage, it was the final failure of the joint with the load dropping rapidly. For the previous ones, only the 1st and 3rd stage can be observed in Fig. 6(a) due to the failure mode of rivet fracture. Fig. 7 presents the failure modes in quasi-static tests and the schematic illustrations. For the previous joints, the failure mode of rivet fracture occurred for all specimens under tensile-shear loading. As shown in Fig. 7(a), the upper sheet was slightly deformed during the quasi-static test, while the lower sheet including the joint bottom was barely deformed. It is noticed that the separated part of the upper sheet was left in the interlock structure due to the action of the broken rivet tail, as marked in the schematic illustration. The brittle rupture features was detected on the rivet fracture surface using SEM technique, which led to the sudden failure of the previous joints in quasi-static tests. For the optimised joints in Fig. 7(b), all specimens failed due to the rivet tail pulled-out from the lower sheet, which is consistent with the previous researches by Sun et al. [25] and Porcaro et al. [10]. The separated part of the upper sheet was also pulled out for the action of the rivet leg. Both sheets presented obvious deformation during the separation of the
3.3. Fatigue property and failure The fatigue test results are summarised in Table 1. Six fatigue load amplitudes were adopted to evaluate the fatigue life of the optimised joints. The dispersion of fatigue cycles for different joints increases with the decrease of fatigue load amplitudes. Besides, two failure modes occurred in fatigue tests: failure mode A: upper sheet fracture and failure mode B: lower sheet fracture, which were quite different from the previous joints. For a comparison analysis, the fatigue data and the fatigue load-fatigue life (F-N) curves were plotted in Fig. 8 for both joints. It is noticed that three interesting regions are presented for the optimised joints: Region Ⅰ: mainly failure mode A; Region Ⅱ: mainly failure mode B; Region III: no failure over 2 million cycles. The region distribution could help to predict the fatigue lives and failure modes of the optimised joints at different fatigue loads. However, it is found that the previous joints mainly failed in the lower sheet when the fatigue cycles were between 105 and 106, while failed in the rivet when the fatigue cycles were below 105 or over 106. The distribution feature of the fatigue data is obviously different for the previous and optimised joints. In addition, it is observed from Fig. 8, while attaining the same fatigue cycles, the fatigue load amplitudes of the optimised joints are
Fig. 5. The Vickers hardness for the cross sections. 4
International Journal of Fatigue 134 (2020) 105465
X. Zhang, et al.
Fig. 6. Quasi-static lap shear test results for the joints in titanium plates: (a) load-displacement curves for the previous joints, and (b) load-displacement curves for the optimised joints.
recorded by a digital camera (Sony ILCE-A7RM2). As shown in Fig. 10(a), the fatigue crack firstly appeared at the ring of the rivet head and extended from the pierced hole to both sides of the upper sheet. In Fig. 10(b), the fatigue crack was observed in the lower sheet and near the joint bottom. It is thought that the fatigue crack was initiated near the riveted part and extended along the sheet width direction. Meanwhile, the cross sections of the failed specimens were fabricated to explore the failure paths along the sheet thickness direction in both plates, and the partial enlarged drawings of cross sections D1 and D2 are presented in Fig. 10(c) and (d), respectively. The upper sheet fracture path occurred underneath the edge of the rivet head for failure mode A, which is consistent to the conclusion by Khanna et al. [27]. The lower sheet fracture path existed under the rivet tail, which was the region of the remaining bottom.
