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A study of fatigue and fracture response of cantilevered luminaire structures made from aluminum alloy 6063
Materials Science and Engineering A 527 (2010) 4680–4686
Contents lists available at ScienceDirect
Materials Science and Engineering A journal homepage: www.elsevier.com/locate/msea
A study of fatigue and fracture response of cantilevered luminaire structures made from aluminum alloy 6063 Craig C. Menzemer a , Diya Azzam a,b,1 , T.S. Srivatsan c,∗ a b c
Department of Civil Engineering, The University of Akron, Akron, OH 443265, USA California Department of Transportation (Caltrans), Bridge Structure Design (Branch, 10) Los Angeles Projects, 1801 30th Street, Sacramento, CA 95816, USA Division of Materials Science and Engineering, Department of Mechanical Engineering, The University of Akron, Akron, OH 44325-3903, USA
a r t i c l e
i n f o
Article history: Received 30 January 2010 Received in revised form 19 March 2010 Accepted 22 March 2010
Keywords: Fatigue life prediction R ratio Plastic zone Equivalent SIF range Stress range
a b s t r a c t In the experimental results elegantly and exhaustively elaborated upon in this paper the local stresses, obtained from finite element analysis, was used to develop estimates of the stress intensity factor (SIF). In combination with crack growth data, the fatigue lives of both the through-plate and an integrally stiffened socket connection were estimated using software developed by the U.S. Air Force (and referred to as AFGROW). The fatigue life estimates correlated well with the test results provided the crack growth rate data was obtained under conditions of minimal closure at higher stress ratios (of the order R = 0.7). In an attempt to establish the fatigue lives in the high cycle regime, the measured residual stresses had to be included in the analysis. For identical stress ranges, the 25 mm thick through-plate socket connection exhibited noticeably lower fatigue lives when compared to the integrally stiffened shoebase structure. Scanning electron microscopy observations revealed pockets of well-defined striations consistent with stable growth of the crack through the microstructure prior to the onset of unstable crack growth culminating in catastrophic fracture. In the slow growth region, the fracture surface revealed pockets of shallow, well-defined striations that were uniformly spaced indicative of the occurrence of localized microplastic deformation. © 2010 Elsevier B.V. All rights reserved.
1. Introduction Fatigue life of welded connections is typically governed by the detail type or geometry and the applied stress range [1–4]. Through the years of sustained research and development efforts at studying and rationalizing the fatigue behavior of structural materials the stress range has been used as the primary parameter governing fatigue life as a direct result of the existence of tensile residual stresses immediately adjacent to the welded detail [5,6]. Welding results in localized heating around the joint. As the weldment and immediate adjacent area is the last to cool, tensile residual stresses develop as a result of the constraints on shrinkage imposed by the material farther away from the connection. Because of the existence of tensile residual stresses, early stages of crack growth and damage accumulation tend to occur under conditions of minimal closure. The discipline of fracture mechanics is both conducive and applicable to the study of welded structures and has been extensively used [7–9].
∗ Corresponding author. Tel.: +1 330 972 6196; fax: +1 330 972 6027. E-mail address: [email protected] (T.S. Srivatsan). 1 Formerly graduate student at The University of Akron. 0921-5093/$ – see front matter © 2010 Elsevier B.V. All rights reserved. doi:10.1016/j.msea.2010.03.085
A number of specimens were fatigue tested as part of an experimental study to examine the fatigue strength and behavior of welded aluminum mono-pole, cantilever supports for luminaires. Two different extruded aluminum poles to support base details were tested. Round extrusions of aluminum alloy 6063 in the condition having a diameter of 250 mm and 6.25 mm thick walls were welded to either an integrally stiffened cast base (referred to as a “shoe-base”) of the cast aluminum alloy A356 or a 25.4 mm thick plate of aluminum alloy 6061 in the T6 condition (temper). The plate supports are referred to as “through-plates” as the extruded tube passes through or is inserted into a hole cut out of the plate. The fillet welds were placed between the tube and the plate at both the top and bottom. For specific details on the shoe-base, an extruded tube is fit to a cast base support, and extends through a hole or socket in the casting. The fillet welds were placed between the tube and the shoe-base at the top-side and bottom-side of the casting. The filler metal was taken to be aluminum alloy 4043. What is interesting to note is the sequence used to manufacture the test specimens. Each extruded pole was inserted into the base, whether a casting or a plate. Fillet welds were placed between the tube and top of the casting or plate. Each sample was subsequently repositioned and the fillet weld between the tube and bottom of the base was placed. Each assembly was then placed in an aging oven to bring the extruded tube and base to a near T6
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Fig. 3. Failed light pole specimen with shoe-base detail. Fig. 1. Shoe-base detail and residual stress measurements.
condition. Upon completion of aging, the specimens were available for mechanical testing. Several samples were used to measure the near weldment residual stresses using the hole drilling method as shown in Fig. 1. Although discussed elsewhere in detail, near surface residual stresses in the tube adjacent to the top fillet weld were consistently in compression for both details and conditions [10]. In essence, shrinkage of the bottom weld attempts to pull the tube into the base thereby, placing the surface of the tube near the joint in compression. The presence of compressive residual stresses at the surface would be expected to alter the fatigue behavior and use of fracture mechanics approach to modeling. The fatigue specimens were 3048 mm long and secured by bolting the base supports to a heavy steel fixture. Each sample was cyclically loaded at the tip of the cantilever as shown in Fig. 2. Loads were applied at a small positive R ratio (minimum stress/maximum stress), and strains immediately adjacent to the weld detail were carefully monitored during the course of each test through the use and application of uniaxial strain gauges. The shoe-base specimens were tested at stress ranges from a low of 24 MPa to a maximum of 60 MPa. The through-plate samples were evaluated at stress ranges from 6.2 MPa to 31 MPa. The fatigue cracks formed on the tension side of the cantilever, along the weld toe between the casting and tube or the plate and tube, are shown in Fig. 3. Failure was defined as the development of a crack that penetrated the thickness of
Fig. 2. Light poles loaded at tip of cantilever.
the tube. Additional details of the testing program can be found elsewhere [11].
