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Fatigue strength assessment of TIG-dressed welded steel joints by local approaches
International Journal of Fatigue 126 (2019) 72–78
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Fatigue strength assessment of TIG-dressed welded steel joints by local approaches
T
⁎
Jörg Baumgartnera, , Halid Can Yıldırımb, Zuheir Barsoumc a
Fraunhofer Institute for Structural Durability and System Reliability LBF, Darmstadt, Germany Aarhus University, Department of Engineering, Aarhus, Denmark c KTH-Royal Institute of Technology, Department of Aeronautical & Vehicle Engineering, Stockholm, Sweden b
A R T I C LE I N FO
A B S T R A C T
Keywords: Fatigue design Finite elements Size effects Welded joints TIG-dressing
Fatigue strength assessment methods by local approaches are widely used in the literature. This paper provides a comprehensive evaluation of published data for welded steel joints improved by TIG dressing methods. Fatigue classes for the local assessment methods with the available fatigue data are recommended. The available fatigue data extracted for transverse non-load carrying welds, cruciform joints as well as butt joints. In total, 17 published test series of weld details with various yield strengths and stress ratios are presented. Fatigue strength assessment is performed by considering the weld profile geometry within Finite Element models and taking the resulting stress gradients as basis for the evaluation. In addition, the influence of the steel grade is included. The most reliable results are derived by using the critical distance approach. Fatigue classes and critical distances are recommended as a result of the evaluations.
1. Introduction Fatigue strength assessment methods are being used widely in engineering applications in order to utilise the full potential of high strength steel grades in structural design. On one hand, some of them are applied during the welding process, on the other hand, some others are performed after welding. The former one may include weld profile control and/or use of special electrodes which may help to result with compressive residual stresses around the weld toe region. The latter improvement methods, however, need to be performed separately and, in the literature, they are generally known as fatigue strength improvement by post-weld treatment methods [1,2]. Post-weld treatment methods can be classified into two groups, (i) weld profile modification methods and (ii) weld residual stress modification methods. In the weld profile modification methods, the aim is to reduce the local stress concentration due to the weld profile by achieving a smooth transition between the main plate and the weld flank. This can be achieved by increasing the weld toe radii and decreasing the flank angle. In addition to this, possible weld toe flaws may also be removed. This leads subsequently in an extended crack initiation phase and a longer fatigue life. For the residual stress modification methods the goal is to eliminate
⁎
the high tensile residual stress in the weld toe region and to induce compressive residual stresses at the weld toe. Hammer and needle peening are two of the well-known traditional methods. A recent application, which is called High frequency mechanical impact treatment (HFMI), has been developed and it has increased its popularity to a great extend in the last decade. As a result, the Commission XIII: Fatigue of Welded Component and Structures of the International Institute of Welding (IIW) has decided to publish a separate recommendation for HFMI improvement methods after extensive and careful examinations [3]. The most well-known weld profile modification methods include machining or grinding of weld seam and toe, as well as re-melting the weld toe by Tungsten Inert Gas (TIG), plasma or laser. In the literature, TIG dressing, among the others, has been considered as one of the most practically applicable ones. The method requires the standard TIG welding equipment to re-melt the area around the weld toe. The application is performed without any addition of filler material. The most important benefit of the re-melting process is that a smooth transition from the weld face to the base material is achieved [4]. Additionally, weld flaws such as micro cracks or cold laps are reduced. Additionally, the original residual stress states is changed. In 2007, the Commission XIII of the IIW completed a
Corresponding author. E-mail address: [email protected] (J. Baumgartner). URL: http://www.lbf.fraunhofer.de (J. Baumgartner).
https://doi.org/10.1016/j.ijfatigue.2019.04.038 Received 18 December 2018; Received in revised form 25 April 2019; Accepted 26 April 2019 Available online 27 April 2019 0142-1123/ © 2019 Elsevier Ltd. All rights reserved.
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suggestion allows larger improvement factors in comparison to the IIW recommendations [6], where only a two (2) and three (3) fatigue class increase for 355 MPa < f y and f y > 355 MPa with a slope of m1 = 3 are allowed, respectively. Additionally, FAT 112 and FAT 125 are recommended for the use of structural hot spot stress for 355 MPa < f y and f y > 355 MPa welds, respectively, whereas a design suggestion for mild weld notches with large radii [7] is given when it comes to analyses in the effective notch stress system. However, the suggestions by Yıldırım [5] were only limited for the nominal stress system, which means that a more closer analyses on local assessment methods are still needed for better understanding of the interpretation of weld toe response. This has raised discussions on carrying out investigations by using local approaches within the Working Group 2: Techniques for improving the fatigue strength of welded components and structures, and Working Group 3: Stress analysis. The aim of this current work has been to develop appropriate guidelines for the fatigue assessment of welds improved by TIG dressing method based on the local stresses. In total, 17 published test series for weld details with various yield strengths (355 MPa ⩽ f y ⩽ 1100 MPa) and stress ratios (0 ⩽ R ⩽ 0.2) are presented and analysed.
recommendation for implementing four common post-weld fatigue strength improvement methods.
• burr grinding • TIG-dressing • hammer peening and • needle peening for welded steel and aluminium structures [1]. The guideline includes procedures for applying the treatment methods, quality assurance measures and recommendations for implementing the methods in the fatigue design of structures. In the IIW recommendations on methods for improving the fatigue lives of welded joints [1], the slope of m1 = 3 of the S-N curve results in conservative design curves in the high cycle fatigue regime but less conservative or even non-conservative results for lower cycles to failure, e.g. N = 10 4 . Individual experimental studies for improved welds also show that the slope of the best-fit line through the S-N data is typically greater than m1 = 3. Therefore, in order to calculate the bestfit S-N curve and the degree of fatigue strength improvement, Yıldırım [5] performed fatigue strength assessment for the test results of welded joints improved by the TIG dressing method in 2015. The study verified the use of m1 = 4 is more appropriate after analysing 311 fatigue data points for longitudinal attachments, transverse non-load carrying welds, and butt and T-joints with varying yield strengths and main plate thickness. The presented fatigue resistant curves in [5] for TIG dressing method were recommended based on the nominal stress (NS) method only. The IIW system gives the fatigue strength as a function of fatigue classes (called FAT class or simply FAT) which are defined by S-N curves. These each fatigue class are determined based on statistical analyses and they show the stress range in MPa corresponding to 95% survival probability at 2 × 106 cycles to failure. Therefore it was possible to propose a set of FAT values for the investigations for TIG dressing. The proposals were only done in terms of the NS method, in which the characteristic fatigue strength values depended on the material yield strength. More clearly, it means that one (1) fatigue class increase in strength is allowed for every 200 MPa increase in static yield strength. This so-called stepwise increase is shown in Fig. 1 with the reference to the same as-welded improved weld conditions. This approach was chosen to be consistent with the available IIW recommendations both with the ones in the as-welded and improved states. In short, the proposal gives a four (4) fatigue class increase in strength for joints manufactured from 235 MPa < f y ⩽ 355 MPa steel with respect to the nominal fatigue class in the as-welded condition. The increase in fatigue strength was also extended up to a six (6) fatigue class when it comes to f y > 950 MPa steel grades. In the study [5], fatigue classes were defined for a fixed S-N slope of m1 = 4 after statistical analyses of welds improved by TIG dressing method. The
2. Assessment approaches 2.1. IIW-recommendations There are various approaches available for the fatigue assessment of welded structures. The most common ones – the nominal, structural and effective notch stress approach – are included in the IIW-recommendations [6] for fatigue design of welded joints and components. In the recommendations, as mentioned previously, FAT-values are given that represent the endurable stress range at N = 2·106 load cycles for a survival probability of PS = 97.7% and high mean resp. residual stresses (R-ratio of R ≈ 0.5). The course of the design S-N curve, a slope of k = 3 and a knee point at Nk = 1·107 cycles, is the same for all approaches. Within the IIW-recommendations [6] fatigue improvement methods are generally considered by upgrading the FAT-class by multiplying the class with a factor f. The factor varies in a range 1.3 ⩽ f ⩽ 1.5. For TIGdressing a factor of f = 1.3 is recommended. Even though it is known that most fatigue improvement methods result in a more shallow slope of the S-N curve, the slope of k = 3 is still recommended for use. For the nominal stress approach the application of the assessment is in principle clear since the actual weld geometry is not considered. Within the structural stress approach the weld geometry is not considered in detail. Subsequently, the increase in the transition radius and in weld leg length due to the TIG-dressing is not included. If the extrapolation is conducted towards the original position of the weld toe (without TIG-dressing) higher stresses will be derived. This should lead to a conservative assessment. For local stress approaches, like the effective notch stress approach with a radius of r = 1 mm , the application is in contrast not as clear. This is discussed in the following. 2.2. Local approaches Before addressing the application of local fatigue assessment approaches on TIG-dressed joints the basic idea behind the effective notch stress approaches will be described and further approaches are presented. A in-depth description can be found in a recent work done by Baumgartner [8]. Within the effective notch stress, as it is used typically in numerous recommendations and guidelines [6], the notches are rounded with a (fictitious) radius of r = 1.0 mm . The radius r = 1.0 mm was introduced by Radaj [9]. Olivier and Seeger [10,11] used this method for the fatigue assessment of welded steel joints. Based on the evaluation of various SN-curves of T-joints and skewed T-joints endurable notch stresses have been derived [12] and the class of FAT225 was
Fig. 1. Proposed maximum increases in the number of FAT classes as a function of f y in the nominal stress method [5]. 73
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in addition with micro-notches in the transition. Second, cracks might start from the weld metal in the fillet radius rfillet (CP2 in Fig. 2). This is the location where the highest stress concentration occurs. But only in a few cases cracks start from the weld metal fillet. Subsequently, the stresses along both possible crack locations and crack paths need to be evaluated from Finite-Element analyses. In order to perform a fatigue assessment, also the necessary material parameters have to be identified. For the critical distance approach, that is in focus of this work, the critical distance a has to be identified. Additionally, design (i.e. FAT-) values have to be available and a course of the S-N curve that is adapted to the characteristics of TIG-dressed joints.
