Three-Point Bending Fatigue Test of TiAl6V4 Titanium Alloy at Room Temperature

A polycrystalline alpha-beta TiAl6V4 alloy in the annealed condition was used for the three-point bending fatigue test at frequency f∼100Hz. *e static preload Fstat. � − 15 kN and variable dynamic force Fdyn. � − 7 kN to − 13.5 kN were set as fatigue test loading parameters. *e fatigue life S-N curve presented the stress amplitude σa as a function of a number of cycles to fracture Nf. A limiting number of cycles to run out of 2.0×10 cycles were chosen for the 3-point fatigue tests of rectangular specimens. In addition, the Smith diagram was used to predict the fatigue life. *e alpha lamellae width has a significant influence on fatigue life. It is assumed that the increasing width of alpha lamellae decreases fatigue life. A comparison of fatigue results with given alpha lamellae width in our material to the results of other researchers was performed. *e SEM fractography was performed with an accent to reveal the initiation sites of crack at low and high load stresses and mechanism of crack propagation for the fatigue part of fracture.


Introduction
e TiAl6V4 alloy has a relatively high resistance in cyclic loading [1][2][3][4]. However, the course of fatigue damage, however, depends on the content of the additive elements, microstructure, surface treatment, and size and type of applied stresses. e fatigue strength value for smooth samples is more than 50% of the tensile strength, but it is largely dependent on surface quality. e greater the roughness of the surface, or the surface saturated with oxygen or nitrogen, the worse the fatigue properties of titanium. e presence of a notch causes a fatigue reduction of 30-35%. e fatigue range for titanium alloys is reached at 10 6 -10 7 cycles, but it is dependent on the load frequency [5].
Generally, the grain size of materials is a very important characteristic of fatigue life. Leyens and Peters in their work [2] present results that the reduction of grain size from 110 to 6 μm for commercially pure titanium increases the fatigue strength from 160 to 210 MPa. Likewise, the fatigue strength increases after increasing strength via work hardening. In addition, Gao et al. [6] and Moussaoui et al. [7] have performed an intensive study about influence α grain size, degree of age hardening, and oxygen content on fatigue life of TiAl6V4. ey show that the fatigue properties of twophase near − α and α + β alloys are strongly influenced by the morphology and arrangements of the two phases α and β (β grain size, colony size of α and β lamellae, and the width of α lamellae in lamellar microstructures). Another research was performed by Peters et al. and Wagner et al. [8,9] report that reducing off the α lamellae width from 10 to 0.50 μm in lamellar microstructures raises the fatigue strength from 480 to 675 MPa. ey also show that reducing the grain size from 12 to 2 μm in equiaxed microstructures increases fatigue strength from 560 to 720 MPa [8]. For duplex structures, reducing the α lamellae width in a lamellar matrix from 1 to 0.5 μm leads to an increase in fatigue strength from 480 to 575 MPa [9].
For the fatigue process as itself, it is a well-known fact that most of the fatigue cracks initiate at free surface (due to surface roughness, oxide presence, or carbide particles). e fatigue process may also start from so-called "fish-eye," what is a typical structural phenomenon when fatigue crack initiates close to free surface on inclusions (oxide or carbide particles).
is phenomenon occurs only in special cases when a very high-cycle fatigue test (over 10 9 cycles to failure) is performed at low-stress amplitude. Zuo et al. [10] have confirmed this phenomenon on this type of alloy with bimodal and basket-weave microstructure. is kind of initiation was not confirmed at low-cycle or high-cycle fatigue of TiAl6V4 alloy. Wagner and Lütjering discussed three various fatigue crack initiation sites in their work [11]. ey have shown that fatigue crack initiation sites are related to microstructure. e first type of the fatigue crack initiation site is for lamellar microstructure where fatigue cracks initiate at slip bands of the α lamellae as itself or at connection areas between α lamellae and primary β grain boundaries. e α lamellae width has significant effect on slowing down dislocation movement which is closely related to fatigue crack initiation, and on the contrary, it has influence on fatigue strength and yield stress too. e second type of the fatigue crack initiation site is for equiaxed structures. e fatigue cracks initiate in concert with slip bands inside α grains at these types of microstructures. erefore, fatigue strength is affected by the grain size and on grain size is yield stress dependent as well. e third type of the fatigue crack initiation site is for duplex structures. e fatigue cracks may initiate in the lamellar matrix (interface the lamellar matrix/the primary α phase), or in the primary α phase as itself. Kuhlman [12] showed that the fatigue crack initiation site also depends on the cooling rate. Boyer and Puschnik et al. [13,14] in their works have discussed the effect of the volume fraction and size of the primary α phase on fatigue crack initiation sites. e influence of microstructure on fatigue properties of Ti6Al4V alloys at high-cycle fatigue is discussed in the works of Wu et al. [15] and Crupi et al. [16]. eir results prove the fact that various microstructures decrease (from bimodal, lamellar to equiaxed) high-cycle fatigue strength. About Ti6Al4V bimodal structures, the high-cycle fatigue strength is strongly affected by the primary α phase volume and grain size. In the initial stage, the high-cycle fatigue strength increases. However, it gradually decreases with increasing amount and size of the primary alpha phase. A similar effect of decreasing high-cycle fatigue strength can also be observed in the case of equiaxed or lamellar microstructures of Ti6Al4V alloy. Essentially, it can be generalized that the fatigue strength of the Ti6Al4V alloy decreases, either due to the increase in the alpha-phase grain or by increasing the width of the alpha-phase lamellae.
