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Viscoelastic properties of fused filament fabrication parts
Additive Manufacturing 28 (2019) 704–710
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Additive Manufacturing journal homepage: www.elsevier.com/locate/addma
Viscoelastic properties of fused filament fabrication parts a,⁎
a
a
T b
José Luis Colón Quintana , Alec Redmann , Gerardo A. Mazzei Capote , Angel Pérez-Irizarry , Abrahan Becharaa, Tim A. Osswalda, Roderic Lakesc a b c
Polymer Engineering Center, Department of Mechanical Engineering, University of Wisconsin-Madison, Madison, WI, 53706-1691, USA Department of Civil & Environmental Engineering, University of Wisconsin-Madison, Madison, WI, 53706, USA Department of Engineering Physics, University of Wisconsin-Madison, Madison, WI, 53706, USA
A R T I C LE I N FO
A B S T R A C T
Keywords: Viscoelastic DMA Ultrasonic Loss tangent FFF
Parts made by fused filament fabrication differ in their mechanical properties from the parent material. To investigate the effect of the manufacturing process on the mechanical properties of 3D-printed parts, a series of experiments including Dynamic Mechanical Analysis (DMA) and ultrasonic wave propagation were conducted. For this purpose, printed parts were made from custom ABS filament and were printed using a rectangular bead shape to minimize porosity. The main properties investigated included the elastic, loss and storage moduli, and the material loss tangent (tan δ). Results indicate that the elastic modulus of the printed material was somewhat lower than that of the parent material, about 2 GPa for frequencies 0.1 Hz–100 Hz. Furthermore, tan δ was largest for the parent material at about 0.03 compared with 0.01–0.02 for the printed material. Ultrasonic longitudinal wave measurements at 1 MHz on printed specimens with bead angles of 0°, 45° and 90°, revealed minimal anisotropy and, consistent with DMA results, tan δ was also largest for the parent material.
1. Introduction Fused Filament Fabrication (FFF) has been on the forefront of Additive Manufacturing (AM) due to low costs and broad availability of machines. Furthermore, FFF relies on thermoplastic materials exhibiting pronounced viscoelastic behavior. However, parts produced by FFF technology tend to be highly anisotropic, due in part to the junction formed between adjacent beads that behaves akin to a non-optimal, thermally-driven polymeric weld [1]. These bead junctions tend to be weaker and more compliant than the solid polymer. Moreover, the introduction of material discontinuities that stem from the junction of elliptical plastic strands, results in parts containing voids that can act as stress concentrators [2]. It is well known that the final mechanical properties of FFF parts are extremely sensitive to build parameters -such as bead orientation, build temperature, air gap, solidity ratio, among others [3–7]. These factors imply that understanding the mechanical performance of FFF parts is a daunting task, yet extremely important if final user grade performance is to be expected. Unfortunately, research about viscoelastic behavior of FFF parts is scarce. Hence, the work reported herein aimed to understand the effect of bead orientation on the viscoelastic properties of ABS FFF parts, through a variety of dynamic loading experimental techniques, i.e., ultrasonic testing and Dynamic Mechanical Analysis (DMA).
⁎
The storage and loss moduli, and the phase angle (δ) were measured and compared between each raster orientation and benchmarked against a sample of the parent material. 2. Materials and procedure 2.1. Material The first step of the experimental work involved development of a custom thermoplastic filament for the FFF process. The reasoning behind this decision was two-fold. First, the use of an off-the-shelf, commercial thermoplastic filament generally does not guarantee that two different spools were produced under the same processing conditions, or even using the same parent material. Secondly, the results from Koch et al. [4] show that fluctuations in the filament diameter have an impact in the mechanical properties of FFF parts due to improper volumetric output at the nozzle. The Cycolac®️ MG94 material produced by SABIC corporation [8] was chosen for this work. This is an Acrylonitrile Butadiene Styrene (ABS) based material traditionally used for injection molding thin walled parts, as well as extrusion of FFF filament. With a reported Melt Flow Index (MFI) of 11.7 g/10 min, it is an ideal material for both the FFF and extrusion processes [8]. The Melt Flow Index is defined as the
Corresponding author. E-mail address: [email protected] (J.L. Colón Quintana).
https://doi.org/10.1016/j.addma.2019.06.003 Received 19 February 2019; Received in revised form 3 May 2019; Accepted 6 June 2019 Available online 15 June 2019 2214-8604/ © 2019 Elsevier B.V. All rights reserved.
Additive Manufacturing 28 (2019) 704–710
J.L. Colón Quintana, et al.
Fig. 1. Extrusion Setup. Taken from [9].
parameter from 0.63 to 1.0. It was shown that by using an extrusion factor of 0.97 and a layer height of 0.1 mm, it was possible to reach 98% of the strength of an injection molded part subjected to tensile load. Pfeifer et al. [7]. performed experiments varying the EF from 0.75 and 1.08. By performing Computer Tomography (CT) scans and mechanical testing of dogbone samples, it was demonstrated that lower solidity ratios, stemming from lower EF, resulted in poor mechanical properties bearing 24–52% of the highest solidity ratio. Previous research on the mechanical performance of FFF have varied the bead geometry of the samples by assigning positive and negative air gap values to the printing parameters [3,11,12], as shown in Fig. 3. By allowing air gaps, the morphology of the beads and the voids is changed in a manner similar to using EF manipulation: a positive air gap produces beads of more rounded nature, while a negative air gap reduces the void content, being comparable to using an EF approaching 1. It was shown that anisotropy was significant, as the mechanical properties of the samples varied considerably between raster orientations, for all air gap configurations. However, a drastic change in mechanical properties was observed between analogous samples of the same bead orientation, but different air gap values: the failure stress of samples with negative airgaps was considerably higher than samples with null or positive air gaps -with the latter having extremely low values: only between 17–45% of its corresponding air gap sample. To minimize the problems associated with using lower EFs, this parameter was fixed to a value of 1, producing beads with a theoretical rectangular cross section exclusively. It should be noted that due to variations in the diameter of the filament and imperfections in the build process, the final bead geometry is not a perfect rectangle, and the junction of contiguous beads still produces voids characteristic of FFF parts.