recommended [13]. The approach using fictitious radii is based on the idea to account for so-called support effects by enlarging the actual radii. The notch stresses calculated with this enlarged radius can directly be used as “fatigue–effective” in a fatigue assessment. Radaj [9] applied this approach to the assessment of welded structures, assuming a worst case radius of rreal = 0 mm and a sharp notch (notch opening angle of ω = 0°). Based on some theoretical considerations, a fictitious radius of
r = rreal + ρ∗ ·s
(1)
was derived. With a micro-structural support length of ρ∗ = 0.4 mm and a factor of s = 2.5 a fictitious radius of r = 1.0 mm was introduced. In many papers, a reliable assessment was documented for the effective notch stress approach applied at specimens in as-welded state [14]. Only at thin (t ⩽ 5 mm ) sheets [15] or welds with a mild weld toe notch [16] (resulting e.g. from a post-weld treatment such as grinding) the application might lead to uncertainties. In these cases, approaches such as the stress averaging or critical distance approach can be applied that lead to a reliable assessment. The theory behind the non-local approaches is to evaluate the stress field or stress gradients in the notch ligament as it is done within the stress averaging approach according to Neuber [17] or the critical distance approach according to Peterson [18]. Both approaches have been successfully used to evaluate fatigue data of various thin- and thick-walled specimens in as-welded condition [19]. A class FAT160 was identified for both approaches. The evaluation was based on results from FE-models with a constant notch radius of r = 0.05 mm .
3. Fatigue of TIG-dressed joints In literature, many papers have been published that deal with the fatigue strength of TIG-dressed welded joints. A collection of papers and a subsequent evaluation of endurable stresses within the nominal stress approach was conducted in [5]. In order to apply local stress approaches, high-quality data needs to be available. This comprises fatigue data itself, geometry of the weld and the specimen, failure location, material data and information on residual stresses. The local geometry of the weld, i.e. the weld profile and the penetration depth, has to be available. This could be provided by a micrograph or a detailed description of the weld geometry. 3D-scan data is very helpful in order to get information on an average geometry. Also information on the axial and angular misalignment is helpful to interpret fatigue data. TIG-dressed joints may fail from various locations, Fig. 3 and Table 1. Therefore, information on the failure location is necessary since the endurable local stresses are dependent on the failure critical detail. In an evaluation, only specimens with failure that origins close to the TIG-layer is considered (No. 1, 2a, 2b, 6 and 9 in Table 1). Specimens with failure starting at other locations are regarded as run-out. The material name, the yield and ultimate strength are given in the majority of publications. But the mechanical properties of the material are highly influenced by the welding process. So the hardness close to the failure critical detail, as measure for the materials strength, should be reported, too. Especially for post-weld treated joints this is important information since the fatigue strength should be correlated at least to some extend to the materials strength. Finally, also information on the residual stress state close to the critical weld detail is of importance. The available data has been gathered and summarised in Table 2. Only fatigue data of butt joints, transverse stiffeners, and cruciform joints has been regarded. The data contains overall 17 different test series. For the subsequent evaluations the individual S-N curves of tests were statistically derived. Only test data with load cycles below the knee-point of the individual S-N curve have been regarded. All fatigue tests with failure locations outside the TIG-pass (No. 3, 4, 5, 7 and 8 in Table 1) have been excluded. All tests have been conducted with similar R-values in the range of 0 ⩽ R ⩽ 0.2 . Therefore, no correction of mean stresses was performed.
2.3. Application at TIG-dressed joints The application of the (effective) notch stress approach is in principle clear for standard welds with a certain defined flank angle, produced, for exampled, by a MAG process. The resulting weld toe notches rtoe are rounded with the 1 mm-notch, Fig. 2, left. For TIG-dressed welds instead, the application is unclear. The reasons can be identified by looking at the typical geometry of such welds, Fig. 2, right. Due to the TIG-dressing a round fillet with a large radius rfillet is produced between the base plate and the weld metal of the original weld. In most case, there is a smooth transition between weld metal and heat affected zone. The notch radius to be enlarged within the effective notch stress approach rtoe can not be defined. The question that arise is, how a fatigue assessment can be conducted bases on local stresses. Different local approaches such as the stress averaging or the critical distance approach can be used for an assessment of TIG-dressed joints. As mentioned above, these approach consider the stress gradients and subsequently the occurring support effects. The stress gradients can easily evaluated within a Finite-Element model. Due to the different failure locations identified in the fatigue tests it is not quite clear, at which positions the stresses should be evaluated. Two failure critical locations need to be assessed. First, the transition between TIG-weld and heat affected zone has to be considered (CP1 in Fig. 2). Even if there is no clear geometrical stress concentration (i.e. notch), the majority of cyclically loaded TIG-dressed welds fail from that position. The reasons are likely to be found the metallurgical notch (difference between weld metal and heat affected zone), maybe
Fig. 2. Local geometry and crack initiation locations at standard and TIGdressed welds.
Fig. 3. different failure locations at TIG-dressed welds. 74
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Table 1 Failure locations for TIG-dressed joints. No.
Failure Location
Description
1
Weld toe
2a 2b 3 4 5 6 7 8 9
Weld metal Weld metal Weld root Base metal Clamping Weld transition Base metal Inner irregularities Base metal
Interface/notch between weld metal of TIG-layer and heat affected zone In the fillet formed by the TIG-dressed layer In the undercut of the TIG-dressed layer Root of the original weld Base material between weld and clamping In the clamping of the specimen Transition between TIG and original weld layer, etc. Base metal below stiffener (only for trans. stiffeners) Inner irregularities like pores, lack of fusion, etc. Heat affected zone
Fig. 4. Evaluation of effective stresses at the weld toe (No. 1) and the position of maximum stress (No. 2).
The following evaluations have been conducted with the critical distance approach. Effective stresses σeff are calculated in a critical distance of a away from the surface.
σeff = σ (a) Table 2 Overview on available fatigue data with detailed weld geometry, clear. Reference
Acronym
Material
Loading
Transverse stiffener S355 3-P. bend. S700 4-P. bend. 304L Axial S31803 Axial S700 Axial S355 Axial S460 Axial S690 Axial
R
Fail. Loc.
