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Mechanical and microstructural testing of wire and arc additivelymanufactured sheet material
Pinelopi Kyvelou a,b,⁎, Harry Slack a, Dafni Daskalaki Mountanou c, M. Ahmer Wadee a, T. Ben Britton c,Craig Buchanan a,b, Leroy Gardner a,b
a Department of Civil and Environmental Engineering, Imperial College London, UKb Data Centric Engineering Programme, The Alan Turing Institute, London, UKc Department of Materials, Imperial College London, UK
H I G H L I G H T S G R A P H I C A L A B S T R A C T
• Results of tensile tests on WAAM stain-less steel coupons are presented.
• Degree of anisotropy and influence ofgeometric variability are examined.
• Non-contact measurement methods areused to determine the geometry and de-formations.
• Microstructural analysis of the samplesreveals a strong crystallographic tex-ture.
• Effective mechanical properties definedfor as-built material based on simplegeometrical measures.
Article history:Received 24 February 2020Received in revised form 20 March 2020Accepted 24 March 2020Available online 29 March 2020
Keywords:Metal 3D printingMaterial anisotropyMechanical propertiesMicrostructural analysisStainless steelTensile coupon tests
Wire and arc additivemanufacturing (WAAM) is a method of 3D printing that enables large elements to be built,with reasonable printing times and costs. There are, however, uncertainties relating to the structural perfor-mance of WAAM material, including the basic mechanical properties, the degree of anisotropy, the influence ofthe as-built geometry and the variability in response. Towards addressing this knowledge gap, a comprehensiveseries of tensile tests onWAAM stainless steel was conducted; the results are presented herein. As-built andma-chined coupons were tested to investigate the influence of the geometrical irregularity on the stress-strain char-acteristics, while material anisotropy was explored by testing coupons produced at different angles to theprinting orientation. Non-contact measurement techniques were employed to determine the geometric proper-ties and deformation fields of the specimens, while sophisticated analysismethodswere used for post processingthe test data. The material response revealed a significant degree of anisotropy, explained by the existence of astrong crystallographic texture, uncovered bymeans of electron backscatter diffraction. Finally, the effectiveme-chanical properties of the as-built material were shown to be strongly dependent on the geometric variability;simple geometric measures were therefore developed to characterise the key aspects of the observed behaviour.
Table 1Chemical composition of 308LSi austenitic stainless steelwire (values in %), as providedbythe manufacturer.
C Mn Si Mo P S Cr Ni Ferrite
0.02 1.80 0.85 0.20 ≤0.025 ≤0.020 20 10 5–10
Table 2Mechanical properties of 308LSi austenitic stainless steel wire, as provided by themanufacturer.
Heat treatment Yield strength(MPa)
Tensile strength(MPa)
ElongationA5 (%)
Impact energyISO – V (J) 20 °C
As-welded ≥350 ≥520 ≥35 ≥47
2 P. Kyvelou et al. / Materials and Design 192 (2020) 108675
1. Introduction
Additive manufacturing (AM), commonly known as 3D printing, is amanufacturing process whereby a component is built-up layer by layer,as defined by a 3D digitalmodel [1]. The potential for a reduction inma-terial consumption and wastage, geometric freedom and enhancedcustomisation when compared to conventional fabrication methods[2], in conjunction with the feasibility of its application across a widerange ofmaterials includingmetals, polymers, ceramics and composites[3], has led to rapid growth in the use of AM across multiple industries[4]. The worldwide AM market was estimated to be worth over $7 bil-lion in 2018, with market trends indicating further upcoming growth[5].
According to ISO/ASTM 52900 [6], the principal types of metal AMare sheet lamination, powder bed fusion (PBF) and directed energy de-position (DED), requiring little to no post-processing, with the two lat-ter techniques being deemed the most appropriate for the constructionindustry [2,3]. Wire and arc additive manufacturing (WAAM) is amethod of DED that uses conventional welding technology [7,8],allowing the production of large-scale metallic components [9,10] in atimely and cost-effective manner [11]. However, there are a numberof prominent challenges, many of which relate to uncertainties and var-iability in the basic mechanical properties [12].
The focus of much of the previous research on metallic AMmaterialreported in the literature has been on alloys and geometries suitable foraerospace and mechanical engineering applications [13–17]; studiesinto the behaviour of AM stainless steel have been conducted byYadollahi et al. [18], Mower and Long [16], Röttger et al. [19], Wanget al. [20], and Liverani et al. [21]. The material behaviour of both PBFand WAAM steels has been shown to be generally anisotropic [22–27]while residuals stresses have also been found to be prominent [28–30].
Studies into themicrostructure andmechanical properties ofWAAMstainless steel, although limited in number, have indicated the potentialcorrelation between the employed printing strategy and the exhibitedmaterial properties [31–33]. Eriksson et al. [34] showed that increasingthe heat input during printing results in increased ductility and reducedyield and tensile strengths, while Ji et al. [35] found that specimensloaded parallel to their printing orientation exhibited higher ultimatestrength and less ductility than specimens with the load acting perpen-dicular to their printing orientation. However, it should be noted thatother researchers did not observe such anisotropic behaviour [36–38].
The aims of the present paper are to broaden the pool of experimen-tal data on WAAM stainless steel and to assess the dependency of thematerial properties on the employed printing strategy. The undulatingas-built geometry, which is inherent to the WAAM process, can be ma-chined smooth, but this additional operationwill add time and cost andmay not be feasible with certain printed geometries. If the material isleft in the as-built state, an important question is: what is the influenceof the inherent geometric undulation on the effective mechanical prop-erties? This issue is also addressed herein and simple design rules forpredicting the key mechanical properties are established.
2. Manufacturing of test specimens
Flat plates of 3.5 mm and 8.0 mm nominal thickness were cut fromoval tubes (with flat sides), printed using Grade 308LSi austenitic stain-less steel wire of 1 mm thickness. The chemical composition of theemployed parent stainless steel wire, as well as its mechanical proper-ties, as provided in the manufacturer's data sheet, are presented inTables 1 and 2, respectively. The specimens were printed by the Dutchstart-up company MX3D, using their proprietary multi-axis roboticWAAM technology [39]. Following fabrication, the specimens weresandblasted with glass beads, the size of which was sufficiently coarseto clean the surface, but sufficiently fine not to affect their geometry.