higher than that of the previous joints. The F-N curve of the optimised joints is completely above that of the previous joints at all fatigue phases. Consequently, it means that the fatigue property of the titanium joints is improved significantly with the optimised parameters. Based on the above analysis, the previous joints had a more serious working harden effect on the lower sheet and an inferior remaining bottom thickness of 0.3 mm comparing with the optimised joints. This indicates that the lower sheet and rivet of the previous joint are hard to deform effectively and easily destroyed under cyclic fatigue loading, which made the optimised joints failed in the lower sheet exhibit the higher fatigue lives than the previous joints failed in the same type of lower sheet. Particularly, due to the modified forming structure obtained using the optimised parameters, the failure modes of rivet fracture were successfully avoided for the optimised joints, and the failure modes were changed from the rivet fracture to the upper sheet fracture. Fig. 9 presents the failures at different fatigue load amplitudes of the optimised joints in fatigue tests, and the failures are similar to the results from Iyer et al. [26]. It is worthy to note that the appearances of the specimens in failure mode A seem quite similar and all the specimens are fractured close to the rivet in Fig. 9(a). The visible plastic deformation could be observed on both sides (near the sheet edges) of the fractured plates. For failure mode B, the appearances are a bit different and the visible cracks (marked in red arrow in Fig. 9(b)) present in the upper sheet at high loads of 3.60 kN and 2.88 kN, which may be caused by the differential stress distribution on the faying interface. Unlike the failure mode A, the visible plastic deformation only occurred on one side of the fractured plates. Further, the locations of fatigue cracks during the fatigue tests were
3.4. Fatigue crack propagation Considering the failure mode A, Fig. 11 presents the characteristics of the typical upper sheet fracture surface. The corresponding enlarged views of the marked parts in Fig. 11a are shown in Fig. 11(b)–(e), respectively. In terms of the morphological features, the fracture surface could be divided into three zones: fatigue source (zone i), fatigue crack propagation (zone ii) and final fracture (zone iii), which have been marked in the enlarged views. The river patterns on the fracture surface indicate the fatigue crack initiation location in zone i. The fatigue crack initiated on the faying surface of the upper sheet and propagated outward as the yellow arrow marked in Fig. 11(d). Fig. 12(c) presents the micro morphology of the crack initiation location. The cleavage step
Rivet tail
Joint bottom
Rivet
Separated part
(a)
Separated part
(b)
Fig. 7. Quasi-static failure modes: (a) rivet fracture for the previous joints, and (b) rivet tail pulled-out for the optimised joints. 5
International Journal of Fatigue 134 (2020) 105465
X. Zhang, et al.
Table 1 Fatigue test results of the optimised joints. Specimen Nos.
Fatigue load amplitude/kN
Fatigue cycles
Failure mode
Specimen Nos.
Fatigue load amplitude/kN
Fatigue cycles
Failure mode
01 02 05 20 28 03 04 11 12 26 07 09 19 23 24
3.60 3.60 3.60 3.60 3.60 2.88 2.88 2.88 2.88 2.88 2.52 2.52 2.52 2.52 2.52
54,923 71,447 75,724 80,352 73,096 198,723 215,729 154,960 174,313 145,010 368,039 317,213 383,087 205,165 304,332
A A A A B B A A B A B A A A A
08 10 13 22 25 16 17 18 21 27 14 15
2.16 2.16 2.16 2.16 2.16 1.94 1.94 1.94 1.94 1.94 1.80 1.80
591,330 707,532 601,146 896,506 1,082,870 1,029,673 1,381,648 1,297,992 924,917 1,051,983 2,172,564 2,269,167
A B B B B B B B B B No failure No failure
and both lines lean towards the crack initiation location. This indicates that the fatigue cracks propagate faster in the sheet width direction than in the sheet thickness direction, which is consistent with the standpoint by Li et al. [28]. Considering the failure mode B, the features of the lower sheet fracture surface are presented in Fig. 13, and Fig. 13(b)–(e) are the enlarged views of the marked parts in the macro fracture in Fig. 13(a), respectively. The three zones as the upper sheet fracture also present on the lower sheet fracture. In the fatigue source region (zone i), a ridge could be observed at the crack initiation location, marked with a blue arrow in Fig. 13(d). It means that the fatigue crack initiates at the faying interface between the rivet and the lower sheet and near the bending point of the lower sheet, and then propagates along the direction marked with a yellow arrow. The micro morphology of the marked area in a red rectangle is presented in Fig. 14(c). The tiny fatigue striations and step patterns present the transgranular fracture in the fatigue source region. The crack propagation direction is indicated in zone ii, marked with the yellow arrows. The fatigue striations and the river marks can be seen in a more obvious perspective than in the fatigue source region. However, the appearances of zone ii look quite different for the far side (Fig. 13(b) and (c)) and close side (Fig. 13(e)) of zone i, and their micro morphology is shown in Fig. 14(b) and (d), respectively. The fatigue striations and ridges in Fig. 14(d), which are the features of the cleavage fracture, indicate that the first stage of the fatigue crack propagation is to extend to the close side of zone i, as the visible initial crack indicated in Fig. 10(b). The river marks and secondary cracks in Fig. 14(b) proved that the joint bottom failed due to the transgranular fracture. The far side of zone i is the second stage of the fatigue crack propagation. Finally, the completed fracture occurred when the residual structure of the joint could no longer bear the continuous cyclic fatigue loading. The micropores and dimples in the microstructure (Fig. 14(a)) indicated that the residual structure of the joint failed due to the ductile fracture. Besides, the crack propagation
Fig. 8. Fatigue load-fatigue life (F-N) curves for the joints in titanium plates.
