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Understanding the mechanical response of built-up welded beams made from commercially pure titanium and a titanium alloy
Materials Science & Engineering A 590 (2014) 390–400
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Understanding the mechanical response of built-up welded beams made from commercially pure titanium and a titanium alloy Anil K. Patnaik a,n, Narendra Poondla a,b, Craig C. Menzemer a, T.S. Srivatsan b a b
Department of Civil Engineering, The University of Akron, Akron, OH 44325, USA Department of Mechanical Engineering, The University of Akron, Akron, OH 44325, USA
art ic l e i nf o
a b s t r a c t
Article history: Received 4 September 2013 Accepted 15 October 2013 Available online 24 October 2013
During the last two decades, titanium has gradually grown in stature, strength and significance to take on the recognition of being a modern and high performance metal that is noticeably stronger and concurrently lighter than the most widely chosen and used steels in a spectrum of industrial applications. Technological innovations have necessitated reduction of part weight, cost and lead time, including concurrent enhancement of performance of structural parts and components made using titanium and its alloys. This has provided the impetus to develop economically viable structural design methodologies and specifications, while at the same time bringing forth innovative and economically affordable manufacturing and fabricating techniques with the primary purpose of both producing and promoting the use of cost-effective titanium structures. The experimental results of a recent study on built-up welded beams are presented in this paper with the primary objective of enabling design, facilitating fabrication, and implementation of large structural members for potential applications in the structural and defense-industry. & 2013 Elsevier B.V. All rights reserved.
Keywords: Commercially pure titanium Titanium alloys Built-up welded beams Static response Structural design
1. Introduction Commendable advances in the domains spanning extraction, processing, fabrication coupled with the methods both related and used for joining have made possible technically viable, structurally efficient and economically affordable applications of both pure titanium metal and its alloys for selection and use in large structural applications in the sectors spanning aerospace, construction and defense. There continues to exist a growing need and incentive to reduce part weight, lower the cost and decrease the time required for manufacturing, while concurrently enabling enhanced performance of structural parts made from both the older as well as the emerging generation of titanium alloys. An overview of alloys suitable for use in a spectrum of performance-critical applications in the domain of both aerospace and non-aerospace related applications, to include corrosion protection, armor, orthopedic, and sports equipment is presented and discussed in some recent papers by the authors [1–4]. Over the years, titanium has gradually grown both in strength and stature to gain for itself the significance of being recognized as a modern and high performance metal. Many titanium alloys are noticeably stronger and concurrently lighter than the most widely chosen and used steels for use in a spectrum of relevant and
n
Corresponding author. E-mail address: [email protected] (A.K. Patnaik).
0921-5093/$ - see front matter & 2013 Elsevier B.V. All rights reserved. http://dx.doi.org/10.1016/j.msea.2013.10.041
appropriate industrial applications. Importantly, two of the most attractive properties of pure titanium are its resistance to corrosion and superior ballistic properties [1,4,5]. In recent years, alloys of titanium have been frequently chosen for use as one of the main candidates for both aircraft structural parts and components of aircraft engines on account of its (a) primarily high strength-toweight ratio [s/ρ], and (b) high stiffness-to-weight ratio [E/ρ]. Despite its superior mechanical properties, the selection and use of pure titanium metal and its alloy counterparts have often been limited to high performance structures like: (i) aircraft wing, (ii) skin stiffened panels, (iii) armor structure, (iv) structural parts in navy vessels, and several other structures and components that find use in the defense industry that are essentially performancecritical, and not necessarily cost-limited [2,5]. It is generally believed that the high cost of both pure titanium metal and its alloy counterparts makes it difficult to implement its exhaustive use for those applications other than performancecritical structures. A one-to-one substitution of steel with an alloy of titanium is not easily possible based entirely on mechanical properties. This is because both pure titanium and its alloy counterparts will always work out to be more expensive. The cost of titanium is about 50 to 75 times higher than the cost of steel on a per-pound basis. Further, selection and immediate use of titanium metal is, as of now, not technology ready for its direct implementation in large civil engineering structures. At the same time, there limited design guidelines are available for the structural design of components making it difficult to implement
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in structures other than those specifically tested to verify performance. A primary reason for the high cost of the titanium metal is that it is not easily found in high concentration, or in the pure state, in nature. This has necessitated the need for multistage processing to bring it to a point where it can be chosen and put to use in a spectrum of structural applications. In order to gradually enhance the selection and utilization of both pure titanium metal and its alloy counterparts for both defense-related and non-defense-related applications, a study aimed at evaluating, understanding and rationalizing the strength, endurance and performance of structures made from commercially pure titanium and a titanium alloy (Ti–6Al–4V) was initiated at the University of Akron. This paper presents the results of the study on the structural behavior of built-up welded beams made from commercially pure titanium (Grade 2) and the titanium alloy Ti–6Al–4V under static loading. Other parts of the study such as fatigue behavior of built-up beams, microstructural studies, tensile performance, etc. have been reported elsewhere [6–12].
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a) Evaluate the static tensile strength and static flexural behavior of commercially pure titanium (Grade 2) built-up beams to produce high strength, light weight, corrosion resistant and cost-effective structural beams for the purpose of possible applications in the defense and civilian sector. b) To evaluate Ti–6Al–4V alloy built-up beams fabricated and tested under conditions of static tensile and flexural loading.
3. Materials The materials selected for this study were (i) commercially pure titanium (Grade 2), and (ii) alloy Ti–6Al–4V. The two materials were systematically evaluated from the standpoint characterization of initial microstructure, tensile properties, and response to static bend loading. 3.1. Commercial pure (CP) titanium (Gr. 2)
2. Background The concept of fabricating built-up steel beams (also known as plate girders) from thin-walled plate elements is not new for highway bridges [13]. Since the early 1950s, a range of riveted, bolted, and welded plate girders have been preferentially chosen and used for appropriate applications. For example, three plates can be easily put together by welding to form the flanges and web of an I-shaped plate girder. For structures made from pure titanium metal and its alloy counterpart, it is common practice to machine the required part from large castings or billets. Overall, machining of titanium metal is a challenge by itself. Further, in the alloy form, titanium is harder to machine. A sizeable amount of material that is removed from either a thick plate or billet is often wasted, and this directly contributes to an observable increase in the cost of the end product. Also, the bars having a thicker cross-section are difficult and expensive to make, and are generally not readily available with material suppliers. Machining of titanium also poses fire hazard and therefore, needs special precautions for safety. With notable strides made in the specific domain of welding technology for both pure titanium metal and its alloy counterparts, it is now possible to weld titanium plates to make built-up structural components. For example, in recent years, a team of researchers and engineers at the US Army [Division at Picatinny Arsenal (New Jersey, USA)] have developed a Pulsed Gas Metal Arc Welding process (referred to henceforth as GMAW-P), which essentially involved a modification of the existing procedures so as to be applicable to both pure titanium metal and its alloys [14–17]. Linear weld speeds can be up to ten times faster than the corresponding speed for both tungsten inert gas (TIG) and gas tungsten arc welding (GTAW) processes. Besides, the number of passes when using the technique of GMAW-P can be reduced by a factor of three. The new process developed by the engineers and researchers at the US Army uses 100% helium shielding gas that not only facilitates good penetration but also ensures a double pulse template. The welding of titanium is standardized in a recently released welding specification put forth by the American Welding Society (AWS) designated as AWS D1.9/D1.9M [18]. It has therefore become feasible in the recent past to make welded built-up beams using titanium plates. The objective of this paper is to describe the results of a recent investigation on structural performance of built-up welded beams made from: (i) commercially pure (Grade 2) titanium, and (ii) the preferred and widely chosen and used alloy, Ti–6Al–4V, under static loading. The overall objectives were to:
The CP titanium (Gr. 2) plates were provided by TICO Titanium (Wixom, MI) in the thicknesses of 3/8 in. (9.5 mm) and 1/8 in. (3.2 mm). The edges of the plates and sheets were sheared. The nominal chemical composition of the as-provided material is summarized in Table 1. 3.2. Ti–6Al–4V alloy Plates made from this alloy were provided by Allegheny Technologies ATI Wah Chang (Albany, OR) in the thickness of 1/4 in. (6.35 mm). The nominal chemical composition of this alloy is summarized in Table 2.
