materials
Article
Optimization of Manufacturing Parameters and Tensile
Specimen Geometry for Fused Deposition Modeling (FDM)
3D-Printed PETG
Arda Özen 1,* , Dietmar Auhl 1, Christina Völlmecke 2, Josef Kiendl 3and Bilen Emek Abali 4
Citation: Özen, A.; Auhl, D.;
Völlmecke, C.; Kiendl, J.; Abali, B.E.
Optimization of Manufacturing
Parameters and Tensile Specimen
Geometry for Fused Deposition
Modeling (FDM) 3D-Printed PETG.
Materials 2021,14, 2556. https://
doi.org/10.3390/ma14102556
Academic Editor: Antonino Recca
Received: 17 April 2021
Accepted: 5 May 2021
Published: 14 May 2021
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Attribution (CC BY) license (https://
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1Chair of Polymer Materials Science and Technologies, Institute of Material Science and Technology,
2Stability and Failure of Functionally Optimized Structures Group, Institute of Mechanics,
3Department of Civil Engineering and Environmental Sciences, Institute of Engineering Mechanics and
Structural Analysis, Universität der Bundeswehr München, Werner-Heisenberg-Weg 39,
85577 Neubiberg, Germany; josef.kiendl@unibw.de
4Division of Applied Mechanics, Department of Materials Science and Engineering, Uppsala University,
*Correspondence: ar[email protected]
Abstract:
Additive manufacturing provides high design flexibility, but its use is restricted by limited
mechanical properties compared to conventional production methods. As technology is still emerg-
ing, several approaches exist in the literature for quantifying and improving mechanical properties.
In this study, we investigate characterizing materials’ response of additive manufactured structures,
specifically by fused deposition modeling (FDM). A comparative analysis is achieved for four differ-
ent tensile test specimens for polymers based on ASTM D3039 and ISO 527-2 standards. Comparison
of specimen geometries is studied with the aid of computations based on the Finite Element Method
(FEM). Uniaxial tensile tests are carried out, after a careful examination of different slicing approaches
for 3D printing. We emphasize the effects of the chosen slicer parameters on the position of failures
in the specimens and propose a simple formalism for measuring effective mechanical properties of
3D-printed structures.
Keywords:
additive manufacturing; 3D printing; mechanics; slicing approach; polymers; finite
element method
1. Introduction
Additive manufacturing is a production process relying on building structures layer-
by-layer by utilizing the following strategy. First, the computer-aided design (CAD) of
a structure is created by a design software. Second, this CAD model is converted by
software, called a slicer, to a 3D printing code, where process information is supplied for
manufacturing [
1
,
2
]. This strategy is taken for granted and used in different branches of the
industry, such as automotive, aeronautics, and biomechanical [
3
,
4
], as well as in research,
especially for studying metamaterials [
5
–
7
]. Additive manufacturing processes, according
to ISO/ASTM52900-15, are categorized in one of the following seven groups: material
extrusion, material jetting, binder jetting, sheet lamination, vat photopolymerization,
powder bed fusion, and directed energy deposition [8–10].
Fused deposition modeling (FDM) is an additive manufacturing method based on
materials extrusion [
1
]. For thermoplastic polymers in the form of spools of filaments,
one filament is pushed through a nozzle at slightly over the melting temperature. In this
way, the layer is deposited as a viscous fluid, solidifying by decreasing its temperature
under convective heat transfer on the surface. This procedure is repeated layer-by-layer,
allowing almost any shape. Properties and the final quality depends on material chem-
istry [
11
], as well as chosen process parameters in manufacturing [
12
,
13
]. Affected by
Materials 2021,14, 2556. https://doi.org/10.3390/ma14102556 https://www.mdpi.com/journal/materials
Materials 2021,14, 2556 2 of 19
the layer-by-layer production, an inner structure occurs in the final product, leading to
structure-related anisotropic properties [
14
,
15
], as well as higher-order (so-called size)
effects [
16
–
18
]. Estimating mechanical behavior of materials is of interest especially re-
garding production parameters [
19
,
20
] and used algorithms for depositing layers. For the
mechanical characterization of parts manufactured by the FDM, no standard methods are
established [21,22].
Ample studies have been done for modeling polymers manufactured by the FDM.
Among others, structurally dependent elasto-plastic behavior has been modeled [
23
]. The-
oretical models are proposed for determining tensile strength and Young’s modulus with
different raster angles and layer thicknesses [
24
]. Different length scales and their inter-
actions have been investigated—mainly based on the classical laminate theory (CLT) [
25
].