0 0.1 0.1 0.1 0.2 0.1 0.1 0.1
1 1∗ 2a∗ 2a∗ Unclear 2b,4,5 2b 2b
[20] [21] [22] [22] [23,24] [25,26] [25,26] [25,26]
RAM PED EUR EUR LEF KUH1 KUH2 KUH3
[27,28,16] [27,28,16] [29–31] [29–31] [29–31] [29–31] [29–31] [29–31]
MEL1 MEL1 ES1 ES2 ES3 ES4 ES5 ES6
Butt joint S960 Axial S1100 Axial S460 Axial S690 Axial S1100 Axial S460 Axial S690 Axial S890 Axial
0.1 0.1 ≈0 ≈0 ≈0 ≈0 ≈0 ≈0
1,2b 1,2b 1 & 2a 2a 1,2a 1,2a 2a 2a
[32]
SCH1
Cruciform joint S355M Axial
0
1
∗)
(2)
The critical distance a is unknown for TIG-dressed joints. In order to identify a recommendable value, it was changed in the range 0 ⩽ a ⩽ 2 mm with a step length of 0.02 mm. For every value of a the scatter of all test data was calculated. As already mentioned, the evaluation was performed at two location, namely at the weld toe and the position in the fillet with σ = σmax . For the calculation of a reference S-N curve a slope needs to be derived. Since an evaluation of the slope by linear regression does in many cases lead to incorrect or illogical results for a heterogeneous database the slope was calculated by averaging and weighting the slope of all 17 test series, see [19]. This evaluation lead to a slope of k = 3.9. Since this slope is quite close to the slope of k = 4.0 as proposed in [5] k = 4.0 was used as pre-set value. The scatter TS of the reference S-N curve can be used as indication of the assessment reliability. The minimum scatter for the evaluation at the position of the weld toe could be determined for a critical distance of a = 0.7 mm , Fig. 5. The reference S-N curve has a scatter of TS = 1: 1.42 . By the evaluation at the location of maximum stress the scatter a slightly lower scatter of TS = 1: 1.39 is derived, Fig. 6. For both reference S-N curves FAT values have been derived by calculating the stresses for a 97.5% survival probability and reducing this value by an additional 10% [34] to account for possible high tensile residual stresses (transformation to R = 0.5). This results in a class FAT160 for the failure at the weld toe and FAT180 for the failure in the fillet. A comparison of the S-N curves for both evaluation positions does not show a major difference. This is likely to be linked to the fact that the stresses at the weld toe and at the position of maximum stress have a similar factor between them. Even though the scatter of the S-N curve is already remarkably low a further reduction might be possible if the material strength is additionally considered. This behaviour was identified by Yıldırım in [5] based on the evaluations of the available fatigue data in the nominal stress system and it should also be present in the local evaluations. To visualise the influence of the material strength on the fatigue strength the results are plotted in Fig. 7 with markers according to the static strength range. With this plot it is possible to discuss that there is a correlation between the fatigue and the material strength. Theoretically, the materials strength has influence on both the support effects and the endurable nominal stresses. The support effects, i.e. the value of a within the critical distance approach, is reduced with increasing materials strength whereas the fatigue strength of the material itself increases with increasing materials strength. Typically, only the latter influence is considered within the nominal stress approach, for example for TIG dressing [5] or HFMI treatment [3]. Nevertheless, a more detailed assessment can be performed within local approaches. In order to include both factors within the evaluation, two modification factors are used. First a factor f to consider the influence of the materials strength on the fatigue strength and second a factor fa for the influence of the materials strength on the support effects. With the factor fa the critical distance a is modified for the evaluation. An inverse linear relationship between a∗ and material strength
not 100%
Comment
GFT-data type “c” type “c” type “c” type “v” type “v” type “v”
4. Calculation of local stresses Finite-Element (FE) models have been created for all specimens with the available fatigue data points. The main geometrical parameters have been regarded as given in micro-sections or in tabled measurement data. In all cases, the TIG-layer was represented by a fillet, that means a circular rounding between original weld and sheet. Especially for welds with a steep flank angle this representation leads to a very good agreement between the existing geometry and the corresponding modelled geometry. For specimens with a small flank angle the surface of the TIG weld had sometimes a varying curvature. Nevertheless, since this exact curvature is unknown prior to the manufacturing, the above-described simplified modelling was used. 2D-models with quadratic plain-strain elements were used in the models since all specimens had a constant cross section. In the vicinity of welds, a structured mesh according to the recommendation in [33] ensures converged stresses. For all the models, a nominal stress of σn = 1 MPa was applied. The stresses were derived in a linear-elastic calculation. A rigid clamping (no rotation, only translation in load direction) was used. Since the exact failure locations are not known for the majority of fatigue tests, the stresses were evaluated at two positions, Fig. 4. Firstly, No. 1 at the weld toe position and secondly, No. 2 at the location of maximum stress. The Maximum Principal stress hypothesis was used for evaluation. Next to the maximum values of stress also the stress gradient on a path perpendicular to the weld surface was taken from the FE-analysis. 75
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Fig. 5. Reference S-N curve for a = 0.7 mm and evaluation at the weld toe (left), dependence of the scatter on the value of a (right).
is assumed, since also nearly linear relationship between the value of ρ∗ ( ρ ∗ ≈ 4 × a , see [35]) and the yield stress was identified by Neuber [36].
a∗=
fa 355 MPa
·Rm + a0 − fa
(3)
The value a 0 is the critical distance for a material strength of 355 MPa . This value is fixed for every evaluation. The factor fa is varied in the range − 0.45 ⩽ fa ⩽ 0 . The influence of the material strength on the fatigue strength is assumed to be linear to the material strength. The factor fRm is varied in the range 0 ⩽ fRm ⩽ 0.5.
f=
fRm 355 MPa
·Rm + 1 − fRm
(4)
The scatter of the reference S-N curve was derived for overall three values of a0 : a0 = 1.0 mm, a0 = 0.8 mm and a 0 = 0.6 mm and for both positions, weld toe and position of maximum stress. In all cases, the scatter could be reduced down to 1: 1.34 ⩽ TS ⩽ 1: 1.37 . The lowest scatter is derived for the value a 0 = 0.6 mm at the
Fig. 7. Reference S-N curve for a = 0.6 mm and evaluation at the position of max. stress.
Fig. 6. Reference S-N curve for a = 0.6 mm and evaluation at the position of max. stress (left), dependence of the scatter on the value of a (right ) . 76
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with large fillet radii and a nearly tangential transition between heat affected zone and weld metal this radius cannot be used. Also from a theoretical point of view there is no need to introduce a fictitious radius for TIG-dressed welds as used for as-welded specimens. From the TIGdressing results a fairly large radius that can be identified, e.g. experimentally or guessed based on experience, and used in the calculation models. In principle, the fatigue strength of different failure critical locations as the weld toe, the heat affected zone or the weld metal in the fillet has to be evaluated separately. All of these locations have been reported at least once in the literature data used for evaluation. Unfortunately, an exact description of each test with its failure location was only reported in few publications. Therefore, a pragmatic approach was followed to evaluate all test data with stresses at the weld toe and all data with the maximum stresses in the fillet. The critical distance approach as easy applicable and reliable approach was used for all evaluations. For a critical distance of a = 0 mm also the use of maximum stresses on the surface can assessed. The evaluation results in a fairly small scatter of TS = 1: 1.42 at the position of the weld toe and TS = 1: 1.39 for the maximum stresses in the fillet. This small scatter surprises to some extend, since the scatter of larger data sets are typically quite high. In [19] a minimum scatter of TS = 1: 2.38 could be derived for as-welded joints with the critical distance approach. Explanation might be that the all tests have been conducted with a similar R-ratio, that the variation in sheet thickness was smaller and that the overall amount of individual tests was smaller. The derived critical distances of a = 0.7 mm for the evaluation at the weld toe and a = 0.56 mm for maximum stresses lie between known distances in literature. In [37] a distance of a = 1.0 mm and in [19] a distance of a = 0.1 mm of was recommend for the assessment of joints in as-welded state. Since the failure critical locations differ (TIG: smooth fillet or metallurgical notch, as-welded: sharp geometrical notch) differences in the critical distance might be expected. The value of a is quite insensitive to a change due to a similarity of the stress gradients. Even with a value of a = 1.0 mm or a = 0.1 mm still a fair assessment would result. The resulting reference S-N curve has a scatter below TS = 1: 1.5, see the correlation between TS and the critical distance a in Figs. 6 and 5. Using the position of maximum stresses in the fillet results to somewhat smaller scatter and subsequently to a higher assessment reliability. But since the difference in scatter is quite low, no
Fig. 8. Dependence of the scatter on the values fRm and fa .
position of maximum stress. For a value of fRm = 0.1 and fa = 0 the highest reduction in scatter can be identified in Fig. 8. A view on the reference S-N curve shows that there is no influence of the materials strength on the fatigue strength visible anymore, Fig. 9. The parameters of the reference S-N curve are: Δσtcd, a, mP, N = 2e6 = 233 MPa, k = 4, 0 and tS = 1: 1.34 . From this reference S-N curve a class FAT160 was derived. 5. Discussion Fatigue assessment with local approach at welded joints is most commonly associated with the effective notch stress approach with the reference radius rref = 1.0 mm . This radius can be used for as-welded joints with sharp weld toe or weld root notches. For TIG-dressed welds
Fig. 9. Reference S-N curve for the evaluation at the position of maximum stress, a = 0.6 mm, fRm = 0.1 and fa = 0 . 77
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recommendation for any of both methods can be given at the moment. For an automated assessment the location at the weld toe might be favoured since the position is already determined during pre-processing of a FE-model. In the evaluation of TIG-dressed joints using nominal stresses [5], an influence of the materials strength on the fatigue strength was identified. This influence could also be determined for the evaluations with local stresses even though, with only 10%, to a much less degree. The reason for this deviation is not yet clear. It might be explained if the fillet radii are depending on the materials strength. However this effects could not be confirmed. An pronounced influence of the materials strength in the support effects resp. the critical distance could not be identified. The reasons for that might be that the stress gradients are quite shallow and subsequently the support factors are quite low.