The as-built surface of WAAM elements is inherently undulating,resulting in geometrical variations across their profile and hence non-
uniform deformations under load. In order to assess the influence ofthe as-built geometry on the mechanical response, tensile coupontests were performed on both the as-built and machined material; inthe case of the latter, the surface undulations of the as-built materialof 8.0 mm nominal thickness were removed by an end mill, as shownin Fig. 1(a). The resultingmachined coupons were prismatic and of uni-form thickness; a comparison of the as-built and machined surface of atypical plate is shown in Fig. 1(b). Coupons were extracted from boththe as-built and machined material at 0°, 45° and 90° to the printlayer orientation, as defined in Fig. 2, in order to investigate anymaterialanisotropy. Typical as-built and machined coupons are illustrated inFig. 3.
Prior to testing, themachined couponswere X-ray scanned to deter-mine the location and extent of any internal voids; a typical scan isshown in Fig. 4. Minor internal defects were identified, with void sizesof up to 1 mm (in the plane of the X-ray image), though there was noobvious correlation between the void locations and the position of frac-ture or noticeable influence on the measured strain field; the presenceof theminor internal defects was therefore deemed not to have any sig-nificant influence on the monotonic deformation response of thecoupons.
The adopted specimen labelling system begins with the letter M forthe machined coupons and AB for the as-built coupons; this is followedby the nominal thickness of the plate inmm, the angle to the print layerorientation in degrees as defined in Fig. 2, the plate fromwhich the cou-pon was cut and finally the number of the specimen originating fromthe plate. For instance, coupon AB-3.5-0°-C4 is an as-built couponwith a nominal thickness of 3.5 mm, aligned at a 0° angle to the printlayer orientation and was the fourth coupon from the plate labelled“C”. In total, 51 coupons of different thicknesses, orientations and fin-ishes were tested.
3. Geometric measurements
The geometricmeasurements of themachined tensile couponsweredetermined prior to testing using Vernier callipers. However, the undu-lating surface of the as-built WAAM coupons rendered the use of con-ventional measuring techniques impractical. Laser scanning wastherefore employed to record the full geometry of the printed speci-mens digitally and hence obtain detailed and accurate geometric mea-surements, as described in this section.
3.1. Methods and techniques
A FARO ScanARM, capable of measuring up to 600,000 points persecond to an accuracy of 0.1mm,was used to scan each as-built coupon.Laser scans were taken at several different angles of incidence to thecoupon surface in order to capture all of the geometric undulations.The scans were subsequently merged using the software Geomagic
(a)
Machined
As-built
(b)
Fig. 1. (a) Milling of WAAM plate to produce machined (flat) coupons and (b) comparison of as-built and machined surfaces.
3P. Kyvelou et al. / Materials and Design 192 (2020) 108675
Wrap [40], represented as a point cloud – see Fig. 5, and inter-connectedto form a polygonal mesh and subsequently a 3D CAD model.
The 3D CAD models created in Geomagic Wrap [40] were importedinto Rhino 3D [41] as STLfiles, and accurately alignedwith the global co-ordinate system using principal component analysis [42,43], with theoverall centroid of the coupon coinciding with the origin of the globalcoordinate system and with the longitudinal axis of the coupon beingparallel to the global axis x - see Fig. 6. Following alignment, contouringof each specimen at regular intervals dx along the longitudinal axis wasconducted to determine the geometric properties at each cross-sectional cut. The cross-sectional area Ai and the eccentricities ey,i andez,i of the cross-sectional centroid Pi, along the global y and z axes re-spectively, relative to the overall centroid of the coupon (coincidingwith the origin of the global coordinate system), were calculated ateach cut. The average cross-sectional thickness ti was calculated bycontouring the coupon in the transverse direction (along the global yaxis) with the same spacing used for the longitudinal contouring. Aschematic illustration of a typical coupon processed in Rhino 3D, aswell as the resulting cross-sectional area Ai, thickness ti and eccentrici-ties ey,i and ez,i at a typical cross-section, are illustrated in Fig. 6.
A sensitivity study was undertaken to determine the most suitablecontour spacing that would allow sufficient measurements to be takento ensure an accurate replication of the geometrical variation alongthe length of the specimen, while achieving computational efficiency.Twelve as-built specimens were chosen (two for each build directionand thickness), the geometric measurements of which were obtainedat contour spacings of 0.05 mm, 0.10 mm, 0.50 mm, 1.00 mm and2.00 mm. Note that a square mesh was employed for all different spac-ings with dx = dy. It should be mentioned that the considered contourspacings were all below the typical WAAM deposition width w, shownin Fig. 7, which was found to vary between about 3 mm and 5 mm forthe studied samples; a similar value of 4 mm was reported by Dinget al. [44].
= 90° = 0°
= 45°
Print layer orientation
Fig. 2. Orientation of tensile coupons extracted from WAAM plate relative to print layerorientation.
A summary of the results obtained from the conducted sensitivitystudy is presented in Fig. 8, where the mean and minimum measure-ments of area (A and Amin respectively) and thickness (t and tmin respec-tively) and the mean and maximum centroid eccentricity (ez and ez,max
respectively) for each contour spacing are normalised against the equiv-alent values corresponding to dx=0.05 mm. As expected, it can be ob-served that the extreme values of all measurements (namely Amin, tmin
and ez,max) are more sensitive to the contour spacing compared totheir equivalent mean values (namely A, t and ez respectively). Overall,the measurements obtained using a spacing of 0.10 mm were almostidentical to these obtained using a spacing of 0.05mm; hence a contourspacing of 0.10 mmwas adopted.