However, the classical laminate theory uses assumptions with limited validity for FDM,
such as perfect bonding [
26
]. Some FDM materials are compared by performing tensile
experiments [
27
]. Even differences between characterization methods for polymers (ten-
sile testing) are investigated [
28
]. Effects of different tensile test specimen geometries
taken from ASTM D638 (Standard Test Method for Tensile Properties of Plastics) on the
anisotropy have been studied regarding processing parameters, such as the raster pattern,
print orientation, and tensile specimen dimensions [
29
]. Layer thickness and build ori-
entation are analyzed by tensile, flexural, and impact tests [
30
]. Polymer composites for
FDM are proposed and characterized with polypropylene (PP) and natural fibers, such as
hemp (Cannabis sativa) or harakeke (Phormium tenax) [
31
]. Different layer orientations of
acrylonitrile butadiene styrene (ABS) polymer [
32
] demonstrate increased tensile strength
along fibers in the FDM [
33
]. It is not only layer thickness, but also orientation angle and air
gap that affects the mechanical properties of polymers greatly [
34
]; this phenomenon has
been identified [
35
], and for details we refer to [
19
]. A correlation between 3D-printing time
and dimensional accuracy is established by parameter optimization [
36
]. For achieving
high visual quality and fast 3D printing, the slicing algorithm is optimized [
37
]. The effect
of printing time on mechanical properties of FDM 3D-printed Poly(Lactic Acid) (PLA)[
38
]
and PLA/Graphene composites are investigated [39].
Determination of material properties by using a uniaxial tensile test is challenging
in 3D-printed materials. ASTM D638 is designed for plastics, but the suggested design
(topology) causes a premature fail [
14
]. Moreover, a standard feature called infill pat-
terns manipulates the materials’ response, owing to an inner substructure [
40
–
42
]. This
substructure-related response deviation is examined in Polylactic Acid (PLA) parts with
five different infill patterns [
43
]. Additionally, other process parameters, such as raster
layup, including raster angle and width, as well as contour width are investigated [
44
–
46
]
for their effects on the toughness and strength leading to interlocking mechanisms [
47
].
Build orientation, layer thickness, and feed rate are discussed on 3D-printed PLA samples
in [
48
], and layer thickness and raster angle parameters for PLA and ABS in [
49
]. In order
to model the mechanical response of additive manufactured polymers [
50
–
52
], well-known
homogenization techniques are used in composite materials [
53
], for example by using a
variation of carbon-fiber content in thermoplastic matrix-based composites built by the
FDM [
54
] and also for identifying substructure-related anisotropic properties [
55
] to be
used in computations [56,57].
As we aim for developing a consistent approach for characterizing ultimate tensile
strength under uniaxial loading, we emphasize the importance of the selected specimen
topology, as well as the process parameters in slicer settings. We study and demonstrate
how to obtain adequate repeatability, and thus consistent material parameters. In order to
determine the materials’ response, ASTM D638 and ISO 527-2 (determination of tensile
properties for molding and extrusion plastics) have been utilized, where some problems
were reported [
14
,
31
,
45
] as mainly being effected by prescribed curvatures in the specimen
structures and being very challenging to manufacture in 3D printers. As a remedy, ASTM
D3039 (Standard Test Method for Tensile Properties of Polymer Matrix Composite Mate-
rials) was suggested [
14
,
25
,
58
]. We stress that these issues are partly because of process
Materials 2021,14, 2556 3 of 19
parameters, and examine different tensile test specimen geometries experimentally, as well
as numerically, by Finite Element Method (FEM)-based simulations. Four different speci-
men configurations have been prepared, namely, two from ASTM D3039 and two based
on ISO 527-2. Reliable results with a low standard deviation are obtained by “fine-tuning”
the process parameters and modifying the specimen structure, nevertheless still using the
suggested ISO standard.
2. Materials and Methods
Four different tensile specimen geometries were investigated: ASTM D3039, ASTM
D3039 angle, ISO 527-2, and ISO-modified (based on ISO 527-2). The specimen specifica-
tions are compiled in Table 1and their drawings are depicted in Figure 1.
Table 1. Specimen specifications.
Description ASTM D3039 ASTM D3039 Angle ISO527-2 ISO-Modified
Tab length in mm 30 20 21.4 21.8
Tab thickness in mm 2.8 2.8 - -
Thickness in mm 3.2 3.2 6 6
Length in mm 150 150 150 188
Width in mm 15 15 21.7 29.7
Gauge length in mm 90 90 60 60
Angle - 31.28° R60 R105
(a) (b)
(c) (d)
Figure 1.
