Jia-Ruey
Ai
a,
Seokhoon
Jang
a,
Wyatt
Fink
b,
Seong H.
Kim
a and
Bryan D.
Vogt
*a
aDepartment of Chemical Engineering, The Pennsylvania State University, University Park, PA 16802, USA. E-mail: bdv5051@psu.edu
bDepartment of Engineering Science, The Pennsylvania State University, University Park, PA 16802, USA
First published on 16th January 2024
Multi-component structured filaments offer the potential for enhanced mechanical performance in 3D printed plastics. Here, the interactions between filament components in the core (polycarbonate, PC)–shell (polypropylene, PP) geometry are manipulated by light maleation (1%) of PP to understand how the inclusion of favorable polar interactions and potential grafting reactions at the core–shell interface impact the mechanical performance of the 3D printed parts. The elastic modulus of the 3D printed tensile bars is essentially independent of the shell selection for the fully isotactic PP (iPP) or maleated PP (miPP), but the strain at break is generally significantly improved with the miPP shell to increase the toughness of the printed parts for both flat and stand-on build orientations. This is counter to compression molded specimens where iPP is more ductile than miPP. The mechanical behavior in the flat orientation is consistent with long fiber composites, where the PC core essentially acts as fiber-reinforcement. Tribo-testing results indicate increased friction between miPP and PC through the interaction of the maleate anhydride group with the carbonate relative to the iPP with PC. This small increase in the interfacial interaction between the core and shell polymers with miPP increases the work required to pull out fibers of the stiffer PC from the PP matrix for the flat build orientation and more energy is required to delaminate the core from the shell, which is the loci of failure, when the stand-on build orientation is stretched. The subtle change in chemistry with a maleation of 1% of PP leads to a larger strain at failure and tougher parts due to the interaction with PC. These results illustrate that the selection of the polymers in structured filaments needs to also consider their potential intermolecular interactions including the potential for grafting reactions to best enhance the mechanical response of 3D printed parts.
An alternative route to enhance the properties of the interfaces between printed beads is to introduce crosslinks between the printed layers during printing.6 This leads to more isotropic properties of the printed part than those typically obtained with MEAM, but requires modifications to the printer and print process flow. Instead of covalent crosslinks, transient interactions associated with hydrogen bonding19 can also be used to improve the interfaces between printed beads without requiring large scale polymer diffusion needed to form entanglements. Nanoparticles are known to enhance adhesion generally20,21 and nanoparticle modified filaments provide significantly increased strength and higher modulus of 3D printed parts.22 However, the mechanical properties in the ‘z’ direction were not determined. An alternative route to improve the properties through engineering of the filaments involves structuring of polymers within the filament to provide enhanced flow at the surface of the filament, while the center rapidly becomes rigid to provide structural stability to the printed object.23–27 These structured filaments have most commonly been presented in terms of core–shell geometries,23–26 where a lower Tg polymer comprises the shell to enable improved formation of the interface between printed beads. Recent work has demonstrated a custom hotend die that can print core–shell structures from two commercial plastic filaments.28 This core–shell approach also enabled printing highly filled plastics where the shell is unfilled to provide higher mobility for interface development.29 Numerical simulations have been used to provide physical insights into process control for printing core–shell materials with MEAM.30 More complex geometries have been demonstrated using active mixing31 and custom dies.27 Post-print annealing of these structured filaments provides an opportunity to overcome the typical trade-off between dimensional accuracy and mechanical properties.27 However the variety of polymers used in these filaments have been generally immiscible, and this structuring of the filaments produces an additional interface within the printed part that arises from the two components within the filament. The role of the interaction between the core and shell polymers in the mechanical performance has not been examined in detail. For core–shell filaments, this interface has been shown to be effectively weak for most examined polymer pairs to lead to effective crack propagation along the core-shell interface when stresses are applied in the ‘z’ direction.25
In this work, the role of the interactions between the polymers in core–shell filaments on the mechanical properties of 3D printed parts is investigated experimentally while printing in flat (XY) and stand-on (YZ) build orientations. The goal of this work is to provide an understanding of how the adhesion between the core and shell polymers in these filaments impacts the performance of 3D printed parts to provide guidance for the selection of polymers in multi-component filaments. Here, a polycarbonate (PC) core and a polypropylene (PP) shell enable this investigation where the interaction between materials can be modestly impacted by maleation of PP (miPP). This change does not significantly impact the temperature for solidification of the shell material to provide a direct comparison. Maleation is known to improve the compatibility of polymer blends32 to improve their mechanical performance,33 including PC and PP where PC can be grafted to PP through maleic anhydride.34 Maleation improves the compatibility of PP with more polar polymers, such as poly(vinyl alcohol),35 typically through grafting with alcohols or amines that decreases the size of the phase segregated domains. The effect of maleation of polyolefins on 3D printed parts has not been investigated. Here, we quantify the influence of maleation (1 mol% of PP) on the adhesion between PC and PP via tribology measurements to rationalize differences in the mechanical performance of 3D printed objects using iPP or miPP shells in PC core–shell filaments. The insights gained from this work provide additional concepts for the design of multiple component structured filaments to enhance the toughness of 3D printed parts through polar interactions (maleic anhydride and carbonate).