mass extruded through a capillary in a 10 min period while applying a constant load. By having a material with a high MFI, it allows more material to be printed, which results in faster printing time. The MG94 material was extruded in a setup that consisted of a single screw extruder (Extrudex EDN 45X30D, Germany) with a 45 mm screw diameter and a screw length to screw diameter ratio (L/D) of 30. The hot melt was extruded at 205 °C through a circular die with a 5.8 mm diameter, and then guided through a pre-skinner into a vacuum-assisted, heated water bath (Conair, USA) to cool the extrudate whilst minimizing void formation. The solidified filament was then passed through a 3-axis laser micrometer (LaserLinc, USA) and a belt puller (Conair, USA) configured in a control loop. The dimensions of the filament were controlled by automatically adjusting the speed of the puller if the readings from the micrometer were out of specification, in this case a diameter of 2.85 mm with a tolerance of ± 0.02 mm. Finally, the product was wound onto spools using a filament winder. A schematic of the extrusion setup can be seen in Fig. 1. Prior to any usage in a printer, the filament was dried in a silo (Novatec, USA) at 82 °C for 3 h. Since the ABS material is amorphous and the temperature at which it is subjected is under the glass transition temperature, no significant thermal or structural change is induced in the material. 2.2. Tool path generation for FFF specimens All test specimens were produced using a TAZ 5 (Lulzbot, USA) FFF printer with a 0.5 mm nozzle. Tool pathing, a set of printing commands that tells the printer where to move, how fast it moves, and how much material need to extrude, was performed using SciSlice: a customized slicing engine developed in-house [9]. A conscious effort was made to maintain as many print parameters as possible constant across all prints. These slicing engine parameters are listed in Table 1. Note that Extrusion Factor (EF) is a user parameter that represents a ratio of the true area occupied by the cross section of a bead (Abead) divided by the product of the bead width (Wbead) with the layer height (Hlayer). This factor is used by SciSlice to calculate the length of filament required to print a bead of known dimensions by using Eq. (1).
L filament *Afilament = EF*Wbead*Hlayer Lbead
2.3. DMA procedure Dynamic Mechanical Analysis (DMA) tests were conducted using the tensile configuration of the RSA 3 (TA Instruments) DMA at ambient temperature. Three rectangular, single layer samples were produced with a thickness of 0.2 mm and dimensions of 35 mm by 6 mm. The samples were manufactured in three bead orientations: 0°, 45°, and 90° where the angle is measured parallel to the load direction, as shown in Fig. 4. Additionally, a sample of the parent material was compression molded with the same geometry and used for comparison. The parent material will be used to compare parts used in FFF applications against parts manufactured with other processes where the final product is a solid part, without induced anisotropy by the bead orientation. Subsequently, it is assumed that parts manufactured using these processes will have the same mechanical response as the compression molded sample. The same comparison was done for samples of the ultrasonic test. Strain sweeps were conducted on all samples at 1 Hz from strains of 10−3–100 percent with five points each decade to identify the range for linear viscoelastic behavior. Once the linear viscoelastic region was identified, frequency sweeps were conducted using the linear viscoelastic strain percentage as the strain amplitude. This value was used assuming the material will remain in the linear viscoelastic region through the frequency sweep. Frequencies from 0.01 Hz to 100 Hz were used for the DMA test.
(1)
The extrusion factor controls the amount of material extruded through the volumetric balance shown in Eq. (1). For example, an EF of π/4 would produce beads of elliptic cross section, while an EF of 1 would produce beads of rectangular nature. Beads of circular or square cross section are possible if the layer height and the bead width have equal dimensions. See Fig. 2 for a visual representation of the impact of this parameter on the geometry of printed beads. Previous research by Koch et al. [4] showed the effect of the Extrusion Factor upon the mechanical properties of FFF, by varying this Table 1 Constant slicing engine parameters. Parameter
Value
Parameter
Value
Extrusion Factor (EF) Filament diameter Path width Layer height Printing temperature
1 2.85 mm 0.5 mm 0.2 mm 230 °C
Bed temperature Print speed Rapid speed Retraction distance Retraction speed
100 °C 2000 mm/min 4000 mm/min 1.5 mm 500 mm/min
2.4. Ultrasound procedure Cubic specimens of 30 mm were produced as described in the 705
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Fig. 2. Sets of beads produced with an extrusion factor EF of a) π/4 and b) 1. Image c) are micrographs of cross sections of printed parts [10]. Fig. 3. Air gap application. Adapted from Rayegani et al. [11].
preceding discussion, with bead orientations of 0°, 45° and 90° with respect to the direction of ultrasonic wave propagation. For comparison, a cylindrical sample of parent material with a 21.8 mm diameter and a length of 20.6 mm was manufactured using a heated chamber, where MG94 pellets were melted and pushed to a cylindrical mold by a pneumatic piston. A processing temperature of 240 °C was used for melting the material. Once the heaters were turned on, 11 min were given to allow the system to reach the target temperature and properly melt the material. A pressure of 689 kPa (100 psi) was used to actuate the piston and make the polymer flow towards the cavity, filling the mold. For more information of the manufacturing procedure of the sample used refer to [13]. To ensure a proper contact with the ultrasonic transducer, the surface of all samples was finely polished using a rotary metallographic polisher with abrasive discs of grit sizes of 800 and 1200. The experimental setup consisted of FFF produced cubes with three bead orientations and cylindrical samples of the parent material, each held in between two longitudinal 1 MHz ultrasonic transducers (Panametrics V102) connected to a pulse generator (Panametrics 500PR). An oscilloscope (Siglent SDS 1052DL) was connected to the pulse generator to monitor the pulse wave as it traveled through the specimen. To stabilize the signal, a weight of mass 3.8 kg was added on top of the uppermost transducer. A schematic of the experimental setup is shown in Fig. 5. To ensure proper contact with the transducers and a stronger signal on the
Fig. 4. DMA sample orientation. Angle was measure with respect to the loading direction.
Fig. 5. Schematic of the ultrasound experimental setup. 706
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Table 2 Density of parent material and bead angle.
Table 3 μ-CT Scan parameters.
Material
0°
45°
90°
PM
Parameter
Value
Parameter
Value
ρ (kg/m )
1036.31
1021.13
1032.06
1032.84
Voltage (kV) Current (A) Integration time (s) Gain
70 110 1 8.0 x
Image averaging Binning Number of projections X-ray intensity correction
Off 1×1 1800 9000–10,000
3
oscilloscope, a thin layer of glycerin was used as a coupling agent. For this series of tests, four sets of 30 mm cubes were printed for the bead orientations of 0° and 45° for a total of 8 samples. The samples at 0° were used for the 0° and 90° test. Each material sample was measured and weighed using a caliper and a high-precision scale (OHAUS®, Explorer EX125D), respectively, to calculate the density of each sample tested as shown in Table 2. Note that the parent material sample did not possess the highest density. The low parent material density is attributed to the material processing, particularly the creation of voids in the samples as ABS was injected into the mold. Since an extrusion factor of 1 was used for the 3D printed parts, the density of the 0° and 90° samples were close to the parent material. The low parent material density is believed to be due to a higher volume of voids in the parent material than the 0° and 90° samples. In contrast, the sample with bead angle of 45° had the lowest density. The slightly lower density could be a result of the tool path required to produce beads in 45°. The specimen length and the wave propagation time were used to compute the wave velocity; then the wave velocity (vl) and the material density (ρ) were used to compute the tensorial modulus as shown in Eq. (2), where C1111 is the modulus component in the wave/loading direction.