patterns indicated the fatigue crack initiated through the grains. In zone ii, the fatigue striations and river patterns point out the crack propagation direction, marked with the yellow arrows. The micro morphology of zone ii is presented in Fig. 12(b) and (d) which are located at the left and right side of the crack initiation location, respectively. The morphological features seem to be quite similar for the both sides. This indicated that the fatigue crack propagated to zone ii simultaneously, as the fatigue crack indicated in Fig. 10(a). The typical fatigue striations and secondary cracks can be seen obviously in zone ii, which is the transgranular fracture. Fig. 11(a) and (d) present the micro fracture features of the final fracture in zone iii caused by the loading amplitude exceeded the fracture limit of the residual structure of the joint. From the micro morphology in Fig. 12(a), many dimples and tearing ridges could be found, which shows the typical ductile fracture feature in zone iii. Particularly, the crack propagation front lines can be observed clearly on both sides of the upper sheet fracture in Fig. 11(b) and (e),
Deformation
Reverse side 3.60 kN
2.88 kN
Deformation
Crack
2.52 kN
2.16 kN
(a)
Crack 3.60 kN 2.88 kN 2.52 kN
2.16 kN 1.88 kN
(b)
Fig. 9. Fatigue failures for the optimised joints: (a) the fatigue failure mode A, and (b) the fatigue failure mode B. 6
International Journal of Fatigue 134 (2020) 105465
X. Zhang, et al.
D1
Crack
Joint bottom Upper sheet
Rivet Rivet
Crack
(a)
Rivet
Upper sheet Lower sheet Crack Lower sheet
D2
Crack
(b)
(c)
(d)
Fig. 10. Fatigue failure cracks for the optimised joints: (a) crack in the upper sheet, (b) crack in the lower sheet, (c) fracture path for failure mode A, and (d) fracture path for failure mode B.
length at high load amplitude of 3.60 kN is shorter than those at the other amplitudes, which could be caused by the fatigue crack grown faster at high load amplitude. However, on the lower sheet fracture (failure mode B), the length at low load amplitude of 1.94 kN is shorter than those at the other amplitudes, which is considered to be affected by a lower fatigue crack growth rate at low load amplitude. This indicates that the fatigue crack growth rate plays an important role on the crack growth region length. Due to the obvious plastic deformation on the lower sheet and different propagation stages on the lower sheet fracture, the crack growth region length decreases with the fatigue load increasing, which is totally in contrast with the upper sheet fracture. In addition, the crack growth region lengths are generally higher on the upper sheet fracture than on the lower sheet fracture, which could be also caused by the differences in the fatigue crack propagation stages and the plastic deformation.
front lines also prove that the fatigue cracks propagate faster in the sheet width direction than in the sheet thickness direction. As the discussion above, the fatigue crack initiated at the faying interface between both sheets, and simultaneously propagated outward to both sides of the upper sheet for the failure mode A; while the fatigue crack initiated at the faying interface between the rivet and the lower sheet, and firstly propagated to the close side of the lower sheet, then extended to the far side of the lower sheet for the failure mode B. Until the residual structure of the corresponding plate could not withstand the fatigue load amplitude, the final fracture happened. The schematic diagrams of the fatigue crack propagation are illustrated in Fig. 15, and the propagation directions are indicated using the yellow arrows. Further, it shows during the fatigue tests that the fatigue source zone i and fatigue crack propagation zone ii sustain most of the fatigue cycles. The length of the crack growth region, included zone i and zone ii, is measured using an industrial microscope (ZW-H3000, China) with a measuring software. The detailed values are listed in Table 2, and Fig. 16 presents the average crack growth region length at different load amplitudes. On the upper sheet fracture (failure mode A), the
3.5. Fretting fatigue behaviour The fretting fatigue behaviour existed precisely and was highly
(a)
Crack propagation front line
(b)
(c)
Crack initiation location
Crack propagation front line
(d)
(e)
Fig. 11. Fatigue failure analysis for failure mode A: (a) macro view of the upper sheet fracture (b) enlarged view of Part 1, (c) enlarged view of Part 2, (d) enlarged view of Part 3, and (e) enlarged view of Part 4. 7