2. Fatigue life estimates using fracture mechanics Fatigue life estimates were prepared for both types of details using fracture mechanics techniques. Development of stress intensity factors for both the shoe-base and through-plate geometries were facilitated by use of the finite element method (FEM). The finite element method (FEM) was employed to determine the stress gradient local to the welded region. Local stresses from the different FEM analyses were combined with the measured residual stresses, published crack growth data for the 6XXX-series alloys, and knowledge of typical initial flaw sizes for welded aluminum construction as input to AFGROW [12–14]. The Air Force crack growth prediction software tool that enables the user to assess the fatigue life (Nf ) of structures is AFGROW. An extensive library within AFGROW provides models for a range of different crack geometries. Of equal importance is the intrinsic ability within AFGROW to input local stresses that may include residual stresses. A parametric study of both types of details was conducted as part of this research investigation in an attempt to identify some of the important geometric parameters that influence the stresses local to the base—tube welded connections. ANSYS was employed in all of the analyses conducted. Several models are shown with results for the as-tested geometries exemplified in Fig. 4 [15]. An extensive description of the parametric study is provided elsewhere [16]. However, one of the important parameters was found to be flexural stiffness of the support base. While the shoe-base is relatively stiff, the 25.4 mm through-plate is not. The models mimicked the test geometries and were loaded at the tip of the cantilever. Distortion of the base plate of the through-plate detail results in additional bending through the tube wall. The fatigue tests clearly showed that the elevated stresses near the weld toe joining the tube to the base plate resulted in significantly lower fatigue lives when compared to poles utilizing the shoe-base support. In order to investigate stiffness of the base and fatigue life, analyses were conducted for the as-tested shoe-base and through-plate test specimen geometry as well as a through-plate detail having a base thickness of 75.2 mm. The local stresses from the three different models were used for the prediction of fatigue life. For purposes of this study, previously developed methodologies were used to evaluate a correction factor for input to AFGROW. As described in an earlier study by Zettlemoyer [17], the stress
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Fig. 4. Finite element models—mesh and sample results.
intensity factor is numerically evaluated using the relationship: √ (1) K = Fe Fw Fs Fg Sr a In this relationship, K is the stress intensity factor range; Fe is the correction factor for crack shape; Fw is the correction factor for finite width; Fs is the correction factor for free surface; Fg is the stress gradient correction factor that makes use of the FEM Model results; Sr is the stress range, and a is the crack size. For direct input to AFGROW, the correction factor, i.e., the stress multiplication factor, was estimated using the relationship SMF = Fe Fw Fs Fg Sr
(2)
In this expression SMF stands for the Stress Multiplication Factor. A successful application of fatigue life prediction using fracture mechanics depends upon the conjoint and mutually interactive influences of (i) similitude, and (ii) the appropriate choice of crack growth data or development of the data under conditions consistent with laboratory tests. The crack growth rate data for the extruded aluminum alloy was represented by an expression developed by Paris and Erdogan [18]: da m = C(K) dN
(3)
In this expression, da/dN is the crack growth rate with a the crack size and N the number of cycles, K is the stress intensity factor, and C and m are constants determined from the measured experimental data that depend on the following: (i) material, (ii) specimen dimensions, (iii) test conditions, and (iv) nature of loading. In this study, fatigue life prediction made use of crack growth rate data at three different R ratios including 0.0, 0.5 and 0.7. For a typical welded construction, crack growth data at an elevated R ratio would be used since tensile residual stresses would tend to minimize crack closure [19]. The crack growth data obtained at an R ratio of 0.0 would be expected to contain some closure and tend to the fatigue test conditions for the light poles that were loaded at a very small positive load ratio. The estimates of fatigue life were initially prepared for the through-plate socket connection having a 25.4 mm thick base plate. For the through-thickness plate specimen, failure due to fatigue is defined and accepted as the number of cycles required to grow a crack through the thickness of the light pole tube. In Fig. 5 is shown the fatigue life estimates obtained from crack growth data obtained for a stress ratio R (minimum / maximum ) of 0.0, both with and without input of the measured residual stresses, when compared to the test results. There exists a significant scatter in the
Fig. 5. Fatigue life prediction with and without residual stress compared to the experimental results for the through-plate socket connection.
experimental data obtained in the laboratory. While inclusion of the residual stresses does increase the estimated fatigue life, it is clear that the predictions fail to correlate with the observed test results. The results of fatigue life estimates for the shoe-base connection developed using crack growth data taken at a stress ratio (R = minimum / maximum ) of 0.0 is shown in Fig. 6. The residual stresses were included in the analyses. In general, the predicted fatigue lives were larger when compared to the experimentally determined test data. This was particularly evident at the stress ranges () between 42 MPa and 65 MPa. However, fatigue life predictions improved as the stress range () was gradually reduced. Further, it is observed that there is good agreement between the predicted fatigue lives and the experimentally determined fatigue life for stress ranges between 24 MPa and 38 MPa. As previously discussed, careful measurements revealed the presence of compressive residual stresses on the surface of the luminaire tube immediately adjacent to the weld. The presence of compressive residual stress did complicate the prediction of fatigue life. In an attempt to account for the intrinsic influence of residual stress, the fatigue crack growth (da/dN) data obtained at stress ratios of R = 0.5 and R = 0.7 were utilized. The occurrence of crack closure at the higher values of the stress ratio, that is, R = 0.5 and R = 0.7, is minimal and crack growth through the microstructure of aluminum alloy 6063 is governed by the presence of intrinsic
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Fig. 6. Fatigue life predictions compared to experimental data for the shoe-base connection using fatigue crack growth data for R = 0.0 with residual stresses included.