C1111 = (VI)2*ρ
3. Results and discussion 3.1. DMA results DMA measurements were performed upon printed material at three different orientations as well as the parent material. Fig. 6 shows the complex modulus (E*) and the tan δ values as a function of the strain percentage. As previously mentioned, a strain sweep was performed on the samples to determine the proportionality limit strain for each bead orientation and the parent material. Table 4 shows the approximate linear viscoelastic strain limit percentage for each orientation. The criteria used for the selection of these values was the strain limit percentage at which the tan δ is not strain dependent. Since the tan δ give the viscoelastic response of the material, it considers the change of the storage and loss modulus. The increase in complex modulus shown in Fig. 6 is due to the nature of the additive manufacturing process of bead placed next to each other. The creation of voids between bead and in some occasions the non-contact between beads induces a lower modulus at the lower strain percentages. As the material starts to elongate, the beads come closer together increasing the resistance to movement, thus increasing the modulus. Higher tan δ values coincide with a response regime where the loss modulus was predominant; however, increases in loss modulus can indicate yield or nonlinearity prior to yielding. The exact cause of the observed response for tan δ values was not determined in this study. Once the strain percentage was determined as previously shown in Table 4, a frequency sweep from 0.01 Hz to 100 Hz was performed on the samples. Fig. 7 shows the result for the frequency sweep. As observed below, the storage modulus increases with increasing frequency while the loss tangent does not differ much in the range 0.1 Hz–100 Hz. Such behavior is typical in the glassy regime [14]. The tan δ values for the printed parts falls within 0.01 and 0.02 while the parent material possesses a value of 0.03 in this frequency range. Values of storage modulus and tan δ found in literature coincides with experimental data acquired for both compression molded samples and samples made with different bead orientations [15,16]. The storage and loss moduli for all bead orientations were comparable. However, in the case of the parent material, the loss modulus was considerably higher than the 3D printed samples. This results in a tan δ trend that indicates higher damping capabilities for the parent material. This behavior, although counter-intuitive, can have its root in multiple factors, including the presence of voids in the parent material specimens. Another potential contribution could have stemmed from the single layer thickness of the printed samples, and the bed temperature being close to the glass transition temperature (Tg) of the material used: as shown in Table 1, the bed temperature was 100 °C, while the glass transition temperature for ABS is 103.8 °C [17]. Since these temperatures are very close to each other, it is suggested that heat transfer through the thickness of the sample is sufficient to facilitate proper bonding of the beads. Also, due to the nature of the manufacturing process, bead overlap could have resulted from the toolpath, thus increasing the strength of the printed samples in all orientations.
(2)
To determine the loss tangent of the FFF and injection molded ABS samples, the amplitude of the stress wave after traveling through the material was measured using the oscilloscope. After the first measurement, the samples were cut into smaller lengths and new measurements were taken to quantify the attenuation of the ultrasonic waves (Eq. (3)). Three lengths were used to determine the stress wave attenuation.
A1 = A 0*e−αz
(3)
With the measured wave amplitudes, the attenuation parameter (α) was determined by fitting an exponential curve to the wave amplitudes as a function of traveled distance. Furthermore, attenuation coefficients were also calculated by relating the measured amplitudes and the traveled distances as shown in Eq. (4). A
α=
ln( A1 ) 2
(Z2 − Z1)
(4)
where Ai and Z i are the measured wave amplitude and corresponding wave travel distance. Finally, the average attenuation coefficient was used to determine the loss tangent (tan δ) for the 3D-printed samples and parent material using Eq. (5).
2VI*α V *α = I ≈ tanδ ω π*f
(5)
where ω is the angular frequency in radians per second and f is the ordinary frequency in Hertz. 2.5. Computed tomography Objects produced using FFF tend to have process induced defects that stem from either the toolpath used during part production, machine limitations or material inconsistencies. Given that the ultrasonic testing is particularly sensitive to defects in the test geometry, computed tomography was used to obtain images of the internal structure of the cubic samples produced for the ultrasonic measurements. A Zeiss Metrotom 800 (Zeiss, Germany) was used to scan the parts, using parameters shown in Table 3 to obtain high quality images. 707
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Fig. 6. Viscoelastic properties as a function of strain percentage of parent material and material printed at angles 0°, 45°, and 90°.
of the filament necessary to prevent dripping of material, was not enough and thus undesired material was added to the part as the print head moved above the object to start a different layer. Note that both images reveal voids. These are introduced every time there is a bead or layer junction due to the circular nozzle orifice of the printer. The two first problems were corrected through manipulation of the slicing engine parameters, but the presence of voids is characteristic to the FFF printing process and thus cannot be avoided. Small voids in FFF parts can be minimized but not eliminated. Due to the cylindrical shape of the nozzle, the beads were placed layer by layer with an elliptical or rectangular shape as shown in Fig. 2. The shape of the bead depends on the extrusion factor, volumetric flow rate of the material, and nozzle diameter to name a few.
Table 4 Strain limit percentage for linear viscoelastic region as a function of bead orientation. Orientation
Approximate Linear Viscoelastic Strain Limit (%)
0° 45° 90° PM
3 × 10−1 5 × 10−2 1 × 10−2 6 × 10−2
3.2. Computed tomography results As mentioned in the previous section, computed tomography was used to obtain images of the internal structure of the cubic samples produced for the ultrasonic measurements. Fig. 8 shows on the left (Fig. 8a) a sample printed using a raster orientation of 45°. Due to bias on the tool pathing algorithm, the raster was printed in a discontinuous manner, thus leaving a sizeable defect that ran across the entire height of the sample. On the right (Fig. 8b), filament retraction, the movement
3.3. Ultrasonic results The stress wave velocity and the density of the samples were used to determine the tensorial modulus along the stress wave path. As result of the manufacturing process of 3D printed parts, voids are commonly found in the final part [18] which resulted in peaks when measuring the
Fig. 7. Viscoelastic properties as a function of frequency of parent material and material printed at angles 0°, 45°, and 90°. 708
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Fig. 8. Printed sample with (a) no excess material and (b) excess material. Scale bar, 5 mm.