Another factor affecting the fatigue strength of the Ti6Al4V alloy is the load frequency. is issue was studied by Furuya and Takeuchi [17]. Based on their work, it is possible to state that the results of fatigue tests at ultrasonic frequency are similar, respectively, comparable to results performed at 100 Hz load frequency. However, this is only true if the fatigue crack has initiated below the alloy surface. In the case of the Ti6Al4V alloy, when the fatigue crack initiated on the surface, it was shown that the fatigue strength was higher in the tests performed at the ultrasonic frequency than in the conventional 100 Hz frequency. e mean stress effect on fatigue strength is commonly evaluated by a modified Goodman's or Smith diagram. For the Ti6Al4V alloy, according to modified Goodman's (or Smith) diagram, the fatigue strength is considered safe when reaching 10 7 at conventional load frequencies.
Morrissey and Nicholas [18] also studied the impact of load frequency on strain rate or temperature increase due to internal damping. ey compare data from ultrasonic tests, servohydraulic test systems (∼60 Hz), and electromagnetic shaker systems (400 Hz). According to their data comparison, there are no frequency effects. e comparison of S-N results at ultrasonic and conventional frequencies is shown in Figure 1. e aim of this paper is to provide information about fatigue resistance of the TiAl6V4 alloy in the annealed condition at the three-point bending loading test at frequency f∼100.0 Hz [19,20] with the run-out number of cycles N f in range 10 6 to 2.0 × 10 7 and to compare it with results in references [8,9,15] to show how the bending load and microstructure (α lamellae length and width or primary α (α p ) grain size) affect the fatigue life.
ere is an assumption that bending load should decrease fatigue life of alloy due to more complex stress course in the specimen. e fatigue test frequency influence or sample heating due to internal damping at 100 Hz is not expected according to references [17,18].

Materials and Methods
e α + β mixed TiAl6V4 (GRADE 5) titanium alloy was used as experimental material. e alloy chemical composition and selected mechanical properties according to the certificate of quality and weight (the alloy was supplied by BIBUS Metals AG, CZ, with heat No. HX-032) are given in Table 1. e alloy was in the annealed condition. Boyer had described in his work [13,21] four common heat-treatments for α + β alloys, namely, Ti6Al4V. According to its work, the most suitable treatment for increasing the fatigue properties and achieving the reasonable fatigue crack growth is mill annealing (MA or A). After this treatment, the strength of about 896 MPa and moderate fracture toughness 66 MPa·m − 2 are achieved. e microstructure of experimental material is shown in Figure 2. It consists of α-phase lamellae which is considered as the hexagonal close-packed phase (HCP) presented at lower temperatures (up to 886°C) situated in β grains (body centred cubic, BCC, for temperatures from 886°C to melting temperature 1660°C) (Figure 2(a)). e arrangement of α-phase lamellae is presented more in detail in Figure 2(b), and as is obvious, it creates the "envelop" of aluminium solid solution in the base titanium matrix.
Samples for the three-point bending test were supplied by BIBUS METALS AG. Samples were cut to 11 × 10 × 50 mm blocks. e specimen surface was sanded using a LaboPol-25 double-disk grinder, where the sample was ground on the grinding disk with a grain size of 400 (grinding in direction 1) and 600 (grinding in a direction perpendicular to direction 1) at 250 rpm. e sample was next rinsed in warm water and alcohol and dried after each step.