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Fig. 8. Comparison of experimental data for the through-plate and shoe-base connections with fatigue life predictions for a 76.2 mm through-plate connection using fatigue crack growth data with R = 0.7 with residual stresses included.
Fig. 7. Fatigue life predictions for the through-plate socket connection compared to experimental data using fatigue crack growth data with R = 0.5 and R = 0.7 including residual stresses.
Fig. 9. Fatigue life predictions for through-plate socket connections using fatigue crack growth data with R = 0.5 and R = 0.7 including the effects of residual stresses.
microstructural features and their specific role and contribution in governing fatigue deformation. The crack growth data (da/dN) obtained at the higher values of the stress ratio (R) coupled with the measured residual stresses were used to simulate the conditions that are likely to exist in the fatigue test specimen. The predicted fatigue lives for the through-plate socket connection having a 25.4 mm thick plate is compared with the experimental test data and is shown in Fig. 7. In general, it is clear that the predicted fatigue lives tend to overestimate the test results obtained from the experiments. A comparison of the fatigue life predictions for the through-plate socket connection having a 76.2 mm thick base plate with the test results obtained from experiments on specimens with both through-plate and shoebase connections is shown in Fig. 8. The observed behavior of the through-plate socket connection having a 76.2 mm thick base is similar to that of the integrally stiffened shoe-base joint as the flexural stiffness of each base is similar, as are the stress fields adjacent to the weldments. The 76.2 mm plate thickness minimizes distortion and through thickness bending of the tube wall is relatively small. The fatigue life estimates for the shoe-base socket connection along with corresponding data obtained from actual tests conducted in the laboratory is shown in Fig. 9. The fatigue life prediction was noticeably improved when crack growth data was obtained at the higher stress ratios (R) coupled with measured residual stresses.
At a high load ratio (R > 0.6) the fatigue crack growth data minimizes the effects of closure and is often used in the fracture mechanics study of welded structures to account for the presence of tensile residual stresses. Use of high load ratio (R) data with measured compressive residual stresses for the tested cantilever specimen’s accounts for the observed fatigue behavior. A low load ratio [R] test data combined with the measured compressive residual stresses tends to over compensate or retard fatigue crack growth beyond what was observed experimentally. The predicted fatigue lives for the shoe-base connection with fatigue test data obtained from the experimental tests using the current AASHTO design stress (S) versus fatigue life (N) curves for both the Category ‘D’ and Category ‘E’ is compared in Fig. 10. The categories are results of full-scale fatigue tests undertaken to categorize each welded detail. Samples in the database used to develop the stress (S) versus fatigue life (N) curves were tested under constant amplitude load conditions. For fatigue lives in excess of 105 cycles, Category D provides a reasonable estimate of the lower bound fatigue strength. Further, the predicted fatigue lives are in reasonable agreement with the laboratory test results for the variation of stress (S) with fatigue life (Nf ) for Category D. The predicted fatigue lives for the through-plate socket connection for the 25.4 mm thick and 76.2 mm thick base plates along with the experimental test results are shown in Fig. 11. Also shown is
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Fig. 10. Design curves for Categories D and E details compared to shoe-base laboratory fatigue data as well as fatigue life prediction using fatigue crack growth data with R = 0.7.
the companion stress (S) versus fatigue life (Nf ) curves put forth by AASHTO for Category D, Category E and Category E’. Much of the experimental test data falls well below the lower bound defined by Category E’ ( endurance = 7 MPa). The predicted fatigue
Fig. 11. Experimental results for the through-plates socket connection compared to AASHTO design Categories D, E and E’ as well as life predictions.
lives tend to be in reasonable agreement with Category E’. However, the predicted fatigue lives for the 76.2 mm thick base plate are significantly improved and quite consistent with the Category ‘D’ details.
Fig. 12. Scanning electron micrographs of the aluminum alloy light pole sample that was cyclically deformed at a stress range of 37 MPa at 1 Hz resulting in 1.6 × 106 cycles to failure, showing: (a) overall morphology showing distinctly the region of early crack initiation and early crack growth. (b) High magnification of (a) showing fine striation-like features in the region of stable crack growth. (c) High magnification of the region of early crack initiation showing the overlap of striations to form a ripple-like pattern. (d) The region of slow and stable crack growth revealing shallow striations reminiscent of localized microplastic deformation.
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Fig. 13. Scanning electron micrographs of the aluminum alloy light pole sample that was cyclically deformed at a stress range of 37 MPa at 1 Hz resulting in 1.6 × 106 cycles to failure, showing: (a) a non-linear fatigue crack traversing the recrystallized grain boundary in the region of early crack growth. (b) Cracking along the grain boundary separating the regions of early crack growth and stable crack growth. (c) An array of fine microscopic cracks in the region of stable crack growth.