Fig. 9. Average Moduli at 1 MHz of parent material and 3D-printed ABS Parts.
Fig. 11. Average Loss Tangent at 1 MHz for parent material and 3D-printed ABS Parts. Table 5 Summary of ultrasonic results at 1 MHz. Bead Angle
(VL) avg (m/ s)
(C1111) (GPa)
0° 45° 90° PM
2190 2182 2177 2186
4.91 4.80 4.85 5.06
a
avg
E (GPa)
αavg (neper/ mm)
(tanδ)
3.65 3.57 3.60 3.76
0.0222 0.0347a 0.0226 0.0659a
0.0155 0.0244a 0.0157 0.0459a
avg
Only one sample used for calculation.
the calculated moduli are 7.6%, 4.0%, 3.3% and 3.0% for the parent material, 0°, 45°, and 90° bead angles, respectively. The ultrasonic C moduli were calculated in the direction of propagation. Furthermore, assuming isotropy and a typical Poisson’s ratio of 0.3, the Young’s modulus (E) corresponding to C = 5 GPa, is E = 0.743 C = 3.71 GPa, at 1 MHz. This is about a factor of 1.8 greater than the Young's modulus obtained in the range 0.1–100 Hz. However, the difference is understandable in the context of viscoelastic dispersion of the modulus with frequency [19]. Using the measured stress wave amplitudes, attenuation coefficients (α) for the parent material and the FFF printed parts were calculated. For this purpose, a total of three different travel lengths (z) and two samples of each type (i.e. Parent material, 0°, 45°, and 90° bead angles) were tested. The calculated attenuation coefficients including the maximum and minimum errors for each average value are shown in Fig. 10. As expected, the result suggests that the parent material exhibits higher attenuation than the 3D-printed parts. However, a larger data set is required to validate the findings and establish robust
Fig. 10. Average Attenuation Coefficients at 1 MHz for parent material and 3Dprinted ABS Parts.
time it takes the waves to propagates through the sample. If a major void was present inside the sample, a perturbation (peak) on the signal would arise as result of the discontinuity of the wave at that location. Between 7 and 10 measurements were taken on the polished samples and the pieces cut from the original cubes for conducting the attenuation experiment. As shown in Fig. 9, the parent material exhibits a marginally larger modulus than the 3D-printed parts. However, the results reported herein suggests that the bead angle does not significantly influence the longitudinal compression modulus of the 3Dprinted ABS parts using an extrusion factor equal to 1.0. The error bars shown in Fig. 9, indicate the maximum and minimum errors relative to the average moduli. Furthermore, the coefficient of variation (COV) for 709
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Fig. 12. Complex modulus and tan δ in the frequency spectrum.
References
conclusions. The loss tangent of materials with small loss angles can be approximated without significant loss in accuracy by Eq. (5). As shown in Fig. 11, at a 1 MHz frequency, the parent material exhibited a loss tangent 2–3 times larger than the 3D-Printed parts. Furthermore, the tan δ of the FFF sample with a bead angle of 45° was approximately 60% larger than that of the samples with bead angles of 0° and 90°. Notice that the loss tangent was effectively the same for bead angles of 0° and 90° (Table 5). Future work will include a wider range of frequencies to develop loss tangent and modulus master curves for 3D printed parts with various bead angles. Additional ultrasonic tests may be performed to further validate the attenuation results. These results presented will help engineers better understand the impact that manufacturing processes have on the mechanical behavior of their designs. While these experiments focus on FFF exclusively, the work can be expanded to any additive manufacturing technique in polymers or metals.
[1] J. Rodríguez, J. Thomas, J. Renaud, Mechanical behavior of acrylonitrile butadiene styrene fused deposition materials modeling, Rapid Prototyp. J. 9 (4) (2003) 219–230. [2] A. Garg, A. Bhattacharya, A. Batish, Chemical vapor treatment of ABS parts built by FDM: analysis of surface finish and mechanical strength, Int. J. Adv. Manuf. Technol. 89 (2017) 2175–2191. [3] S.-H. Ahn, M. Montero, D. Odell, S. Roundy, P.K. Wright, Anisotropic material properties of fused deposition modeling ABS, Rapid Prototyp. J. 8 (4) (2002) 248–257. [4] C. Koch, L. Van Hulle, N. Rudolph, Investigation of mechanical anisotropy of the fused filament fabrication process via customized tool path generation, Addit. Manuf. 16 (2017) 138–145. [5] A. Bellini, S. Güçeri, Mechanical characterization of parts fabricated using fused deposition modeling, Rapid Prototyp. J. 9 (4) (2003) 252–264. [6] B. Rankouhi, S. Javadpour, F. Delfanian, Failure analysis and mechanical characterization of 3D printed ABS with respect to layer thickness and orientation, J. Fail. Anal. Prev. 16 (3) (2016) 467–481. [7] T. Pfeifer, C. Koch, L. Van Hulle, G.A. Mazzei Capote, N. Rudolph, Optimization of the FDM™ additive manufacturing process, Proceedings of the Annual Technical Conference (ANTEC) of the Society of Plastics Engineers (2016) 22–29. [8] Sabic, "CYCOLACTM RESIN MG94," [Online]. Available: https://www.sabic.com/ en/products/popup.html?q=MG94&grades=113eaad7-e0d8-e611-819b06b69393ae39. (Accessed 6 July 2018). [9] VanHulleOne, "SciSlice - GitHub," GitHub Universe, [Online]. Available: https:// github.com/VanHulleOne/SciSlice. (Accessed 6 July 2018). [10] G.A. Mazzei Capote, A. Redmann, C. Koch, N. Rudolph, Towards a robust production of FFF end-user parts with improved tensile properties, Solid Freeform Fabrication Symposium – An Additive Manufacturing Conference (2017). [11] F. Rayegani, G.C. Onwubolu, Fused deposition modelling (FDM) process parameter prediction and optimization using group method for data handling (GMDH) and differential evolution (DE), Int. J. Adv. Manuf. Technol. 73 (2014) 509–519. [12] B. Huang, S. Singamneni, Raster angle mechanics in fused deposition modelling, J. Compos. Mater. 49 (3) (2015) 363–383. [13] J.L. Colón Quintana, T. Heckner, A. Chrupala, J. Pollock, S. Goris, T. Osswald, Experimental study of particle migration in polymer processing, Polym. Compos. 40 (6) (2018) 2165–2177. [14] A.F. Yee, M.T. Takemori, Dynamic bulk and shear relaxation in glassy polymers. I. Experimental techniques and results on PMMA, J. Polym. Sci.: Polym. Phys. Ed. 20 (1982) 205–224. [15] S. Dul, L. Fambri, A. Pegoretti, Fused deposition modelling with ABS–graphene nanocomposites, Composites: Part A 85 (2016) 181–191. [16] A. Arivazhagan, S.H. Masood, Dynamic mechanical properties of ABS material processed by fused deposition modelling, Int. J. Eng. Res. Appl. (IJERA) 2 (3) (2012) 2009–2014. [17] T.-W. Liu, S.K. Gaggar, S. Agarwal, A. Al-Mulla, R.K. Gupta, Novel "Eco-FR" Styrenic Polymers, Antec, Chicago, Illinois, USA, 2009. [18] A.R. Torrado Perez, D.A. Roberson, R.B. Wicker, Fracture surface analysis of 3Dprinted tensile specimens of novel ABS-based materials, J. Fail. Anal. Prev. 14 (3) (2014) 343–353. [19] R. Lakes, Viscoelastic Materials, Cambridge University Press, NY, USA, 2009.