International Journal of Fatigue 134 (2020) 105465
X. Zhang, et al.
Striations Striations Dimples
Step pattern Secondary crack
Ridges Secondary crack
(a)
(b)
(c)
(d)
Fig. 12. Micro morphology at different zones marked with red rectangles for failure mode A: (a) at zone iii marked in Fig. 11b, (b) at zone ii marked in Fig. 11c, (c) at zone ii marked in Fig. 11d, and (d) at zone ii marked in Fig. 11e. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
(marked with the dot-dashed line), the constituent of the fretting debris could be inferred to be one of the titanium oxides. Therefore, it is considered that due to the fretting wear in the faying interface between both sheets, the elements on the sheet surface are scratched out and oxidized. The fatigue crack initiate in the fretting region S at the bottom surface of the upper sheet fracture because of the continuous cyclic loads and micro-vibrations. According to the previous analysis for failure mode B, the typical rivet of the failed joints is detected using micro test methods, and the upward view of the rivet is presented in Fig. 18(b). There is no any signs of wear on the bottom surface of rivet tail, and the separated part of the upper sheet is unharmed. It means that for the forming structure of the optimised joints, no fretting occur in the contact interface between the bottom of rivet tail and the lower sheet. The front view of the rivet is shown in Fig. 18(c). The fretting debris, especially area β, could be observed on the fretting region R. Probably, this is the reason of the fatigue crack initiation, which is consistent with the conclusions by Sun et al. [25]. Besides, Fig. 18(d) shows the enlarged view of area γ on the
correlated with the fatigue failures of the optimised joints. Due to the optimised riveting process, the fretting behaviour of the optimised joints is quite different from that of the previous joints which has been discussed in Zhao et al. [16]. The fracture paths and fretting regions in different fatigue failure modes for the optimised joints are illustrated and marked in Fig. 17 for a better understanding. Particularly, the fretting regions R (rivet) and S (sheets) indicate the faying interfaces between the rivet tail and lower sheet and between both sheets, respectively. It was found that the heavy fretting and fracture path always occurred at the side along the direction of cyclic loading for all failed optimised joints. Fig. 18(a) presents the upward view of the upper sheet fracture (failure mode A). The fretting phenomenon near the fatigue crack initiation point and the fatigue crack in fretting region S can be observed clearly. Further, the area α at the edge of the fretting region was analysed using EDS, as shown in Fig. 18(e). The energy spectrum of area α contains some Ti, O, N and other elements in an infinitesimal quantity. Concerning the distributions of the fretting debris and oxygen element
(a)
Crack propagation front line
(b)
(c)
Crack initiation location
Crack propagation front line
(d)
(e)
Fig. 13. Fatigue failure analysis for failure mode B: (a) macro view of the lower sheet fracture, (b) enlarged view of Part 1, (c) enlarged view of Part 2, (d) enlarged view of Part 3, and (e) enlarged view of Part 4. 8
International Journal of Fatigue 134 (2020) 105465
X. Zhang, et al.
Ridges
Striations Striations Secondary crack Step pattern
Micropores
(a)
(b)
(c)
(d)
Fig. 14. Micro morphology at different zones marked with red rectangles for failure mode B: (a) at zone iii marked in Fig. 13b, (b) at zone ii marked in Fig. 13c, (c) at zone i marked in Fig. 13d, and (d) at zone ii marked in Fig. 13e. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
Final fracture
Crack initiation
Final fracture
Crack growth region length Crack initiation
Final fracture
Crack growth region length
(a)
Final fracture
(b)
Fig. 15. Schematic diagrams for the macro fractures in different failure modes: (a) the upper sheet fracture surface, and (b) the lower sheet fracture surface. Fracture path for failure mode A
Table 2 Statistics for the fatigue crack growth region length at different load amplitudes. The upper sheet fracture (Failure mode A) Specimen Nos.
Cyclic load
The lower sheet fracture (Failure mode B)
Fatigue load amplitude/ kN
Crack growth region length/ mm
Specimen Nos.