Tensile test specimen geometries and their specifications in mm: (
a
) ASTM D3039; (
b
) ISO527-2; (
c
) ASTM D3039
angle and (d) ISO-modified.
2.1. Fused Deposition Modeling
The samples were produced by an FDM-type 3D printer, namely Ultimaker 3 Extended
(Ultimaker B.V., Geldermalsen, The Netherlands). White-PETG filaments were purchased
from Materials4Print GmbH & Co. KG (Bad Oeynhausen, Germany). The tensile specimens’
CAD models were achieved in open-source platform Salome 9.3 and exported as stl files
leading to G-codes prepared by Ultimaker Cura 4.3.0 (Ultimaker B.V., Geldermalsen, The
Netherlands) with the aid of selecting process parameters, such as slicing speed, layer
thickness, temperature, and so forth. These parameters are of utmost importance, and we
provide them in Table 2.
Materials 2021,14, 2556 4 of 19
Table 2. Process parameters of 3D printing.
Parameter Value Unit
Layer thickness 0.3 mm
Layer width 0.4 mm
Print speed 55 mm/s
Initial layer speed 40 mm/s
Print acceleration 4000 mm/s2
Printing temperature 250 °C
Printing temperature initial layer 255 °C
Final printing temperature 240 °C
Bed temperature 70 °C
All specimens were printed layer-by-layer by choosing the maximum possible filling.
All contours are avoided in 3D printing in order to eliminate any destruction of the
unidirectional structure. All layers were unidirectional with 0
°
orientation. For printing the
structures, we used line patterns. We emphasize that the fiber is used in the jargon of FDM
denoting the 3D-printed line pattern. The line pattern steers the layers and thus generates
fibers. These fibers were not connected to the endpoints—the connecting infill lines setting
was off, because it could destruct the unidirectionality. The material properties of PETG
used for printing are given in Table 3. We assume that these parameters (supplied by the
manufacturer) are determined by using an injection mold specimen such that the porosity
is expected to be nearly zero (ideal case). In FDM, between filaments, depending on the
parameters, voids occur, leading to a porous structure. Hence, the values supplied by the
manufacturer are understood as an upper threshold of “effective” parameters obtained
from the specimens printed by FDM-based manufacturing.
Table 3. Material properties of PETG.
Value Unit Method
Mass density 1.27 g/cm3ASTM D792
Elongation at break 70 % ASTM D638
Tensile strength at break 26 MPa ASTM D638
Flexural modulus 2150 MPa ASTM D790
Melting point 200 −230 °C ASTM D3418
Heat distortion temperature 74 °C ASTM D648
2.1.1. Slicing Approach
By selecting default settings in the slicer, we observed problems in producing a
unidirectional (UD) structure. Hence, we propose two particular changes: development of
new travel paths and optimization of slicing sequences.
Optimization of Excess Travel Paths
Travel settings are one of the key process parameters in slicer settings. We demonstrate
how to adjust these settings for the uniaxial tensile specimen of a UD structure in order
to prevent any excess travel lines in specimens. Excess travel lines may cause premature
failures, leading to stress concentrations as observed in experiments. In order to make
the role of this parameter obvious, we start off with the default travel configuration in
Cura (slicer). In default, the slicer tries to minimize the print time such that the continuous
production may end up with additional (excess) layers deposited while the nozzle travels.
These travel lines are pathways for the nozzle so as to reach a specific position. The default
configuration in Cura is generating travel lines inadequate for the UD structure aimed
for uniaxial testing, and the suggested layer deposition by the default configuration is
illustrated in Figure 2a, making it obvious that the UD structure is difficult to maintain.
Materials 2021,14, 2556 5 of 19
(a)
(b)
Figure 2.
Simulation of production with two travel configuration setups by using arrows visualizing the travel of the nozzle.
(
a
) Default setup with parameters from Table 4, black lines are travel lines, also depositing material, indicating weak points
in the specimen. (b) Optimized setup, the travel lines are along the outer surface of the specimen.
Mostly, a standard configuration is obtained by an algorithm optimizing speed or
weight. Resulting inner travel lines disturb the aimed for UD structure. Several specimens
are manufactured simultaneously in one production run; therefore, inner travel lines
are different in each specimen, making a comparative analysis unreliable. Moreover, as
seen in Figure 2a, the black lines indicate that the nozzle introduces weak spots within
the specimen, and along these inner travel lines, we expect a stress localization and a
premature failure of the structure. Obviously, such a failure is not representative of the
material itself. An optimized set of settings, specifically for the UD structure, is compiled
in Table 4for establishing a new travel path approach.