4. Conclusion The complex modulus and tanδ values were found for a wide frequency spectrum utilizing DMA and ultrasonic methods. Test were performed at both low and high frequency to determine the viscoelastic response of the material. As shown in Fig. 12, the complex modulus increases with frequency. A distinction of magnitude for parts with different bead orientation is clearly observed in the range of 0.1–100 Hz. However, the complex modulus tends to converge at 1 MHz with small differences between the measured values. Nonlinear behavior was observed in the complex modulus for the higher amplitudes in the DMA test. 3D printed parts made with rectangular beads corresponding to extrusion factor 1 exhibited some oriented porosity but minimal anisotropy of modulus as expected from samples with lower void content and prominent bead overlap. The viscoelastic damping tan δ was greater for the parent material than for printed material at low frequency 0.1 Hz–100 Hz and at ultrasonic frequency 1 MHz. Declaration of Conflict of Interest The author(s) declare(s) that there is no conflict of interest regarding the publication of this paper.
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Additive Manufacturing journal homepage: www.elsevier.com/locate/addma
Viscoelastic properties of fused filament fabrication parts a,⁎
a
a
T b
José Luis Colón Quintana , Alec Redmann , Gerardo A. Mazzei Capote , Angel Pérez-Irizarry , Abrahan Becharaa, Tim A. Osswalda, Roderic Lakesc a b c
Polymer Engineering Center, Department of Mechanical Engineering, University of Wisconsin-Madison, Madison, WI, 53706-1691, USA Department of Civil & Environmental Engineering, University of Wisconsin-Madison, Madison, WI, 53706, USA Department of Engineering Physics, University of Wisconsin-Madison, Madison, WI, 53706, USA
A R T I C LE I N FO
A B S T R A C T
Keywords: Viscoelastic DMA Ultrasonic Loss tangent FFF
Parts made by fused filament fabrication differ in their mechanical properties from the parent material. To investigate the effect of the manufacturing process on the mechanical properties of 3D-printed parts, a series of experiments including Dynamic Mechanical Analysis (DMA) and ultrasonic wave propagation were conducted. For this purpose, printed parts were made from custom ABS filament and were printed using a rectangular bead shape to minimize porosity. The main properties investigated included the elastic, loss and storage moduli, and the material loss tangent (tan δ). Results indicate that the elastic modulus of the printed material was somewhat lower than that of the parent material, about 2 GPa for frequencies 0.1 Hz–100 Hz. Furthermore, tan δ was largest for the parent material at about 0.03 compared with 0.01–0.02 for the printed material. Ultrasonic longitudinal wave measurements at 1 MHz on printed specimens with bead angles of 0°, 45° and 90°, revealed minimal anisotropy and, consistent with DMA results, tan δ was also largest for the parent material.
1. Introduction Fused Filament Fabrication (FFF) has been on the forefront of Additive Manufacturing (AM) due to low costs and broad availability of machines. Furthermore, FFF relies on thermoplastic materials exhibiting pronounced viscoelastic behavior. However, parts produced by FFF technology tend to be highly anisotropic, due in part to the junction formed between adjacent beads that behaves akin to a non-optimal, thermally-driven polymeric weld [1]. These bead junctions tend to be weaker and more compliant than the solid polymer. Moreover, the introduction of material discontinuities that stem from the junction of elliptical plastic strands, results in parts containing voids that can act as stress concentrators [2]. It is well known that the final mechanical properties of FFF parts are extremely sensitive to build parameters -such as bead orientation, build temperature, air gap, solidity ratio, among others [3–7]. These factors imply that understanding the mechanical performance of FFF parts is a daunting task, yet extremely important if final user grade performance is to be expected. Unfortunately, research about viscoelastic behavior of FFF parts is scarce. Hence, the work reported herein aimed to understand the effect of bead orientation on the viscoelastic properties of ABS FFF parts, through a variety of dynamic loading experimental techniques, i.e., ultrasonic testing and Dynamic Mechanical Analysis (DMA).
⁎
The storage and loss moduli, and the phase angle (δ) were measured and compared between each raster orientation and benchmarked against a sample of the parent material. 2. Materials and procedure 2.1. Material The first step of the experimental work involved development of a custom thermoplastic filament for the FFF process. The reasoning behind this decision was two-fold. First, the use of an off-the-shelf, commercial thermoplastic filament generally does not guarantee that two different spools were produced under the same processing conditions, or even using the same parent material. Secondly, the results from Koch et al. [4] show that fluctuations in the filament diameter have an impact in the mechanical properties of FFF parts due to improper volumetric output at the nozzle. The Cycolac®️ MG94 material produced by SABIC corporation [8] was chosen for this work. This is an Acrylonitrile Butadiene Styrene (ABS) based material traditionally used for injection molding thin walled parts, as well as extrusion of FFF filament. With a reported Melt Flow Index (MFI) of 11.7 g/10 min, it is an ideal material for both the FFF and extrusion processes [8]. The Melt Flow Index is defined as the
Corresponding author. E-mail address: [email protected] (J.L. Colón Quintana).
https://doi.org/10.1016/j.addma.2019.06.003 Received 19 February 2019; Received in revised form 3 May 2019; Accepted 6 June 2019 Available online 15 June 2019 2214-8604/ © 2019 Elsevier B.V. All rights reserved.