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Fig. 1 Optical micrographs of cross-sections of the core–shell filaments. The core ratio is (a) 44.3% for PC-iPP and (b) 49.2% for PC-miPP for these images. |
The thermal properties were examined by Differential Scanning Calorimetry (DSC, TA Instruments Q2000). Thermograms were obtained under N2 with a flow rate of 50 ml min−1 to determine the processing conditions for the core–shell structured filaments. The resin samples (4–6 mg) were placed in an aluminum DSC pan (DSC Consumables, Inc.) and then hermetically sealed. The thermograms were recorded between 40 °C and 240 °C upon heating and cooling at 10 °C min−1 for 2 cycles. The first heating cycle removed the thermal history associated with resin manufacture conditions and aging. The thermal properties were reported from the cooling and 2nd heating branch as shown in Fig. S3.†
For comparison, tensile specimens of the individual components (PP, iPP and miPP) were prepared by compression molding with a heat press (Carver, model 3912) and an ASTM D638 type V die with 1.5 mm thickness. Prior to compression molding, a release agent (Stoner, silicon out mold release) was applied to the internal surface of the die to enable facile removal. Sufficient polymer was added to the die and a constant force (2200 N), equivalent to ≈1730 MPa, was applied for 5 min. The molding temperature was selected based on the polymer (PC 240 °C, iPP 200 °C and miPP 200 °C). The specimens were allowed to cool for 30 min prior to removal from the mold. Additional iPP and miPP sheets were also prepared with a 1.5 mm square mold (50 mm × 50 mm). PC spherical balls were fabricated with a customized mold (Fig. S3†). The PC pellets were placed in the socket and then molded following the same procedure for the tensile specimen.
The tensile properties of the printed specimens were measured using an MTS 50kN load frame system (Criterion model 43) with a 5kN load cell and video extensometer. The tensile tests were performed at 10 mm min−1. This rate is within the recommended crosshead rates for ASTM D638 testing (5–500 mm min−1). The true strain was captured using a video extensometer. For each print condition, a minimum of three specimens were tested. The results are reported as the average with an uncertainty associated with the standard error of the measurement determined by the variance associated with multiple measurements.
As shown in Fig. 2a, the combination of a relatively high Tg value, limited volume change upon solidification, and adequate stiffness of the glass for PC leads to good dimensional accuracy overall from the 3D scan. This contrasts with the iPP print (Fig. 2b) where there is clear warpage and bowing of the overall part with the front right and back left corners being bent upwards. Moreover, there is clear deformation of the overhang print at 75° where the upper part of the artifact is thinner than the digital design. This poorer print capability of iPP is not surprising given the reported challenges with printing polypropylene due to its large volume change upon crystallization40 and the variety of strategies implemented to reduce the deformation.41–43 Here, the iPP filaments were not formulated to improve the 3D printing performance. For the core–shell filaments that contain approximately ½PC and ½iPP shown in Fig. 2c, the dimensional accuracy is similar to pure PC except for the unsupported print rails where there is some added deformation. Visual inspection of the surface illustrates that the printed lines are not as apparent in the iPP/PC in comparison with the PC. There is some degradation in the dimensional accuracy when switching to miPP with the core–shell filaments (Fig. 2d). This is attributed to the increased flow of miPP as a result of the lower viscosity (Fig. S5†) and its lower modulus in the solid state. As the maleation extent is low (ca. 1%), it is not expected to significantly alter the flow at the core–shell interface from interactions between the maleic anhydride in miPP and the carbonyl in PC, although grafting reactions would broaden the interfacial width.