3.2. Geometric properties
The average geometric properties of the as-built coupons are pre-sented in Table 3, where tnom is the nominal thickness of each specimen,θ is the orientation of the coupon relative to the print layer orientation,as defined in Fig. 2, A is the mean cross-sectional area, t, tmin and tsd arethe mean, minimum and standard deviation values of the thickness re-spectively, ey, ey,max and ey,sd are themean, maximum and standard de-viation values of the eccentricity along the y axis respectively, ez, ez,max
and ez,sd are the mean, maximum and standard deviation values of theeccentricity along the z axis respectively andΔ is thedifference betweenthe actual length of the as-built coupon and its corresponding devel-oped length (i.e. the length of a line passing through each cross-sectional centroid, within the parallel length).
The θ = 0° coupons exhibited the least variation in thickness alongtheir length, as indicated by the tsd values and the values of the tmin/tratio being close to unity; this can be explainedwith reference to the di-rection of the contours relative to the print layer orientation. For theθ= 0° specimens, as illustrated in Fig. 9(a), the contour planes are per-pendicular to the direction of the deposition paths, and all cross-sections are similar; hence, the variation in the values of ti between dif-ferent cross-sections is expected to be minor, leading to low values ofstandard deviation tsd. In contrast, since for the θ = 90° coupons thecontour planes are parallel to the direction of the deposition paths(see Fig. 9(b)), some cross-section cuts are expected to contain onlytrough regions (leading to values of tmin being substantially lower
Fig. 3. As-built (top) and machined (bottom) tensile coupons.
= 0° = 90° = 45°
A C I K Q S
B D J L T R
Fig. 4. X-ray scan of machined coupons.
4 P. Kyvelou et al. / Materials and Design 192 (2020) 108675
than t) while others only crest regions, leading to amuch greater spreadin ti, and, thus, higher values of tsd.
A histogram of themeasured thicknesses, grouped by coupon orien-tation relative to the print layer direction, is presented in Fig. 10, witheach cross-sectional thicknessmeasurement ti normalised by the corre-sponding average coupon thickness t. For the 0° specimens, thedistribu-tion is approximately symmetric, but, for the other two types ofspecimen, the distributions exhibit an increasingly positive skew (i.e.themass of the distribution is concentrated on the left of the histogram)as θ increases, particularly for the 90° specimens.
Finally, according to the values reported in Table 3, the maximumcentroid eccentricities, as well as their standard deviations, increasewith increasing values of θ. This can be also observed in Fig. 11, wherethe normalised eccentricities along the y and z axes (ey/t and ez/t respec-tively) are plotted against the position along the coupon length for threetypical specimens with θ = 0°, 45° and 90°. As expected, the eccentric-ities measured in the y-direction ey were generally significantly lowerthan those in the z-direction ez.
4. Tensile coupon testing
4.1. Test set-up
In order to obtain the monotonic stress-strain response of theWAAM material, both machined and as-built coupons were subjectedto tensile testing at room temperature. All tests were conducted in com-pliance with EN ISO 6892-1 [45] in the Structures Laboratory of the De-partment of Civil and Environmental Engineering at Imperial CollegeLondon. The cross-sectional area of the machined coupons A was mea-sured using a mechanical micrometer of 0.001 mm accuracy, and stan-dard gauge lengths of 5.65
ffiffiffiA
pwere marked onto their surfaces for the
Fig. 5. Point cloud representation of typical specimen in Geomagic Wrap [40].
calculation of fracture strains εf. The geometric properties of the as-built coupons were determined by laser scanning, as described in theprevious section. Strains at fracture were again determined over thestandard gauge length, calculated based on the mean cross-sectionalarea.
For the machined specimens, two electrical resistance strain gaugeswere attached at the mid-height of each coupon, one on either side, torecord longitudinal strains in the early stages of testing (up to the 0.2%proof stress σ0.2) while, for both the machined and as-built specimens,a four-camera LaVision digital image correlation (DIC) system wasemployed to provide highly accurate measurements of the surfacestrain field on both sides of the coupon.
Prior to testing, the parallel length of all coupons was painted whiteand then spray-painted with a random black speckle pattern, in orderfor the strains to be calculated over the full area of the parallel length.The DIC cameraswere positioned around the specimen, with each cam-era pair monitoring opposite faces, while the angle between the pairs ofcameras was kept at approximately 60° as a compromise between alarger angle to increase the stereo effect and hence the accuracy of thecomputed out-of-plane deformations and a smaller angle to ensurethat the area-of-interest remained in sharp focus. The DIC setup isshown in Fig. 12. The use of DIC is particularly important forWAAMma-terial since the strain field under macroscopic uniaxial loading is not asuniform as for conventionally producedmaterial. The use of single pointstrain gauge measurements, which can only reflect the localised re-sponse of the sample, can lead to inaccurate results.
An overview of the tensile test setup is presented in Fig. 13. A 250 kNInstron 8802 testingmachinewas used for the application of the tensileload, which was measured by a load cell within the actuator. Both axialload and strain gauge measurements were recorded at a frequency of2 Hz using an in-house developed data logger, while the DIC system re-corded the tensile force through an analogue to digital converter and ac-quired images at a frequency of 1 Hz. The acquired images wereprocessed in the software DaVis [46]. A longitudinal strain box wasdrawn over the full parallel length of both sides of each coupon, withinwhich the relative displacements of the surface patterns were analysed.An average stress-strain curve was then produced.
Load was applied using strain control at a strain rate of 0.00007 s−1
prior to the 0.2% proof stress σ0.2, while a strain rate of 0.00025 s-1 wasemployed beyond σ0.2 and until failure, in accordance with EN ISO6892-1 [45]. A gradual transition from the initial to the final strain ratewas achieved by using three intermediate strain rates.
4.2. Results
4.2.1. GeneralStainless steel exhibits a rounded stress-strain response with no
sharply defined yield stress and significant strain hardening. Severalmaterial models have been devised over the years to describe the
Grid points
zx
y
dx
Cross-section at x = xi of area Ai
z
y
Centroid of coupon
Centroid Pi of
cross-sectionti
ey
ezdy
dy
Fig. 6. Typical coupon geometry processed in Rhino 3D [41].