joints using reference radii. Int J Fatigue 2017;101:459–68. [9] Radaj D, editor. Design and analysis of fatigue resistant welded structures. Cambridge: Woodhead Publishing; 1990. [10] Olivier R, Koettgen VB. Schweißverbindungen I - Schwingfestigkeitsnachweis für Schweißverbindungen auf der Grundlage örtlicher Beanspruchungen, Abschlussbericht, Number 105. Forschungkuratorium Maschinenbau (FKM); 1989. [11] Olivier R, Koettgen VB. Schweißverbindungen II - Untersuchung zur Einbindung eines neuartigen Zeit- und Dauerschwingfestigkeitsnachweises von Schweißverbindungen aus Stahl in Regelwerke, Number 128. Forschungkuratorium Maschinenbau (FKM); 1991. p. 389. [12] Koettgen VB, Olivier R, Seeger T. Fatigue analysis of welded connections based on local stresses. IIW - Commission XIII, XIII-1408-91; 1992. [13] Hobbacher Adolf. Database for the effective notch stress method at steel. Int Inst Weld, JWG-XIII-XV-197-08; 2008. [14] Pedersen MM, Mouritsen O, Hansen MR, Andersen JG, Wenderby J. Re-analysis of fatigue data of welded joints using the notch stress approach. Int J Fatigue 2010;32(10):1620–6. [15] Bruder T, Störzel K, Baumgartner J, Hanselka H. Evaluation of nominal and local stress based approaches for the fatigue assessment of seam welds. Int J Fatigue 2012;34(1):86–102. [16] Möller Benjamin, Baumgartner Jörg, Wagener Rainer, Kaufmann Heinz, Melz Tobias. Low cycle fatigue life assessment of welded high-strength structural steels based on nominal and local design concepts. Int J Fatigue 2017;101:192–208. [17] Neuber H. Kerbspannungslehre. 2nd ed. Springer Verlag; 1958. ISBN 9783540135586. [18] Peterson RE. Metal fatigue. McGraw-Hill Book Company, Inc.; 1959. [19] Baumgartner J, Schmidt H, Ince E, Melz T, Dilger K. Fatigue assessment of welded joints using stress averaging and critical distance approaches. Weld World 2015;59(5):731–42. [20] Ramalho Armando L, Ferreira José AM, Branco Carlos AGM. Fatigue behaviour of t welded joints rehabilitated by tungsten inert gas and plasma dressing. Mater Des 2011;32(10):4705–13. [21] Pedersen Mikkel Melters, Mouritsen Ole Østergaard, Hansen Michael Rygaard, Andersen Jes Grøn, Wenderby Jimmi. Comparison of post-weld treatment of highstrength steel welded joints in medium cycle fatigue. Weld World 2010;54(78):R208–17. [22] Maddox SJ, Hopkin GR, Holy A, Moura Branco CA, Infante V, Baptista R, et al. Improving the fatigue performance of welded stainless steels, Technical report, EUR 22809 EN, European Commission; 2007. [23] Lagerqvist O, Clarin M, Gozzi J, Völling B, Pak D, Stötzl J, et al. Efficient lifting equipment with extra high-strength steel. Technical report, European Commission; 2005. [24] Lieurade HP, Huther I, Lefebvre F. Effect of weld quality and postweld improvement techniques on the fatigue resistance of extra high strength steels. Weld World 2008;52(7-8):106–15. [25] Dürr A. Zur Ermüdungsfestigkeit von Schweißkonstruktionen aus höherfesten Baustählen bei Anwendung von UIT-Nachbehandlung, PhD thesis, Universität Stuttgart; 2007. [26] Kuhlmann U, Dürr A, Bergmann J, Thumser R. FOSTA Forschungsvorhaben P 620: Effizienter Stahlbau aus höherfesten Stählen unter Ermüdungsbeanspruchung, Verlag und Vertriebsgesellschaft mbH, Düsseldorf; 2006. [27] Melz Tobias, Möller Benjamin, Baumgartner Jörg, Ummenhofer Thomas, Herion Stefan, Hrabowski Jennifer, et al. FOSTA P900: Erweiterung des örtlichen Konzeptes zur Bemessung von LCF-beanspruchten geschweißten Kranstrukturen aus hochfesten Stählen. Technical report, Forschungsvereinigung Stahlanwendung e.V.; 2015. [28] Hrabowski Jennifer, Herion Stefan, Ummenhofer Thomas. Kurzzeitfestigkeit von Schweißverbindungen aus höchst- und ultrahochfesten stählen. [29] van Es SHJ. Effect of TIG-dressing on fatigue strength and weld toe geometry of butt welded connections in high strength steel. Master’s thesis, Delft University of Technology; 2012. [30] van Es SHJ, Kolstein MH, Pijpers RJM, Bijlaard FSK. TIG-dressing of high strength steel butt welded connections – part 1: weld toe geometry and local hardness. Proc Eng 2013;66:216–25. [31] van Es SHJ, Kolstein MH, Pijpers RJM, Bijlaard FSK. TIG-dressing of high strength butt welded connections – part 2: Physical testing and modelling. Proc Eng 2013;66:126–37. [32] Schliebner. Experimentelle und numerische Untersuchungen zur Bewertung der Schwingfestigkeit von Hybridschweißverbindungen, PhD thesis, Technische Universität Darmstadt; 2005. [33] Baumgartner J, Bruder T. An efficient meshing approach for the calculation of notch stresses. Weld World 2013;57:137–45. [34] Sonsino Cetin Morris. A consideration of allowable equivalent stresses for fatigue design of welded joints according to the notch stress concept with reference radii rref = 1.00 and 0.05 mm. Weld World 2009;53:64–75. [35] Taylor D, Hoey D. High cycle fatigue of welded joints: the tcd experience. Int J Fatigue 2009;31:20–7. [36] Neuber H. Über die Berücksichtigung der Spannungskonzentration bei Festigkeitsberechnungen. Konstruktion 1968;20(7):245–51. [37] Xiao Zhi-Gang, Yamada Kentaro. A method of determining geometric stress for fatigue strength evaluation of steel welded joints. Int J Fatigue 2004;26:1277–93.
6. Conclusions A fatigue assessment approach for TIG-dressed joints was derived in this paper. For the application only a simplified modeling of the joint, regarding only the rough weld profile and the fillet radius of the TIGweld, and a linear-elastic calculation is necessary. For an assessment between two locations for evaluating stresses can be chosen.
• Position of the weld toe: The stresses in a distance of a = 0.7 mm •
below the surface should be used. These can be evaluated against the stress range Δσtcd, a, mP = 242 MPa (Survival probability: Ps = 50%) or a class FAT160. Position of the maximum stress: The stresses in a distance of a = 0.6 mm below the surface should be used. These can be evaluated against the stress range Δσtcd, a, mP = 253 MPa (Survival probability: Ps = 50%) or a class FAT180.
A slight improvement in the assessment reliability can be achieved if the materials strength is included in the evaluation. For the evaluation at the fillet a class FAT160 can be used that is multiplied by the factor f given in Eq. (4). For a further validation and improvement of these recommendations, high quality fatigue data for TIG-dressed joints (statistically ensured weld geometry, exact failure location, hardness distribution close to the failure location, residual stress, etc.) and also further fatigue tests on other structural details (joint types next to butt welds, transverse stiffeners and cruciform joints) needs to be available. With these future data an assessment would be enabled that relies to an higher extent on physically based descriptions than on rough empirical assumptions. Researchers are encouraged to document all tests in detail and provide data for further evaluations. References [1] Haagensen PJ, Maddox S. IIW recommendations on methods for improving the fatigue lives of welded joints. Cambridge: Woodhead Publishing Ltd.; 2013. [2] Kirkhope KJ, Bell R, Caron L, Basu RI, Ma K-T. Weld detail fatigue life improvement techniques. Part 1: review. Mar Struct 1999;12:447–74. [3] Marquis Gary B, Barsoum Zuheir. IIW Recommendations on High Frequency Mechanical Impact (HFMI) Treatment for Improving the Fatigue Strength of Welded Joints. Singapore, Singapore: Springer; 2016. ISBN 978-981-10-2504-4. pp. 1–34. [4] Watkinson F, Bodger PH, Harrison JD. The fatigue strength of welded joints in high strength steels an methods for its improvement. UK: The Welding Institute; 1971. [5] Yıldırım Halid Can. Review of fatigue data for welds improved by tungsten inert gas dressing. Int J Fatigue 2015;79:36–45. [6] Hobbacher AF. Recommendations for fatigue design of welded joints and components. Springer International Publishing; 2016. [7] Fricke Wolfram. IIW recommendations for the fatigue assessment of welded structures by notch stress analysis. Woodhead Publishing Ltd., Cambridge. International Institute of Welding, Paris; 2012. [8] Baumgartner J. Review and considerations on the fatigue assessment of welded
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Contents lists available at ScienceDirect
International Journal of Fatigue journal homepage: www.elsevier.com/locate/ijfatigue
Fatigue strength assessment of TIG-dressed welded steel joints by local approaches
T
⁎
Jörg Baumgartnera, , Halid Can Yıldırımb, Zuheir Barsoumc a
Fraunhofer Institute for Structural Durability and System Reliability LBF, Darmstadt, Germany Aarhus University, Department of Engineering, Aarhus, Denmark c KTH-Royal Institute of Technology, Department of Aeronautical & Vehicle Engineering, Stockholm, Sweden b
A R T I C LE I N FO
A B S T R A C T
Keywords: Fatigue design Finite elements Size effects Welded joints TIG-dressing
Fatigue strength assessment methods by local approaches are widely used in the literature. This paper provides a comprehensive evaluation of published data for welded steel joints improved by TIG dressing methods. Fatigue classes for the local assessment methods with the available fatigue data are recommended. The available fatigue data extracted for transverse non-load carrying welds, cruciform joints as well as butt joints. In total, 17 published test series of weld details with various yield strengths and stress ratios are presented. Fatigue strength assessment is performed by considering the weld profile geometry within Finite Element models and taking the resulting stress gradients as basis for the evaluation. In addition, the influence of the steel grade is included. The most reliable results are derived by using the critical distance approach. Fatigue classes and critical distances are recommended as a result of the evaluations.