4. Experimental procedures 4.1. Sample preparation for mechanical testing Tensile tests were conducted on specimens that were precision machined from both the longitudinal (L) and transverse (T) orientation of the as-provided annealed plate stock. The longitudinal (L) specimens were machined with the major stress axis parallel to the rolling direction of the Ti–6Al–4V alloy and CP (Grade 2) titanium plates, while the transverse (T) specimens were machined with the major stress axis perpendicular to the rolling direction of the two plates, i.e., Ti–6Al–4V and CP titanium (Grade 2). The round test specimens conformed to specifications outlined in the standard ASTM E-8 and had threaded ends. At the gage section, the test specimens measured 1/8 in. (3.2 mm) in diameter and 1/2 in. (12.5 mm) in length. To minimize the effects and/or contributions from both surface irregularities and finish, final surface preparation was achieved by mechanically polishing the gage section of the machined test specimens with progressively finer grades of silicon carbide impregnated emery paper (320 grit, 400 grit and 600 grit) to remove any and all circumferential scratches and surface machine marks. 4.2. Mechanical (tension) testing Uniaxial tensile tests were performed on a fully-automated, closed-loop servo-hydraulic mechanical test machine [INSTRON8500 Plus] using a 100 kN load cell. The tests were conducted at room temperature (300 K) and in the laboratory air (Relative Humidity of 55 pct.) environment. The test specimens were deformed at a constant strain rate of 0.0001/s. An axial 12.5 mm gage length clip-on type extensometer was attached to the test
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Table 1 Specifications, composition and average properties of the as provided CP Ti (Grade 2). Description
10 mm thick plate
3 mm thick sheet
Specification
ASTM B-265-06/ASME SB-265-A06 GR. 2EN 10024:2004 type 3.1
ASTM B-265-07/ASME SB-265-7/ASTM F-67-06GR. 2EN 10204:2004 type 3.1
0.13 0.16 0.002 0.012 Residual element (each) less than 0.10 Residual element (total) less than 0.40 Final product hydrogen 0.0032 Titanium remainder 507 MPa 360 MPa 33 pct 59 pct Yes
0.08 0.1 0.0105 0.014 Residual element (each) less than 0.10 Residual element (total) less than 0.4 Final product hydrogen 11/10 PPM Titanium remainder 484 MPa 341 MPa 25.75 pct – 1370 F, HGA 3 to 11 min Pass
Chemical composition Fe iron Oxygen N Nitrogen C Carbon
Average tensile strength, psi Yield 0.2% offset Elongation % RA % Anneal condition Guided bend test
Table 2 Nominal chemical composition of Ti–6Al–4V alloy (in wt%). Material
Ti
Al
N
V
C
Fe
H
O
Ti–6Al–4V
90.0
6.0
0.05
4.0
0.1
0.4
0.02
0.20
specimen at the gage section. The stress and strain measurements, parallel to the line of loading, and the resultant mechanical properties, such as, (i) stiffness, (ii) strength (yield strength and ultimate tensile strength), (iii) failure stress, and (iv) ductility (strain-to-failure) was provided as a computer output by the control unit of the test machine. 4.3. Static bend tests of titanium alloy beams The test specimens were fabricated using GMAW-P welding process. The dimensional details of the beams are given in Fig. 1a and b. Complete details of the fabrication can be found elsewhere [17,19]. Two of the beams (designated as B1 and B4) were tested with the primary purpose of studying the flexural performance when bent in single curvature under static loading. Beam B1 was made from the Ti–6Al–4V alloy while beam B4 was made from commercially pure titanium (Gr. 2). The span of the two test beams measured 610 mm. Another beam (denoted as B5) was also made from commercially pure titanium (Gr. 2) and deformed under conditions of fatigue loading as described elsewhere [6]. The fatigue test was stopped following one million cycles as a consequence of load at the one-millionth cycle. It was then tested under static loading after the beam was subjected to one million cycles of fatigue loading. Details of the static load test on beam B5 are also presented and discussed in this paper. 4.3.1. The test set-up The test set-up for the static bend tests on the beams is shown in Fig. 2a and b. The test beams were loaded on an INSTRON 5300 machine with a maximum load capacity of 1000 kN (225 kips). The beams were loaded, using a four-point loading arrangement, over a simple span of 610 mm (24 in.). The total length of each specimen was 685 mm. The test beams were supported at either end on rollers that allowed the beam to essentially behave as “simply” supported. Also, the supports at either end allowed for free
rotation of the beam while concurrently permitting free sliding to occur. Bearing plates, measuring 12.5 mm in thickness and 50 mm in width, were provided over the roller supports. The compression flange of each test beam was restrained from any lateral displacement at the supports. The restraint was provided by a fixture that was specifically designed and fabricated for the static tests. The two fixtures that were provided at the two ends of the test beam are as shown in Fig. 2a. This figure shows a front view of the fixtures at the two ends. A side view of the fixture at one end (support) is shown in Fig. 2b. Lateral displacement of the compression flange was prevented by two threaded rods that were provided in each fixture. Complete details of the test fixture are presented and discussed elsewhere [19].
4.3.2. Instrumentation The development of strains during the entire duration of static loading was recorded with the help of strain gages positioned at seven different locations. A Digital Image Correlation (DIC) system (ARAMIS) was used to capture the strain fields of each test beam during the entire duration of each static test. The deflections occurring at the mid-span were recorded with the aid of a digital dial gage. The failure mode was captured in the form of digital images in the DIC system. During the entire duration of the test, the data was programmed to be collected by: (i) the control unit of the INSTRON test machine, and (ii) an independent stand-alone Data Acquisition System (DAQ). The typical strain gage layout for each test beam (i.e., B1 and B4) is as shown in Fig. 3. The strain gages were attached in the region of the midspan on both the top flange and the bottom flange, one on each side of the web. Additionally, two strain gages were attached in a 451 pattern on the web, at a quarter span locations, on each side of the points of loading.
4.3.3. Test procedure A preload of 2.25 kN (0.5 kips) was initially applied to ensure firm contact between the beam and its supports. (a) A loading rate of 4.45 kN (1000 lbs.) per minute was used for loading beams B4 and B5 both of which were made from commercially pure titanium (Gr. 2). In case of beam B5, static
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Fig. 1. (a) Details of commercially pure titanium beam B4 (Beam B5 is similar) and (b) details of commercially Ti–6Al–4V beam B1 (all dimensions are in mm).
load test was conducted to find the residual strength after fatigue loading. (b) On account of its higher load carrying capability, the loading rate was tripled for the beam made from alloy Ti–6Al–4V. The images were captured by the DIC system at the same rate as the deflection readings and frequency of the data acquisition system (DAQ).
5. Results and discussion 5.1. Initial microstructure The microstructure of the as-provided material is an important factor that determines or governs its response to an external load or mechanical stimulus. In essence, the microstructure governs the tensile properties, fracture toughness, compressive response, fatigue resistance of the material chosen for the beams, and the overall kinetics of fracture at both the macroscopic and microscopic level. Optical microstructure of the as-received commercially pure titanium (Gr 2) plate is shown in Fig. 4 at two magnifications. At low magnification the overall microstructure revealed the primary alpha (α) grains to be intermingled with the
small yet noticeable pockets of the beta (β) grains. At a higher magnification, very fine alpha (α) phase lamellae dispersed well within the beta (β) grains was evident (Fig. 4b). The optical microstructure of the Ti–6Al–4V alloy is shown for two different magnifications in Fig. 5. The two micrographs in this figure reveal the as-received, undeformed microstructure of the titanium alloy specifically the volume fraction, morphology, size and distribution of the intrinsic micro-constituents through the microstructure. Over the range of magnifications spanning very low to high revealed a duplex microstructure consisting of near equiaxed alpha (α) and transformed beta (β) phases. The primary nearequiaxed shaped alpha (α) grains (light in color) were well distributed in a lamellar matrix with transformed beta (dark in color). The presence of trace amounts of aluminum and oxygen in this Ti–6Al–4V alloy contributes to strengthening and stabilizing the alpha (α) phase, which is beneficial for enhancing hardenability, increasing strength and improving the response kinetics of the alloy to heat treatment. 5.2. Tensile response The room temperature tensile properties of both CP (Grade 2) and the Ti–6Al–4V alloy are summarized in Table 3. The results reported are the mean values based on duplicate tests. The elastic
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Fig. 2. (a) The test set-up and (b) a side view showing lateral supports.
modulus (E), yield strength (sYS), ultimate tensile strength (sUTS), elongation-to-failure (εf) and strength at failure (fracture) (sfr) were recorded in a data acquisition system. The yield strength was determined by identifying the stress at a point on the engineering stress versus engineering strain curve where a straight line drawn parallel to the elastic portion of the stress versus strain curve at 0.2% offset intersects the curve. The ductility is reported as elongation-to-failure over a gage length of 12.7 mm. This elongation was measured using a clip-on extensometer that was attached to the gage section of the test specimen.
5.2.1. Ti–6Al–4V alloy A representative engineering stress versus engineering strain curve for the two orientations of the as-received plate, i.e., longitudinal (L) and transverse (T), is shown in Fig. 6. The elastic modulus of the alloy is 126 GPa (18,300 ksi) in the longitudinal (L) orientation and 137 GPa (19,900 ksi) in the transverse (T) orientation. The yield strength (sYS), of the alloy in the annealed condition is 948 MPa (137 ksi) in the longitudinal (L) orientation and 1047 MPa (152 ksi) in the transverse (T) orientation. The ultimate tensile strength (sUTS), of the alloy is 1060 MPa (154 ksi) in the
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longitudinal (L) orientation and 1181 MPa (171 ksi) in the transverse (T) orientation. Both the yield strength and tensile strength are higher in the transverse orientation than in the longitudinal orientation. The ultimate tensile strength of the Ti–6Al–4V alloy is only marginally higher than the yield strength, i.e., about 10%, indicating a low work hardening rate beyond yield. The elongation-to-failure (εf) of the alloy is 7.8% in the longitudinal (L) orientation and 11.5% in the transverse (T) orientation. The reduction-in-area of the alloy was 24% in the longitudinal (L) orientation and 22% in the transverse (T) orientation and agrees reasonably well with the annealed condition of the alloy microstructure. The yield strength and tensile strength values of the alloy conform well to the values obtained and reported by the manufacturer (ATI Wah Chang) and recorded in ASM Metals Handbook and by others [20–22]. Variation of stress with plastic strain reveals the Ti–6Al–4V to be marginally harder in the transverse (T) orientation when compared to the longitudinal (L)
Fig. 3. Schematic showing layout of the strain gages.