5P. Kyvelou et al. / Materials and Design 192 (2020) 108675
response of nonlinear metallic materials, the vast majority of which arebased on the Ramberg-Osgood expression [47,48]. Nowadays, the mostfrequently employed material model is a two-stage version of the orig-inal expression, as given by Eqs. (1) and (2) [49–52]:
ε ¼ σEþ 0:002
σσ0:2
� �n
; for σ ≤σ0:2 ð1Þ
ε ¼ σ−σ0:2
E0:2þ εu−ε0:2−
σu−σ0:2
E0:2
� �σ−σ0:2
σu−σ0:2
� �mu
þ ε0:2; for σ0:2bσ ≤σu
ð2Þ
where σ and ε are the engineering stress and strain respectively, E is theYoung's modulus, σu and σ0.2 are the ultimate and 0.2% proof stressesrespectively, E0.2 is the tangent modulus of the stress-strain curve atσ0.2, as given in Eq. (3), εu and ε0.2 are the total strains correspondingto the ultimate and 0.2% proof stresses respectively while n and mu arestrain hardening exponents determining the degree of roundedness ofthe stress-strain curve. An alternative two-stage Ramberg-Osgoodmodel, in which the 1% proof stress σ1.0 features in the second stage inplace of the ultimate tensile stress σu, was proposed by Gardner andAshraf [53]. In this model, the strain hardening exponent for the secondstage is denoted m1.0.
E0:2 ¼ E
1þ 0:002nE
σ0:2
ð3Þ
Fig. 7. Deposition p
The two-stage Ramberg-Osgoodmodel given by Eqs. (1) and (2) hasbeen employed herein to describe the measured stress-strain curves ofthe WAAM material. The process by which the key parameters werefitted to the measured data is described below.
Particular attention was given to the determination of the Young'smodulus E for each of the examined coupons; a typical example of theadopted automated process is presented in Fig. 14. First, the numericalderivative of the stress-strain curve was computed. Since the use of a fi-nite difference approach was deemed to be inappropriate as the stress-strain curve is expected to contain experimental noise, amoving regres-sion filter was passed over the curve and, for each point, ordinary leastsquares regression (OLSR) analysis was employed to calculate E. Then,as shown in Fig. 14(b), the calculated values of E were plotted againstthe corresponding values of ε (calculated as the mean strains acrossthe regression window). In Fig. 14(b), five regions can be identified:(i) noise, though barely visible, prior to the start of the test, (ii) a rampup in E as the test begins, (iii) a plateau of constant E, correspondingto the elastic region of the material response, (iv) a ramp down in Edue to the gradual introduction of plasticity and (v) a final plateau cor-responding to the plastic region of the material response. Only region(iii) is required for calculating the value of the Young'smodulus E; how-ever, its identification and isolation in a consistent and automatedman-ner is difficult. To overcome this difficulty, the Haar wavelet transform[54] was used to detect the position of the ramp up and ramp downevents in regions (ii) and (iv) respectively, from which a window con-taining region (iii) (defined in between bands A and B in Fig. 14(b))can be inferred. An S-shaped logistic curve, similar to Richards curve[55], was then fitted to the data between bands A and B using the
w
ath width w.
0.96
0.98
1.00
1.02
1.04
0.0 0.5 1.0 1.5 2.0
Contour spacing dx (mm)
AB-3.5-0-C2
AB-8.0-45-C1
AB-8.0-90-B2
0°
45°
90°0.94
0.96
0.98
1.00
1.02
1.04
1.06
0.0 0.5 1.0 1.5 2.0
Contour spacing dx (mm)
AB-3.5-0-C2
AB-8.0-45-C1
AB-8.0-90-B2
0°
45°
90°
0.99
1.00
1.01
0.0 0.5 1.0 1.5 2.0
Contour spacing dx (mm)
AB-3.5-0-C2
AB-8.0-45-C1
AB-8.0-90-B2
0°
45°
90°
0.96
0.97
0.98
0.99
1.00
1.01
1.02
1.03
1.04
0.0 0.5 1.0 1.5 2.0
Contour spacing dx (mm)
AB-3.5-0-C2AB-8.0-45-C1AB-8.0-90-B2
0°
45°
90°
0.99
1.00
1.01
0.0 0.5 1.0 1.5 2.0
Contour spacing dx (mm)
AB-3.5-0-C2AB-8.0-45-C1AB-8.0-90-B2
0°
45°
90°0.92
0.94
0.96
0.98
1.00
1.02
1.04
1.06
1.08
1.10
0.0 0.5 1.0 1.5 2.0
Contour spacing dx (mm)
AB-3.5-0-C2AB-8.0-45-C1AB-8.0-90-B2
0°
45°
90°
Fig. 8. Results of sensitivity study on contour spacing.
6 P. Kyvelou et al. / Materials and Design 192 (2020) 108675
Levenberg-Marquardt nonlinear least squares regression algorithm[56,57]. The Young's modulus E was then calculated as the value of thefitted logistic curve at the strain associated with band A (point C inFig. 14(b)).
4.2.2. Machined couponsThe full stress-strain curves of the machined tensile coupons are
plotted in Fig. 15(a), while the elastic region and initial yielding areshown more clearly in Fig. 15(b). Note that only the curves derivedusing the DIC data are reported herein since they were almost identicalto these calculated based on the strain gauge measurements and aredeemed to be more accurate. A summary of the average material prop-erties arranged by direction of testing relative to the print layer
Table 3Average geometric properties of the as-built coupons.
tnom (mm) θ (°) A (mm2) t (mm) tmin
ttsdt
3.50 71.3 3.57 0.94 0.03
45 70.1 3.74 0.90 0.0590 74.5 3.76 0.86 0.09
8.00 140.5 7.31 0.98 0.01
45 139.4 7.19 0.95 0.0390 137.5 7.15 0.87 0.09
orientation (i.e. θ = 0°, 45° and 90°) is presented in Table 4, while thematerial properties of all coupons are reported in Table 5. In Tables 4and 5, θ is the direction of testing relative to the print layer orientationas defined in Fig. 2, E is the Young's modulus, σ0.2 and σ1.0 are the 0.2%and 1.0% proof stresses respectively, σu is the ultimate tensile stress, εuis the strain at the ultimate tensile stress, εf is the fracture strain mea-sured over the standard gauge length [45] and n, m1.0 and mu are thestrain hardening exponents of the two-stage Ramberg-Osgoodmaterialmodel.