3. Failure-damage analysis: characterization of the fracture surface Careful and comprehensive scanning electron microscopy (SEM) observations of the fatigue fracture surface were made on the test samples taken from a shoe-base connection that was cyclically deformed at a stress range of 36.5 MPa and having a resultant fatigue life of 1.6 × 106 cycles. An examination of the fracture surfaces of the deformed and failed fatigue specimens in a JEOL SEM was done at: (a) Low magnifications to identify the precise size, shape and location of microscopic crack initiation and early crack growth, and the progress of stable crack growth through the alloy microstructure, culminating in unstable crack growth and failure due to an overload. (b) Higher magnifications in the regions of early crack growth, stable crack growth and unstable crack growth to identify the following: (i) Nature and severity of damage initiation. (ii) Nature of microscopic crack growth. Overall, the fatigue fracture surface revealed features that are typical of fatigue fracture. The fracture surface was essentially a flat topography at the chosen value of stress range. On a micro-
scopic scale, the nature and morphology of the intrinsic features on the fatigue fracture surfaces was found to depend on the level of maximum stress, at the chosen load ratio. Only representative fractographs of the fatigue fracture surfaces are shown in Figs. 12 and 13. The presence of a flat fracture surface area having a distinct region that identifies well-defined crack growth and sharp transfer to unstable crack growth culminating in catastrophic fracture is sufficient for identification of failure due to fatigue. Overall morphology of the fracture surface revealed the distinct regions of crack initiation and early microscopic crack growth (Fig. 12a). High magnification observations in the region of early crack growth revealed striation-like features (Fig. 12b) reminiscent of localized microplastic deformation. At higher allowable magnifications of the SEM the region of crack initiation and early crack growth revealed an overlap of the pockets of striations creating a ripple-like pattern on the fracture surface (Fig. 12c). The shallow nature of the striations was distinctly evident in the region of slow and stable crack growth (Fig. 12d). Also evident on the fracture surface of the 6063 alloy sample were isolated cracks, both macroscopic and fine microscopic, that were essentially non-linear in nature. Several of these cracks were found traversing through the recrystallized grain boundaries (Fig. 13a). There are a few situations that favor the occurrence of cracking along the recrystallized grain boundaries. For this aluminum alloy (i.e., 6063) of interest these include the following:
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(i) Precipitation and presence of brittle second-phase particles both at and along the grain boundary region. (ii) Hydrogen embrittlement and liquid metal embrittlement depending on the severity of the environment the alloy is exposed to. (iii) Environment assisted cracking (EAC). (iv) Intergranular corrosion. (v) Grain boundary cavitations and cracking at high temperatures.
6. Fracture surface examination was conducted for the failed shoebase connection in a scanning electron microscope. The fracture surface revealed fine and shallow striation-like features in the region of slow and stable growth and a combination of widely spaced striations and cracks in the region of unstable growth approaching overload.
The presence of fine second-phase particles along the grain alloy boundary regions coupled with local stress concentration effects and their synergistic interactions is responsible for the observed cracking that is few and far in-between. At low magnification the macroscopic cracks were evident in large numbers in the region separating early crack growth and stable crack growth (Fig. 13b). At the higher allowable magnifications the region of stable crack growth also revealed fine microscopic cracks (Fig. 13c). The microscopic cracks though few were well defined and distinct during the early stages of unstable crack growth.
The authors extend most sincere thanks and appreciation to the unknown reviewer for his many useful comments and corrections pertaining to the text and relevance of this technical manuscript.
4. Conclusions Based on a fairly exhaustive examination of the fatigue behavior of cantilevered luminaire structures the following are the key observations: 1. The fatigue test results correlate well with US Air Force AFGROW prediction of life provided the crack growth rate data used was obtained under conditions of minimal closure and at a high stress ratio (R) of 0.7. 2. At identical stress ranges, the through-plate connection having a base plate thickness of 25.4 mm exhibited lower predicted fatigue life when compared to the same detail having a 76.2 mm thick base plate. 3. The fatigue life of the through-plate socket connection having a 76.2 mm thick base plate was similar to the light pole having a shoe-base joint. 4. The fatigue strength of through-plate connection having a 25.4 mm thick base was below the stress-fatigue curve predicted by AASHTO Category E’. 5. For fatigue lives in excess of 105 cycles, the shoe-base connection behaved in a manner quite consistent with the stress-fatigue life curve put forth by AASHTO for Category D.