Two single-disk MTH polishers were used for polishing. In the first step, diamond paste D2 with a diamond grain size of 2 μm was added to the sample using a Mol roll at 300 rpm. e disk was moistened with alcohol during polishing, and the sample was polished against the counterclockwise rotation of the disc. In the second step, the D07 paste with a diamond grain size of 0.7 μm was applied using a Nap spin at 300 rpm. e disk was moistened with alcohol during polishing, and the sample was polished against the counterclockwise rotation of the disc. e sample was rinsed in warm water and alcohol and dried after each step. One side of the specimen was over polished due to the good observation of fatigue crack propagation. e surface roughness was measured by the MITUTOYO-ABSOLUTE-DIGIMATIC-HEIGHTGAGE device to compare how the surface roughness affects the fatigue life.
e fatigue test at the three-point bending was performed on a ZWICK/ROELL Amsler 150HFP 5100 resonance pulsator (Figure 3(a)) on 10 experimental specimens at room temperature 22°C ± 5.0°C. e specimens were numbered from 1 to 10. Vibrophore Amsler 150HFP 5100 is for fatigue testing of materials or components by applying sinusoid loads using the resonance principle (testing frequency range 35-300 Hz) with constant or variable amplitude (maximum force amplitude range is ±75 kN) and mean load (maximum mean load is ±150 kN). It provides fatigue testing of materials and components, e.g., fatigue tests according to DIN 50100 (S-N curve) for tensile stress and compressive and alternating stress ranges. Testing can be performed either force controlled (precision force measurement through calibration according to DIN 51 221 and US MIL Std. 1312 B) or strain controlled (two measurement channels for additional extensometers-force and strain; optional extensible through two measuring inputs). Tests can also be carried out under various environmental conditions, e.g., temperature range from cryogenic temperatures in a liquid nitrogen atmosphere to high-temperature testing up to 1200°C (for pushpull loading only). In addition, torsion and bending tests can also be carried out. e parameters of the test were set as follows: the static preload force F stat. � − 15.0 kN; the dynamic load was represented by dynamic force varied F dyn. � − 7.00 kN to − 13.5 kN; frequency during the test was f � 82.50 Hz-108.6 Hz; stress cycle asymmetry R < 1; the number of cycles was set on value 2.0 × 10 7 representing the run-out. For titanium alloys, this value is considered as the fatigue limit (if the specimens withstand the 2.0 × 10 7 cycles at set stress amplitude without break, then the fatigue limit is reached). e sample (of the size 10 mm × 11 mm × 50 mm) was positioned as shown in Figure 3(b), and it means that the center of the sample was loaded by the main force. To prevent specimens heating during the fatigue test, the specimens were cooled by an external fan. e stress amplitude σ a is calculated according to the following equation: where F is the applied dynamic force [N], L is the distance of supports which is 30 mm, b is the sample width which is 10 mm, h is the sample height which is 11 mm, and σ a is the maximum amplitude (MPa). e S-N curve was drawn. e Smith diagram discusses the relation between mean stress σ m and stress amplitude σ a and provides information about the secure area of loading at various values of mean stress σ m . e three-point bending loading is not so common way to obtain fatigue life values in comparison to push-pull loading with a coefficient of asymmetry R � − 1. ree-point bending fatigue loading provides more complex loading of specimens due to the shifting of mean stress to negative values with a coefficient of asymmetry R < 1. It means that the specimen is preloaded by negative static force with a higher value than the stress amplitude which results in more complex loading in the center of the specimen. ere is an expectation that fatigue life shifts to lower values due to more complex loading compared to push-pull results.
Metallography specimens were prepared by cutting with MTH Micron 3000 precise saw and then mounting into bakelite mixture in Struers CitoPress 1 and finally ground and polished using Struers TegraSystem (TegraPol-15 and TegraForce-1) and a special program for titanium alloys. Grinding and polishing consist of a few steps: grinding with SiC sandpaper No. 320, followed by fine grinding with Largo grinding disk and emulsion Allegro Largo with 9 μm grains. e first step of polishing is performed with Dac polishing disk with an emulsion of Diap 3 μm grains and fine polishing with OP-S lubricant with 0.25 μm grains. e sample was rinsed in warm water and alcohol and dried after each step.
is procedure was applied on specimens for light microscopy (LM) and for scanning electron microscopy (SEM) analysis as well.