The results from the three tested orientations demonstrate an inher-ent material anisotropy, as apparent from Fig. 15 and Table 4. It was ob-served that, despite removing the surface undulations prior to testing togive prismatic coupons of nominally the same geometry, the coupon
Fig. 9. Contour plane relative to deposition path for coupons of: (a) 0° and (b) 90°.
7P. Kyvelou et al. / Materials and Design 192 (2020) 108675
surfaces exhibited deformations during testing that matched their builddirection – see Fig. 16. Furthermore, the pattern of the longitudinalsurface strain field clearly remained influenced by their print layer ori-entation; a typical example of this is shown in Fig. 17, where the longi-tudinal surface strain fields of machined coupons tested in the three
0.0
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
0.9
1.0
0.8 0.9 1.0 1.1
Normalised t
Norm
alis
ed f
requen
cy
0.0
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
0.9
1.0
0.8 0.9 1.0 1.1
Normalised t
Norm
alis
ed f
requen
cy
0.0
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
0.9
1.0
0.8 0.9 1.0 1.1
Normalised
Norm
alis
ed f
requen
cy
Fig. 10. Distribution of normalised thi
orientations are plotted just after attainment of their 1% proof stressσ1.0. The reasons behind this are explained in Section 5.
The coupons tested in the 90° orientation had the lowest 0.2% proofstress σ0.2, 1.0% proof stress σ1.0 and ultimate tensile strength σu, whichwould be expected since the material is being loaded perpendicular toits individual layers. The coupons tested in the 0° and 90° directionshad the lowest Young's moduli, with the values observed being approx-imately 30% lower than the typically observed value of E=200,000MPafor stainless steel [49,58], while the coupons tested in the 45° orienta-tion exhibited a significantly higher Young's modulus and slightlyhigher strength than the other tested orientations.
4.2.3. As-built couponsThe full stress-strain curves of the as-built tensile coupons of 3.5mm
and 8.0 mm nominal thickness are plotted in Figs. 18(a) and 19(a) re-spectively, with the early portion of the curves response shown inFigs. 18(b) and 19(b), respectively. The as-built material propertiesare described as ‘effective’ due to their dependence on the variabilityof the built geometry. A summary of the average effectivematerial prop-erties arranged by direction of testing relative to the print layer orienta-tion (i.e. θ=0°, 45° and 90°) and by nominal thickness tnom (i.e. 3.5mmand 8.0 mm) is presented in Table 6, while the effective material
1.2 1.3 1.4 1.5
= 0°
hickness ti/t
1.2 1.3 1.4 1.5
= 45°
hickness ti/t
1.2 1.3 1.4 1.5
= 90°
thickness ti/t
cknesses ti/t of as-built coupons.
Fig. 11. Eccentricity measurements along y axis and z axis of three typical specimens ofdifferent build directions.
Tensile
coupon
Fig. 13. Tensile test setup.
8 P. Kyvelou et al. / Materials and Design 192 (2020) 108675
properties of all 3.5 mm and 8.0 mm as-built coupons are presented inTables 7 and 8 respectively, where the employed notation is the sameas that used in Tables 4 and 5. Note that the material responses of theas-built coupons were determined based on the mean cross-sectionalarea over the parallel length of each coupon, as measured by the 3Dlaser scanning and the subsequent geometric analysis outlined inSection 3. It should be also mentioned that, other than the 0° coupons,the majority of the as-built coupons fractured at the cross-section ofminimum thickness (i.e. cross-section of minimum area), as shown inFig. 20, immediately after attainment of the ultimate stress σu, resulting
Strain gauge
Camera Masked
light
Tensile coupon
with speckle pattern
60°
Aluminium bar used to secure
camera positions
Fig. 12. Plan view of DIC tensile coupon setup.
in the fracture-to-ultimate strain ratios εf/εu being approximately equalto unity.
The results of the tensile tests on the as-built coupons followed asomewhat similar trend to the underlying material properties of themachined coupons. The as-built coupons tested in the 90° orientationhad the lowest effective values of Eeff, σ0.2,eff, σ1.0,eff and σu,eff, with Eeffbeing up to 55% lower than the typically assumed value of E =200,000 MPa for stainless steel. Although the values of strength werenot substantially affected by the thickness of the coupons, the valuesof the Young's modulus in all orientationswere found to be consistentlylower for the thinner material; this can be explained with reference tothe magnitude of the geometric undulations relative to the couponthickness (i.e. the ez/t values), which were higher for the 3.5 mm thanthe 8.0 mm material, reflecting the fact that positional control duringprinting is essentially absolute, rather than relative to the thickness.The higher relative eccentricities induce bending in the coupons duringloading, which has a detrimental effect on the effective materialresponse.
In terms of ductility, the as-built coupons tested in the 0° orientationwere the most ductile due to the load acting parallel to the depositionpaths. Typical as-built coupons are shown before and after testing inFig. 21, where the differences in fracture strain for the different orienta-tions are clearly visible. A post-test inspection of the coupons revealed alocal lack of fusion in the AB-8.0-90°-B4 specimen, which explains thelower strength and ductility relative to the other specimens – seeFig. 19(a).
4.3. Comparison between as-built and machined coupons
In this section, comparisons between the results of the as-built andmachined coupon tests are presented in order to quantify the influenceof the geometric variability that is inherent in theWAAMprocess on thematerial response – see Fig. 22 and Table 9, where the effectivematerialproperties of the as-built coupons are labelledwith the subscript ‘eff’. Asexpected, the as-built specimens exhibited a drop in effective Young'smodulus, 0.2% proof stress, ultimate stress, and ultimate strain (withthe exception of the 8.0 mm 0° coupons on average) when comparedto their equivalent machined specimens, with the 3.5 mm 90° couponsshowing themost severe reductions (up to about 35%, 20%, 20% and 60%reductions in E, σ0.2, σu and εu, respectively). It should be mentionedthat one of themachined coupons (M-8.0-0°-H4) showed lower ductil-ity than the others and that a εu,eff/εu ratio of 1.0 is found if this individ-ual result is excluded.