Optimization of Slicing Sequence
By considering the production time as well, we have succeeded in defining a new
configuration setup with parameters given in Table 4that substantially increases the inner
structure by placing travel lines to the outer contour, as demonstrated in Figure 2b. Note
that all layers should be produced by the top layer configuration except the bottom layer.
In this context, the slicing sequence means the order of the printing areas. The specimens
were printed with two different slicing approaches, called Slicing A and Slicing B. In
one process, many specimens were manufactured. Their positioning, called layout, is of
importance for the Ultimaker Cura slicing algorithm. Figure 3shows the Slicing A and
Slicing B results figuratively.
Table 4. Standard and modified travel configurations in Cura.
Parameter Standard Modified
Combining mode All Not in Skin
Max comb. distance with no retract 0 mm 100 mm
Avoid printed parts when traveling X X
Travel avoid distance 3 mm 10 mm
Layer Start X 213.0 mm 200.0 mm
Layer Start Y 198.0 mm 200.0 mm
Z hop when retracted X X
Z hop only over printed parts X X
Z hop height 2 mm 5 mm
Materials 2021,14, 2556 6 of 19
(a)
(b)
Figure 3. Simulation by using two different slicing approaches. (a) Newly optimized Slicing A and (b) Current Slicing B.
In Figure 3, the yellow areas present the already deposited material, whereas the
gray places show yet to-be-deposited sections. In Slicing B (Figure 3b), the 3D printer
deposits in the order given by 1, then 2, and 3. When Area 4 is being manufactured, Areas
1 and 2 nearly attain the room temperature such that the bond in the contact lines of the
areas between 1 and 4, as well as 2 and 4, may differ from the rest. Observed premature
failures in experiments in these possibly weak contact areas are justified by this deficiency
in thermal fusion. Hence, we understand that slicer settings may affect the onset of failure
in 3D-printed parts as a consequence of heterogeneous properties caused by the printing
strategy. Therefore, we have developed Slicing A as follows.
In Slicing A (see Figure 3a), the area denoted by 1 is initially produced. It is important
that the top-right and left-bottom edges are sliced without pausing in Area 1. When Area 1
has been completed, there are two open edges (Areas 2 and 3) in opposite locations. Still,
we expect weaker bonds in contact lines between Areas 1 and 2 as well as 1 and 3. However,
these weak regions are divided on both ends, contrary to Slicing B, where weak regions are
located at the same end. Therefore, we expect that Slicing A will perform better by using
the layout given in Table 5.
Table 5. Specimen coordinates of Slicing A.
Specimen Positions x y
First specimen −11.5385 −67.8167
Second specimen −11.5385 −2.9837
Third specimen −11.5385 67.8167
A new travel path can be used in every geometric shape and even in complex shapes.
However, the slicing sequence depends on the geometry. Each geometric shape is to be
optimized separately, and default parameters were obtained by an optimization pro-
cedure to minimize the printing time by ignoring process-related anisotropy and its
possible consequences.
2.2. Unidirectional Tensile Tests
Prior to mechanical tests, all specimens were preserved at a 40
°
C vacuum oven
against the water uptake. Uniaxial tensile tests were performed with a Zwick 1446 (Zwick,
Materials 2021,14, 2556 7 of 19
Ulm, Germany). testing machine. A mechanical extensometer was utilized to measure
strain. We provide the tensile equipment and the test set-up in Figure 4. Experiments were
conducted and steered by displacement with a ramp speed of 2mm/min on ISO527-2,
ASTM D3039, and ASTM D3039 angle specimens; and with a ramp speed of 2.5mm/min
on ISO-modified specimens. In order to study the strain rate sensitivity (viscoelasticity),
additional experiments were conducted for a sizeable range of strain rates from 0.01 to
100s
−1
. Postprocessing was performed by the corresponding software leading to values of
the ultimate tensile strength (UTS) and Young’s modulus. We provide the number of 3D-
printed and tested specimens in Table 6for assessing the reliability. In total, 48 specimens
were 3D-printed and tested.
Table 6.
Experimental design with different specimen types and slicing approaches, all produced
and tested on six specimens to obtain a statistical confidence interval.
Standard Name Slicing A Slicing B
ISO-modified 6 specimens 6 specimens
ISO 527-2 6 specimens 6 specimens
ASTM D3039 6 specimens 6 specimens
ASTM D3039 angle 6 specimens 6 specimens
Figure 4.