Acknowledgements
References [1] J.W. Fisher, Bridge Fatigue Guide—Design and Details, American Institute of Steel Construction, Chicago, 1977. [2] J.W. Fisher, M.A. Hirt, B.M. McNamee, Effect of weldments on the fatigue strength of steel beams, National Cooperative Highway Research Program, NCHRP Report 102, Transportation Research Board, Washington, DC, 1970. [3] D. Kosteas, K. Polas, U. Graf, Proceedings of the 3rd International Aluminum Conference, International Aluminum Institute, 1985. [4] A. Hobbacher, Fatigue Design of Welded Joints and Components, International Institute of Welding IIW, Abingdon Publishing, Cambridge, 1996. [5] S.J. Maddox, Influence of Tensile Residual Stresses on the Fatigue Behavior of Joints in Steel: Residual Stress Effects in Fatigue, ASTM STP 776, 1982. [6] T.R. Gurney, Fatigue of Welded Structures, 2nd ed., Cambridge University Press, 1979. [7] S. Maddox, Proceedings of the Sixth International Conference on Aluminum Weldments, American Welding Society, Cleveland, OH, 1995. [8] R. Jaccard, Proceedings, 2nd International Conference on Aluminum Weldments, Technical University of Munich, 1982. [9] R. Wei, Journal of Engineering Materials and Technology 100 (1978). [10] D. Azzam, C.C. Menzemer, Journal of Structural Engineering-ASCE 132 (10) (2006) 1603–1610. [11] D. Azzam, The fatigue behavior of highway welded aluminum light pole support structures, Doctor of Philosophy Thesis, The University of Akron, May 2006. [12] AFGROW, Version 4.0010.13, Air Force Research Laboratory, Wright Patterson Air Force Base, 2005. [13] C. Menzemer, Fatigue behavior of welded aluminum structures, Doctor of Philosophy Thesis, Lehigh University, 1992. [14] R. Jaccard, Fatigue Crack Propagation in Aluminum, International Institute of Welding, 1990 (document XIII-1377-90). [15] Swanson Analysis Systems, ANSYS Release 9.0, Swanson Analysis Systems, 2004. [16] D. Azzam, C.C. Menzemer, Journal of Structural Engineering-ASCE 134 (February (2)) (2008) 173–180. [17] N. Zettlemoyer, Stress concentration and fatigue of welded details, Doctor of Philosophy Thesis, Lehigh University, 1976. [18] P.C. Paris, F. Erdogan, ASME Trans, Journal of Basic Engineering 85 (1963). [19] W.A. Herman, R. Hertzberg, R. Jaccard, Fatigue and Fracture of Engineering Materials and Structures 11 (4) (1988).
Contents lists available at ScienceDirect
Materials Science and Engineering A journal homepage: www.elsevier.com/locate/msea
A study of fatigue and fracture response of cantilevered luminaire structures made from aluminum alloy 6063 Craig C. Menzemer a , Diya Azzam a,b,1 , T.S. Srivatsan c,∗ a b c
Department of Civil Engineering, The University of Akron, Akron, OH 443265, USA California Department of Transportation (Caltrans), Bridge Structure Design (Branch, 10) Los Angeles Projects, 1801 30th Street, Sacramento, CA 95816, USA Division of Materials Science and Engineering, Department of Mechanical Engineering, The University of Akron, Akron, OH 44325-3903, USA
a r t i c l e
i n f o
Article history: Received 30 January 2010 Received in revised form 19 March 2010 Accepted 22 March 2010
Keywords: Fatigue life prediction R ratio Plastic zone Equivalent SIF range Stress range
a b s t r a c t In the experimental results elegantly and exhaustively elaborated upon in this paper the local stresses, obtained from finite element analysis, was used to develop estimates of the stress intensity factor (SIF). In combination with crack growth data, the fatigue lives of both the through-plate and an integrally stiffened socket connection were estimated using software developed by the U.S. Air Force (and referred to as AFGROW). The fatigue life estimates correlated well with the test results provided the crack growth rate data was obtained under conditions of minimal closure at higher stress ratios (of the order R = 0.7). In an attempt to establish the fatigue lives in the high cycle regime, the measured residual stresses had to be included in the analysis. For identical stress ranges, the 25 mm thick through-plate socket connection exhibited noticeably lower fatigue lives when compared to the integrally stiffened shoebase structure. Scanning electron microscopy observations revealed pockets of well-defined striations consistent with stable growth of the crack through the microstructure prior to the onset of unstable crack growth culminating in catastrophic fracture. In the slow growth region, the fracture surface revealed pockets of shallow, well-defined striations that were uniformly spaced indicative of the occurrence of localized microplastic deformation. © 2010 Elsevier B.V. All rights reserved.
1. Introduction Fatigue life of welded connections is typically governed by the detail type or geometry and the applied stress range [1–4]. Through the years of sustained research and development efforts at studying and rationalizing the fatigue behavior of structural materials the stress range has been used as the primary parameter governing fatigue life as a direct result of the existence of tensile residual stresses immediately adjacent to the welded detail [5,6]. Welding results in localized heating around the joint. As the weldment and immediate adjacent area is the last to cool, tensile residual stresses develop as a result of the constraints on shrinkage imposed by the material farther away from the connection. Because of the existence of tensile residual stresses, early stages of crack growth and damage accumulation tend to occur under conditions of minimal closure. The discipline of fracture mechanics is both conducive and applicable to the study of welded structures and has been extensively used [7–9].
∗ Corresponding author. Tel.: +1 330 972 6196; fax: +1 330 972 6027. E-mail address: [email protected] (T.S. Srivatsan). 1 Formerly graduate student at The University of Akron. 0921-5093/$ – see front matter © 2010 Elsevier B.V. All rights reserved. doi:10.1016/j.msea.2010.03.085
A number of specimens were fatigue tested as part of an experimental study to examine the fatigue strength and behavior of welded aluminum mono-pole, cantilever supports for luminaires. Two different extruded aluminum poles to support base details were tested. Round extrusions of aluminum alloy 6063 in the condition having a diameter of 250 mm and 6.25 mm thick walls were welded to either an integrally stiffened cast base (referred to as a “shoe-base”) of the cast aluminum alloy A356 or a 25.4 mm thick plate of aluminum alloy 6061 in the T6 condition (temper). The plate supports are referred to as “through-plates” as the extruded tube passes through or is inserted into a hole cut out of the plate. The fillet welds were placed between the tube and the plate at both the top and bottom. For specific details on the shoe-base, an extruded tube is fit to a cast base support, and extends through a hole or socket in the casting. The fillet welds were placed between the tube and the shoe-base at the top-side and bottom-side of the casting. The filler metal was taken to be aluminum alloy 4043. What is interesting to note is the sequence used to manufacture the test specimens. Each extruded pole was inserted into the base, whether a casting or a plate. Fillet welds were placed between the tube and top of the casting or plate. Each sample was subsequently repositioned and the fillet weld between the tube and bottom of the base was placed. Each assembly was then placed in an aging oven to bring the extruded tube and base to a near T6
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Fig. 3. Failed light pole specimen with shoe-base detail. Fig. 1. Shoe-base detail and residual stress measurements.