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Fig. 1. Extrusion Setup. Taken from [9].
parameter from 0.63 to 1.0. It was shown that by using an extrusion factor of 0.97 and a layer height of 0.1 mm, it was possible to reach 98% of the strength of an injection molded part subjected to tensile load. Pfeifer et al. [7]. performed experiments varying the EF from 0.75 and 1.08. By performing Computer Tomography (CT) scans and mechanical testing of dogbone samples, it was demonstrated that lower solidity ratios, stemming from lower EF, resulted in poor mechanical properties bearing 24–52% of the highest solidity ratio. Previous research on the mechanical performance of FFF have varied the bead geometry of the samples by assigning positive and negative air gap values to the printing parameters [3,11,12], as shown in Fig. 3. By allowing air gaps, the morphology of the beads and the voids is changed in a manner similar to using EF manipulation: a positive air gap produces beads of more rounded nature, while a negative air gap reduces the void content, being comparable to using an EF approaching 1. It was shown that anisotropy was significant, as the mechanical properties of the samples varied considerably between raster orientations, for all air gap configurations. However, a drastic change in mechanical properties was observed between analogous samples of the same bead orientation, but different air gap values: the failure stress of samples with negative airgaps was considerably higher than samples with null or positive air gaps -with the latter having extremely low values: only between 17–45% of its corresponding air gap sample. To minimize the problems associated with using lower EFs, this parameter was fixed to a value of 1, producing beads with a theoretical rectangular cross section exclusively. It should be noted that due to variations in the diameter of the filament and imperfections in the build process, the final bead geometry is not a perfect rectangle, and the junction of contiguous beads still produces voids characteristic of FFF parts.
mass extruded through a capillary in a 10 min period while applying a constant load. By having a material with a high MFI, it allows more material to be printed, which results in faster printing time. The MG94 material was extruded in a setup that consisted of a single screw extruder (Extrudex EDN 45X30D, Germany) with a 45 mm screw diameter and a screw length to screw diameter ratio (L/D) of 30. The hot melt was extruded at 205 °C through a circular die with a 5.8 mm diameter, and then guided through a pre-skinner into a vacuum-assisted, heated water bath (Conair, USA) to cool the extrudate whilst minimizing void formation. The solidified filament was then passed through a 3-axis laser micrometer (LaserLinc, USA) and a belt puller (Conair, USA) configured in a control loop. The dimensions of the filament were controlled by automatically adjusting the speed of the puller if the readings from the micrometer were out of specification, in this case a diameter of 2.85 mm with a tolerance of ± 0.02 mm. Finally, the product was wound onto spools using a filament winder. A schematic of the extrusion setup can be seen in Fig. 1. Prior to any usage in a printer, the filament was dried in a silo (Novatec, USA) at 82 °C for 3 h. Since the ABS material is amorphous and the temperature at which it is subjected is under the glass transition temperature, no significant thermal or structural change is induced in the material. 2.2. Tool path generation for FFF specimens All test specimens were produced using a TAZ 5 (Lulzbot, USA) FFF printer with a 0.5 mm nozzle. Tool pathing, a set of printing commands that tells the printer where to move, how fast it moves, and how much material need to extrude, was performed using SciSlice: a customized slicing engine developed in-house [9]. A conscious effort was made to maintain as many print parameters as possible constant across all prints. These slicing engine parameters are listed in Table 1. Note that Extrusion Factor (EF) is a user parameter that represents a ratio of the true area occupied by the cross section of a bead (Abead) divided by the product of the bead width (Wbead) with the layer height (Hlayer). This factor is used by SciSlice to calculate the length of filament required to print a bead of known dimensions by using Eq. (1).
L filament *Afilament = EF*Wbead*Hlayer Lbead
2.3. DMA procedure Dynamic Mechanical Analysis (DMA) tests were conducted using the tensile configuration of the RSA 3 (TA Instruments) DMA at ambient temperature. Three rectangular, single layer samples were produced with a thickness of 0.2 mm and dimensions of 35 mm by 6 mm. The samples were manufactured in three bead orientations: 0°, 45°, and 90° where the angle is measured parallel to the load direction, as shown in Fig. 4. Additionally, a sample of the parent material was compression molded with the same geometry and used for comparison. The parent material will be used to compare parts used in FFF applications against parts manufactured with other processes where the final product is a solid part, without induced anisotropy by the bead orientation. Subsequently, it is assumed that parts manufactured using these processes will have the same mechanical response as the compression molded sample. The same comparison was done for samples of the ultrasonic test. Strain sweeps were conducted on all samples at 1 Hz from strains of 10−3–100 percent with five points each decade to identify the range for linear viscoelastic behavior. Once the linear viscoelastic region was identified, frequency sweeps were conducted using the linear viscoelastic strain percentage as the strain amplitude. This value was used assuming the material will remain in the linear viscoelastic region through the frequency sweep. Frequencies from 0.01 Hz to 100 Hz were used for the DMA test.
(1)
The extrusion factor controls the amount of material extruded through the volumetric balance shown in Eq. (1). For example, an EF of π/4 would produce beads of elliptic cross section, while an EF of 1 would produce beads of rectangular nature. Beads of circular or square cross section are possible if the layer height and the bead width have equal dimensions. See Fig. 2 for a visual representation of the impact of this parameter on the geometry of printed beads. Previous research by Koch et al. [4] showed the effect of the Extrusion Factor upon the mechanical properties of FFF, by varying this Table 1 Constant slicing engine parameters. Parameter
Value
Parameter
Value
Extrusion Factor (EF) Filament diameter Path width Layer height Printing temperature
1 2.85 mm 0.5 mm 0.2 mm 230 °C
Bed temperature Print speed Rapid speed Retraction distance Retraction speed
100 °C 2000 mm/min 4000 mm/min 1.5 mm 500 mm/min
2.4. Ultrasound procedure Cubic specimens of 30 mm were produced as described in the 705
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Fig. 2. Sets of beads produced with an extrusion factor EF of a) π/4 and b) 1. Image c) are micrographs of cross sections of printed parts [10]. Fig. 3. Air gap application. Adapted from Rayegani et al. [11].
preceding discussion, with bead orientations of 0°, 45° and 90° with respect to the direction of ultrasonic wave propagation. For comparison, a cylindrical sample of parent material with a 21.8 mm diameter and a length of 20.6 mm was manufactured using a heated chamber, where MG94 pellets were melted and pushed to a cylindrical mold by a pneumatic piston. A processing temperature of 240 °C was used for melting the material. Once the heaters were turned on, 11 min were given to allow the system to reach the target temperature and properly melt the material. A pressure of 689 kPa (100 psi) was used to actuate the piston and make the polymer flow towards the cavity, filling the mold. For more information of the manufacturing procedure of the sample used refer to [13]. To ensure a proper contact with the ultrasonic transducer, the surface of all samples was finely polished using a rotary metallographic polisher with abrasive discs of grit sizes of 800 and 1200. The experimental setup consisted of FFF produced cubes with three bead orientations and cylindrical samples of the parent material, each held in between two longitudinal 1 MHz ultrasonic transducers (Panametrics V102) connected to a pulse generator (Panametrics 500PR). An oscilloscope (Siglent SDS 1052DL) was connected to the pulse generator to monitor the pulse wave as it traveled through the specimen. To stabilize the signal, a weight of mass 3.8 kg was added on top of the uppermost transducer. A schematic of the experimental setup is shown in Fig. 5. To ensure proper contact with the transducers and a stronger signal on the
Fig. 4. DMA sample orientation. Angle was measure with respect to the loading direction.