Initial state of an ISO 527-2 specimen clamped into a Zwick 1446 testing device with an
extensometer (left) and a loading cell (top). One mounting side is fixed (bottom), while displacement
is applied vertically through the upper mounting side.
In order to remove slack from the specimen and test equipment, preloading is rec-
ommended prior to the tensile test. Therefore, we applied up to 0.1 MPa at a low rate
(quasistatic condition). Strain was corrected by initially setting it to zero. However, we
did not correct the stress results, as the modulus calculation involved stress differences.
We determined that the material was elastic without any significant rate-dependency
Materials 2021,14, 2556 8 of 19
at room temperature. We circumvented from determining Poisson’s ratio and used the
manufacturer’s value.
2.3. Computation by Using the Finite Element Method
Accompanying the experiments, simulations have been utilized for understanding
the structural effect on the test results. A standard numerical implementation was used
based on [
59
]. The finite element method (FEM) was employed by solving the so-called
weak form: ZΩ
σijδuj,idV=Z∂ΩN
ˆ
tiδuidA, (1)
on a computational domain,
Ω
, with its closure,
∂Ω
, as the boundary, which is the image
of the underlying continuum body. We understand the Einstein summation convention
over repeated indices, where all Latin indices
i
,
j. . .
run from 1 to 3 (
x
to
z
) in Cartesian
coordinates. The test function,
δui
, is chosen from the same Hilbertian Sobolev space
as the displacement field,
ui
, known as the isoparametric Galerkin procedure [
60
–
63
].
As the deformation on a tensile test is small, we simplify the system and use the linear
strain measure,
εij =1
2ui,j+uj,i. (2)
Observed elasticity without rate effects justifies the use of Hooke’s law,
σij =Cijklεkl , (3)
where the stress tensor,
σij
, is linearly related to the strain tensor,
εij
, by the stiffness ten-
sor of rank four,
Cijkl
. As we want to distinguish topology (specimen geometry) effects
from the slicer caused anisotropy, we model the material as isotropic and homogeneous
(ideal bond between fibers and layers) by using the same engineering constants, Young’s
modulus,
E=
2150MPa ,and Poisson’s ratio,
ν=
0.35. For the space discretization, we im-
plement linear form functions and tetrahedron elements with a suitable mesh acquired by
an
h
-convergence analysis. All preprocessing steps, that is, CAD generation, boundary con-
ditions marking, and triangulation, have been accomplished in Salome 9.3. Computation is
achieved with the aid of open-source codes developed under the FEniCS project [
64
,
65
].
This implementation has been verified by closed-form solutions in [66,67].
Standard uniaxial testing simulation was performed for different geometries. On one
end,
x1=
0, the specimen is clamped along the axis,
u1=
0. This boundary is of Dirichlet
type. On the other end, the specimen is pulled by a given force per area, traction
ˆ
ti
, which
is a Neumann-type boundary condition applied on
ΩN
. An illustration is depicted in
Figure 5.
Figure 5. An illustration of the ISO 527-2 CAD model with boundary conditions.
Since linear elements for displacement were used, stresses are constant within the
elements. Postprocessing was done in ParaView 5.6 where the stresses were smoothed
due to use of the same mesh for the sake of better visualization. The accuracy of the
computation was obtained by an a posteriori error analysis based on the aforementioned
convergence analysis. Affected by these computations, preparation of the experiments was
made possible, as discussed in the following.
3. Results and Discussion
For a comparative study of structure-related effects on stress distribution during
a uniaxial tensile test along the
x
axis, we performed simulations with aforementioned
Materials 2021,14, 2556 9 of 19
geometries, namely, ISO527-2, modified-ISO, ASTM D3039, and ASTM D3039 angles.
Normal stress,
σxx
, is used for a qualitative analysis. The stress distribution is expected to
be constant within the part of the specimen used for assessment. Results are depicted in
Figure6.
Figure 6. Simulation of uniaxial tensile tests, stress distribution, σxx, (in color) resulted in all investigated geometries.
All geometries achieved, as expected, a constant stress within the gauge, away from
the clamped ends (tabs). Curvature in the ISO geometry allows for smoother transition by
preventing a stress concentration near the edges. Therefore, we designed ISO-modified
geometries with greater curvatures. This minor modification helps to manufacture a
specimen, where a failure within the transition zone is prevented. After a complete tensile
experiment until failure, a macroscopic crack is initiated in the appropriate region (gauge
region) as seen in Figure7a. In ASTM D3039 geometries, stress localization between the
tab and gauge sections causes a premature failure, as demonstrated in Figure7b.