1. Introduction Fatigue strength assessment methods are being used widely in engineering applications in order to utilise the full potential of high strength steel grades in structural design. On one hand, some of them are applied during the welding process, on the other hand, some others are performed after welding. The former one may include weld profile control and/or use of special electrodes which may help to result with compressive residual stresses around the weld toe region. The latter improvement methods, however, need to be performed separately and, in the literature, they are generally known as fatigue strength improvement by post-weld treatment methods [1,2]. Post-weld treatment methods can be classified into two groups, (i) weld profile modification methods and (ii) weld residual stress modification methods. In the weld profile modification methods, the aim is to reduce the local stress concentration due to the weld profile by achieving a smooth transition between the main plate and the weld flank. This can be achieved by increasing the weld toe radii and decreasing the flank angle. In addition to this, possible weld toe flaws may also be removed. This leads subsequently in an extended crack initiation phase and a longer fatigue life. For the residual stress modification methods the goal is to eliminate
⁎
the high tensile residual stress in the weld toe region and to induce compressive residual stresses at the weld toe. Hammer and needle peening are two of the well-known traditional methods. A recent application, which is called High frequency mechanical impact treatment (HFMI), has been developed and it has increased its popularity to a great extend in the last decade. As a result, the Commission XIII: Fatigue of Welded Component and Structures of the International Institute of Welding (IIW) has decided to publish a separate recommendation for HFMI improvement methods after extensive and careful examinations [3]. The most well-known weld profile modification methods include machining or grinding of weld seam and toe, as well as re-melting the weld toe by Tungsten Inert Gas (TIG), plasma or laser. In the literature, TIG dressing, among the others, has been considered as one of the most practically applicable ones. The method requires the standard TIG welding equipment to re-melt the area around the weld toe. The application is performed without any addition of filler material. The most important benefit of the re-melting process is that a smooth transition from the weld face to the base material is achieved [4]. Additionally, weld flaws such as micro cracks or cold laps are reduced. Additionally, the original residual stress states is changed. In 2007, the Commission XIII of the IIW completed a
Corresponding author. E-mail address: [email protected] (J. Baumgartner). URL: http://www.lbf.fraunhofer.de (J. Baumgartner).
https://doi.org/10.1016/j.ijfatigue.2019.04.038 Received 18 December 2018; Received in revised form 25 April 2019; Accepted 26 April 2019 Available online 27 April 2019 0142-1123/ © 2019 Elsevier Ltd. All rights reserved.
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suggestion allows larger improvement factors in comparison to the IIW recommendations [6], where only a two (2) and three (3) fatigue class increase for 355 MPa < f y and f y > 355 MPa with a slope of m1 = 3 are allowed, respectively. Additionally, FAT 112 and FAT 125 are recommended for the use of structural hot spot stress for 355 MPa < f y and f y > 355 MPa welds, respectively, whereas a design suggestion for mild weld notches with large radii [7] is given when it comes to analyses in the effective notch stress system. However, the suggestions by Yıldırım [5] were only limited for the nominal stress system, which means that a more closer analyses on local assessment methods are still needed for better understanding of the interpretation of weld toe response. This has raised discussions on carrying out investigations by using local approaches within the Working Group 2: Techniques for improving the fatigue strength of welded components and structures, and Working Group 3: Stress analysis. The aim of this current work has been to develop appropriate guidelines for the fatigue assessment of welds improved by TIG dressing method based on the local stresses. In total, 17 published test series for weld details with various yield strengths (355 MPa ⩽ f y ⩽ 1100 MPa) and stress ratios (0 ⩽ R ⩽ 0.2) are presented and analysed.
recommendation for implementing four common post-weld fatigue strength improvement methods.
• burr grinding • TIG-dressing • hammer peening and • needle peening for welded steel and aluminium structures [1]. The guideline includes procedures for applying the treatment methods, quality assurance measures and recommendations for implementing the methods in the fatigue design of structures. In the IIW recommendations on methods for improving the fatigue lives of welded joints [1], the slope of m1 = 3 of the S-N curve results in conservative design curves in the high cycle fatigue regime but less conservative or even non-conservative results for lower cycles to failure, e.g. N = 10 4 . Individual experimental studies for improved welds also show that the slope of the best-fit line through the S-N data is typically greater than m1 = 3. Therefore, in order to calculate the bestfit S-N curve and the degree of fatigue strength improvement, Yıldırım [5] performed fatigue strength assessment for the test results of welded joints improved by the TIG dressing method in 2015. The study verified the use of m1 = 4 is more appropriate after analysing 311 fatigue data points for longitudinal attachments, transverse non-load carrying welds, and butt and T-joints with varying yield strengths and main plate thickness. The presented fatigue resistant curves in [5] for TIG dressing method were recommended based on the nominal stress (NS) method only. The IIW system gives the fatigue strength as a function of fatigue classes (called FAT class or simply FAT) which are defined by S-N curves. These each fatigue class are determined based on statistical analyses and they show the stress range in MPa corresponding to 95% survival probability at 2 × 106 cycles to failure. Therefore it was possible to propose a set of FAT values for the investigations for TIG dressing. The proposals were only done in terms of the NS method, in which the characteristic fatigue strength values depended on the material yield strength. More clearly, it means that one (1) fatigue class increase in strength is allowed for every 200 MPa increase in static yield strength. This so-called stepwise increase is shown in Fig. 1 with the reference to the same as-welded improved weld conditions. This approach was chosen to be consistent with the available IIW recommendations both with the ones in the as-welded and improved states. In short, the proposal gives a four (4) fatigue class increase in strength for joints manufactured from 235 MPa < f y ⩽ 355 MPa steel with respect to the nominal fatigue class in the as-welded condition. The increase in fatigue strength was also extended up to a six (6) fatigue class when it comes to f y > 950 MPa steel grades. In the study [5], fatigue classes were defined for a fixed S-N slope of m1 = 4 after statistical analyses of welds improved by TIG dressing method. The
2. Assessment approaches 2.1. IIW-recommendations There are various approaches available for the fatigue assessment of welded structures. The most common ones – the nominal, structural and effective notch stress approach – are included in the IIW-recommendations [6] for fatigue design of welded joints and components. In the recommendations, as mentioned previously, FAT-values are given that represent the endurable stress range at N = 2·106 load cycles for a survival probability of PS = 97.7% and high mean resp. residual stresses (R-ratio of R ≈ 0.5). The course of the design S-N curve, a slope of k = 3 and a knee point at Nk = 1·107 cycles, is the same for all approaches. Within the IIW-recommendations [6] fatigue improvement methods are generally considered by upgrading the FAT-class by multiplying the class with a factor f. The factor varies in a range 1.3 ⩽ f ⩽ 1.5. For TIGdressing a factor of f = 1.3 is recommended. Even though it is known that most fatigue improvement methods result in a more shallow slope of the S-N curve, the slope of k = 3 is still recommended for use. For the nominal stress approach the application of the assessment is in principle clear since the actual weld geometry is not considered. Within the structural stress approach the weld geometry is not considered in detail. Subsequently, the increase in the transition radius and in weld leg length due to the TIG-dressing is not included. If the extrapolation is conducted towards the original position of the weld toe (without TIG-dressing) higher stresses will be derived. This should lead to a conservative assessment. For local stress approaches, like the effective notch stress approach with a radius of r = 1 mm , the application is in contrast not as clear. This is discussed in the following. 2.2. Local approaches Before addressing the application of local fatigue assessment approaches on TIG-dressed joints the basic idea behind the effective notch stress approaches will be described and further approaches are presented. A in-depth description can be found in a recent work done by Baumgartner [8]. Within the effective notch stress, as it is used typically in numerous recommendations and guidelines [6], the notches are rounded with a (fictitious) radius of r = 1.0 mm . The radius r = 1.0 mm was introduced by Radaj [9]. Olivier and Seeger [10,11] used this method for the fatigue assessment of welded steel joints. Based on the evaluation of various SN-curves of T-joints and skewed T-joints endurable notch stresses have been derived [12] and the class of FAT225 was
Fig. 1. Proposed maximum increases in the number of FAT classes as a function of f y in the nominal stress method [5]. 73
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in addition with micro-notches in the transition. Second, cracks might start from the weld metal in the fillet radius rfillet (CP2 in Fig. 2). This is the location where the highest stress concentration occurs. But only in a few cases cracks start from the weld metal fillet. Subsequently, the stresses along both possible crack locations and crack paths need to be evaluated from Finite-Element analyses. In order to perform a fatigue assessment, also the necessary material parameters have to be identified. For the critical distance approach, that is in focus of this work, the critical distance a has to be identified. Additionally, design (i.e. FAT-) values have to be available and a course of the S-N curve that is adapted to the characteristics of TIG-dressed joints.