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orientation. The strain hardening exponent (n) of this alloy in the two orientations is 0.014 for the transverse (T) orientation and 0.012 for the longitudinal (L) orientation.
5.2.2. Commercially pure titanium (Grade 2) The elastic modulus of this material is 115 GPa (16,700 ksi) in the longitudinal (L) orientation and 108 GPa (15,700 ksi) in the transverse (T) orientation. For the CP (Grade 2) titanium, the yield strength is 432 MPa (63 ksi) in the longitudinal (L) orientation and 15 pct. higher than in the transverse (T) orientation (382 MPa or 55 ksi). The ultimate tensile strength is 561 MPa (81 ksi) in the longitudinal (L) orientation and 465 MPa (67 ksi) in the transverse (T) orientation, a noticeable difference of 20 pct. The ultimate tensile strength is higher than the yield strength indicating a noticeable work hardening beyond yield. The elongation of CP (Grade 2) was 14.7% in the longitudinal (L) orientation and 18.9% in the transverse (T) orientation. The reduction-in-area was 43.4% in the longitudinal (L) orientation and 47% in the transverse (T) orientation. The engineering stress versus engineering strain curves for the material in both the longitudinal (L) and transverse (T) orientations are shown in Fig. 7. Variation of stress with plastic strain reveals the CP (Grade 2) to be harder in the longitudinal (L) orientation when compared to the transverse (T) orientation. The strain hardening exponent (n) of this material in the two orientations is 0.069 for the longitudinal (L) orientation and 0.054 for the transverse (T) orientation. In the longitudinal orientation of the as-provided annealed plates, the Ti–6Al–4V alloy has an elastic modulus of 125 GPa (18,100 ksi), while the commercially pure (CP: Grade 2) titanium has an average elastic modulus of 115 GPa (16,700 ksi).
Fig. 4. Optical micrographs showing the key micro-constituents in the commercially pure titanium (Gr 2) plate at two different magnifications.
Fig. 5. Optical micrographs showing the key micro-constituents in the Ti–6Al–4V alloy plate at two different magnifications.
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Table 3 Room temperature tensile properties of Ti–6Al–4V and commercially pure (Grade 2) [Results are mean values based on duplicate tests.]. Material
Ti–6Al–4V Commercially pure (Grade 2)
Orientation
Longitudinal Transverse Longitudinal Transverse
Elastic modulus
Yield strength
UTS
msi
Gpa
Ksi
MPa
Ksi
MPa
18 20 16 15
126 137 115 108
137 152 63 55
948 1047 432 382
154 171 81 67
1060 1181 561 465
Elongation GL¼ 0.5″(%)
Reduction in area (%)
Tensile ductility In (Aa/Af) (%)
7.8 11.5 14.7 18.9
23.8 21.7 43.4 47.0
27.0 25.0 57.0 63.0
5.3.1. Beam B4 – commercially pure titanium (Gr. 2) The failure mode of beam B4 is shown in Fig. 8a. The failure mode captured by the Digital Image Correlation (DIC) system (ARAMIS) is shown in Fig. 8b. Failure of the beam was initiated by shear buckling of the web, followed by excessive deflection and yielding of the flange. A careful physical inspection of the welds between the web and the flange and the connecting stiffeners revealed that the joints were intact with no evidence of failure. Also, there was no evidence of visible cracking. The beam maintained its structural integrity during the entire range of loading.
Fig. 6. Influence of test specimen orientation on engineering stress versus engineering strain curve of Ti–6Al–4V alloy.
Fig. 7. Influence of test specimen orientation on engineering stress versus engineering strain curve for commercially pure (Grade 2) titanium.
5.3. The static bend tests The two test beams B1 and B4 failed at a static load greater than the predicted value obtained using the design methodologies developed in an earlier study [19]. The predicted methods were developed based on the 14th edition of the prevalent AISC steel design specifications [23]. Beam B4 made from commercially pure titanium (Grade 2) failed at a load of 187 kN (42 kips). The residual static strength (i.e., the strength determined from static bend test following one million cycles of fatigue loading) of beam B5 made from commercially pure titanium (Gr 2) at failure was determined to be 205 kN (46 kips). The failure load predicted for this test beam using the proposed design methodology was 160 kN (36 kips) [19]. Failure for both beams occurred by a combination of shear buckling of the web and excessive deflection of the beam with yielding of the flanges.
5.3.2. Beam B5 (following fatigue loading) – commercially pure titanium (Gr. 2) Beam B5 that was subject to static loading subsequent to cyclic loading for one million cycles, failed in a manner quite similar to the failure mode experienced by beam B4. Failure of this beam was initiated by shear buckling of the web, followed by excessive deflection and yielding of the flange. Deflections were recorded over the entire range of static loading. Failure of beam B5 by web shear buckling is shown in Fig. 9. 5.3.3. Beam B1 – Ti–6Al–4V alloy Beam B1, which is made from Ti–6Al–4V alloy failed at a load of 583 kN (131 kips). The predicted failure load for this beam was 494 kN (111 kips). Failure of the beam occurred by a combination of excessive deflection (Fig. 10) with buckling of the compression flange between the two loading points (Fig. 11). There was visible evidence of complete plasticization of the length of the beam between the two loading points. Both the straightening and stretching of the bottom flange, as seen in Fig. 10, is evidence of tensile yielding of the bottom flange, while the top flange in compression completely yielded and buckled to a multiple wave form (Fig. 11). The proposed design method described in detail elsewhere by the authors [19] predicted the failure mode to be plasticization of both the compression flange and the tension flange. Therefore, the observed failure mode was as expected. There was no visible evidence of cracking of the welds in the test beam. Further, the welds were intact over the entire range of static loading. However, there occurred permanent deformation of the beam due to excessive stressing of both the compression flange and the tension flange. 5.4. The load versus deflection curves The load versus deflection curves were obtained from tests on each test beam. One was obtained from the digital dial gage while the other was obtained from readings of the test machine [INSTRON]. The load versus deflection curves for beam B1 (Ti– 6Al–4V beam) are as shown in Fig. 12. The shape of the load versus deflection curves was quite similar to the stress versus strain curve obtained for the material using small round tensile specimens that were prepared in conformance with ASTM Standard E-8. The initial part of the load-deflection curve is linear, up to about
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Fig. 8. (a) Failure mode of the commercially pure titanium (Gr. 2) beam (B4) by web shear buckling. (b) Failure mode of commercially pure titanium (Gr. 2) beam (B4) by web shear buckling as captured by the DIC system (ARAMIS). The strain contours are also shown.
Fig. 10. Failure of beam B1 (Ti–6Al–4V) as captured by DIC system (ARAMIS).
Fig. 9. Failure of beam B5 by web shear buckling.
400 kN (90 kip) load, following which the curve reveals the initiation of non-linearity indicating the onset of yielding. The load versus deflection curve demonstrates the ductile nature of the built-up beam under the conditions of static loading. The overall shape of the curve can be classified as being quite
similar to that of a classic ductile beam that reveals adequate reserve strength coupled with deflection capability to provide the required warning prior to catastrophic failure. Fig. 12 also shows the deflection curve predicted using the classical elastic beam bending theory. The elastic deflection method was found to marginally under-estimate the overall deflection of the beam. This figure reveals the existence of a match between the predicted deflection and the deflections measured during the test up to the onset of the elastic limit.
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Fig. 11. Compression flange (top flange) local buckling of beam B1 (Ti–6Al–4V).
Fig. 14. The load versus deflection curve of beam B5 (commercially pure titanium (Gr. 2)) following the fatigue test.
Fig. 12. The load versus deflection curve obtained for beam B1 (Ti–6Al–4V).
Fig. 15. Load versus strain curves for beam B4 (commercially pure titanium (Gr. 2)). Outside strains on the top flange.
Fig. 13. The load versus deflection curve of beam B4 made from commercially pure titanium (Gr. 2).
elastic shear buckling limit into the post-buckling region. The deflection curve obtained was quite similar to the behavior shown by a classic welded steel beam [13]. Both Figs. 13 and 14 also include the deflection curves predicted using the classic elastic beam bending theory. The method of elastic deflection was found to under-estimate the actual deflection experienced by the titanium beams. During mechanical testing, the beams revealed noticeably greater deflection than the value predicted using classic elastic beam bending theory. At a level of the service load, that was taken to be 50% of the failure load, the actual deflections experienced by the beam were found to be about 50 pct. to 75 pct. greater than the predicted values. This behavior is more pronounced for beam B5. However, the softening may have occurred due to the fatigue loading the beam was subjected to prior to the static test.