An analogous procedure then follows by adding a curvature to ASTM D3039 speci-
mens. ASTM D3039 angle specimens perform better; however, many tested specimens did
fail by a macroscopic crack initiated within the transition zone due to stress concentration
between the tab and gauge sections.
(a)
Figure 7. Cont.
Materials 2021,14, 2556 10 of 19
(b)
Figure 7.
Experimental results (top) showing a brittle failure at a position directly related to the
stress distribution verified by computations (bottom). (
a
) For the ISO-modified specimen. (
b
) For the
ASTM D3039 specimen (zoom to the clamped end).
3.1. Material Characterization
The material was expected to be elastic. We provide the stress-strain curve of ISO
527-2 (performed with a ramp speed of 2mm/min) in Figure8. A linear elastic stress-
strain response is observed up to 20MPa and between 1 to 1.5% strain. A bare-eye visual
inspection qualifies the cross-section of failure as a brittle crack. No plasticity occurred in
the specimens. In order to determine its possible viscoelastic character, several uniaxial tests
were performed on ISO 527-2 specimens with the aforementioned (displacement controlled)
equipment. The maximum force was chosen by performing several measurements, and
we show herein results up to 20MPa equivalent stress. The strain rate dependencies are
compiled in Table 7with their corresponding stress-strain curves in Figure 9.
Figure 8.
Stress-strain curves recorded in uniaxial tensile tests until failure with a ramp speed
of 2mm/min.
Materials 2021,14, 2556 11 of 19
Table 7.
Strain rate-dependency, uniaxial tensile test results, normal stress at two normal strain
values,
σ0.05 =σ(ε=
0.05%
)
and
σ0.25 =σ(ε=
0.25%
)
, are used for determining the Young’s
modulus, Et.
σ(0–20MPa)
Strain-Rate σ0.05 in MPa σ0.25 in MPa Etin MPa
0.1 0.992 4.228 1796.373
0.5 1.109 4.458 1751.313
1.0 0.958 4.300 1702.541
2.0 1.041 4.280 1729.479
4.0 1.066 4.339 1738.880
8.0 0.970 4.379 1742.779
10.0 1.057 4.360 1738.423
16.0 0.949 4.331 1725.279
50.0 0.928 4.351 1735.809
100.0 1.080 4.263 1714.809
Figure 9. Stress-strain curves recorded in uniaxial tensile tests until 20MPa.
Here, all specimens were printed by the Slicing B parameters and all tests were carried
out at room temperature. Evidently, no rate effects are visible at room temperature for a
great range of strain rates in several orders. Therefore, we explain the stress-strain behavior
after 1.5% of strain by geometric nonlinearities under large displacement.
For determining the linear elastic properties, additional uniaxial tensile tests were
performed until failure. For all geometries with both slicing approaches, qualitatively,
results indicate that the material shows a linear elastic response until failure at the room
temperature. For a quantitative assessment, we used the slope of the stress-strain curve
between 0.05 and 0.25% to determine Young’s modulus, ultimate tensile strength, and for
all tests performed for different geometries are demonstrated in Figure10.
Relatively small errors denote an adequate repeatability owing to the chosen process
parameters. Below, we provide the equation for the error assessments,
R=S
m×100 =q∑n
i=1(xi−˜
x)
n−1
∑n
i=1xi/n×100, (4)
where the relative standard error,
R
in percent, is obtained by the standard deviation,
S
,
and the arithmetic mean of all samples,
m
. By
xi
we denote the value of the
ith
point in the
Materials 2021,14, 2556 12 of 19
data set, where
˜
x
is the mean value of all samples and
n
is the total number of data points.
Relative errors of the elasticity modulus and maximum stress are quantified in
Table 8
.
By comparing the Slicing A and B settings, obviously the Slicing A performs better than
the Slicing B regarding the obtained errors less than 2%. Figure 10 (top) and
Table 8(left)
show that deviation is small in determined Young’s moduli among all specimen types.
However, there are slightly higher deviations in maximum stress results with more errors
in Figure
10 (bottom)
and Table 8(right). As seen from Table 8, slicing settings have a more
significant impact on the accuracy than the chosen specimen type according to ASTM
or ISO. Hence, we conclude that all specimen types are adequate for obtaining Young’s
modulus. However, this fact changes in the case of the ultimate strength, as the error range
is affected by the specimen type as well as the slicing approach. As expected, the effective
mechanical properties of 3D-printed polymers are less than molded structures [
68
,
69
].