3.60 3.60 3.60 3.60 2.88 2.88 2.88 2.52 2.52 2.52 2.52 2.16
15.04 15.06 15.47 15.11 16.41 16.20 16.39 16.11 16.68 16.66 16.80 16.64
28 03 12 07 10 13 22 25 16 17 18 21 27
Fatigue load amplitude/ kN
Crack growth region length/ mm
3.60 2.88 2.88 2.52 2.16 2.16 2.16 2.16 1.94 1.94 1.94 1.94 1.94
15.41 15.35 15.72 14.54 15.94 15.68 15.10 15.44 14.46 14.43 14.77 14.41 14.22
Cyclic load Fretting region S Fretting region R Fracture path for failure mode B
Fig. 17. Schematic illustration for the fatigue failures of the optimised joints. 01 02 05 20 04 11 26 09 19 23 24 08
lower sheet fracture which belongs to the fatigue crack propagation zone (Fig. 13(e)). The obvious fretting layer, thickness over 50 μm, exists on the sheet surface. It is thought that due to the continuous cyclic loads and micro-vibrations between both sheets, the fretting wear on the lower sheet fracture is generated in the course of the fatigue crack propagation, and it accelerates the propagation of the fatigue cracks. 4. Conclusions The paper is highlighted to the mechanical improvement, fatigue failure, crack propagation and fretting behaviour of self-piercing riveted joints in titanium plates. The main conclusions are summarised as following: 1. The forming feature of self-piercing riveting in titanium plates can be optimised in this work using a riveting pressure of 270 bar, rivet length of 5.5 mm and rivet hardness of 50 ± 2 HRC; 2. The static and fatigue properties are both improved effectively, and the rivet fracture in the previous joints is avoided in the optimised joints. The failure in the quasi-static tests presents the pull-out of the rivet tail from the lower sheet, and two failure modes occurred in the fatigue tests: upper sheet fracture and lower sheet fracture. 3. For the upper sheet fracture, the fatigue crack initiates at the faying interface between both sheets, and then propagates to both sides of the upper sheet simultaneously. For the lower sheet fracture, the fatigue crack initiates at the faying interface between the rivet and the lower sheet, and propagates to the close side of the lower sheet at the first stage, extends to the far side at the second stage. The final
Fig. 16. Average crack growth region length of the fractures at different fatigue load amplitudes.
9
International Journal of Fatigue 134 (2020) 105465
X. Zhang, et al.
Fig. 18. Fretting in the optimised joints: (a) upward view of the upper sheet fracture for failure mode A, (b) upward view of the rivet for failure mode B, (c) front view of the rivet for failure mode B, (d) enlarged view of area γ in Fig. 13(e) for failure mode B, and (e) energy spectrum of area α in Fig. 18(a) for failure mode A.
fracture occurs until the residual structure could no longer bear the continuous fatigue loading. 4. The fretting fatigue behaviours occurred precisely at the faying interfaces between the rivet tail and lower sheet and between both sheets where the fatigue crack initiated. The fretting was highly correlated with the fatigue failures for self-piercing riveted joints.
[8] Di Franco G, Fratini L, Pasta A. Influence of the distance between rivets in self-piercing riveting bonded joints made of carbon fiber panels and AA2024 blanks. Mater Des 2012;35:342–9. [9] Zhang X, He X, Gu F, Ball A. Self-piercing riveting of aluminium–lithium alloy sheet materials. J Mater Process Technol 2019;268:192–200. [10] Porcaro R, Hanssen A, Langseth M, Aalberg A. An experimental investigation on the behaviour of self-piercing riveted connections in aluminium alloy AA6060. Int J Crashworth 2006;11(5):397–417. [11] Mucha J. A study of quality parameters and behaviour of self-piercing riveted aluminium sheets with different joining conditions. Stroj Vestn-J Mech Eng 2011;57(4):323–33. [12] Xie Z, Yan W, Yu C, Mu T, Song L. Experimental investigation of cold-formed steel shear walls with self-piercing riveted connections. Thin Wall Struct 2018;131:1–15. [13] Moraes JFC, Jordon JB, Su X, Barkey ME, Jiang C, Ilieva E. Effect of process deformation history on mechanical performance of AM60B to AA6082 self-pierce riveted joints. Eng Fract Mech 2019;209:92–104. [14] He X, Wang Y, Lu Y, Zeng K, Gu