recommended [13]. The approach using fictitious radii is based on the idea to account for so-called support effects by enlarging the actual radii. The notch stresses calculated with this enlarged radius can directly be used as “fatigue–effective” in a fatigue assessment. Radaj [9] applied this approach to the assessment of welded structures, assuming a worst case radius of rreal = 0 mm and a sharp notch (notch opening angle of ω = 0°). Based on some theoretical considerations, a fictitious radius of
r = rreal + ρ∗ ·s
(1)
was derived. With a micro-structural support length of ρ∗ = 0.4 mm and a factor of s = 2.5 a fictitious radius of r = 1.0 mm was introduced. In many papers, a reliable assessment was documented for the effective notch stress approach applied at specimens in as-welded state [14]. Only at thin (t ⩽ 5 mm ) sheets [15] or welds with a mild weld toe notch [16] (resulting e.g. from a post-weld treatment such as grinding) the application might lead to uncertainties. In these cases, approaches such as the stress averaging or critical distance approach can be applied that lead to a reliable assessment. The theory behind the non-local approaches is to evaluate the stress field or stress gradients in the notch ligament as it is done within the stress averaging approach according to Neuber [17] or the critical distance approach according to Peterson [18]. Both approaches have been successfully used to evaluate fatigue data of various thin- and thick-walled specimens in as-welded condition [19]. A class FAT160 was identified for both approaches. The evaluation was based on results from FE-models with a constant notch radius of r = 0.05 mm .
3. Fatigue of TIG-dressed joints In literature, many papers have been published that deal with the fatigue strength of TIG-dressed welded joints. A collection of papers and a subsequent evaluation of endurable stresses within the nominal stress approach was conducted in [5]. In order to apply local stress approaches, high-quality data needs to be available. This comprises fatigue data itself, geometry of the weld and the specimen, failure location, material data and information on residual stresses. The local geometry of the weld, i.e. the weld profile and the penetration depth, has to be available. This could be provided by a micrograph or a detailed description of the weld geometry. 3D-scan data is very helpful in order to get information on an average geometry. Also information on the axial and angular misalignment is helpful to interpret fatigue data. TIG-dressed joints may fail from various locations, Fig. 3 and Table 1. Therefore, information on the failure location is necessary since the endurable local stresses are dependent on the failure critical detail. In an evaluation, only specimens with failure that origins close to the TIG-layer is considered (No. 1, 2a, 2b, 6 and 9 in Table 1). Specimens with failure starting at other locations are regarded as run-out. The material name, the yield and ultimate strength are given in the majority of publications. But the mechanical properties of the material are highly influenced by the welding process. So the hardness close to the failure critical detail, as measure for the materials strength, should be reported, too. Especially for post-weld treated joints this is important information since the fatigue strength should be correlated at least to some extend to the materials strength. Finally, also information on the residual stress state close to the critical weld detail is of importance. The available data has been gathered and summarised in Table 2. Only fatigue data of butt joints, transverse stiffeners, and cruciform joints has been regarded. The data contains overall 17 different test series. For the subsequent evaluations the individual S-N curves of tests were statistically derived. Only test data with load cycles below the knee-point of the individual S-N curve have been regarded. All fatigue tests with failure locations outside the TIG-pass (No. 3, 4, 5, 7 and 8 in Table 1) have been excluded. All tests have been conducted with similar R-values in the range of 0 ⩽ R ⩽ 0.2 . Therefore, no correction of mean stresses was performed.
2.3. Application at TIG-dressed joints The application of the (effective) notch stress approach is in principle clear for standard welds with a certain defined flank angle, produced, for exampled, by a MAG process. The resulting weld toe notches rtoe are rounded with the 1 mm-notch, Fig. 2, left. For TIG-dressed welds instead, the application is unclear. The reasons can be identified by looking at the typical geometry of such welds, Fig. 2, right. Due to the TIG-dressing a round fillet with a large radius rfillet is produced between the base plate and the weld metal of the original weld. In most case, there is a smooth transition between weld metal and heat affected zone. The notch radius to be enlarged within the effective notch stress approach rtoe can not be defined. The question that arise is, how a fatigue assessment can be conducted bases on local stresses. Different local approaches such as the stress averaging or the critical distance approach can be used for an assessment of TIG-dressed joints. As mentioned above, these approach consider the stress gradients and subsequently the occurring support effects. The stress gradients can easily evaluated within a Finite-Element model. Due to the different failure locations identified in the fatigue tests it is not quite clear, at which positions the stresses should be evaluated. Two failure critical locations need to be assessed. First, the transition between TIG-weld and heat affected zone has to be considered (CP1 in Fig. 2). Even if there is no clear geometrical stress concentration (i.e. notch), the majority of cyclically loaded TIG-dressed welds fail from that position. The reasons are likely to be found the metallurgical notch (difference between weld metal and heat affected zone), maybe
Fig. 2. Local geometry and crack initiation locations at standard and TIGdressed welds.
Fig. 3. different failure locations at TIG-dressed welds. 74
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Table 1 Failure locations for TIG-dressed joints. No.
Failure Location
Description
1
Weld toe
2a 2b 3 4 5 6 7 8 9
Weld metal Weld metal Weld root Base metal Clamping Weld transition Base metal Inner irregularities Base metal
Interface/notch between weld metal of TIG-layer and heat affected zone In the fillet formed by the TIG-dressed layer In the undercut of the TIG-dressed layer Root of the original weld Base material between weld and clamping In the clamping of the specimen Transition between TIG and original weld layer, etc. Base metal below stiffener (only for trans. stiffeners) Inner irregularities like pores, lack of fusion, etc. Heat affected zone
Fig. 4. Evaluation of effective stresses at the weld toe (No. 1) and the position of maximum stress (No. 2).
The following evaluations have been conducted with the critical distance approach. Effective stresses σeff are calculated in a critical distance of a away from the surface.
σeff = σ (a) Table 2 Overview on available fatigue data with detailed weld geometry, clear. Reference
Acronym
Material
Loading
Transverse stiffener S355 3-P. bend. S700 4-P. bend. 304L Axial S31803 Axial S700 Axial S355 Axial S460 Axial S690 Axial
R
Fail. Loc.