condition. Upon completion of aging, the specimens were available for mechanical testing. Several samples were used to measure the near weldment residual stresses using the hole drilling method as shown in Fig. 1. Although discussed elsewhere in detail, near surface residual stresses in the tube adjacent to the top fillet weld were consistently in compression for both details and conditions [10]. In essence, shrinkage of the bottom weld attempts to pull the tube into the base thereby, placing the surface of the tube near the joint in compression. The presence of compressive residual stresses at the surface would be expected to alter the fatigue behavior and use of fracture mechanics approach to modeling. The fatigue specimens were 3048 mm long and secured by bolting the base supports to a heavy steel fixture. Each sample was cyclically loaded at the tip of the cantilever as shown in Fig. 2. Loads were applied at a small positive R ratio (minimum stress/maximum stress), and strains immediately adjacent to the weld detail were carefully monitored during the course of each test through the use and application of uniaxial strain gauges. The shoe-base specimens were tested at stress ranges from a low of 24 MPa to a maximum of 60 MPa. The through-plate samples were evaluated at stress ranges from 6.2 MPa to 31 MPa. The fatigue cracks formed on the tension side of the cantilever, along the weld toe between the casting and tube or the plate and tube, are shown in Fig. 3. Failure was defined as the development of a crack that penetrated the thickness of
Fig. 2. Light poles loaded at tip of cantilever.
the tube. Additional details of the testing program can be found elsewhere [11].
2. Fatigue life estimates using fracture mechanics Fatigue life estimates were prepared for both types of details using fracture mechanics techniques. Development of stress intensity factors for both the shoe-base and through-plate geometries were facilitated by use of the finite element method (FEM). The finite element method (FEM) was employed to determine the stress gradient local to the welded region. Local stresses from the different FEM analyses were combined with the measured residual stresses, published crack growth data for the 6XXX-series alloys, and knowledge of typical initial flaw sizes for welded aluminum construction as input to AFGROW [12–14]. The Air Force crack growth prediction software tool that enables the user to assess the fatigue life (Nf ) of structures is AFGROW. An extensive library within AFGROW provides models for a range of different crack geometries. Of equal importance is the intrinsic ability within AFGROW to input local stresses that may include residual stresses. A parametric study of both types of details was conducted as part of this research investigation in an attempt to identify some of the important geometric parameters that influence the stresses local to the base—tube welded connections. ANSYS was employed in all of the analyses conducted. Several models are shown with results for the as-tested geometries exemplified in Fig. 4 [15]. An extensive description of the parametric study is provided elsewhere [16]. However, one of the important parameters was found to be flexural stiffness of the support base. While the shoe-base is relatively stiff, the 25.4 mm through-plate is not. The models mimicked the test geometries and were loaded at the tip of the cantilever. Distortion of the base plate of the through-plate detail results in additional bending through the tube wall. The fatigue tests clearly showed that the elevated stresses near the weld toe joining the tube to the base plate resulted in significantly lower fatigue lives when compared to poles utilizing the shoe-base support. In order to investigate stiffness of the base and fatigue life, analyses were conducted for the as-tested shoe-base and through-plate test specimen geometry as well as a through-plate detail having a base thickness of 75.2 mm. The local stresses from the three different models were used for the prediction of fatigue life. For purposes of this study, previously developed methodologies were used to evaluate a correction factor for input to AFGROW. As described in an earlier study by Zettlemoyer [17], the stress
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Fig. 4. Finite element models—mesh and sample results.
intensity factor is numerically evaluated using the relationship: √ (1) K = Fe Fw Fs Fg Sr a In this relationship, K is the stress intensity factor range; Fe is the correction factor for crack shape; Fw is the correction factor for finite width; Fs is the correction factor for free surface; Fg is the stress gradient correction factor that makes use of the FEM Model results; Sr is the stress range, and a is the crack size. For direct input to AFGROW, the correction factor, i.e., the stress multiplication factor, was estimated using the relationship SMF = Fe Fw Fs Fg Sr
(2)
In this expression SMF stands for the Stress Multiplication Factor. A successful application of fatigue life prediction using fracture mechanics depends upon the conjoint and mutually interactive influences of (i) similitude, and (ii) the appropriate choice of crack growth data or development of the data under conditions consistent with laboratory tests. The crack growth rate data for the extruded aluminum alloy was represented by an expression developed by Paris and Erdogan [18]: da m = C(K) dN
(3)
In this expression, da/dN is the crack growth rate with a the crack size and N the number of cycles, K is the stress intensity factor, and C and m are constants determined from the measured experimental data that depend on the following: (i) material, (ii) specimen dimensions, (iii) test conditions, and (iv) nature of loading. In this study, fatigue life prediction made use of crack growth rate data at three different R ratios including 0.0, 0.5 and 0.7. For a typical welded construction, crack growth data at an elevated R ratio would be used since tensile residual stresses would tend to minimize crack closure [19]. The crack growth data obtained at an R ratio of 0.0 would be expected to contain some closure and tend to the fatigue test conditions for the light poles that were loaded at a very small positive load ratio. The estimates of fatigue life were initially prepared for the through-plate socket connection having a 25.4 mm thick base plate. For the through-thickness plate specimen, failure due to fatigue is defined and accepted as the number of cycles required to grow a crack through the thickness of the light pole tube. In Fig. 5 is shown the fatigue life estimates obtained from crack growth data obtained for a stress ratio R (minimum / maximum ) of 0.0, both with and without input of the measured residual stresses, when compared to the test results. There exists a significant scatter in the
Fig. 5. Fatigue life prediction with and without residual stress compared to the experimental results for the through-plate socket connection.