To better quantify differences in the printability of these different filaments, the hole printing in the standard specimens was examined. Fig. 3 illustrates the areal accuracy of the 12 different holes in the test specimens for the different filaments. These holes are associated with three different shapes, circles, squares and hexagons, as well as three nominal sizes (1, 2, 3, 4 mm) for each shape. The images used to calculate the hole areas are shown in Fig. S6.† For circular holes (Fig. 3a), there is good agreement between the model and print with PC for the 3 and 4 mm holes (>98%), while the holes printed with the other filaments tend to be appreciably smaller with ∼90% of the area at 4 mm and ∼85% of the area of 3 mm. This decreased ability to accurately print holes with polyolefin containing filaments may be associated with their lower surface energy (typically 31–32 mN m−1 in comparison with ∼40 mN m−1 for PC) and the volumetric change upon melting of the PP crystals that leads to an initially larger bead. However decreasing the hole size leads to a significant degradation in the hole accuracy for all of the filaments examined. At 1 mm design, core–shell filaments perform worse than the iPP or PC alone. This difference may be associated with the higher print temperature with the core–shell filaments than for iPP, which provides more opportunity for flow. In particular, the hole printed with the miPP/PC filament is nearly closed. The force for closure of the hole would be surface tension, but the viscosity of the polymer would act to limit the closure. The higher extrusion temperatures with the core–shell filament increase the time that the shell in the melt can flow. The differences in the hole size between the iPP and miPP shells are likely associated with differences in the viscosity where the melt flow index (MFI) of miPP (7.7 g per 10 min) is nearly 3 times as high as the MFI of iPP (2.7 g per 10 min) from the data sheets for the plastics. This is also consistent with the rheology from the small amplitude oscillatory shear (SAOS) as shown in Fig. S6† where the storage modulus for iPP is approximately double that of miPP at 0.01 rad s−1 and 200 °C. These data are consistent with a lower average molecular weight for miPP; reduction in the molecular weight of PP during maleation is common due to simultaneous beta scission occurring with the desired reaction.44
Similarly with a square hole (Fig. 3b), the relative accuracy of the hole size decreases as the nominal hole size decreases. However, the iPP filament in general leads to better areal accuracy. This difference is likely associated with the print path changes between a circle and a square. The largest effect of the filament selection is observed with the smallest (1 mm) hole, where the miPP/PC filament does not perform as well as the iPP/PC core shell filament. This behavior is consistent across the three hole shapes examined for the 1 mm hole, which points towards the viscosity of the shell being important to maintaining small holes in printed structures especially considering the higher printing temperatures.
As the viscosity difference between miPP and iPP appears to impact the accuracy of printed holes, the surface finish may be impacted by this selection as well. Fig. S7a† illustrates a representative block printed with the two core–shell filaments. Examination of this surface by optical profilometry reveals a periodic surface associated with the layering of the printed beads. Fig. 4 illustrates the surface morphology of the printed core–shell filaments at the different print temperatures examined. Visually the surfaces are similar with clear degradation in the aligned parallel bead structure at 300 °C when using miPP/PC. Although the full optical profilometry scans can quantify the surface morphology through average roughness (Ra), this averaging loses the directionality associated with the print. As the flow of the polymer will reduce the peak to valley amplitude between adjacent beads, examination of the profiles solely in this direction will provide insights into the polymer flow that occurs transversely to the print direction. The Ra value in this case will be predominately influenced by the interbead height variation associated with the print path.