0 0.1 0.1 0.1 0.2 0.1 0.1 0.1
1 1∗ 2a∗ 2a∗ Unclear 2b,4,5 2b 2b
[20] [21] [22] [22] [23,24] [25,26] [25,26] [25,26]
RAM PED EUR EUR LEF KUH1 KUH2 KUH3
[27,28,16] [27,28,16] [29–31] [29–31] [29–31] [29–31] [29–31] [29–31]
MEL1 MEL1 ES1 ES2 ES3 ES4 ES5 ES6
Butt joint S960 Axial S1100 Axial S460 Axial S690 Axial S1100 Axial S460 Axial S690 Axial S890 Axial
0.1 0.1 ≈0 ≈0 ≈0 ≈0 ≈0 ≈0
1,2b 1,2b 1 & 2a 2a 1,2a 1,2a 2a 2a
[32]
SCH1
Cruciform joint S355M Axial
0
1
∗)
(2)
The critical distance a is unknown for TIG-dressed joints. In order to identify a recommendable value, it was changed in the range 0 ⩽ a ⩽ 2 mm with a step length of 0.02 mm. For every value of a the scatter of all test data was calculated. As already mentioned, the evaluation was performed at two location, namely at the weld toe and the position in the fillet with σ = σmax . For the calculation of a reference S-N curve a slope needs to be derived. Since an evaluation of the slope by linear regression does in many cases lead to incorrect or illogical results for a heterogeneous database the slope was calculated by averaging and weighting the slope of all 17 test series, see [19]. This evaluation lead to a slope of k = 3.9. Since this slope is quite close to the slope of k = 4.0 as proposed in [5] k = 4.0 was used as pre-set value. The scatter TS of the reference S-N curve can be used as indication of the assessment reliability. The minimum scatter for the evaluation at the position of the weld toe could be determined for a critical distance of a = 0.7 mm , Fig. 5. The reference S-N curve has a scatter of TS = 1: 1.42 . By the evaluation at the location of maximum stress the scatter a slightly lower scatter of TS = 1: 1.39 is derived, Fig. 6. For both reference S-N curves FAT values have been derived by calculating the stresses for a 97.5% survival probability and reducing this value by an additional 10% [34] to account for possible high tensile residual stresses (transformation to R = 0.5). This results in a class FAT160 for the failure at the weld toe and FAT180 for the failure in the fillet. A comparison of the S-N curves for both evaluation positions does not show a major difference. This is likely to be linked to the fact that the stresses at the weld toe and at the position of maximum stress have a similar factor between them. Even though the scatter of the S-N curve is already remarkably low a further reduction might be possible if the material strength is additionally considered. This behaviour was identified by Yıldırım in [5] based on the evaluations of the available fatigue data in the nominal stress system and it should also be present in the local evaluations. To visualise the influence of the material strength on the fatigue strength the results are plotted in Fig. 7 with markers according to the static strength range. With this plot it is possible to discuss that there is a correlation between the fatigue and the material strength. Theoretically, the materials strength has influence on both the support effects and the endurable nominal stresses. The support effects, i.e. the value of a within the critical distance approach, is reduced with increasing materials strength whereas the fatigue strength of the material itself increases with increasing materials strength. Typically, only the latter influence is considered within the nominal stress approach, for example for TIG dressing [5] or HFMI treatment [3]. Nevertheless, a more detailed assessment can be performed within local approaches. In order to include both factors within the evaluation, two modification factors are used. First a factor f to consider the influence of the materials strength on the fatigue strength and second a factor fa for the influence of the materials strength on the support effects. With the factor fa the critical distance a is modified for the evaluation. An inverse linear relationship between a∗ and material strength
not 100%
Comment
GFT-data type “c” type “c” type “c” type “v” type “v” type “v”
4. Calculation of local stresses Finite-Element (FE) models have been created for all specimens with the available fatigue data points. The main geometrical parameters have been regarded as given in micro-sections or in tabled measurement data. In all cases, the TIG-layer was represented by a fillet, that means a circular rounding between original weld and sheet. Especially for welds with a steep flank angle this representation leads to a very good agreement between the existing geometry and the corresponding modelled geometry. For specimens with a small flank angle the surface of the TIG weld had sometimes a varying curvature. Nevertheless, since this exact curvature is unknown prior to the manufacturing, the above-described simplified modelling was used. 2D-models with quadratic plain-strain elements were used in the models since all specimens had a constant cross section. In the vicinity of welds, a structured mesh according to the recommendation in [33] ensures converged stresses. For all the models, a nominal stress of σn = 1 MPa was applied. The stresses were derived in a linear-elastic calculation. A rigid clamping (no rotation, only translation in load direction) was used. Since the exact failure locations are not known for the majority of fatigue tests, the stresses were evaluated at two positions, Fig. 4. Firstly, No. 1 at the weld toe position and secondly, No. 2 at the location of maximum stress. The Maximum Principal stress hypothesis was used for evaluation. Next to the maximum values of stress also the stress gradient on a path perpendicular to the weld surface was taken from the FE-analysis. 75
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Fig. 5. Reference S-N curve for a = 0.7 mm and evaluation at the weld toe (left), dependence of the scatter on the value of a (right).
is assumed, since also nearly linear relationship between the value of ρ∗ ( ρ ∗ ≈ 4 × a , see [35]) and the yield stress was identified by Neuber [36].
a∗=
fa 355 MPa
·Rm + a0 − fa
(3)
The value a 0 is the critical distance for a material strength of 355 MPa . This value is fixed for every evaluation. The factor fa is varied in the range − 0.45 ⩽ fa ⩽ 0 . The influence of the material strength on the fatigue strength is assumed to be linear to the material strength. The factor fRm is varied in the range 0 ⩽ fRm ⩽ 0.5.
f=
fRm 355 MPa
·Rm + 1 − fRm
(4)
The scatter of the reference S-N curve was derived for overall three values of a0 : a0 = 1.0 mm, a0 = 0.8 mm and a 0 = 0.6 mm and for both positions, weld toe and position of maximum stress. In all cases, the scatter could be reduced down to 1: 1.34 ⩽ TS ⩽ 1: 1.37 . The lowest scatter is derived for the value a 0 = 0.6 mm at the
Fig. 7. Reference S-N curve for a = 0.6 mm and evaluation at the position of max. stress.
Fig. 6. Reference S-N curve for a = 0.6 mm and evaluation at the position of max. stress (left), dependence of the scatter on the value of a (right ) . 76
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with large fillet radii and a nearly tangential transition between heat affected zone and weld metal this radius cannot be used. Also from a theoretical point of view there is no need to introduce a fictitious radius for TIG-dressed welds as used for as-welded specimens. From the TIGdressing results a fairly large radius that can be identified, e.g. experimentally or guessed based on experience, and used in the calculation models. In principle, the fatigue strength of different failure critical locations as the weld toe, the heat affected zone or the weld metal in the fillet has to be evaluated separately. All of these locations have been reported at least once in the literature data used for evaluation. Unfortunately, an exact description of each test with its failure location was only reported in few publications. Therefore, a pragmatic approach was followed to evaluate all test data with stresses at the weld toe and all data with the maximum stresses in the fillet. The critical distance approach as easy applicable and reliable approach was used for all evaluations. For a critical distance of a = 0 mm also the use of maximum stresses on the surface can assessed. The evaluation results in a fairly small scatter of TS = 1: 1.42 at the position of the weld toe and TS = 1: 1.39 for the maximum stresses in the fillet. This small scatter surprises to some extend, since the scatter of larger data sets are typically quite high. In [19] a minimum scatter of TS = 1: 2.38 could be derived for as-welded joints with the critical distance approach. Explanation might be that the all tests have been conducted with a similar R-ratio, that the variation in sheet thickness was smaller and that the overall amount of individual tests was smaller. The derived critical distances of a = 0.7 mm for the evaluation at the weld toe and a = 0.56 mm for maximum stresses lie between known distances in literature. In [37] a distance of a = 1.0 mm and in [19] a distance of a = 0.1 mm of was recommend for the assessment of joints in as-welded state. Since the failure critical locations differ (TIG: smooth fillet or metallurgical notch, as-welded: sharp geometrical notch) differences in the critical distance might be expected. The value of a is quite insensitive to a change due to a similarity of the stress gradients. Even with a value of a = 1.0 mm or a = 0.1 mm still a fair assessment would result. The resulting reference S-N curve has a scatter below TS = 1: 1.5, see the correlation between TS and the critical distance a in Figs. 6 and 5. Using the position of maximum stresses in the fillet results to somewhat smaller scatter and subsequently to a higher assessment reliability. But since the difference in scatter is quite low, no
Fig. 8. Dependence of the scatter on the values fRm and fa .
position of maximum stress. For a value of fRm = 0.1 and fa = 0 the highest reduction in scatter can be identified in Fig. 8. A view on the reference S-N curve shows that there is no influence of the materials strength on the fatigue strength visible anymore, Fig. 9. The parameters of the reference S-N curve are: Δσtcd, a, mP, N = 2e6 = 233 MPa, k = 4, 0 and tS = 1: 1.34 . From this reference S-N curve a class FAT160 was derived. 5. Discussion Fatigue assessment with local approach at welded joints is most commonly associated with the effective notch stress approach with the reference radius rref = 1.0 mm . This radius can be used for as-welded joints with sharp weld toe or weld root notches. For TIG-dressed welds
Fig. 9. Reference S-N curve for the evaluation at the position of maximum stress, a = 0.6 mm, fRm = 0.1 and fa = 0 . 77
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recommendation for any of both methods can be given at the moment. For an automated assessment the location at the weld toe might be favoured since the position is already determined during pre-processing of a FE-model. In the evaluation of TIG-dressed joints using nominal stresses [5], an influence of the materials strength on the fatigue strength was identified. This influence could also be determined for the evaluations with local stresses even though, with only 10%, to a much less degree. The reason for this deviation is not yet clear. It might be explained if the fillet radii are depending on the materials strength. However this effects could not be confirmed. An pronounced influence of the materials strength in the support effects resp. the critical distance could not be identified. The reasons for that might be that the stress gradients are quite shallow and subsequently the support factors are quite low.