experimental data obtained in the laboratory. While inclusion of the residual stresses does increase the estimated fatigue life, it is clear that the predictions fail to correlate with the observed test results. The results of fatigue life estimates for the shoe-base connection developed using crack growth data taken at a stress ratio (R = minimum / maximum ) of 0.0 is shown in Fig. 6. The residual stresses were included in the analyses. In general, the predicted fatigue lives were larger when compared to the experimentally determined test data. This was particularly evident at the stress ranges () between 42 MPa and 65 MPa. However, fatigue life predictions improved as the stress range () was gradually reduced. Further, it is observed that there is good agreement between the predicted fatigue lives and the experimentally determined fatigue life for stress ranges between 24 MPa and 38 MPa. As previously discussed, careful measurements revealed the presence of compressive residual stresses on the surface of the luminaire tube immediately adjacent to the weld. The presence of compressive residual stress did complicate the prediction of fatigue life. In an attempt to account for the intrinsic influence of residual stress, the fatigue crack growth (da/dN) data obtained at stress ratios of R = 0.5 and R = 0.7 were utilized. The occurrence of crack closure at the higher values of the stress ratio, that is, R = 0.5 and R = 0.7, is minimal and crack growth through the microstructure of aluminum alloy 6063 is governed by the presence of intrinsic
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Fig. 6. Fatigue life predictions compared to experimental data for the shoe-base connection using fatigue crack growth data for R = 0.0 with residual stresses included.
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Fig. 8. Comparison of experimental data for the through-plate and shoe-base connections with fatigue life predictions for a 76.2 mm through-plate connection using fatigue crack growth data with R = 0.7 with residual stresses included.
Fig. 7. Fatigue life predictions for the through-plate socket connection compared to experimental data using fatigue crack growth data with R = 0.5 and R = 0.7 including residual stresses.
Fig. 9. Fatigue life predictions for through-plate socket connections using fatigue crack growth data with R = 0.5 and R = 0.7 including the effects of residual stresses.
microstructural features and their specific role and contribution in governing fatigue deformation. The crack growth data (da/dN) obtained at the higher values of the stress ratio (R) coupled with the measured residual stresses were used to simulate the conditions that are likely to exist in the fatigue test specimen. The predicted fatigue lives for the through-plate socket connection having a 25.4 mm thick plate is compared with the experimental test data and is shown in Fig. 7. In general, it is clear that the predicted fatigue lives tend to overestimate the test results obtained from the experiments. A comparison of the fatigue life predictions for the through-plate socket connection having a 76.2 mm thick base plate with the test results obtained from experiments on specimens with both through-plate and shoebase connections is shown in Fig. 8. The observed behavior of the through-plate socket connection having a 76.2 mm thick base is similar to that of the integrally stiffened shoe-base joint as the flexural stiffness of each base is similar, as are the stress fields adjacent to the weldments. The 76.2 mm plate thickness minimizes distortion and through thickness bending of the tube wall is relatively small. The fatigue life estimates for the shoe-base socket connection along with corresponding data obtained from actual tests conducted in the laboratory is shown in Fig. 9. The fatigue life prediction was noticeably improved when crack growth data was obtained at the higher stress ratios (R) coupled with measured residual stresses.
At a high load ratio (R > 0.6) the fatigue crack growth data minimizes the effects of closure and is often used in the fracture mechanics study of welded structures to account for the presence of tensile residual stresses. Use of high load ratio (R) data with measured compressive residual stresses for the tested cantilever specimen’s accounts for the observed fatigue behavior. A low load ratio [R] test data combined with the measured compressive residual stresses tends to over compensate or retard fatigue crack growth beyond what was observed experimentally. The predicted fatigue lives for the shoe-base connection with fatigue test data obtained from the experimental tests using the current AASHTO design stress (S) versus fatigue life (N) curves for both the Category ‘D’ and Category ‘E’ is compared in Fig. 10. The categories are results of full-scale fatigue tests undertaken to categorize each welded detail. Samples in the database used to develop the stress (S) versus fatigue life (N) curves were tested under constant amplitude load conditions. For fatigue lives in excess of 105 cycles, Category D provides a reasonable estimate of the lower bound fatigue strength. Further, the predicted fatigue lives are in reasonable agreement with the laboratory test results for the variation of stress (S) with fatigue life (Nf ) for Category D. The predicted fatigue lives for the through-plate socket connection for the 25.4 mm thick and 76.2 mm thick base plates along with the experimental test results are shown in Fig. 11. Also shown is
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Fig. 10. Design curves for Categories D and E details compared to shoe-base laboratory fatigue data as well as fatigue life prediction using fatigue crack growth data with R = 0.7.
the companion stress (S) versus fatigue life (Nf ) curves put forth by AASHTO for Category D, Category E and Category E’. Much of the experimental test data falls well below the lower bound defined by Category E’ ( endurance = 7 MPa). The predicted fatigue
Fig. 11. Experimental results for the through-plates socket connection compared to AASHTO design Categories D, E and E’ as well as life predictions.
lives tend to be in reasonable agreement with Category E’. However, the predicted fatigue lives for the 76.2 mm thick base plate are significantly improved and quite consistent with the Category ‘D’ details.