5.5. The load versus strain relation The load versus deflection curves of beams B4 and B5 (B5 tested following run out during the fatigue test) are shown in Figs. 13 and 14, respectively. These load versus deflection curves demonstrate evidence of distinct softening of the beam on reaching: (a) a load of 111 kN (25 kips) for beam B4 (Fig. 13), and (b) a load of 165 kN (37 kips) for beam B5 (Fig. 14). The observed softening is rationalized from the viewpoint of the initiation and occurrence of inelastic shear buckling. Softening of the beam occurred when the web reached its limit for the elastic shear buckling load, while the beam carried a load well beyond the
The load versus strain data collected from the data acquisition system (DAQ) was plotted and the relationship is shown in Fig. 15 for beam B4 (commercially pure titanium (Gr. 2) beam). Also, Fig. 16 shows the strain curves for the bottom flange at the midspan location. Both the theoretical strains and the strains recorded during the static test of the beam agree well for this beam (B4) for low values of flange strain up until the elastic limit. However, at a load level corresponding to the initiation of elastic web shear buckling (approximately 111 kN–133 kN (25–30 kips)) the curves
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Fig. 16. Load versus strain curves for beam B4 made from commercially pure titanium (Gr. 2): outside strains on the bottom flange.
Fig. 17. The load versus strain curves for beam B1 (Ti–6Al–4V): outside strains on the top flange.
obtained from the two strain gages during the static test, start to deviate from the theoretical strain. In Fig. 17 is shown the strain curves on the top flange taken at the mid-span location. While Fig. 18 reveals the corresponding curves for one web of the beam. Both these figures also show the theoretical strain curves that were obtained from theoretical analysis. For the case of the flange, both the theoretical strains and the strains recorded during the static test of the beam accord reasonably well revealing that the strains obtained from the test results match reasonably well for the two strain gages used.
399
Fig. 18. Load versus strain curves for beam B1 (Ti–6Al–4V): web strains (451).
The load versus deflection curves clearly demonstrate the predicted deflection curve for the beam made from Ti–6Al–4V (i.e., B1) to accord well with the actual deflection curve obtained from the test. Also, both beam B4 and beam B5 demonstrate a distinct softening of the beam upon reaching a load that corresponds to the onset of elastic shear buckling. The elastic deflections determined for the two beams underestimates the deflection when compared one-on-one with the values obtained from the static bend tests. The load versus strain response determined from tests conducted on both beam B1 (Ti–6Al–4V) and B4 (CP Ti – Gr. 2) match reasonably well with those determined by elastic beam bending theory for the initial part of the loading regime. The observed difference between the theoretical strain values and the experimental strain values was noticeable once the load exceeded the yield limit of the material. The strains obtained from the experimental tests and measured on the web panels deviated appreciably from the theoretical values. Both the commercially pure titanium (Grade 2) beam and the Ti–6Al–4V alloy test beam failed in modes that were predicted using the methodologies developed in an earlier study by the authors [19]. Further, the test beams demonstrated significant reserve strength following initial yielding. The experimental deflection curves along with the predicted deflection curves are presented and briefly discussed. These curves along with a comparison of load versus strain relationship obtained from the experimental tests with those predicted using theoretical methods demonstrate a reasonably close match between the theoretical predictions and the experimental test results up until the elastic limit of the material.
5.6. Analysis of results of the static bend tests on the welded beams
6. Conclusions
A summary of the failure loads that were determined from the experimental tests on the beams along with the corresponding predicted load is provided in Table 4. The failure load obtained for each of the three beams is greater than the predicted value of the failure load. Also, the actual strength obtained from the experimental tests was greater than the predicted value of strength by 16–28%. This suggests the existence of significant reserve strength in the test beams when they are designed in conformance with the methods used in this research study. Also, both the predicted failure mode and actual failure mode accord reasonably well. However, the occurrence of noticeably large deflections at failure does limit the usable strength to a value that is close to the predicted theoretical load in order to limit deflection.
With the latest GMAW-P welding technology developed at US Army Picatinny Arsenal, it is feasible to fabricate large welded built-up titanium beams by welding plates together to achieve structural performance that is quite comparable with an identical monolithic beam that is produced by machining from either thick plates or billets. The study demonstrated that welded built-up beams made from both sheets and plates of commercially pure titanium (Gr. 2) and Ti–6Al–4V can be fabricated with relative ease. The welds produced by the new GMAW-P method are sound and revealed an absence of visible cracks. (a) The results obtained in this study reveal the absence of a deleterious influence of welding on structural performance of
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Table 4 A summary of failure load and predicted strength of the beams studied. Beam
Material
B1
Ti–6Al–4V
B4
CP Ti (Gr. 2)
B5
CP Ti (Gr. 2)
(b)
(c)
(d)
(e)
Predicted failure load Predicted
Actual
494 kN (111 kips) 160 kN (36 kips) 160 kN (36 kips)
583 kN (131 kips) 187 kN (42 kips) 205 kN (46 kips)
built-up welded beams of commercially pure titanium (Grade 2) and Ti–6Al–4V titanium alloy. The welded beam concept used in this study worked well for the titanium beams that were tested. It is anticipated that the concept of welded built-up beams can be applied to structural members made from both commercially pure titanium (Grade 2) and the alloy (Ti–6Al–4V). This is a cost saving alternative to fabricating large structural elements and members by machining the parts from either a thick plate or billet. Currently, prevalent AISC steel design specifications (14th edition, 2011) were modified to suit the material properties of commercially pure titanium and its alloy counterparts. With suitable modifications, these specifications can be used for the design of welded built-up beams made from both commercially pure titanium and its alloys. This approach was found to be successful. However, additional research work is needed to refine the design approach. The failure mode, the failure load, deflections, and the strains induced in the welded built-up titanium beams are predictable to a reasonable and acceptable level of accuracy. The test beams demonstrated significant reserve strength following the onset of yielding. The experimental results versus predicted response demonstrate a reasonably close match between the theoretical predictions and the experimental test results up until the elastic limit of the material. The study reiterated that elastic stress analysis based on classical beam bending theory is adequate to predict the stresses, strains and deflections that are developed under the influence of a static load up until the initiation of yield in the material. The ultimate behavior is also predictable to a reasonable degree of accuracy in a conservative manner.
[4] [5] [6] [7]
[8] [9] [10] [11] [12]
[13] [14]
[15]
[16]
[17] [18]
References [1] A.K. Patnaik, C. Menzemer, T.S. Srivatsan, On the use of titanium alloys for aerospace and non-aerospace applications, in: N. Bhatnagar, T.S. Srivatsan (Eds.), Proceedings of the 17th PAFM XVII, December 2008, pp. 3–22. [2] A. Patnaik, T.S. Srivatsan, C.C., Menzemer, Structural static performance of welded built-up beams made from titanium and a titanium alloy, Proceedings of 2011 DoD Corrosion Conference, NACE International, Palm Springs, CA, July– August 2011, 10 pp. [3] A.K. Patnaik, N. Poondla, U. Bathini, T.S. Srivatsan, An overview of large structures fabricated from titanium and titanium alloys, in: M. Niimoni,
[19]
[20]
[21] [22] [23]
Difference
Failure mode both predicted and actual
þ18.0%
Bending failure leading to excessive deflection
þ16.7
Shear failure by web buckling
þ27.8%
Shear failure by web buckling
M. Morinaga, M. Nakai, N. Bhatnagar, T.S. Srivatsan (Eds.), Proceedings of the Processing and Fabrication of Advanced Materials—XVIII, vol. 2, Japan, December 2009, pp. 831–48. A. Patnaik, X. Shan, M. Adams, T.S. Srivatsan, C.C. Menzemer, J. Payer, Adv. Steel Constr. 10 (2) (2014) 8. V.N. Moiseyev, Metal Sci. Heat Treat. 46 (3-4) (2004) 115–120. A.K. Patnaik, N. Poondla, U. Bathini, T.S. Srivatsan, J. Mater. Eng. Performance 20 (7) (2011) 1247–1255. A.K. Patnaik, N. Poondla, U. Bathini, T.S. Srivatsan, Investigating and understanding the fatigue response and failure of built-up welded beams of a titanium alloy, Fatigue of Materials: Advances and Emergence in Understanding Materials Science and Technology (MS&T 10), Houston, TX, October 2010, pp. 195–216. U. Bathini, T.S. Srivatsan, A. Patnaik, C.C. Menzemer, J. Aerosp. Eng., ASCE 24 (4) (2011) 415–424. T.S. Srivatsan, U. Bathini, A. Patnaik, T. Quick, Mater. Sci. Eng. – A 527 (2010) 6649–6659. U. Bathini, T.S. Srivatsan, A. Patnaik, T. Quick, J. Mater. Eng. Performance 19 (8) (2010) 1172–1182. N. Poondla, T.S. Srivatsan, A. Patnaik, M. Petraroli, Journal of Alloys and Compounds 486 (2009) 162–167. U. Bathini, T.S. Srivatsan, A. Patnaik, C.C. Menzemer, T. Quick, The cyclic fatigue and final fracture behavior of a titanium alloy taken from weldments: influence of load ratio and orientation, fatigue of materials, Advances and Emergence in Understanding Materials Science and Technology (MS&T 10), Houston, TX, October 2010, pp. 355–380. C.G. Salmon, J.E. Johnson, F.A. Malhas, Steel Structures, 5th Ed., Prentice Hall, Pearson, 2009. S. Luckowski, J. Schutz, New Titanium Armor Application Provides Protection in Lightweight Kit, Proceedings of the Titanium 2008–24th Annual Conference Proceedings, International Titanium Association. S. Luckowski, Applications of titanium in armor solutions, Proceedings of the Titanium 2006–22nd Annual Conference Proceedings, International Titanium Association, New York, USA. S. Luckowski, Titanium – Protecting the Soldiers of Operation Iraqi Freedom, Proceedings of the Titanium 2005-21st Annual Conference Proceedings, International Titanium Association. AWS D1.9/D1.9M, Structural Welding Code – Titanium, The American Welding Society, July 2007. A.K. Patnaik, N. Poondla, U. Bathini, T.S. Srivatsan, Mater. Manuf. Process. 26 (2) (2011) 311–318. AWS D1.9/D 1.9M: Structural Welding Code for Titanium, The American Welding Society, July 2007. A.K. Patnaik, T.S. Srivatsan, A Study Aimed at Evaluating, Understanding, and Rationalizing the Strength, Endurance and Performance of Structures Made from Titanium and Titanium Alloy, Final Report, US Army and DMTC, October 2009, 256 pp. ASM International, ASM Handbook, Volume 2: Properties and Selection: Nonferrous Alloys and Special-Purpose Materials Publisher: ASM International, 10th ed. November 1990, ISBN-10: 0871703785, 586–591. P. Manfred, K. Jorg, C.H. Ward, C. Leyens, Adv. Eng. Mater. 5 (6) (2003) 419–427. C. Leyens, M. Peters (Eds.), Wiley-VCH, 2003. Steel Construction Manual (14th Edition), American Institute of Steel Construction, Chicago, Illinois, 2011.