joints using reference radii. Int J Fatigue 2017;101:459–68. [9] Radaj D, editor. Design and analysis of fatigue resistant welded structures. Cambridge: Woodhead Publishing; 1990. [10] Olivier R, Koettgen VB. Schweißverbindungen I - Schwingfestigkeitsnachweis für Schweißverbindungen auf der Grundlage örtlicher Beanspruchungen, Abschlussbericht, Number 105. Forschungkuratorium Maschinenbau (FKM); 1989. [11] Olivier R, Koettgen VB. Schweißverbindungen II - Untersuchung zur Einbindung eines neuartigen Zeit- und Dauerschwingfestigkeitsnachweises von Schweißverbindungen aus Stahl in Regelwerke, Number 128. Forschungkuratorium Maschinenbau (FKM); 1991. p. 389. [12] Koettgen VB, Olivier R, Seeger T. Fatigue analysis of welded connections based on local stresses. IIW - Commission XIII, XIII-1408-91; 1992. [13] Hobbacher Adolf. Database for the effective notch stress method at steel. Int Inst Weld, JWG-XIII-XV-197-08; 2008. [14] Pedersen MM, Mouritsen O, Hansen MR, Andersen JG, Wenderby J. Re-analysis of fatigue data of welded joints using the notch stress approach. Int J Fatigue 2010;32(10):1620–6. [15] Bruder T, Störzel K, Baumgartner J, Hanselka H. Evaluation of nominal and local stress based approaches for the fatigue assessment of seam welds. Int J Fatigue 2012;34(1):86–102. [16] Möller Benjamin, Baumgartner Jörg, Wagener Rainer, Kaufmann Heinz, Melz Tobias. Low cycle fatigue life assessment of welded high-strength structural steels based on nominal and local design concepts. Int J Fatigue 2017;101:192–208. [17] Neuber H. Kerbspannungslehre. 2nd ed. Springer Verlag; 1958. ISBN 9783540135586. [18] Peterson RE. Metal fatigue. McGraw-Hill Book Company, Inc.; 1959. [19] Baumgartner J, Schmidt H, Ince E, Melz T, Dilger K. Fatigue assessment of welded joints using stress averaging and critical distance approaches. Weld World 2015;59(5):731–42. [20] Ramalho Armando L, Ferreira José AM, Branco Carlos AGM. Fatigue behaviour of t welded joints rehabilitated by tungsten inert gas and plasma dressing. Mater Des 2011;32(10):4705–13. [21] Pedersen Mikkel Melters, Mouritsen Ole Østergaard, Hansen Michael Rygaard, Andersen Jes Grøn, Wenderby Jimmi. Comparison of post-weld treatment of highstrength steel welded joints in medium cycle fatigue. Weld World 2010;54(78):R208–17. [22] Maddox SJ, Hopkin GR, Holy A, Moura Branco CA, Infante V, Baptista R, et al. Improving the fatigue performance of welded stainless steels, Technical report, EUR 22809 EN, European Commission; 2007. [23] Lagerqvist O, Clarin M, Gozzi J, Völling B, Pak D, Stötzl J, et al. Efficient lifting equipment with extra high-strength steel. Technical report, European Commission; 2005. [24] Lieurade HP, Huther I, Lefebvre F. Effect of weld quality and postweld improvement techniques on the fatigue resistance of extra high strength steels. Weld World 2008;52(7-8):106–15. [25] Dürr A. Zur Ermüdungsfestigkeit von Schweißkonstruktionen aus höherfesten Baustählen bei Anwendung von UIT-Nachbehandlung, PhD thesis, Universität Stuttgart; 2007. [26] Kuhlmann U, Dürr A, Bergmann J, Thumser R. FOSTA Forschungsvorhaben P 620: Effizienter Stahlbau aus höherfesten Stählen unter Ermüdungsbeanspruchung, Verlag und Vertriebsgesellschaft mbH, Düsseldorf; 2006. [27] Melz Tobias, Möller Benjamin, Baumgartner Jörg, Ummenhofer Thomas, Herion Stefan, Hrabowski Jennifer, et al. FOSTA P900: Erweiterung des örtlichen Konzeptes zur Bemessung von LCF-beanspruchten geschweißten Kranstrukturen aus hochfesten Stählen. Technical report, Forschungsvereinigung Stahlanwendung e.V.; 2015. [28] Hrabowski Jennifer, Herion Stefan, Ummenhofer Thomas. Kurzzeitfestigkeit von Schweißverbindungen aus höchst- und ultrahochfesten stählen. [29] van Es SHJ. Effect of TIG-dressing on fatigue strength and weld toe geometry of butt welded connections in high strength steel. Master’s thesis, Delft University of Technology; 2012. [30] van Es SHJ, Kolstein MH, Pijpers RJM, Bijlaard FSK. TIG-dressing of high strength steel butt welded connections – part 1: weld toe geometry and local hardness. Proc Eng 2013;66:216–25. [31] van Es SHJ, Kolstein MH, Pijpers RJM, Bijlaard FSK. TIG-dressing of high strength butt welded connections – part 2: Physical testing and modelling. Proc Eng 2013;66:126–37. [32] Schliebner. Experimentelle und numerische Untersuchungen zur Bewertung der Schwingfestigkeit von Hybridschweißverbindungen, PhD thesis, Technische Universität Darmstadt; 2005. [33] Baumgartner J, Bruder T. An efficient meshing approach for the calculation of notch stresses. Weld World 2013;57:137–45. [34] Sonsino Cetin Morris. A consideration of allowable equivalent stresses for fatigue design of welded joints according to the notch stress concept with reference radii rref = 1.00 and 0.05 mm. Weld World 2009;53:64–75. [35] Taylor D, Hoey D. High cycle fatigue of welded joints: the tcd experience. Int J Fatigue 2009;31:20–7. [36] Neuber H. Über die Berücksichtigung der Spannungskonzentration bei Festigkeitsberechnungen. Konstruktion 1968;20(7):245–51. [37] Xiao Zhi-Gang, Yamada Kentaro. A method of determining geometric stress for fatigue strength evaluation of steel welded joints. Int J Fatigue 2004;26:1277–93.
6. Conclusions A fatigue assessment approach for TIG-dressed joints was derived in this paper. For the application only a simplified modeling of the joint, regarding only the rough weld profile and the fillet radius of the TIGweld, and a linear-elastic calculation is necessary. For an assessment between two locations for evaluating stresses can be chosen.
• Position of the weld toe: The stresses in a distance of a = 0.7 mm •
below the surface should be used. These can be evaluated against the stress range Δσtcd, a, mP = 242 MPa (Survival probability: Ps = 50%) or a class FAT160. Position of the maximum stress: The stresses in a distance of a = 0.6 mm below the surface should be used. These can be evaluated against the stress range Δσtcd, a, mP = 253 MPa (Survival probability: Ps = 50%) or a class FAT180.
A slight improvement in the assessment reliability can be achieved if the materials strength is included in the evaluation. For the evaluation at the fillet a class FAT160 can be used that is multiplied by the factor f given in Eq. (4). For a further validation and improvement of these recommendations, high quality fatigue data for TIG-dressed joints (statistically ensured weld geometry, exact failure location, hardness distribution close to the failure location, residual stress, etc.) and also further fatigue tests on other structural details (joint types next to butt welds, transverse stiffeners and cruciform joints) needs to be available. With these future data an assessment would be enabled that relies to an higher extent on physically based descriptions than on rough empirical assumptions. Researchers are encouraged to document all tests in detail and provide data for further evaluations. References [1] Haagensen PJ, Maddox S. IIW recommendations on methods for improving the fatigue lives of welded joints. Cambridge: Woodhead Publishing Ltd.; 2013. [2] Kirkhope KJ, Bell R, Caron L, Basu RI, Ma K-T. Weld detail fatigue life improvement techniques. Part 1: review. Mar Struct 1999;12:447–74. [3] Marquis Gary B, Barsoum Zuheir. IIW Recommendations on High Frequency Mechanical Impact (HFMI) Treatment for Improving the Fatigue Strength of Welded Joints. Singapore, Singapore: Springer; 2016. ISBN 978-981-10-2504-4. pp. 1–34. [4] Watkinson F, Bodger PH, Harrison JD. The fatigue strength of welded joints in high strength steels an methods for its improvement. UK: The Welding Institute; 1971. [5] Yıldırım Halid Can. Review of fatigue data for welds improved by tungsten inert gas dressing. Int J Fatigue 2015;79:36–45. [6] Hobbacher AF. Recommendations for fatigue design of welded joints and components. Springer International Publishing; 2016. [7] Fricke Wolfram. IIW recommendations for the fatigue assessment of welded structures by notch stress analysis. Woodhead Publishing Ltd., Cambridge. International Institute of Welding, Paris; 2012. [8] Baumgartner J. Review and considerations on the fatigue assessment of welded
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