5. Correlation between microstructure and mechanical properties
Rapid solidification of additively manufactured metals is a commoncharacteristic of all AM techniques and, although this is expected to af-fect the material microstructure and hence the material properties,
Fig. 14. Process followed for the determination of the Young's modulus E.
9P. Kyvelou et al. / Materials and Design 192 (2020) 108675
research into its influence on different metals is not yet comprehensive[38]. Several researchers have observed variation in the mechanicalproperties of additively manufactured metals and linked this variationto differences in the internal microstructure [20,59–63]. Hence, inorder for the anisotropicmechanical performance of the tested couponspresented herein to be examined from ametallurgical point of view, themicrostructure and crystallographic texture of the printed material hasbeen investigated and compared to that of the parentweldingwire. The
Fig. 15. Stress-strain curves of machined coupons: (a) full range and (b) initial range.
experimentswere carried out at theHarvey Flowers EM Suite in the De-partment of Materials at Imperial College London.
5.1. Employed techniques and methods
In order to examine the microstructure of the investigated WAAMplates, samples for metallographic examination were extracted fromall surfaces of a typical plate, with the axes LD, TD and ND being parallel(i.e. θ= 0°), transverse (i.e. θ= 90°) and normal to the print layer ori-entation, respectively – see Fig. 23.
In preparation for metallographic analysis, grinding and polishing ofall sampleswas performed in linewith ASTM: ACI 301 and 318 [64]. Thespecimens were polished using magnetic polishing pads and dilutedOP-S lubricant, with ultrasonic cleaning undertaken at standard time in-tervals to remove any residual OP-S particles. Following polishing, allsamples were etched to reveal their microstructure, providing informa-tion about the grain size and shape. Since this process relies on thechemical reaction between the metal surface and the etchant, selectionof an acid solution appropriate for austenitic stainless steel is crucial.Since the use of glycerin reagent, recommended by ASTM E407-07[65], was not effective in revealing the microstructure of the examinedmaterial, an acid solution proposed by Fernandes de Lima and Sankare[66] was employed.
The crystallographic texture of the sampleswas examined bymeansof electron backscatter diffraction (EBSD) using pole figures. EBSD is anadvanced technique that can provide qualitative and quantitative mi-crostructure measurements such as the distribution of grain sizes,grain shape and crystal orientation [67], while the resulting pole figureis a 2D stereographic projection revealing the locations and intensitiesof specified average crystal orientations relative to the examined sur-face. Since the quality of the EBSD diffraction patterns is quite sensitiveto the surface quality of the top layer of the examined sample, all spec-imens were prepared until there were no visible scratches or contami-nation under an optical microscope. EBSD analyses were thenperformed at 20 keV on a Quanta Field Emission Gun (FEG) 650 scan-ning electron microscope equipped with a Bryher eFlash HD andanalysedwith Esprit 2.1. Themorphology of themicrostructurewas im-aged using FSD imaging [68].
5.2. Results
5.2.1. Polishing and etchingAfter 30 min of polishing, the manufacturing characteristics of
WAAM were apparent on the surface morphology of all samples, withthe boundaries between the welding layers being clearly distinguish-able, revealing a preferential growth orientation of the grains along
Table 4Average material properties of machined coupons by direction of testing relative to the layer direction.
θ (°) E (MPa) σ0.2 (MPa) σ1.0 (MPa) σu (MPa) εu εf n m1.0 mu
10 P. Kyvelou et al. / Materials and Design 192 (2020) 108675
the TD direction; this observation is consistent with a study carried outby Liverani et al. [21], in which the grains were found to be orientatedpreferentially along the path with the greatest thermal gradient.
equipped within the scanning electron microscope (SEM) were used
(b) 0° coupon after testing
(d) 90° coupon after testing
ed coupons before and after testing.
0.025
0.020
0.015
0.010
0.005
0.000
(c) 90°
0.025
0.020
0.015
0.010
0.005
0.000
(a) 0° (b) 45°
0.025
0.020
0.015
0.010
0.005
0.000
Fig. 17. Typical longitudinal surface strain fields of machined coupons at their 1% proof stress σ1.0.
11P. Kyvelou et al. / Materials and Design 192 (2020) 108675
to generatemicrostructure images of the specimens, revealing themor-phology of the material, where different grains are illustrated with con-trasting colours. A typical FSD image (corresponding to a sample taken
0
100
200
300
400
500
600
0.0 0.1 0.2 0.3
ssertS
σ)a
PM(
Strain ε
AB-3.5-0-A1
AB-3.5-45-B3
AB-3.5-90-D1
θ = 0°
θ = 45°
θ = 90°
0
100
200
300
400
0.000 0.001 0.002 0.003 0.004 0.005
ssertS
σ)a
PM(
Strain ε
AB-3.5-0-A1
AB-3.5-45-B3
AB-3.5-90-D1
θ = 0°
θ = 45°
θ = 90°
E = 200000 MPa
(b) Initial range
(a) Full range
Fig. 18. Stress-strain curves of as-built coupons of 3.5 mm thickness.
from the top surface of the WAAM plate shown in Fig. 23) is illustratedin Fig. 24(a) while the corresponding inverse pole figure map is shownin Fig. 24(b), where the employed colours indicate the dominant crystal
Fig. 19. Stress-strain curves of as-built coupons of 8.0 mm thickness.
Table 7Effective material properties of as-built coupons of 3.5 mm nominal thickness.