F, Ball A. Self-piercing riveting of similar and dissimilar titanium sheet materials. Int J Adv Manuf Technol 2015;80(9–12):2105–15. [15] Zhang X, He X, Xing B, Zhao L, Lu Y, Gu F, et al. Influence of heat treatment on fatigue performances for self-piercing riveting similar and dissimilar titanium, aluminium and copper alloys. Mater Des 2016;97:108–17. [16] Zhao L, He X, Xing B, Zhang X, Cheng Q, Gu F, et al. Fretting behavior of self-piercing riveted joints in titanium sheet materials. J Mater Process Technol 2017;249:246–54. [17] Li D, Han L, Thornton M, Shergold M. Influence of rivet to sheet edge distance on fatigue strength of self-piercing riveted aluminium joints. Mater Sci Eng A 2012;558:242–52. [18] Kang SH, Kim HK. Fatigue strength evaluation of self-piercing riveted Al-5052 joints under different specimen configurations. Int J Fatigue 2015;80:58–68. [19] Moraes JFC, Rao HM, Jordon JB, Barkey ME. High cycle fatigue mechanisms of aluminum self-piercing riveted joints. Fatigue Fract Eng Mater Struct 2018;41(1):57–70. [20] Huang L, Bonnen J, Lasecki J, Guo H, Su X. Fatigue and fretting of mixed metal selfpiercing riveted joint. Int J Fatigue 2016;83:230–9. [21] Chen Y, Han L, Chrysanthou A, O’Sullivan J. Fretting wear in self-piercing riveted aluminium alloy sheet. Wear 2003;255:1463–70. [22] Han L, Chrysanthou A, O’Sullivan J. Fretting behaviour of self-piercing riveted aluminium alloy joints under different interfacial conditions. Mater Des 2006;27:200–8. [23] Szolwinski M, Farris T. Linking riveting process parameters to the fatigue performance of riveted aircraft structures. J Aircraft 2000;37:130–7. [24] Wagle S, Kato H. Ultrasonic detection of fretting fatigue damage at bolt joints of aluminum alloy plates. Int J Fatigue 2009;31:1378–85. [25] Sun X, Khaleel MA. Strength estimation of self-piercing rivets using lower bound limit load analysis. Sci Technol Weld Joi 2005;10(5):624–35. [26] Iyer K, Hu SJ, Brittman FL, Wang PC, Hayden DB, Marin SP. Fatigue of single- and doublerivet self-piercing riveted lap joints. Fatigue Fract. Eng. Mater Struct 2005;28(11):997–1007. [27] Khanna SK, Long X, Krishnamoorthy S, Agrawal HN. Fatigue properties and failure characterisation of self-piercing riveted 6111 aluminium sheet joints. Sci Technol Weld Joi 2006;11(5):544–9. [28] Li D, Han L, Thornton M, Shergold M, Williams G. The influence of fatigue on the stiffness and remaining static strength of self-piercing riveted aluminium joints. Mater Des 2014;54:301–14.
Declaration of Competing Interest The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper. Acknowledgements This work was supported by the National Natural Science Foundation of China (Grant No. 51565023). The authors would like to thank Mr. Yun Lei from Faculty of Materials Science and Engineering of Kunming University of Science and Technology for his support during micro-hardness testing. References [1] Bishoyi B, Sabat R, Sahu J, Sahoo S. Effect of temperature on microstructure and texture evolution during uniaxial tension of commercially pure titanium. Mater Sci Eng A 2017;703:399–412. [2] Gurao N, Kapoor R, Suwas S. Deformation behaviour of commercially pure titanium at extreme strain rates. Acta Mater 2001;59:3431–46. [3] Sridharan N, Wolcott P, Dapino M, Babu S. Microstructure and texture evolution in aluminum and commercially pure titanium dissimilar welds fabricated using ultrasonic additive manufacturing. Scripta Mater 2016;117:1–5. [4] Haque R. Quality of self-piercing riveting (SPR) joints from cross-sectional perspective: a review. Arch Civ Mech Eng 2018;8(1):83–93. [5] Mori K, Kato T, Abe Y, Ravshanbek Y. Plastic joining of ultra high strength steel and aluminium alloy sheets by self piercing rivet. CIRP Ann-Manuf Technol 2006;55(1):283–6. [6] Rao H, Kang J, Huff G, Avery K. Impact of specimen configuration on fatigue properties of self-piercing riveted aluminum to carbon fiber reinforced polymer composite. Int J Fatigue 2018;113:11–22. [7] Calabrese L, Proverbio E, Pollicino E, Galtieri G, Borsellino C. Effect of galvanic corrosion on durability of aluminium/steel self-piercing rivet joints. Corros Eng Sci Technol 2015;50(1):10–7.
10