After polishing, the sample was etched; the specimen surface was immersed into 10% solution of HF for 7 seconds and then rinsed with warm water and alcohol and dried by hot air. e microstructure was observed on light microscope 60Hz (servohydraulic) 20kHz -Not cooled 20kHz -Cooled Neophot 32. e grain size and α lamellae length and width were measured by using NIS-elements 4.20 metallography software.
e fractography analysis of specimens' surface after the fatigue test was performed using the TESCAN VEGA LMU II scanning electron microscope with the aim to detect the   Figures 2 and 4. It consists of elongated grains of α lamellae in the transformed β phase (Figure 3). e average length of the grains is 479.1 μm (Figure 4(a)), and the average width is 159.3 μm (Figure 4(b)). e average length of the α lamellae is 24.42 μm (Figure 4(c)), and the average thickness is 2.530 μm (Figure 4(d)). is microstructure is in good agreement with β heat-treated TiAl6V4 alloy at 1020°C/20 min/FC with Widmanstätten α structure presented on prior β grain boundaries reported in Videhi Arun work [22]. e morphology of the Widmanstätten α phase may change from a colony of similarly aligned α lath to a basket-weave arrangement with an increase in the cooling rate or alloying content. From the present knowledge, the dependence of the fatigue life on the lamellar α phase is known. e larger the lamella dimensions are, the lower the fatigue life of the alloy is. To confirm this phenomenon, it would be advisable to perform tests for a more fine-grained structure and to compare the obtained values.

Fatigue Test Results.
e S-N curve is shown in Figure 5. e experimental data were interpolated and the following coefficients of the Basquin [23, 24] equation were obtained: From the S-N curve of the fatigue life for σ a , it is clear that, after the translation of the resulting values of the number of cycles N f to the failure and the stress amplitude σ a by the power regression curve, the scattering of the obtained values is relatively small. erefore, the maximum stress amplitude σ a in tested material at a given number of N f cycles can be predicted fairly accurately from equation (2). e fatigue stress σ c (defined as the highest stress at which the test bar � specimen is not broken even after the set number of cycles 2.0 × 10 7 has been exceeded) was calculated according to equation (1)   results for all specimens used at the fatigue test with specimen numbering are shown in Table 2.
Results obtained after the three-point bending fatigue test with load cycle asymmetry R < 1 were compared to results of Peters et al. and Wagner et al. [8,9], Wu et al. [15], and Morrissey and Nicholas [18]. ey have used common push-pull loading at R � − 1. e results comparison shows a difference about 30.0 MPa-40.0 MPa lower in fatigue life and stress amplitude at run-out.
is result shows a more complex character of fatigue specimens loading than simple push-pull loading. e three-point bending loading includes a compression loading and a tension loading as well, and the specimen is subjected not only to direct stress but to bending moment with increasing value when approaching the specimen centre. e influence of the mean stress σ m on the fatigue stress σ c is expressed by the Smith diagram [17,25] in Figure 6. e region bounded by red indicates the area of the stress, where there is no break even after the theoretically unlimited number of cycles at the given mean stress σ m and the amplitude of the stress σ a cycles. e diagram is designed for the compressive stress area, and therefore, it is advantageous for its construction to use the values of the yield bearing strength and ultimate bearing strength [26] that involve a more complex load. It can be seen from the diagram that the fatigue resistance of the TiAl6V4 alloy increases with the decreasing amplitude of the stress σ a .
Based on TiAl6V4 titanium alloy analyses with grains with an average length of 479 μm and a width of 159 μm, and    α lamellae with an average length of 24.4 μm and a width of 2.53 μm, it can be concluded that the fatigue interval σ c is strongly dependent on the stress amplitude σ a . According to the established rules obtained from numerous experiments [23,24], the fatigue limit σ c at the load push-pull and R � − 1 corresponds to approximately 50% of ultimate tensile stress. Based on the constructed Smith diagram, it can be assumed that, for σ m � 0 MPa, σ c ∼650 MPa and UTS∼1300 MPa, which corresponds to our calculated value UTS � 1323 MPa. Figure 7 shows the fracture surface where the fatigue region, static break area, and a fatigue crack initiation site are marked (black arrow). e diminution of the fatigue region (Figure 7(b)) with the increasing value of the dynamic loading force and the stress amplitude σ a as well is visible when macro-fractographic images are compared.

SEM Fractography.
In Figure 8, the images of samples No. 9 and 7 of the fatigue crack initiation region are shown (Figures 8(a) and  8(b)). e fatigue crack in both cases initiated at the free surface of the polished samples at the sites of the highest concentration of stress. Major crack propagation proceeded from the free surface by transcrystalline cleavage of TiAl6V4 alloy grains with the cleavage facets observed at the crack initiation site. e cleavage facets of samples 9 and 7 are shown in Figures 8(c) and 8(d).