Contents lists available at ScienceDirect
Materials Science & Engineering A journal homepage: www.elsevier.com/locate/msea
Understanding the mechanical response of built-up welded beams made from commercially pure titanium and a titanium alloy Anil K. Patnaik a,n, Narendra Poondla a,b, Craig C. Menzemer a, T.S. Srivatsan b a b
Department of Civil Engineering, The University of Akron, Akron, OH 44325, USA Department of Mechanical Engineering, The University of Akron, Akron, OH 44325, USA
art ic l e i nf o
a b s t r a c t
Article history: Received 4 September 2013 Accepted 15 October 2013 Available online 24 October 2013
During the last two decades, titanium has gradually grown in stature, strength and significance to take on the recognition of being a modern and high performance metal that is noticeably stronger and concurrently lighter than the most widely chosen and used steels in a spectrum of industrial applications. Technological innovations have necessitated reduction of part weight, cost and lead time, including concurrent enhancement of performance of structural parts and components made using titanium and its alloys. This has provided the impetus to develop economically viable structural design methodologies and specifications, while at the same time bringing forth innovative and economically affordable manufacturing and fabricating techniques with the primary purpose of both producing and promoting the use of cost-effective titanium structures. The experimental results of a recent study on built-up welded beams are presented in this paper with the primary objective of enabling design, facilitating fabrication, and implementation of large structural members for potential applications in the structural and defense-industry. & 2013 Elsevier B.V. All rights reserved.
Keywords: Commercially pure titanium Titanium alloys Built-up welded beams Static response Structural design
1. Introduction Commendable advances in the domains spanning extraction, processing, fabrication coupled with the methods both related and used for joining have made possible technically viable, structurally efficient and economically affordable applications of both pure titanium metal and its alloys for selection and use in large structural applications in the sectors spanning aerospace, construction and defense. There continues to exist a growing need and incentive to reduce part weight, lower the cost and decrease the time required for manufacturing, while concurrently enabling enhanced performance of structural parts made from both the older as well as the emerging generation of titanium alloys. An overview of alloys suitable for use in a spectrum of performance-critical applications in the domain of both aerospace and non-aerospace related applications, to include corrosion protection, armor, orthopedic, and sports equipment is presented and discussed in some recent papers by the authors [1–4]. Over the years, titanium has gradually grown both in strength and stature to gain for itself the significance of being recognized as a modern and high performance metal. Many titanium alloys are noticeably stronger and concurrently lighter than the most widely chosen and used steels for use in a spectrum of relevant and
n
Corresponding author. E-mail address: [email protected] (A.K. Patnaik).
0921-5093/$ - see front matter & 2013 Elsevier B.V. All rights reserved. http://dx.doi.org/10.1016/j.msea.2013.10.041
appropriate industrial applications. Importantly, two of the most attractive properties of pure titanium are its resistance to corrosion and superior ballistic properties [1,4,5]. In recent years, alloys of titanium have been frequently chosen for use as one of the main candidates for both aircraft structural parts and components of aircraft engines on account of its (a) primarily high strength-toweight ratio [s/ρ], and (b) high stiffness-to-weight ratio [E/ρ]. Despite its superior mechanical properties, the selection and use of pure titanium metal and its alloy counterparts have often been limited to high performance structures like: (i) aircraft wing, (ii) skin stiffened panels, (iii) armor structure, (iv) structural parts in navy vessels, and several other structures and components that find use in the defense industry that are essentially performancecritical, and not necessarily cost-limited [2,5]. It is generally believed that the high cost of both pure titanium metal and its alloy counterparts makes it difficult to implement its exhaustive use for those applications other than performancecritical structures. A one-to-one substitution of steel with an alloy of titanium is not easily possible based entirely on mechanical properties. This is because both pure titanium and its alloy counterparts will always work out to be more expensive. The cost of titanium is about 50 to 75 times higher than the cost of steel on a per-pound basis. Further, selection and immediate use of titanium metal is, as of now, not technology ready for its direct implementation in large civil engineering structures. At the same time, there limited design guidelines are available for the structural design of components making it difficult to implement
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in structures other than those specifically tested to verify performance. A primary reason for the high cost of the titanium metal is that it is not easily found in high concentration, or in the pure state, in nature. This has necessitated the need for multistage processing to bring it to a point where it can be chosen and put to use in a spectrum of structural applications. In order to gradually enhance the selection and utilization of both pure titanium metal and its alloy counterparts for both defense-related and non-defense-related applications, a study aimed at evaluating, understanding and rationalizing the strength, endurance and performance of structures made from commercially pure titanium and a titanium alloy (Ti–6Al–4V) was initiated at the University of Akron. This paper presents the results of the study on the structural behavior of built-up welded beams made from commercially pure titanium (Grade 2) and the titanium alloy Ti–6Al–4V under static loading. Other parts of the study such as fatigue behavior of built-up beams, microstructural studies, tensile performance, etc. have been reported elsewhere [6–12].
391
a) Evaluate the static tensile strength and static flexural behavior of commercially pure titanium (Grade 2) built-up beams to produce high strength, light weight, corrosion resistant and cost-effective structural beams for the purpose of possible applications in the defense and civilian sector. b) To evaluate Ti–6Al–4V alloy built-up beams fabricated and tested under conditions of static tensile and flexural loading.
3. Materials The materials selected for this study were (i) commercially pure titanium (Grade 2), and (ii) alloy Ti–6Al–4V. The two materials were systematically evaluated from the standpoint characterization of initial microstructure, tensile properties, and response to static bend loading. 3.1. Commercial pure (CP) titanium (Gr. 2)
2. Background The concept of fabricating built-up steel beams (also known as plate girders) from thin-walled plate elements is not new for highway bridges [13]. Since the early 1950s, a range of riveted, bolted, and welded plate girders have been preferentially chosen and used for appropriate applications. For example, three plates can be easily put together by welding to form the flanges and web of an I-shaped plate girder. For structures made from pure titanium metal and its alloy counterpart, it is common practice to machine the required part from large castings or billets. Overall, machining of titanium metal is a challenge by itself. Further, in the alloy form, titanium is harder to machine. A sizeable amount of material that is removed from either a thick plate or billet is often wasted, and this directly contributes to an observable increase in the cost of the end product. Also, the bars having a thicker cross-section are difficult and expensive to make, and are generally not readily available with material suppliers. Machining of titanium also poses fire hazard and therefore, needs special precautions for safety. With notable strides made in the specific domain of welding technology for both pure titanium metal and its alloy counterparts, it is now possible to weld titanium plates to make built-up structural components. For example, in recent years, a team of researchers and engineers at the US Army [Division at Picatinny Arsenal (New Jersey, USA)] have developed a Pulsed Gas Metal Arc Welding process (referred to henceforth as GMAW-P), which essentially involved a modification of the existing procedures so as to be applicable to both pure titanium metal and its alloys [14–17]. Linear weld speeds can be up to ten times faster than the corresponding speed for both tungsten inert gas (TIG) and gas tungsten arc welding (GTAW) processes. Besides, the number of passes when using the technique of GMAW-P can be reduced by a factor of three. The new process developed by the engineers and researchers at the US Army uses 100% helium shielding gas that not only facilitates good penetration but also ensures a double pulse template. The welding of titanium is standardized in a recently released welding specification put forth by the American Welding Society (AWS) designated as AWS D1.9/D1.9M [18]. It has therefore become feasible in the recent past to make welded built-up beams using titanium plates. The objective of this paper is to describe the results of a recent investigation on structural performance of built-up welded beams made from: (i) commercially pure (Grade 2) titanium, and (ii) the preferred and widely chosen and used alloy, Ti–6Al–4V, under static loading. The overall objectives were to:
The CP titanium (Gr. 2) plates were provided by TICO Titanium (Wixom, MI) in the thicknesses of 3/8 in. (9.5 mm) and 1/8 in. (3.2 mm). The edges of the plates and sheets were sheared. The nominal chemical composition of the as-provided material is summarized in Table 1. 3.2. Ti–6Al–4V alloy Plates made from this alloy were provided by Allegheny Technologies ATI Wah Chang (Albany, OR) in the thickness of 1/4 in. (6.35 mm). The nominal chemical composition of this alloy is summarized in Table 2.