Fig. 12. Scanning electron micrographs of the aluminum alloy light pole sample that was cyclically deformed at a stress range of 37 MPa at 1 Hz resulting in 1.6 × 106 cycles to failure, showing: (a) overall morphology showing distinctly the region of early crack initiation and early crack growth. (b) High magnification of (a) showing fine striation-like features in the region of stable crack growth. (c) High magnification of the region of early crack initiation showing the overlap of striations to form a ripple-like pattern. (d) The region of slow and stable crack growth revealing shallow striations reminiscent of localized microplastic deformation.
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Fig. 13. Scanning electron micrographs of the aluminum alloy light pole sample that was cyclically deformed at a stress range of 37 MPa at 1 Hz resulting in 1.6 × 106 cycles to failure, showing: (a) a non-linear fatigue crack traversing the recrystallized grain boundary in the region of early crack growth. (b) Cracking along the grain boundary separating the regions of early crack growth and stable crack growth. (c) An array of fine microscopic cracks in the region of stable crack growth.
3. Failure-damage analysis: characterization of the fracture surface Careful and comprehensive scanning electron microscopy (SEM) observations of the fatigue fracture surface were made on the test samples taken from a shoe-base connection that was cyclically deformed at a stress range of 36.5 MPa and having a resultant fatigue life of 1.6 × 106 cycles. An examination of the fracture surfaces of the deformed and failed fatigue specimens in a JEOL SEM was done at: (a) Low magnifications to identify the precise size, shape and location of microscopic crack initiation and early crack growth, and the progress of stable crack growth through the alloy microstructure, culminating in unstable crack growth and failure due to an overload. (b) Higher magnifications in the regions of early crack growth, stable crack growth and unstable crack growth to identify the following: (i) Nature and severity of damage initiation. (ii) Nature of microscopic crack growth. Overall, the fatigue fracture surface revealed features that are typical of fatigue fracture. The fracture surface was essentially a flat topography at the chosen value of stress range. On a micro-
scopic scale, the nature and morphology of the intrinsic features on the fatigue fracture surfaces was found to depend on the level of maximum stress, at the chosen load ratio. Only representative fractographs of the fatigue fracture surfaces are shown in Figs. 12 and 13. The presence of a flat fracture surface area having a distinct region that identifies well-defined crack growth and sharp transfer to unstable crack growth culminating in catastrophic fracture is sufficient for identification of failure due to fatigue. Overall morphology of the fracture surface revealed the distinct regions of crack initiation and early microscopic crack growth (Fig. 12a). High magnification observations in the region of early crack growth revealed striation-like features (Fig. 12b) reminiscent of localized microplastic deformation. At higher allowable magnifications of the SEM the region of crack initiation and early crack growth revealed an overlap of the pockets of striations creating a ripple-like pattern on the fracture surface (Fig. 12c). The shallow nature of the striations was distinctly evident in the region of slow and stable crack growth (Fig. 12d). Also evident on the fracture surface of the 6063 alloy sample were isolated cracks, both macroscopic and fine microscopic, that were essentially non-linear in nature. Several of these cracks were found traversing through the recrystallized grain boundaries (Fig. 13a). There are a few situations that favor the occurrence of cracking along the recrystallized grain boundaries. For this aluminum alloy (i.e., 6063) of interest these include the following:
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(i) Precipitation and presence of brittle second-phase particles both at and along the grain boundary region. (ii) Hydrogen embrittlement and liquid metal embrittlement depending on the severity of the environment the alloy is exposed to. (iii) Environment assisted cracking (EAC). (iv) Intergranular corrosion. (v) Grain boundary cavitations and cracking at high temperatures.
6. Fracture surface examination was conducted for the failed shoebase connection in a scanning electron microscope. The fracture surface revealed fine and shallow striation-like features in the region of slow and stable growth and a combination of widely spaced striations and cracks in the region of unstable growth approaching overload.
The presence of fine second-phase particles along the grain alloy boundary regions coupled with local stress concentration effects and their synergistic interactions is responsible for the observed cracking that is few and far in-between. At low magnification the macroscopic cracks were evident in large numbers in the region separating early crack growth and stable crack growth (Fig. 13b). At the higher allowable magnifications the region of stable crack growth also revealed fine microscopic cracks (Fig. 13c). The microscopic cracks though few were well defined and distinct during the early stages of unstable crack growth.
The authors extend most sincere thanks and appreciation to the unknown reviewer for his many useful comments and corrections pertaining to the text and relevance of this technical manuscript.
4. Conclusions Based on a fairly exhaustive examination of the fatigue behavior of cantilevered luminaire structures the following are the key observations: 1. The fatigue test results correlate well with US Air Force AFGROW prediction of life provided the crack growth rate data used was obtained under conditions of minimal closure and at a high stress ratio (R) of 0.7. 2. At identical stress ranges, the through-plate connection having a base plate thickness of 25.4 mm exhibited lower predicted fatigue life when compared to the same detail having a 76.2 mm thick base plate. 3. The fatigue life of the through-plate socket connection having a 76.2 mm thick base plate was similar to the light pole having a shoe-base joint. 4. The fatigue strength of through-plate connection having a 25.4 mm thick base was below the stress-fatigue curve predicted by AASHTO Category E’. 5. For fatigue lives in excess of 105 cycles, the shoe-base connection behaved in a manner quite consistent with the stress-fatigue life curve put forth by AASHTO for Category D.
Acknowledgements
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