12 P. Kyvelou et al. / Materials and Design 192 (2020) 108675
orientation per grain (as also illustrated by the overlayed unit cube ori-entation). Long and columnar grains can be observed, preferentiallyaligned along the TD (i.e. θ=90°) axis. This alignment can be explainedwith reference to the arc welded plate being built layer by layer alongthe TD axis since, during solidification of most metals, grains tend totrack the highest thermal gradient [20]. Typical b100N, b110N andb111N pole figures of the examined surface are shown in Fig. 25,where the scale bar denotes the intensity of specific crystal orientationdistributions in comparison with a random distribution; the pole figurepositions corresponding to the coupons at θ = 0°, 45° and 90° to theprint layer orientation (as defined in Figs. 2 and 23) are marked onthe reference frame. According to the b100N polefigure, theweldedma-terial has a strong crystallographic texture, with a higher intensity ofb100N type directions parallel to the LD (i.e. θ = 0°) and TD (i.e. θ =90°) directions. This indicates that there is a higher proportion of crys-tals, than expected from a random texture, that are being deformed intension along the b100N type directions (i.e. the family of [100], [010],[001], [�1 00], [0�1 0] and [00�1 ] directions) in the θ=0° and θ=90° ten-sile coupons. From the b110N pole figure, it can be seen that for the θ=45° coupons, a higher proportion of crystals are being deformed in the
Table 8Material properties of as-built coupons of 8.0 mm nominal thickness.
b110N type directions (i.e. the family of [110], [011], [101], [�1 10], [0�11] and [�1 01bar101] directions). The implications of this are explainedin Section 5.3.
5.3. Variation in elastic modulus
Based on the obtained results, it has been concluded that the exam-ined WAAM austenitic stainless steel has a strong crystallographic tex-ture. Hence, in order to interpret the macroscopic elastic performanceof this textured material, knowledge of the single crystal elastic moduliis required [69]. The Young's modulus of a single crystal in a specific di-rection depends on the number and strength of its interatomic bonds.The theoretical and experimentalmonocrystal elastic stiffness constantsC11, C12 and C44 of Grade 304 austenitic stainless steel were provided ina studymade by Ledbetter [70]. Hence, for the calculation of the mono-crystal elastic moduli, since the elemental composition of Grade 308LSistainless steel is similar to that of Grade 304 stainless steel, the singlecrystal elastic-stiffness constants of Grade 304 were adopted and usedin conjunction with a set of formulae established by Armstrong et al.
Fig. 20. Typical coupon with fracture location coinciding with the cross-section ofminimum thickness.
13P. Kyvelou et al. / Materials and Design 192 (2020) 108675
[71]. The Young's modulus in the b100N type directions, correspondingto the θ = 0° and θ = 90° coupons was calculated to be 91.6 GPa. TheYoung’s modulus in the b110N type directions, corresponding to the θ=45° couponswas calculated as 195.5GPa. Thesefindings verify the re-sults of the tensile coupon tests, according to which the value of theYoung's modulus for the θ = 45° coupons was higher than the equiva-lent values for θ=0° and θ=90°− see Table 4. It should bementionedthat a greater value of Young’s modulus of 314.2 GPa was calculated for
θ = 0°
θ = 90°
θ = 45°
Fig. 21. Elongation of typical as-built coupons of different orientations at fracture.
the b111N type directions, which were not sampled by the conductedcoupon tensile tests.
5.4. Variation in yield and ultimate strength
Although to a lesser extent than the Young's modulus, the yield andultimate strengths were also observed to be anisotropic, with the 0.2%proof strength σ0.2 and ultimate strength σu of the θ = 45° couponsbeing higher than the equivalent values for the θ=0° and θ=90° cou-pons – see Table 4. This variation is attributed to a difference in the ef-fective mean free path for dislocations to travel in the differentdirections, where grain boundaries limit the dislocation motion andact as barriers in so called ‘Hall-Petch’ type strengthening mechanism[72–74].
5.5. Comparison with parent feedstock wire
In order to compare themicrostructures of the parent feedstockwireand theweldedmaterial, the cross-section and side face of the feedstockwirewere examined in the SEM. The FSD image of the cross-section andside face of the wire are presented in Fig. 26(a) and 26(b), respectively,while the b100N, b110N and b111N pole figures are shown in Fig. 26(c).It can be observed that the microstructure of the wire is entirely differ-ent from that of the arc welded sheet, with the feedstock material hav-ing a small grain size and a common wire crystallographic texture inwhich the grains are elongated along the rolling direction. The b100Nand b110N polefigures demonstrate an intense crystallographic texture,with the grains within the wire stretched along the extrusion directionfollowed during the forming process. Hence, it may be concluded thatthe microstructure of the WAAMmaterial is substantially altered due tothe printing process, resulting in anisotropic mechanical characteristics.
6. Correlation between effective mechanical properties and as-builtgeometry
With the results of the tensile coupon tests performed on the as-built material showing a clear dependence of the effective mechani-cal properties on the geometric variability, expressions to describethe correlation between the effective mechanical response and thegeometric characteristics of WAAM material are sought in thissection.
Of all the geometric parameters determined from laser scanningof the coupons and examined in Section 3, the mean and standarddeviation values of the coupon thickness (t and tsd respectively)were found to be the most influential and suitable for establishinga correlation between effective mechanical properties and geometricvariability. One additional simple geometric parameter, denotedvmax20, was also introduced to enable effective mechanical proper-ties to be estimated without the reliance on laser-scanning technol-ogy. The parameter vmax20 is defined as the maximum deviationbetween the examined WAAM surface and a straight edge (e.g. aruler of approximately 1 mm thickness) parallel to the direction ofloading (i.e. along the length of the coupon), as shown in Fig. 27(a), and measuring the maximum distance between the straightedge and the coupon over a 20 cm length, as shown in Fig. 27(b).This measurement should be undertaken using a slip gauge at a se-ries of points over the surface.
Ordinary least squares regression analysis was employed tofit linearfunctions to the test data, providingpredictions for the key effectivema-terial properties (i.e. Eeff, σ0.2,eff, σ1.0,eff, σu,eff and εu,eff). The results of theregression analysis are graphically illustrated in Fig. 28, where the ex-amined as-built coupons are grouped by their angle θ relative to theprint layer orientation, while the effective material properties of eachas-built coupon are normalised by the material properties of the corre-spondingmachined coupon (as reported in Table 4) and plotted againstthe examined geometric parameter (namely the standard deviation of
Fig. 22. Decreases in key material properties of as-built material relative to machined material.