Processing conditions like low interlayer adhesion and occurring porosity within the
structure are understood to be the main reasons for this outcome.
Figure 10.
Quantitative comparison of the geometry effect in the case of tensile tests. Arithmetic
mean Young’s modulus, E, and maximum stress before failure.
Premature failure was observed among almost all specimen types. Failures of ASTM
D3039 and ASTM D3039 angles occurred in the acceptable areas (within the gauge section).
ISO527-2 and ISO-modified specimens had more premature failures in non-acceptable
areas, such as the shoulder (edge) sections. The Slicing A achieved shifting of the position
of the failure to the gauge section, especially in ISO and ISO-modified specimens, but not
Materials 2021,14, 2556 13 of 19
in ASTM D3039 and ASTM D3039 angle specimens. Among them, the best improvement
regarding the position of failure was observed in ISO 527-2 and ISO-modified specimens.
Table 8. Relative errors of elasticity modulus and max stress.
Relative Error of Elasticity Modulus in %Relative Error of Max Stress in %
Specimen Type/Slicing Approach Slicing A Slicing B Slicing A Slicing B
ISO-modified 2.851 3.305 2.179 4.945
ISO 527-2 0.604 4.462 4.331 6.293
ASTM D3039 2.480 4.087 2.566 6.644
ASTM D3039 angle 3.681 6.548 3.560 9.932
Tabs are necessary to provide an adequate clamping and curvatures are of interest
to distribute the stress appropriately. However, slicing may cause some problems in the
transition zone between clamping edges and gauge. As observed in many cases,the slicing
software begins the production from the clamped edges and finishes these areas before
continuing to the rest. This situation is depicted in Figure3b as a result of the Slicing B,
where three edges were produced initially. Obviously, the right side of the specimen had
different characteristics than the left side caused by a heterogeneous solidification. The rods
and layers did not connect to each other very well because of the temperature difference
on the solidified layer and the deposited melt. Thus, molecular diffusion decreased at the
interface between the edges and the main parts due to the low thermal energy. Experimental
results proved that this approach was one of the reasons of premature failures. Usually,
ISO527-2 and ISO-modified specimens were failed in those areas. Failure behavior of
specimens with the Slicing B are depicted in Figure11a,c(top).
(a) Geometries: ISO-modified, ISO 527-2, ASTM
D3039 angle. Failures of Slicing B.
(b) Geometries: ISO-modified, ISO 527-2, ASTM
D3039 angle. Failures of Slicing A.
(
c
) Geometries: ASTM D3039, Top: Slicing B,
Down: Slicing A
Figure 11.
Pictures of failures according to geometry types and slicing approaches. The specimens which have been broken
off out of the extensometer limit are provided only for visual observations. Their results are not included.
Materials 2021,14, 2556 14 of 19
On the contrary, in the Slicing A, the CAD model was sliced from one edge until
the end of the opposite edge in one line. This production route is depicted in Figure3a.
Apparently, the specimens resulted by the Slicing A had better interlayer and interrod
bonding due to homogeneous solidification than the specimens crafted by the Slicing B.
The Slicing B resulted in the onset of failure within (near the end of) the transition zone as
seen in Figure11a. Fortunately, the Slicing A shifted the onset to the gauge regime as seen
in Figure11b.
3.2. Geometry Effect
The geometry of specimens fails to have a significant effect on mechanical properties;
however, we have observed that the position of failure did depend on the chosen structure.
Stress distributions of FEA simulations are provided in Figure12.
(a) Stress distribution, σxx, of ISO 527-2 is measured
between x(0,10.8513,3)and x(150,10.8513,3)
(b) Stress distribution, σxx, of ASTM D3039 is
determined between x(0,4.4,7.5)and x(150,4.4,7.5)
(c) Stress distribution, σxx, of ISO-modified is
computed between x(0,13,3)and x(188,13,3)
(d) Stress distribution, σxx, of ASTM D3039 angle is
determined between x(0,4.4,7.5)and x(150,4.4,7.5)
Figure 12.
Stress results,
σxx
, of uniaxial tensile test simulations of different specimen geometries: (
a
) ISO527-2, (
b
) ASTM
D3039, (c) ISO-modified and (d) ASTM D3039 angle.
As also mentioned in its corresponding standard [
58
], ASTM D3039 and ASTM D3039
angle specimens did fail near the tab sections, although this situation was undesired.
Materials 2021,14, 2556 15 of 19
According to the suggestion in ASTM D3039, more than 40% of failures near the clamping
end set the design in question. The design of ASTM D3039 has been realized for fiber-
reinforced composites such that the localization of stress is different than herein. We
schematically show this situation in Figure13.