Micro-fractographic images of cleavage facets created by transcrystalline cleavage of TiAl6V4 alloy grains along the direction of propagation of the magistral fatigue crack are notable in Figure 9. In Figure 9(a), the transcrystalline cleavage facet with river morphology is visible. e river reliefs on facets are created due to plastic deformation that preceded the crack formation and its growth, presence of grain boundaries, or α-phase lamellae. e origin of the rivers also means energy consumption, which slows down the rate of propagation of the fatigue crack tip. In another case, especially at higher stress amplitude, the transcrystalline cleavage failure of β grains with a higher degree of cleavage has occurred (Figure 9(b)).
In the area of the stable fatigue crack propagation, the striations (Figure 10(a)) are visible on the surface, indicating the position of the crack tip at the given moment and creating ridges spreading from the initiation site. ese ridges are perpendicular to the direction of magistral fatigue crack propagation. Another characteristic feature of the fatigue process is the secondary crack parallel to the advancing fatigue crack front documented in Figure 10(b).
From the comparison of the individual micrographs, it is clear that the amount of secondary cracks increases with the rising value of the stress amplitude. e change in the direction of the striation propagation due to the change in the direction of the magistral fatigue crack growth is shown in Figure 11(a). e area of static failure documented in Figure 11(b) is characterized by a transcrystalline ductile fracture with dimple morphology. A ductile fracture occurs by coalescence of microcells that nucleate at the grain boundaries, secondary-phase particles, or inclusions. e orientation of the holes varies with the orientation of the applied stress and β grain orientation to the applied load as well.
e fractography analysis revealed that the polished surface of samples had a single initiation site just below the surface of the sample where the characteristic of initiation and propagation of the fatigue crack was the transcrystalline cleavage of TiAl6V4 alloy β grains, which is supported by images of transcrystalline cleavage facets in the fatigue crack initiation region. e smaller the area of fatigue fracture was, the higher the amplitude of the stress σ a was. Striations and secondary fatigue cracks, which are features of fatigue fracture, have also been observed in this area. e transcrystalline cleavage failure was caused by loss of coherency at the α lamella interface and the transformed β grain matrix.  e maximum bending stress σ Omax � 1090 MPa which is obviously higher then ultimate tensile stress UTS � 1011 MPa (in the longitudinal direction) reported for this material in the material list was applied at the fatigue test. To explain this phenomenon, the hardness measurements of material were performed and the results of measurement are shown in Table 3. e values are the mean values of four hardness measurements.
ere is a relation between the hardness and the material toughness, and it is expressed by using equation (3), where k is the coefficient that depends on the material type, and for titanium alloys, k � 3-4. e volumes of the k coefficient for Advances in Materials Science and Engineering various material types and their effect on final UTS were discussed more in detail in the work of Tabor et al. [27] and Zhang et al. [28]: For our experimental material TiAl6V4 alloy, k � 3.85 based on calculation from values provided by the supplier. According to this, the calculated ultimate tensile stress UTS � 1323 MPa. is value is over 301 MPa higher than that reported in the material list. is difference in material toughness and hardness is possible due to secondary hardening, which was not reported in the material list.

Conclusions
e titanium alloy TiAl6V4 was subjected to three-point bending fatigue loading at a frequency ∼100 Hz for a high number of cycles 2.0 × 10 7 at room temperature 22°C ± 5.0°C. It was shown that fatigue crack initiates from free surface and has single initiation at lower amplitudes of stress σ a and multiple initiation sites when higher stress amplitude σ a was used.
e fatigue life was reached after a 2.0 × 10 7 number of cycles, and it was set on σ c � 431 MPa. is value is about 30 MPa-40 MPa lower compared to common push-pull fatigue results on this alloy with similar equiaxed microstructure and α lamellae width as reported in experiments of other authors. However, it is necessary to take into count the different loading modes, specimen shape, and surface finishing of samples used at push-pull loading. e comparison with push-pull loading is performed due to the lack of data about the three-point bending fatigue test for this alloy in English.
e results show that a three-point bend is more complex loading and is more suitable for obtaining a fatigue life for advanced materials compared to the commonly used push-pull loads. e three-point bending loading includes both compression loading and tension loading, and a specimen is not subjected to direct stress only for the whole cross section but to the bending moment with increasing value when approaching the specimen center.

Data Availability
e data used to support the findings of this study are available from the corresponding author upon request. e data used for Figure 1 and equation (3) are available in [18] and [27,28], respectively.

Conflicts of Interest
e authors declare that they have no conflicts of interest.