Table 9Decreases in material properties of as-built relative to machined material.
14 P. Kyvelou et al. / Materials and Design 192 (2020) 108675
thickness tsd or the vmax20 measurement) normalised by the averagecoupon thickness t. Note that in the absence of laser scanning data,the average coupon thickness would have to be estimated based onweight or by taking a series of individual calliper measurements.
The devised expressions are presented in Eqs. (4) to (13), while thecoefficients of determination R2 are reported on the respective graphs –
LD
(θ = 0°)
TD
(θ = 90°)
ND
Layer orientation
Fig. 23. Position and labelling of sam
see Fig. 28. Note that all the devised expressions were forced throughthe point (0,1) since the effective and underlying material propertiesshould be equal when tsd or vmax20 are zero.
EeffE
¼ 1−2:9227tsdt
� �ð4Þ
EeffE
¼ 1−0:2598v max20
t
� �ð5Þ
σ0:2;eff
σ0:2¼ 1−2:4960
tsdt
� �ð6Þ
σ0:2;eff
σ0:2¼ 1−0:2375
v max20
t
� �ð7Þ
σ1:0;eff
σ1:0¼ 1−1:7998
tsdt
� �ð8Þ
θ = 90°
θ = 0°
θ = 45°
Layer orientation
TD
ND LD
ples from typical WAAM plate.
Fig. 24. (a) FSD image and (b) inverse pole figure map of examined surface, with colours presented with respect to LD.
<100> <110> <111>Max:
5.2Max:
14Max:
6.5
Min:
0.0075
Min:
0.0072Min:
0.033
2.5
2.0
1.5
1.0
0.5
θ=90°
θ=45°
θ=0°
Reference frame
LD ND
TD
Fig. 25. Pole figures of examined surface.
(a) Cross-section (b) Side face
(c) Pole figures
<100> <110> <111>
2.5
2.0
1.5
1.0
0.5
Min:
0.34Min:
0.33
Min:
0.43
Max:
2.5
Max:
2.1
Max:
4.6
8 μm 80 μm
Fig. 26. SEM examination of feedstock wire: (a) FSD image of cross-section, (b) FSD image of side face and (c) pole figures.
15P. Kyvelou et al. / Materials and Design 192 (2020) 108675
Measurements performed
at a series of points across
coupon width
vmax20 Straight edge
WAAM coupon
(b)
(a)
Fig. 27. vmax20 measured between WAAM surface and a straight edge: (a) overall setupand (b) vmax20 measurement.
16 P. Kyvelou et al. / Materials and Design 192 (2020) 108675
σ1:0;eff
σ1:0¼ 1−0:1725
v max20
t
� �ð9Þ
σu;eff
σu¼ 1−2:3231
tsdt
� �ð10Þ
σu;eff
σu¼ 1−0:2065
v max20
t
� �ð11Þ
εu;effεu
¼ 1−6:7470tsdt
� �ð12Þ
εu;effεu
¼ 1−0:5824v max20
t
� �ð13Þ
The derived models were found to be capable of yielding rea-sonable predictions of the effective mechanical properties of thestudied as-built WAAM material, and can be further verified in fu-ture work.
7. Conclusions
An experimental investigation, conducted to obtain insight into themechanical behaviour of WAAM sheet material, has been presented.Mechanical properties were determined by means of tensile testingand correlation with the underlying microstructure was explored. As-built and machined coupons were tested to investigate the influenceof the geometrical irregularity, which is inherent to theWAAM process,on the effective stress-strain characteristics, while material anisotropywas examined by testing coupons produced at 0°, 45° and 90° to theprint layer orientation. Advanced non-contact measurement tech-niques – laser scanning and digital image correlation –were employed
to determine the geometric properties and deformation fields of thespecimens.
The conducted coupon tests revealed substantial material an-isotropy, with the direction of loading relative to the print layerorientation having a strong influence on the stress-strain charac-teristics. The 45° coupons exhibited the highest values of Young'smodulus and strength while the 90° coupons presented the lowestproperties since the material was loaded across its individuallayers. The as-built specimens showed an effective decrease inYoung's modulus and strength when compared to the machinedspecimens due to the influence of the geometrical undulations;the most severe reductions were observed for the thinner materialand for loading applied perpendicular to the layer orientation (i.e.θ = 90°).
The microstructures of the printed material and its parent metalwire were found to be substantially different, with the printed materialexhibiting a strong crystallographic texture, attributed to rapid solidifi-cation due to theWAAM process, with grains tracking the highest ther-mal gradient. Finally, since the results of the tensile coupon testsshowed a direct dependence of the mechanical properties on the geo-metric variability of the WAAM material, predictive expressions corre-lating key effective material properties to the irregular geometry weredevised.
The authors declare that they have no known competing financialinterests or personal relationships that could have appeared to influ-ence the work reported in this paper.
Acknowledgements
This experimental programme was possible thanks to funding andsupport from the Data Centric Engineering programme at the Alan Tu-ring Institute (ATI), funded by the Lloyd's Register Foundation. This re-search also benefitted from EPSRC funding under grant number EP/R010161/1 and from the UKCRIC Coordination Node EPSRC grant underEP/R017727/1. The authors would like to acknowledge the contributionof MX3D for building the specimens and Gordon Herbert, Yunong Tan,WingWan and Alfredo Olivo at Imperial College London for their assis-tance in the experimental programme. Dafni Daskalaki Mountanou andT. Ben Britton would like to thank Shell Global Solutions for financialsupport. Electron microscopy was conducted in the Harvey Flowers Mi-croscopy Suite on an instrument funded by the Shell AIMS UTC. T. BenBritton would like to thank the Royal Academy of Engineering forfunding his research fellowship.
Data availability
The raw/processed data required to reproduce these findingscannot be shared at this time as the data also forms part of an ongo-ing study.
Fig. 28. Graphical illustration of devised expressions correlating material properties with geometric variability.
17P. Kyvelou et al. / Materials and Design 192 (2020) 108675
18 P. Kyvelou et al. / Materials and Design 192 (2020) 108675
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