Fig. 5A illustrates the differences in the 1D profiles between miPP and iPP shells when printed at 300 °C. Line cuts are perpendicular to the printed beads at a spacing of 150 mm between cuts. There are 10 curves shown in Fig. 5A for each print condition. The amplitude of the oscillations associated with the printed beads decreases with the use of miPP in the shell. This is consistent with increased flow as might be expected due to the higher MFI and lower viscosity. Analogous tensile test profiles associated with the other print conditions are shown in Fig. S7.† The absolute value of the deviation of the height from the average of these 1D profiles provides a roughness metric (Ra) that is biased to the interbead roughness as opposed to intrinsic roughness along the printed bead. Fig. 5B illustrates this roughness as a function of print temperature for the two filaments. At the lowest temperature printed (260 °C), there is no statistical difference in Ra between iPP and miPP shells; this is somewhat surprising given the lower viscosity for miPP, which should promote additional flow. However, rapid cooling during the print likely limits the potential for additional flow. At 280 °C, there is a modest decrease in the average roughness with the miPP shell, while the roughness using the iPP shell is essentially unchanged. This is consistent with the additional flow of miPP due to its lower viscosity to level out the height differences associated with the interbead region. However increasing the temperature to 300 °C leads to an increase in the roughness, most notably for the iPP shell as shown in Fig. 5B. This increase in roughness is counter to expectations that increasing temperature will tend to reduce the interbead roughness. Examination of the 2D profilometry data provides hints of the origins of this increased roughness. There are nodules that appear on the surface of the 300 °C prints (Fig. 4) that are in some cases greater than the peak to valley height between beads. The bead morphology is still clearly distinguishable for the iPP shell, but this is significantly reduced for miPP. Thus, there is a significant increase in roughness with iPP/PC at 300 °C due to the combined effect of the nodules and the interbead morphology. Nonetheless, the similarity in roughness indicates that large scale flow between adjacent beads does not appreciably depend on the melt viscosity of the shell polymer over the short timescales between printing and solidification.
Although the near net shape manufacture, and thus dimensional accuracy, is a key attribute of 3D printing, the performance of plastic parts manufactured by MEAM tends to be limited by their mechanical attributes.2Fig. 6 illustrates representative stress–strain responses for the two core–shell materials as a function of build orientation and extrusion temperature. All of the stress–strain curves for the specimens tested are shown in Fig. S8† for iPP/PC and Fig. S9† for miPP/PC core–shell printed tensile specimens. When printed in the flat build orientation (Fig. 6a), the tensile behavior for the miPP/PC and iPP/PC core–shell materials are qualitatively similar for the different extrusion temperatures examined. There is initially a sharp increase in stress up to approximately 10% strain, which is associated with the elastic modulus. The specimens yield and then extend at lower stress to failure. The effect of miPP versus iPP in the shell appears to be primarily associated with the post-yield behavior with the miPP/PC material being more ductile for most temperatures examined. This difference can be rationalized in terms of the interactions between the shell and core polymers, but the effect of viscosity differences of the shell polymers cannot be discounted out of hand. Incomplete filling of the printed object to leave voids has been used to explain differences in mechanical properties.45 However, prior work on core–shell printed materials has shown a significant reduction in the void fraction when there is a large difference in the solidification temperature of the shell and core, even in cases where the viscosity of the shell is greater than iPP,25 so it is unexpected that there will be significant differences in the void fraction with the iPP vs. miPP shell. At low strain, the elastic response is dominated by the PC fibers that extend through the material, but at high strain the adhesion between the core and shell will be important to the mechanical response as the shell is more ductile while PC provides a rigid component. PC and iPP are immiscible, which provide low adhesion. The maleation of PP provides some favorable polar interactions between PC and miPP as well as the potential for grafting34 to increase the adhesion strength between the core and shell. However, the limited extent of maleation (1%) reduces the probability of these interactions and large differences in the mechanical structure were not expected.
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Fig. 6 Representative stress–strain curves of iPP/PC and miPP/PC 3D printed in (a) flat and (b) stand-on orientations as a function of temperature. |