4. Experimental procedures 4.1. Sample preparation for mechanical testing Tensile tests were conducted on specimens that were precision machined from both the longitudinal (L) and transverse (T) orientation of the as-provided annealed plate stock. The longitudinal (L) specimens were machined with the major stress axis parallel to the rolling direction of the Ti–6Al–4V alloy and CP (Grade 2) titanium plates, while the transverse (T) specimens were machined with the major stress axis perpendicular to the rolling direction of the two plates, i.e., Ti–6Al–4V and CP titanium (Grade 2). The round test specimens conformed to specifications outlined in the standard ASTM E-8 and had threaded ends. At the gage section, the test specimens measured 1/8 in. (3.2 mm) in diameter and 1/2 in. (12.5 mm) in length. To minimize the effects and/or contributions from both surface irregularities and finish, final surface preparation was achieved by mechanically polishing the gage section of the machined test specimens with progressively finer grades of silicon carbide impregnated emery paper (320 grit, 400 grit and 600 grit) to remove any and all circumferential scratches and surface machine marks. 4.2. Mechanical (tension) testing Uniaxial tensile tests were performed on a fully-automated, closed-loop servo-hydraulic mechanical test machine [INSTRON8500 Plus] using a 100 kN load cell. The tests were conducted at room temperature (300 K) and in the laboratory air (Relative Humidity of 55 pct.) environment. The test specimens were deformed at a constant strain rate of 0.0001/s. An axial 12.5 mm gage length clip-on type extensometer was attached to the test
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Table 1 Specifications, composition and average properties of the as provided CP Ti (Grade 2). Description
10 mm thick plate
3 mm thick sheet
Specification
ASTM B-265-06/ASME SB-265-A06 GR. 2EN 10024:2004 type 3.1
ASTM B-265-07/ASME SB-265-7/ASTM F-67-06GR. 2EN 10204:2004 type 3.1
0.13 0.16 0.002 0.012 Residual element (each) less than 0.10 Residual element (total) less than 0.40 Final product hydrogen 0.0032 Titanium remainder 507 MPa 360 MPa 33 pct 59 pct Yes
0.08 0.1 0.0105 0.014 Residual element (each) less than 0.10 Residual element (total) less than 0.4 Final product hydrogen 11/10 PPM Titanium remainder 484 MPa 341 MPa 25.75 pct – 1370 F, HGA 3 to 11 min Pass
Chemical composition Fe iron Oxygen N Nitrogen C Carbon
Average tensile strength, psi Yield 0.2% offset Elongation % RA % Anneal condition Guided bend test
Table 2 Nominal chemical composition of Ti–6Al–4V alloy (in wt%). Material
Ti
Al
N
V
C
Fe
H
O
Ti–6Al–4V
90.0
6.0
0.05
4.0
0.1
0.4
0.02
0.20
specimen at the gage section. The stress and strain measurements, parallel to the line of loading, and the resultant mechanical properties, such as, (i) stiffness, (ii) strength (yield strength and ultimate tensile strength), (iii) failure stress, and (iv) ductility (strain-to-failure) was provided as a computer output by the control unit of the test machine. 4.3. Static bend tests of titanium alloy beams The test specimens were fabricated using GMAW-P welding process. The dimensional details of the beams are given in Fig. 1a and b. Complete details of the fabrication can be found elsewhere [17,19]. Two of the beams (designated as B1 and B4) were tested with the primary purpose of studying the flexural performance when bent in single curvature under static loading. Beam B1 was made from the Ti–6Al–4V alloy while beam B4 was made from commercially pure titanium (Gr. 2). The span of the two test beams measured 610 mm. Another beam (denoted as B5) was also made from commercially pure titanium (Gr. 2) and deformed under conditions of fatigue loading as described elsewhere [6]. The fatigue test was stopped following one million cycles as a consequence of load at the one-millionth cycle. It was then tested under static loading after the beam was subjected to one million cycles of fatigue loading. Details of the static load test on beam B5 are also presented and discussed in this paper. 4.3.1. The test set-up The test set-up for the static bend tests on the beams is shown in Fig. 2a and b. The test beams were loaded on an INSTRON 5300 machine with a maximum load capacity of 1000 kN (225 kips). The beams were loaded, using a four-point loading arrangement, over a simple span of 610 mm (24 in.). The total length of each specimen was 685 mm. The test beams were supported at either end on rollers that allowed the beam to essentially behave as “simply” supported. Also, the supports at either end allowed for free
rotation of the beam while concurrently permitting free sliding to occur. Bearing plates, measuring 12.5 mm in thickness and 50 mm in width, were provided over the roller supports. The compression flange of each test beam was restrained from any lateral displacement at the supports. The restraint was provided by a fixture that was specifically designed and fabricated for the static tests. The two fixtures that were provided at the two ends of the test beam are as shown in Fig. 2a. This figure shows a front view of the fixtures at the two ends. A side view of the fixture at one end (support) is shown in Fig. 2b. Lateral displacement of the compression flange was prevented by two threaded rods that were provided in each fixture. Complete details of the test fixture are presented and discussed elsewhere [19].
4.3.2. Instrumentation The development of strains during the entire duration of static loading was recorded with the help of strain gages positioned at seven different locations. A Digital Image Correlation (DIC) system (ARAMIS) was used to capture the strain fields of each test beam during the entire duration of each static test. The deflections occurring at the mid-span were recorded with the aid of a digital dial gage. The failure mode was captured in the form of digital images in the DIC system. During the entire duration of the test, the data was programmed to be collected by: (i) the control unit of the INSTRON test machine, and (ii) an independent stand-alone Data Acquisition System (DAQ). The typical strain gage layout for each test beam (i.e., B1 and B4) is as shown in Fig. 3. The strain gages were attached in the region of the midspan on both the top flange and the bottom flange, one on each side of the web. Additionally, two strain gages were attached in a 451 pattern on the web, at a quarter span locations, on each side of the points of loading.
4.3.3. Test procedure A preload of 2.25 kN (0.5 kips) was initially applied to ensure firm contact between the beam and its supports. (a) A loading rate of 4.45 kN (1000 lbs.) per minute was used for loading beams B4 and B5 both of which were made from commercially pure titanium (Gr. 2). In case of beam B5, static
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Fig. 1. (a) Details of commercially pure titanium beam B4 (Beam B5 is similar) and (b) details of commercially Ti–6Al–4V beam B1 (all dimensions are in mm).
load test was conducted to find the residual strength after fatigue loading. (b) On account of its higher load carrying capability, the loading rate was tripled for the beam made from alloy Ti–6Al–4V. The images were captured by the DIC system at the same rate as the deflection readings and frequency of the data acquisition system (DAQ).
