Influence of Screw Configuration and Processing Temperature on the Properties of Short Glass Fiber Reinforced Polypropylene Composites.
Over the last decades, the use of glass fiber reinforced composites has been quite high, and it is still continuously growing as the need of structural and lightweight composites and high output manufacturing processes is in great demand in multiple fields of industry. As the fiber length, besides matrix type, volume content, fiber orientation, and interfacial adhesion, is directly correlating with the composite properties [1-3], long glass fiber composites, that is, unidirectional reinforced ones, provide high mechanical performance for various applications. However, they are still complicated and expensive to produce as the production requires high-level automation and complex process steps. On the other hand, short fiber reinforced composites are produced in one or more melt processing steps where heat and shear energy provide mixing and adhesion between fibers and matrix. The semifinished product gained from such production route is a pellet, which can be readily molded into parts or extruded into profiles, just to name some examples.
Looking at the different possible reinforcement fibers available for composite reinforcement, one can find several options. In general, there are natural and synthetic fibers, which can be further subcategorized. In case of the natural fibers, cellulose is often the basis, as these fibers grow as the structural reinforcement in plants. Another variant is protein, as in animal hair fibers . Both of these materials suffer from limited thermal stability, which is a drawback in polymer processing , where temperatures beyond 200[degrees]C are often necessary either due to the polymer type itself or, in the case of polypropylene for example, to shift polymer viscosity to a more favorable lower value due to the temperature increase.
In case of synthetic fibers, also a wide variation of materials can be found. There are different polymeric fibers, for example, polyester or polyamide, which are not very often used as reinforcement due to the fact that these fibers exhibit anisotropic properties in the fiber longitudinal direction and across. Apart from that, they are also more common in textile application, where these have to be crimped, which is unfavorable for composite reinforcement. Furthermore, there are cellulose-based fibers, which are produced via dissolving and spinning the cellulose again, so called man-made fibers. These have some reinforcement applications, for example, Rayon fibers for tires , but again cellulose is a material with limited thermal stability (and is also exhibiting anisotropic behavior). The last group here is inorganic fibers such as glass or carbon fibers. While carbon fibers exceed the properties of nearly all other fibers in the field , these fibers are costly and also show some drawbacks in regard to health-related issues as well as in disposal at the end of their life cycle. The most widespread fiber type is glass, more specifically E-glass fibers, which show a favorable balance between their relatively high properties and reasonable cost. Furthermore, they are isotropic in regard to the fiber properties due to the production process, where an amorphous melt is spun by gravity through a die to form thin fibers. Also, they tolerate thermal load up to the temperatures where the fiber surface treatment, the sizing, starts to degrade, but they suffer from the high shear stress during melt processing and therefore break to shorter lengths. The influence of fiber length on composite properties are well known [8-10], that is why continuous efforts on research and development are concerned with the improvement of existing low-cost production processes of short fiber reinforced plastic parts to preserve the fiber length during the whole manufacturing.
Fiber degradation mechanisms have been treated in many scientific contributions. Early studies show that fiber breakage occurs mostly when fiber bundles start to separate in filaments, that is, in the early stage of the dispersion process of the fibers . Investigations on fiber length distributions at different positions along the screw show that the aspect ratio of the fibers reduces along the extruder . Most of the fracture occurs in the melting zone; however, adding the fibers after the melting zone does not necessarily preserve fiber length, since sufficient shear also takes place in the mixing zone to break the fibers . Two types of fiber breakage mechanisms have been observed in the melting zone : first, in the interface between the solid bed and barrel, fibers that are close to the surface bend and break due to shear forces caused by the molten film. After that, these broken fibers flow with the melt and can experience further damage due to postbuckling deformation. A decrease of fiber length after passing through the die has been observed as well, which indicates a further breakage of fibers in the die flow .
The correlation between process parameters during extrusion and fiber degradation has been widely studied as well. It was found that more severe screws with high intensive mixing  and higher screw speeds [17-19] lead to shorter fiber lengths. Another study concluded that fiber length decreases less than proportionally to the increment on screw speed due to the lower shear stress between the fiber and matrix caused by a less viscous polymer . On the contrary, increasing the temperature profile during extrusion contributed to a slight increase in fiber length . Longer fibers could be observed when the feed rate was increased as well, due to a shorter residence time at a constant screw speed [21, 22]. Other influence is the increase of fiber volume content, which leads to an increase in the extent of fiber breakage due to more intensive fiber-fiber interactions [10, 12, 23, 24].
In order to preserve fiber length other methods rather than the variation of processing conditions have been applied, as for example, the use of a single-step extrusion without prior compounding . In this case, a single screw with a steep compression zone was used, yielding very good homogeneity and less fiber damage compared to the process including a compounding step. However, this method also shows some limitations such as limited glass fiber content and the necessity of an accurate control of the feeding of the screw, otherwise homogeneous fiber distribution cannot be achieved.