Fig. 5. Schematic of the ultrasound experimental setup. 706
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Table 2 Density of parent material and bead angle.
Table 3 μ-CT Scan parameters.
Material
0°
45°
90°
PM
Parameter
Value
Parameter
Value
ρ (kg/m )
1036.31
1021.13
1032.06
1032.84
Voltage (kV) Current (A) Integration time (s) Gain
70 110 1 8.0 x
Image averaging Binning Number of projections X-ray intensity correction
Off 1×1 1800 9000–10,000
3
oscilloscope, a thin layer of glycerin was used as a coupling agent. For this series of tests, four sets of 30 mm cubes were printed for the bead orientations of 0° and 45° for a total of 8 samples. The samples at 0° were used for the 0° and 90° test. Each material sample was measured and weighed using a caliper and a high-precision scale (OHAUS®, Explorer EX125D), respectively, to calculate the density of each sample tested as shown in Table 2. Note that the parent material sample did not possess the highest density. The low parent material density is attributed to the material processing, particularly the creation of voids in the samples as ABS was injected into the mold. Since an extrusion factor of 1 was used for the 3D printed parts, the density of the 0° and 90° samples were close to the parent material. The low parent material density is believed to be due to a higher volume of voids in the parent material than the 0° and 90° samples. In contrast, the sample with bead angle of 45° had the lowest density. The slightly lower density could be a result of the tool path required to produce beads in 45°. The specimen length and the wave propagation time were used to compute the wave velocity; then the wave velocity (vl) and the material density (ρ) were used to compute the tensorial modulus as shown in Eq. (2), where C1111 is the modulus component in the wave/loading direction.
C1111 = (VI)2*ρ
3. Results and discussion 3.1. DMA results DMA measurements were performed upon printed material at three different orientations as well as the parent material. Fig. 6 shows the complex modulus (E*) and the tan δ values as a function of the strain percentage. As previously mentioned, a strain sweep was performed on the samples to determine the proportionality limit strain for each bead orientation and the parent material. Table 4 shows the approximate linear viscoelastic strain limit percentage for each orientation. The criteria used for the selection of these values was the strain limit percentage at which the tan δ is not strain dependent. Since the tan δ give the viscoelastic response of the material, it considers the change of the storage and loss modulus. The increase in complex modulus shown in Fig. 6 is due to the nature of the additive manufacturing process of bead placed next to each other. The creation of voids between bead and in some occasions the non-contact between beads induces a lower modulus at the lower strain percentages. As the material starts to elongate, the beads come closer together increasing the resistance to movement, thus increasing the modulus. Higher tan δ values coincide with a response regime where the loss modulus was predominant; however, increases in loss modulus can indicate yield or nonlinearity prior to yielding. The exact cause of the observed response for tan δ values was not determined in this study. Once the strain percentage was determined as previously shown in Table 4, a frequency sweep from 0.01 Hz to 100 Hz was performed on the samples. Fig. 7 shows the result for the frequency sweep. As observed below, the storage modulus increases with increasing frequency while the loss tangent does not differ much in the range 0.1 Hz–100 Hz. Such behavior is typical in the glassy regime [14]. The tan δ values for the printed parts falls within 0.01 and 0.02 while the parent material possesses a value of 0.03 in this frequency range. Values of storage modulus and tan δ found in literature coincides with experimental data acquired for both compression molded samples and samples made with different bead orientations [15,16]. The storage and loss moduli for all bead orientations were comparable. However, in the case of the parent material, the loss modulus was considerably higher than the 3D printed samples. This results in a tan δ trend that indicates higher damping capabilities for the parent material. This behavior, although counter-intuitive, can have its root in multiple factors, including the presence of voids in the parent material specimens. Another potential contribution could have stemmed from the single layer thickness of the printed samples, and the bed temperature being close to the glass transition temperature (Tg) of the material used: as shown in Table 1, the bed temperature was 100 °C, while the glass transition temperature for ABS is 103.8 °C [17]. Since these temperatures are very close to each other, it is suggested that heat transfer through the thickness of the sample is sufficient to facilitate proper bonding of the beads. Also, due to the nature of the manufacturing process, bead overlap could have resulted from the toolpath, thus increasing the strength of the printed samples in all orientations.
(2)
To determine the loss tangent of the FFF and injection molded ABS samples, the amplitude of the stress wave after traveling through the material was measured using the oscilloscope. After the first measurement, the samples were cut into smaller lengths and new measurements were taken to quantify the attenuation of the ultrasonic waves (Eq. (3)). Three lengths were used to determine the stress wave attenuation.
A1 = A 0*e−αz
(3)
With the measured wave amplitudes, the attenuation parameter (α) was determined by fitting an exponential curve to the wave amplitudes as a function of traveled distance. Furthermore, attenuation coefficients were also calculated by relating the measured amplitudes and the traveled distances as shown in Eq. (4). A
α=
ln( A1 ) 2
(Z2 − Z1)
(4)
where Ai and Z i are the measured wave amplitude and corresponding wave travel distance. Finally, the average attenuation coefficient was used to determine the loss tangent (tan δ) for the 3D-printed samples and parent material using Eq. (5).