For the stand-on orientation (Fig. 6b), there is a marked decrease in the strength of the printed sample for both core–shell materials. Even for single component filaments, the mechanical properties in this print orientation tend to be inferior since the stress is applied almost exclusively through the interfaces between printed beads.2,15,46 This orientation is effectively all welds, which due to the kinetics of welding and the thermal history of 3D printing tends to be relatively weak.6 However in this case, the decreased strength can be rationalized by considering these printed objects as fiber composites. In the flat orientation, the stronger PC core fibers are aligned in the direction of the force to effectively carry the applied load, similar to designs that use continuous fiber composites.47 In the stand-on orientation, the PC fibers are perpendicular to the applied force and thus cannot carry the stress. Unlike most plastics printed in the stand-on build orientation, the failure of the printed parts does not occur at the bead–bead interface, but rather within a bead with the PC core remaining intact on one of the fracture surfaces. This behavior has previously been observed for other immiscible polymer pairs with core–shell filaments when printed in the stand-on orientation.25 The PP matrix, which is significantly weaker than PC (see Fig. S10 and S11†), controls the stress–strain response. As the PC fibers align perpendicular to the tensile direction, these could effectively work as defects in the sample to promote the growth of cracks. We attribute the increased strain at break for the miPP/PC specimens to the improved adhesion of miPP to PC which reduces the probability of catastrophic failure at this interface. Images of the specimens after failure are shown in Fig. S12.†
Fig. 7 summarizes the mechanical properties of the printed objects. The elastic modulus of 3D printed parts is the most common mechanical metric to describe their performance.9,48 Due to the printing temperatures being defined by the higher Tg of the PC core of the filament, the effect of the print temperature on Young's modulus is small as shown in Fig. 6a. Similarly, the selection of the shell (iPP vs. miPP) did not significantly impact the Young's modulus of the printed parts. However, the print orientation had a significant impact on Young's modulus. This can be rationalized in terms of continuous fiber composite theory.49 For the tensile specimens printed in the flat orientation, there are 7 layers of beads that are deposited in an alternating 0° and 90° orientation relative to the elongation direction. Using this composite theory, the predicted elastic modulus of the flat orientation is 1465 MPa for iPP/PC and 1418 MPa for miPP/PC based on the measured properties of the individual components with compression molding (horizontal lines in Fig. 7a). These predicted values agree well with the properties obtained with the core shell filaments, which suggests that there are only limited voids in the printed objects and the difference in viscosity between miPP and iPP are not responsible for the differences in mechanical performance. The actual values are higher than the prediction due to the printed perimeter being effectively aligned at 0° for all layers. In contrast, the PC fibers do not extend in the elongation direction when printed in the stand-on orientation. Thus, the weaker continuous PP phase deforms under the load around the stiff PC fibers. The load bearing capabilities of the PC fibers that run perpendicular to the applied stress are dependent on their coupling with the continuous PP phase. As the adhesion between PP and PC is not expected to be high, the PC effectively acts as a defect to mostly reduce the effective area supporting the load, which is consistent with the modulus of the printed stand-on specimens being lower than the compression molded iPP or miPP. This poor adhesion between PP and PC is consistent with attempts at lap shear joints, which failed upon handling the polymers used in the core–shell filaments examined (Fig. S13†). Despite the lower modulus of miPP, the stand-on printed specimen using miPP tends to exhibit a higher modulus in the stand-on build orientation. This is likely associated with the improved adhesion of the polypropylene matrix to the PC core to increase the ability of the PC core fibers that are perpendicular to the applied stress to hold some of the load.
In addition to the elastic modulus, the yield stress of the printed parts provides another key metric to describe the mechanical performance of these core–shell materials as shown in Fig. 7b. Similar to Young's modulus, the yield stress is not strongly dependent on the extrusion temperature or selection of the shell PP. The build orientation is however strongly coupled to the yield stress, which can be rationalized in the same manner as Young's modulus in terms of the PC forming fibers within the printed part that act like a composite material. However, the yield stress for the flat orientation is not appreciably greater than those of pure PP materials. The ability to withstand an applied load is significantly less when printed in the stand-on orientation, which is likely associated with the PC fibers effectively acting as defects.
The toughness of the specimens was determined from the area under the stress–strain curve prior to failure. As shown in Fig. 7c, the miPP shell tends to lead to enhanced toughness at lower temperatures when printed in the flat build orientation, while the toughness is statistically independent of shell PP selection for extrusion at 300 °C. There is more than an order of magnitude reduction in the toughness of the printed materials when comparing the two build orientations. This is a result of two factors: the decreased Young's modulus (Fig. 6a) and the reduced elongation at break (Fig. 7d). The failure for the flat build orientation occurs at strains between that of pure PC and PP, while the failure occurs at lower strains with the stand-on orientation. This inferior strain performance with the stand-on orientation is consistent with the PC fibers acting essentially as defects where a crack can readily propagate along the PC/PP interface. In general, the miPP shell leads to 3D printed parts with a higher strain at break in both build orientations. This behavior suggests that the adhesion of miPP to PC is better than that of iPP to PC.