5. Results and discussion 5.1. Initial microstructure The microstructure of the as-provided material is an important factor that determines or governs its response to an external load or mechanical stimulus. In essence, the microstructure governs the tensile properties, fracture toughness, compressive response, fatigue resistance of the material chosen for the beams, and the overall kinetics of fracture at both the macroscopic and microscopic level. Optical microstructure of the as-received commercially pure titanium (Gr 2) plate is shown in Fig. 4 at two magnifications. At low magnification the overall microstructure revealed the primary alpha (α) grains to be intermingled with the
small yet noticeable pockets of the beta (β) grains. At a higher magnification, very fine alpha (α) phase lamellae dispersed well within the beta (β) grains was evident (Fig. 4b). The optical microstructure of the Ti–6Al–4V alloy is shown for two different magnifications in Fig. 5. The two micrographs in this figure reveal the as-received, undeformed microstructure of the titanium alloy specifically the volume fraction, morphology, size and distribution of the intrinsic micro-constituents through the microstructure. Over the range of magnifications spanning very low to high revealed a duplex microstructure consisting of near equiaxed alpha (α) and transformed beta (β) phases. The primary nearequiaxed shaped alpha (α) grains (light in color) were well distributed in a lamellar matrix with transformed beta (dark in color). The presence of trace amounts of aluminum and oxygen in this Ti–6Al–4V alloy contributes to strengthening and stabilizing the alpha (α) phase, which is beneficial for enhancing hardenability, increasing strength and improving the response kinetics of the alloy to heat treatment. 5.2. Tensile response The room temperature tensile properties of both CP (Grade 2) and the Ti–6Al–4V alloy are summarized in Table 3. The results reported are the mean values based on duplicate tests. The elastic
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Fig. 2. (a) The test set-up and (b) a side view showing lateral supports.
modulus (E), yield strength (sYS), ultimate tensile strength (sUTS), elongation-to-failure (εf) and strength at failure (fracture) (sfr) were recorded in a data acquisition system. The yield strength was determined by identifying the stress at a point on the engineering stress versus engineering strain curve where a straight line drawn parallel to the elastic portion of the stress versus strain curve at 0.2% offset intersects the curve. The ductility is reported as elongation-to-failure over a gage length of 12.7 mm. This elongation was measured using a clip-on extensometer that was attached to the gage section of the test specimen.
5.2.1. Ti–6Al–4V alloy A representative engineering stress versus engineering strain curve for the two orientations of the as-received plate, i.e., longitudinal (L) and transverse (T), is shown in Fig. 6. The elastic modulus of the alloy is 126 GPa (18,300 ksi) in the longitudinal (L) orientation and 137 GPa (19,900 ksi) in the transverse (T) orientation. The yield strength (sYS), of the alloy in the annealed condition is 948 MPa (137 ksi) in the longitudinal (L) orientation and 1047 MPa (152 ksi) in the transverse (T) orientation. The ultimate tensile strength (sUTS), of the alloy is 1060 MPa (154 ksi) in the
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longitudinal (L) orientation and 1181 MPa (171 ksi) in the transverse (T) orientation. Both the yield strength and tensile strength are higher in the transverse orientation than in the longitudinal orientation. The ultimate tensile strength of the Ti–6Al–4V alloy is only marginally higher than the yield strength, i.e., about 10%, indicating a low work hardening rate beyond yield. The elongation-to-failure (εf) of the alloy is 7.8% in the longitudinal (L) orientation and 11.5% in the transverse (T) orientation. The reduction-in-area of the alloy was 24% in the longitudinal (L) orientation and 22% in the transverse (T) orientation and agrees reasonably well with the annealed condition of the alloy microstructure. The yield strength and tensile strength values of the alloy conform well to the values obtained and reported by the manufacturer (ATI Wah Chang) and recorded in ASM Metals Handbook and by others [20–22]. Variation of stress with plastic strain reveals the Ti–6Al–4V to be marginally harder in the transverse (T) orientation when compared to the longitudinal (L)
Fig. 3. Schematic showing layout of the strain gages.
395
orientation. The strain hardening exponent (n) of this alloy in the two orientations is 0.014 for the transverse (T) orientation and 0.012 for the longitudinal (L) orientation.
5.2.2. Commercially pure titanium (Grade 2) The elastic modulus of this material is 115 GPa (16,700 ksi) in the longitudinal (L) orientation and 108 GPa (15,700 ksi) in the transverse (T) orientation. For the CP (Grade 2) titanium, the yield strength is 432 MPa (63 ksi) in the longitudinal (L) orientation and 15 pct. higher than in the transverse (T) orientation (382 MPa or 55 ksi). The ultimate tensile strength is 561 MPa (81 ksi) in the longitudinal (L) orientation and 465 MPa (67 ksi) in the transverse (T) orientation, a noticeable difference of 20 pct. The ultimate tensile strength is higher than the yield strength indicating a noticeable work hardening beyond yield. The elongation of CP (Grade 2) was 14.7% in the longitudinal (L) orientation and 18.9% in the transverse (T) orientation. The reduction-in-area was 43.4% in the longitudinal (L) orientation and 47% in the transverse (T) orientation. The engineering stress versus engineering strain curves for the material in both the longitudinal (L) and transverse (T) orientations are shown in Fig. 7. Variation of stress with plastic strain reveals the CP (Grade 2) to be harder in the longitudinal (L) orientation when compared to the transverse (T) orientation. The strain hardening exponent (n) of this material in the two orientations is 0.069 for the longitudinal (L) orientation and 0.054 for the transverse (T) orientation. In the longitudinal orientation of the as-provided annealed plates, the Ti–6Al–4V alloy has an elastic modulus of 125 GPa (18,100 ksi), while the commercially pure (CP: Grade 2) titanium has an average elastic modulus of 115 GPa (16,700 ksi).
Fig. 4. Optical micrographs showing the key micro-constituents in the commercially pure titanium (Gr 2) plate at two different magnifications.
Fig. 5. Optical micrographs showing the key micro-constituents in the Ti–6Al–4V alloy plate at two different magnifications.
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Table 3 Room temperature tensile properties of Ti–6Al–4V and commercially pure (Grade 2) [Results are mean values based on duplicate tests.]. Material
Ti–6Al–4V Commercially pure (Grade 2)
Orientation
Longitudinal Transverse Longitudinal Transverse
Elastic modulus
Yield strength
UTS
msi
Gpa
Ksi
MPa
Ksi
MPa
18 20 16 15
126 137 115 108
137 152 63 55
948 1047 432 382
154 171 81 67
1060 1181 561 465
Elongation GL¼ 0.5″(%)
Reduction in area (%)
Tensile ductility In (Aa/Af) (%)
7.8 11.5 14.7 18.9
23.8 21.7 43.4 47.0
27.0 25.0 57.0 63.0
5.3.1. Beam B4 – commercially pure titanium (Gr. 2) The failure mode of beam B4 is shown in Fig. 8a. The failure mode captured by the Digital Image Correlation (DIC) system (ARAMIS) is shown in Fig. 8b. Failure of the beam was initiated by shear buckling of the web, followed by excessive deflection and yielding of the flange. A careful physical inspection of the welds between the web and the flange and the connecting stiffeners revealed that the joints were intact with no evidence of failure. Also, there was no evidence of visible cracking. The beam maintained its structural integrity during the entire range of loading.
Fig. 6. Influence of test specimen orientation on engineering stress versus engineering strain curve of Ti–6Al–4V alloy.
Fig. 7. Influence of test specimen orientation on engineering stress versus engineering strain curve for commercially pure (Grade 2) titanium.
5.3. The static bend tests The two test beams B1 and B4 failed at a static load greater than the predicted value obtained using the design methodologies developed in an earlier study [19]. The predicted methods were developed based on the 14th edition of the prevalent AISC steel design specifications [23]. Beam B4 made from commercially pure titanium (Grade 2) failed at a load of 187 kN (42 kips). The residual static strength (i.e., the strength determined from static bend test following one million cycles of fatigue loading) of beam B5 made from commercially pure titanium (Gr 2) at failure was determined to be 205 kN (46 kips). The failure load predicted for this test beam using the proposed design methodology was 160 kN (36 kips) [19]. Failure for both beams occurred by a combination of shear buckling of the web and excessive deflection of the beam with yielding of the flanges.
5.3.2. Beam B5 (following fatigue loading) – commercially pure titanium (Gr. 2) Beam B5 that was subject to static loading subsequent to cyclic loading for one million cycles, failed in a manner quite similar to the failure mode experienced by beam B4. Failure of this beam was initiated by shear buckling of the web, followed by excessive deflection and yielding of the flange. Deflections were recorded over the entire range of static loading. Failure of beam B5 by web shear buckling is shown in Fig. 9. 5.3.3. Beam B1 – Ti–6Al–4V alloy Beam B1, which is made from Ti–6Al–4V alloy failed at a load of 583 kN (131 kips). The predicted failure load for this beam was 494 kN (111 kips). Failure of the beam occurred by a combination of excessive deflection (Fig. 10) with buckling of the compression flange between the two loading points (Fig. 11). There was visible evidence of complete plasticization of the length of the beam between the two loading points. Both the straightening and stretching of the bottom flange, as seen in Fig. 10, is evidence of tensile yielding of the bottom flange, while the top flange in compression completely yielded and buckled to a multiple wave form (Fig. 11). The proposed design method described in detail elsewhere by the authors [19] predicted the failure mode to be plasticization of both the compression flange and the tension flange. Therefore, the observed failure mode was as expected. There was no visible evidence of cracking of the welds in the test beam. Further, the welds were intact over the entire range of static loading. However, there occurred permanent deformation of the beam due to excessive stressing of both the compression flange and the tension flange. 5.4. The load versus deflection curves The load versus deflection curves were obtained from tests on each test beam. One was obtained from the digital dial gage while the other was obtained from readings of the test machine [INSTRON]. The load versus deflection curves for beam B1 (Ti– 6Al–4V beam) are as shown in Fig. 12. The shape of the load versus deflection curves was quite similar to the stress versus strain curve obtained for the material using small round tensile specimens that were prepared in conformance with ASTM Standard E-8. The initial part of the load-deflection curve is linear, up to about
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Fig. 8. (a) Failure mode of the commercially pure titanium (Gr. 2) beam (B4) by web shear buckling. (b) Failure mode of commercially pure titanium (Gr. 2) beam (B4) by web shear buckling as captured by the DIC system (ARAMIS). The strain contours are also shown.