2VI*α V *α = I ≈ tanδ ω π*f
(5)
where ω is the angular frequency in radians per second and f is the ordinary frequency in Hertz. 2.5. Computed tomography Objects produced using FFF tend to have process induced defects that stem from either the toolpath used during part production, machine limitations or material inconsistencies. Given that the ultrasonic testing is particularly sensitive to defects in the test geometry, computed tomography was used to obtain images of the internal structure of the cubic samples produced for the ultrasonic measurements. A Zeiss Metrotom 800 (Zeiss, Germany) was used to scan the parts, using parameters shown in Table 3 to obtain high quality images. 707
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Fig. 6. Viscoelastic properties as a function of strain percentage of parent material and material printed at angles 0°, 45°, and 90°.
of the filament necessary to prevent dripping of material, was not enough and thus undesired material was added to the part as the print head moved above the object to start a different layer. Note that both images reveal voids. These are introduced every time there is a bead or layer junction due to the circular nozzle orifice of the printer. The two first problems were corrected through manipulation of the slicing engine parameters, but the presence of voids is characteristic to the FFF printing process and thus cannot be avoided. Small voids in FFF parts can be minimized but not eliminated. Due to the cylindrical shape of the nozzle, the beads were placed layer by layer with an elliptical or rectangular shape as shown in Fig. 2. The shape of the bead depends on the extrusion factor, volumetric flow rate of the material, and nozzle diameter to name a few.
Table 4 Strain limit percentage for linear viscoelastic region as a function of bead orientation. Orientation
Approximate Linear Viscoelastic Strain Limit (%)
0° 45° 90° PM
3 × 10−1 5 × 10−2 1 × 10−2 6 × 10−2
3.2. Computed tomography results As mentioned in the previous section, computed tomography was used to obtain images of the internal structure of the cubic samples produced for the ultrasonic measurements. Fig. 8 shows on the left (Fig. 8a) a sample printed using a raster orientation of 45°. Due to bias on the tool pathing algorithm, the raster was printed in a discontinuous manner, thus leaving a sizeable defect that ran across the entire height of the sample. On the right (Fig. 8b), filament retraction, the movement
3.3. Ultrasonic results The stress wave velocity and the density of the samples were used to determine the tensorial modulus along the stress wave path. As result of the manufacturing process of 3D printed parts, voids are commonly found in the final part [18] which resulted in peaks when measuring the
Fig. 7. Viscoelastic properties as a function of frequency of parent material and material printed at angles 0°, 45°, and 90°. 708
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Fig. 8. Printed sample with (a) no excess material and (b) excess material. Scale bar, 5 mm.
Fig. 9. Average Moduli at 1 MHz of parent material and 3D-printed ABS Parts.
Fig. 11. Average Loss Tangent at 1 MHz for parent material and 3D-printed ABS Parts. Table 5 Summary of ultrasonic results at 1 MHz. Bead Angle
(VL) avg (m/ s)
(C1111) (GPa)
0° 45° 90° PM
2190 2182 2177 2186
4.91 4.80 4.85 5.06
a
avg
E (GPa)
αavg (neper/ mm)
(tanδ)
3.65 3.57 3.60 3.76
0.0222 0.0347a 0.0226 0.0659a
0.0155 0.0244a 0.0157 0.0459a
avg
Only one sample used for calculation.
the calculated moduli are 7.6%, 4.0%, 3.3% and 3.0% for the parent material, 0°, 45°, and 90° bead angles, respectively. The ultrasonic C moduli were calculated in the direction of propagation. Furthermore, assuming isotropy and a typical Poisson’s ratio of 0.3, the Young’s modulus (E) corresponding to C = 5 GPa, is E = 0.743 C = 3.71 GPa, at 1 MHz. This is about a factor of 1.8 greater than the Young's modulus obtained in the range 0.1–100 Hz. However, the difference is understandable in the context of viscoelastic dispersion of the modulus with frequency [19]. Using the measured stress wave amplitudes, attenuation coefficients (α) for the parent material and the FFF printed parts were calculated. For this purpose, a total of three different travel lengths (z) and two samples of each type (i.e. Parent material, 0°, 45°, and 90° bead angles) were tested. The calculated attenuation coefficients including the maximum and minimum errors for each average value are shown in Fig. 10. As expected, the result suggests that the parent material exhibits higher attenuation than the 3D-printed parts. However, a larger data set is required to validate the findings and establish robust
Fig. 10. Average Attenuation Coefficients at 1 MHz for parent material and 3Dprinted ABS Parts.
time it takes the waves to propagates through the sample. If a major void was present inside the sample, a perturbation (peak) on the signal would arise as result of the discontinuity of the wave at that location. Between 7 and 10 measurements were taken on the polished samples and the pieces cut from the original cubes for conducting the attenuation experiment. As shown in Fig. 9, the parent material exhibits a marginally larger modulus than the 3D-printed parts. However, the results reported herein suggests that the bead angle does not significantly influence the longitudinal compression modulus of the 3Dprinted ABS parts using an extrusion factor equal to 1.0. The error bars shown in Fig. 9, indicate the maximum and minimum errors relative to the average moduli. Furthermore, the coefficient of variation (COV) for 709
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Fig. 12. Complex modulus and tan δ in the frequency spectrum.
References
conclusions. The loss tangent of materials with small loss angles can be approximated without significant loss in accuracy by Eq. (5). As shown in Fig. 11, at a 1 MHz frequency, the parent material exhibited a loss tangent 2–3 times larger than the 3D-Printed parts. Furthermore, the tan δ of the FFF sample with a bead angle of 45° was approximately 60% larger than that of the samples with bead angles of 0° and 90°. Notice that the loss tangent was effectively the same for bead angles of 0° and 90° (Table 5). Future work will include a wider range of frequencies to develop loss tangent and modulus master curves for 3D printed parts with various bead angles. Additional ultrasonic tests may be performed to further validate the attenuation results. These results presented will help engineers better understand the impact that manufacturing processes have on the mechanical behavior of their designs. While these experiments focus on FFF exclusively, the work can be expanded to any additive manufacturing technique in polymers or metals.
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4. Conclusion The complex modulus and tanδ values were found for a wide frequency spectrum utilizing DMA and ultrasonic methods. Test were performed at both low and high frequency to determine the viscoelastic response of the material. As shown in Fig. 12, the complex modulus increases with frequency. A distinction of magnitude for parts with different bead orientation is clearly observed in the range of 0.1–100 Hz. However, the complex modulus tends to converge at 1 MHz with small differences between the measured values. Nonlinear behavior was observed in the complex modulus for the higher amplitudes in the DMA test. 3D printed parts made with rectangular beads corresponding to extrusion factor 1 exhibited some oriented porosity but minimal anisotropy of modulus as expected from samples with lower void content and prominent bead overlap. The viscoelastic damping tan δ was greater for the parent material than for printed material at low frequency 0.1 Hz–100 Hz and at ultrasonic frequency 1 MHz. Declaration of Conflict of Interest The author(s) declare(s) that there is no conflict of interest regarding the publication of this paper.
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