To better understand the differences, the fracture area of the tensile bars was examined with optical microscopy. Fig. 8 illustrates the differences in the morphology of the failure based on the shell (iPP vs. miPP) and build orientation (flat vs. stand-on) at an extrusion temperature of 280 °C. Additional micrographs for other temperatures are shown in Fig. S14 and S15.† Samples printed in the flat build orientation have continuous lines of the core along the deformation direction. The core acts as a long fiber reinforcement with the PC core materials extending beyond where iPP or miPP failed, but the morphology where the matrix failed is dependent on the shell material. For iPP/PC materials, iPP fails in nearly a straight line across the width of the specimen (Fig. 8A). Conversely, the miPP failure is jagged with changes in the angle occurring near the center of a PC core fiber in the specimen. This difference can be explained in terms of the interaction of the matrix polymer with the core fibers. With iPP and PC, iPP is non-polar with alkyl units (CH, CH2 and CH3) that can only act through van der Waals interactions with PC, which contains polar components within the monomer repeat. Without specific favorable interactions between the highly non-polar iPP and PC, the matrix can fail around the fibers and there is nothing to blunt or slow the crack tip. miPP adds a small fraction (1%) of a polar moiety that has the potential to crosslink with any hydroxyl end groups on PC. The polar interactions between maleic anhydride in miPP and the carbonyl in PC will increase the energy to separate the core PC from the shell miPC. This interaction will increase the force required to separate the matrix polymer from the PC fiber but the low functionalization of miPP likely produces heterogeneity in the local force required for separation and thus cracks will not propagate straight across the specimen. These are consistent with the higher toughness of miPP/PC at 280 °C (Fig. 7c). Conversely, the failure of specimens printed in the stand-on build orientation is a nearly clean fracture across the width. The failure surface is nearly parallel to the printed line with iPP/PC with a small deformation of some fibers near the fracture surface (Fig. 8C), while there is some additional deformation near the fracture with miPP/PC (Fig. 8D). This difference could result from the improved coupling between miPP and PC that allows for the stress near the initial crack to spread over a larger area to deform the fracture surface; this mechanism could modestly increase the strain at break (Fig. 7D), but this change does not appear to be statistically significant.
To better understand the differences in the interfacial interactions with PC that are suggested by the fracture images and mechanical behavior, a ball-on-flat reciprocating tribometer was used to quantify the coefficient of friction (COF) between a PC ball and a flat PP sheet. Based on the modulus of each material (Fig. S7a†), the nominal Hertzian contact pressure, deformation, and diameter are estimated to be 27–29 MPa, 5.4–5.9 μm, 90–94 μm, respectively. Fig. 9 illustrates the difference in the COF. The surface roughness is similar between miPP and iPP as shown in Fig. S16.† The higher COF for the PC-on-miPP pair than the PC-on-iPP pair is indicative of stronger interactions between PC and miPP as a result of maleation. The higher COF must be due to the enhanced interfacial interactions which could originate from chain entanglements at the sliding interface through grafting of PC to the maleic anhydride in miPP or simply due to stronger interactions between polar groups (maleate in miPP and carbonate in PC).50,51 Due to the glassy nature of PC, the low maleation extent (1%), and the reported reaction kinetics for reactions of maleic anhydrides with −OH,52 it is unlikely that the increase in friction is due to chain entanglements at the sliding interface. In any case, the difference in the COF is consistent with the observed mechanical behavior of the core–shell filaments. It should be noted that the adhesion between the PC and PP is weak as indicated by the failure of lap shear specimens prior to testing. Nonetheless, the minor improvement in the adhesion between the core and shell with maleation leads to primary improvements in the extension at break. Along with the COF result, these indicate that some ‘enhanced’ interfacial interaction, likely through polarity, between the core and shell materials is beneficial for the overall mechanical response. The use of partially miscible polymer pairs (PC and ABS) in structured filaments has been shown to lead to improved mechanical properties for the stand-on orientation through post-process annealing at a temperature between the Tg values of the two polymers,15 but the strength of the 3D printed part remains less than that of the injection molded ABS. Prior work with core–shell filaments suggested that poor adhesion between the core and shell polymers could be beneficial for impact energy dissipation when printed in the flat or edge-on build orientation.25 Further investigations are needed to better understand how the adhesion between the components in structured filaments influences the mechanical properties of 3D printed parts, but this work does illustrate the potential advantages of improving the adhesion between the polymeric components.
Footnote |
† Electronic supplementary information (ESI) available: Additional experimental details and data, including a schematic of coextrusion for filament fabrication, DSC thermograms, mold design for tribology measurements, optical micrographs of the surface of printed objects, surface topology scans, stress–strain curves, and images of fractured tensile bars. See DOI: https://doi.org/10.1039/d3lp00168g |
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