Fig. 10. Failure of beam B1 (Ti–6Al–4V) as captured by DIC system (ARAMIS).
Fig. 9. Failure of beam B5 by web shear buckling.
400 kN (90 kip) load, following which the curve reveals the initiation of non-linearity indicating the onset of yielding. The load versus deflection curve demonstrates the ductile nature of the built-up beam under the conditions of static loading. The overall shape of the curve can be classified as being quite
similar to that of a classic ductile beam that reveals adequate reserve strength coupled with deflection capability to provide the required warning prior to catastrophic failure. Fig. 12 also shows the deflection curve predicted using the classical elastic beam bending theory. The elastic deflection method was found to marginally under-estimate the overall deflection of the beam. This figure reveals the existence of a match between the predicted deflection and the deflections measured during the test up to the onset of the elastic limit.
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Fig. 11. Compression flange (top flange) local buckling of beam B1 (Ti–6Al–4V).
Fig. 14. The load versus deflection curve of beam B5 (commercially pure titanium (Gr. 2)) following the fatigue test.
Fig. 12. The load versus deflection curve obtained for beam B1 (Ti–6Al–4V).
Fig. 15. Load versus strain curves for beam B4 (commercially pure titanium (Gr. 2)). Outside strains on the top flange.
Fig. 13. The load versus deflection curve of beam B4 made from commercially pure titanium (Gr. 2).
elastic shear buckling limit into the post-buckling region. The deflection curve obtained was quite similar to the behavior shown by a classic welded steel beam [13]. Both Figs. 13 and 14 also include the deflection curves predicted using the classic elastic beam bending theory. The method of elastic deflection was found to under-estimate the actual deflection experienced by the titanium beams. During mechanical testing, the beams revealed noticeably greater deflection than the value predicted using classic elastic beam bending theory. At a level of the service load, that was taken to be 50% of the failure load, the actual deflections experienced by the beam were found to be about 50 pct. to 75 pct. greater than the predicted values. This behavior is more pronounced for beam B5. However, the softening may have occurred due to the fatigue loading the beam was subjected to prior to the static test.
5.5. The load versus strain relation The load versus deflection curves of beams B4 and B5 (B5 tested following run out during the fatigue test) are shown in Figs. 13 and 14, respectively. These load versus deflection curves demonstrate evidence of distinct softening of the beam on reaching: (a) a load of 111 kN (25 kips) for beam B4 (Fig. 13), and (b) a load of 165 kN (37 kips) for beam B5 (Fig. 14). The observed softening is rationalized from the viewpoint of the initiation and occurrence of inelastic shear buckling. Softening of the beam occurred when the web reached its limit for the elastic shear buckling load, while the beam carried a load well beyond the
The load versus strain data collected from the data acquisition system (DAQ) was plotted and the relationship is shown in Fig. 15 for beam B4 (commercially pure titanium (Gr. 2) beam). Also, Fig. 16 shows the strain curves for the bottom flange at the midspan location. Both the theoretical strains and the strains recorded during the static test of the beam agree well for this beam (B4) for low values of flange strain up until the elastic limit. However, at a load level corresponding to the initiation of elastic web shear buckling (approximately 111 kN–133 kN (25–30 kips)) the curves
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Fig. 16. Load versus strain curves for beam B4 made from commercially pure titanium (Gr. 2): outside strains on the bottom flange.
Fig. 17. The load versus strain curves for beam B1 (Ti–6Al–4V): outside strains on the top flange.
obtained from the two strain gages during the static test, start to deviate from the theoretical strain. In Fig. 17 is shown the strain curves on the top flange taken at the mid-span location. While Fig. 18 reveals the corresponding curves for one web of the beam. Both these figures also show the theoretical strain curves that were obtained from theoretical analysis. For the case of the flange, both the theoretical strains and the strains recorded during the static test of the beam accord reasonably well revealing that the strains obtained from the test results match reasonably well for the two strain gages used.
399
Fig. 18. Load versus strain curves for beam B1 (Ti–6Al–4V): web strains (451).
The load versus deflection curves clearly demonstrate the predicted deflection curve for the beam made from Ti–6Al–4V (i.e., B1) to accord well with the actual deflection curve obtained from the test. Also, both beam B4 and beam B5 demonstrate a distinct softening of the beam upon reaching a load that corresponds to the onset of elastic shear buckling. The elastic deflections determined for the two beams underestimates the deflection when compared one-on-one with the values obtained from the static bend tests. The load versus strain response determined from tests conducted on both beam B1 (Ti–6Al–4V) and B4 (CP Ti – Gr. 2) match reasonably well with those determined by elastic beam bending theory for the initial part of the loading regime. The observed difference between the theoretical strain values and the experimental strain values was noticeable once the load exceeded the yield limit of the material. The strains obtained from the experimental tests and measured on the web panels deviated appreciably from the theoretical values. Both the commercially pure titanium (Grade 2) beam and the Ti–6Al–4V alloy test beam failed in modes that were predicted using the methodologies developed in an earlier study by the authors [19]. Further, the test beams demonstrated significant reserve strength following initial yielding. The experimental deflection curves along with the predicted deflection curves are presented and briefly discussed. These curves along with a comparison of load versus strain relationship obtained from the experimental tests with those predicted using theoretical methods demonstrate a reasonably close match between the theoretical predictions and the experimental test results up until the elastic limit of the material.
5.6. Analysis of results of the static bend tests on the welded beams
6. Conclusions
A summary of the failure loads that were determined from the experimental tests on the beams along with the corresponding predicted load is provided in Table 4. The failure load obtained for each of the three beams is greater than the predicted value of the failure load. Also, the actual strength obtained from the experimental tests was greater than the predicted value of strength by 16–28%. This suggests the existence of significant reserve strength in the test beams when they are designed in conformance with the methods used in this research study. Also, both the predicted failure mode and actual failure mode accord reasonably well. However, the occurrence of noticeably large deflections at failure does limit the usable strength to a value that is close to the predicted theoretical load in order to limit deflection.
With the latest GMAW-P welding technology developed at US Army Picatinny Arsenal, it is feasible to fabricate large welded built-up titanium beams by welding plates together to achieve structural performance that is quite comparable with an identical monolithic beam that is produced by machining from either thick plates or billets. The study demonstrated that welded built-up beams made from both sheets and plates of commercially pure titanium (Gr. 2) and Ti–6Al–4V can be fabricated with relative ease. The welds produced by the new GMAW-P method are sound and revealed an absence of visible cracks. (a) The results obtained in this study reveal the absence of a deleterious influence of welding on structural performance of
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Table 4 A summary of failure load and predicted strength of the beams studied. Beam
Material
B1
Ti–6Al–4V
B4
CP Ti (Gr. 2)
B5
CP Ti (Gr. 2)
(b)
(c)
(d)
(e)
Predicted failure load Predicted
Actual
494 kN (111 kips) 160 kN (36 kips) 160 kN (36 kips)
583 kN (131 kips) 187 kN (42 kips) 205 kN (46 kips)
built-up welded beams of commercially pure titanium (Grade 2) and Ti–6Al–4V titanium alloy. The welded beam concept used in this study worked well for the titanium beams that were tested. It is anticipated that the concept of welded built-up beams can be applied to structural members made from both commercially pure titanium (Grade 2) and the alloy (Ti–6Al–4V). This is a cost saving alternative to fabricating large structural elements and members by machining the parts from either a thick plate or billet. Currently, prevalent AISC steel design specifications (14th edition, 2011) were modified to suit the material properties of commercially pure titanium and its alloy counterparts. With suitable modifications, these specifications can be used for the design of welded built-up beams made from both commercially pure titanium and its alloys. This approach was found to be successful. However, additional research work is needed to refine the design approach. The failure mode, the failure load, deflections, and the strains induced in the welded built-up titanium beams are predictable to a reasonable and acceptable level of accuracy. The test beams demonstrated significant reserve strength following the onset of yielding. The experimental results versus predicted response demonstrate a reasonably close match between the theoretical predictions and the experimental test results up until the elastic limit of the material. The study reiterated that elastic stress analysis based on classical beam bending theory is adequate to predict the stresses, strains and deflections that are developed under the influence of a static load up until the initiation of yield in the material. The ultimate behavior is also predictable to a reasonable degree of accuracy in a conservative manner.
[4] [5] [6] [7]
[8] [9] [10] [11] [12]
[13] [14]
[15]
[16]
[17] [18]
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[19]
[20]
[21] [22] [23]
Difference
Failure mode both predicted and actual
þ18.0%
Bending failure leading to excessive deflection
þ16.7
Shear failure by web buckling
þ27.8%
Shear failure by web buckling
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