ISO 527-2 specimens have been specifically designed for molding and extrusion plas-
tics. In the case of 3D printing, specimens’ curvatures between tabs and gauge lengths
provide a force distribution possibly not aligned with the used slicing approach. Hence,
structure undergoes stress concentration leading to premature failure within the transition
zone. Moreover, if the specimens are produced with contour lines, there may be gaps
between contours and infills. This heterogeneity may lead to stress concentrations, too.
Increased curvatures in ISO-modified specimens resulted more adequate and reliable posi-
tioning of failure in repeated tests. Thus, we conclude that the suggested ISO-modified
geometry is to be used for characterization of mechanical response in structures manufac-
tured by 3D printing based on fused deposition modeling.
Figure 13.
Schematic illustration of the stress concentration on ASTM D3039 specimens. Sharp stress increase is measured
and visualized by finite element simulations (left). This outcome is validated with experimental investigations (right).
4. Conclusions
In this study, the mechanical characterization of polymers has been established in
fused deposition modeling, a common 3D printing technique. Four different types of
geometries were investigated for uniaxial tensile experiments.
The default slicing settings were analyzed and modified in order to obtain a unidirec-
tional structure to be used in a uniaxial tensile test. Two amendments are proposed:
1.
Excess travel lines were eliminated by using a newly developed printing path ap-
proach. The new travel path did not affect the unidirectional configurations of the
specimens. In this way, premature failures have been circumvented. A more homoge-
neous structure has been achieved at the macroscale.
2.
An adequate slicing sequence was developed and examined. The amount of weak
bond areas were reduced by this new approach. In this way, better solidification
and bond formation have been accomplished, leading to a homogeneous structure at
the microscale.
All these improvements have been achieved by modifying process parameters sug-
gested by the slicer software. Different slicing approaches have been compared in order to
quantify the amendment, as follows:
•
We performed tensile tests and found that the used PETG material was linear elastic and
performed a brittle fracture at room temperature (below the glass transition temperature).
•
We emphasize that the slicing technique influences the performance of the final
product significantly, such that the tensile test may lead to inaccurate results caused
by a premature failure. We have discussed possible slicing techniques, especially
for uniaxial tensile testing, and demonstrated the strength of an adequate approach
leading to adequate repeatability, as well as a small deviation in results.
Materials 2021,14, 2556 16 of 19
•
We emphasize the importance of the slicing settings by showing the failure along
printing directions in tensile experiments.
•
For a better understanding of premature failure, a digital twin of laboratory tests has
been realized by finite element method analysis. The same geometries have been
implemented for the uniaxial test, leading to stress concentration around the onset of
the failure. Based on this observation, a simple yet efficient design change is used in
ISO 527-2 and ASTM D3039 specimens.
•
The flat shape suggested in ASTM D3039 is appropriate for 3D printing; however, the
transition causes a stress concentration leading to premature failures, as demonstrated
in both the experimental and FEM results.
•
ISO527-2 and ISO-modified performs better, especially by tuning the slicing approach
as demonstrated herein. By using computations, we have suggested a modification
for a better performance regarding repeatability and accuracy in uniaxial tensile tests.
Author Contributions:
Conceptualization, A.Ö. and D.A.; methodology and software, A.Ö. and
B.E.A.; investigation, A.Ö.; manufacturing of specimens, A.Ö., C.V.; measurements and data curation,
A.Ö.; formal analysis, A.Ö., D.A., J.K., B.E.A.; writing—draft, A.Ö.; writing—review and editing,
D.A., C.V., J.K. and B.E.A. All authors have given approval to the final version of the manuscript.
Funding:
J. Kiendl was supported by the European Research Council through the H2020 ERC
Consolidator Grant 2019 n. 864482 FDM2.
Institutional Review Board Statement: Not Applicable.
Informed Consent Statement: Not Applicable
Data Availability Statement:
Raw data were generated at the TU Berlin, Institute of Material
Science and Technology. Derived data supporting the findings of this study are available from the
corresponding author (Arda Özen) upon request.
Acknowledgments:
We acknowledge support by the German Research Foundation and the Open
Access Publication Fund of TU Berlin. The authors would like to also acknowledge Tobias Bretz for
his valuable contributions on tensile testing, Oliver Löschke and Jens Butzke for helpful discussions.
Conflicts of Interest:
On behalf of all authors, the corresponding author states that there is no conflict
of interest.
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