Knowing the optimum processing parameters is important in terms of cost and energy saving as well as product optimization and structural performance. Therefore, the aim of this work was to carry out an integral study of the influence of screw configuration and process parameters on the mechanical properties and residual fiber length distribution of short glass fiber reinforced polypropylene, to gain further insight on how to preserve the glass fiber length in the process. Furthermore, a micromechanical model was applied for the evaluation of the interface in regard to any processing influences. A potential upscaling from a laboratory extruder to a production machine does not fall within the scope of this article.
MATERIALS & METHODS
To keep the focus on the process parameters all experiments were performed with the same material mixture, in this case a homopolymer polypropylene (Borealis), short glass fiber HP3299 (PPG Fiber Glass Europe GmbH), and an Orevac CA 100 (Arkema) compatibilizer. The detailed properties of the used materials and the proportion of the mixture can be found in Table 1.
From these materials, composites were produced using a co-rotating twin screw extruder (ThermoPrism TSE 24HC) with 24-mm screw diameter and a processing length of 28 L/D, equipped with a gravimetric dosing system, where all three components were dosed individually. While the polymer and the compatibilizer were dosed in the intake of the extruder, the glass fibers were added through a side feeding port at a later position, approximately at 14 L/D into the then already molten polymer. The throughput was kept constant at 10 kg/h with a rotational speed of 400 rpm. The melt strands were cooled by a water bath, cut to granules my means of a strand cutter and dried at 80[degrees]C for at least 2 h. Two series of experiments were performed:
a. To evaluate the influence of the processing temperature the screw configuration was kept the same (standard screw configuration), while the temperature profile along the screw was altered. The applied temperature profiles can be seen in Table 2.
b. To evaluate the influence of the screw configuration, 4 different configurations with different mixing elements in screw zones 5 and 6 were used. These mixing elements are described in Fig. 1, where (f) and (g) are 2-flighted mixing elements with 0[degrees] and 90[degrees] offset, respectively, (h) is a tooth mixing element formed from a pitch-less ring with milled holes around its circumference, and (i) is a 2-flighted conveying element.
The used elements are plugged on a hexagonal shaft in the order depending on the screw configuration according to the different sequences as shown in Fig. 1 (a) the use of two TME elements on mixing zone 5 ensures a shear gentle mixing with neutral conveying properties, (b) The A90[degrees] kneading block provides very shear-intensive mixing also with neutral conveying properties, due to a high stagger angle of 90[degrees] between the individual mixing elements, (c) By lowering the stagger angle to 30[degrees]. a gentler mixing with some conveying is achieved, (d) Here two shorter blocks are combined, namely an A90[degrees] kneading block with a neutral character and a 60[degrees] forward conveying kneading block, (e) Zone were kept the same for all configurations and also combine conveying positive and neutral conveying kneading blocks with different stagger angles (30[degrees]-60[degrees]-90[degrees]).
In order to test the reproducibility of the results with the different screw configurations and temperature variations, some of the trials were repeated and are shown with the suffix "r" in the results.
From the produced granules, specimens were made by means of an injection molding machine (Engel victory 80/330) following ISO 3167. The processing parameters were kept the same for all formulations (according to Table 3) in order to minimize the influence of injection molding on the compounded materials.
Before testing, the samples were stored for at least 88 h under standardized conditions, at a temperature of 23[degrees]C and a relative humidity of 50%. These standard specimens were used for mechanical characterization, in this case tensile test referring ISO 527-4 (Zwick TC-FR020 TH, 3 replicates, specimen type 1A, test speed 5 mm/min, modulus test speed 1 mm/min, modulus and tensile strength in one sample, test temperature 23[degrees]C) and Charpy impact test following EN ISO 179-1 (Zwick 5,113.300, 5 replicates, specimens according to ISO 3167, impact speed 2,93 m/s, break modulus C, test temperature 23[degrees]C, type of notch A). To further evaluate the impact of the temperature on the flow behavior, the mass volume rate (MVR) was measured by means of an MVR test device (Zwick 4,106) with different melt temperatures but constant test parameters (2 replicates, heating time 5 min, load weight 2,16 kg, measure distance 5 mm). Fiber length measurements were carried out after washing the parallel parts of the universal test specimens at 625[degrees]C in a thermogravimetric analyzer Leco Makro-TGA to yield the fibers. About 0.5 g were dispersed in water by ultrasonication in an ultrasonic bath and afterward analyzed by means of a Horiba Partica LA 950 V2 laser diffraction particle size distribution analyzer, where directly before the measurement started, the sample was sonicated again, to investigate the fiber length distributions. For scanning electron microscopy (SEM) observations universal test specimens were notched and broken after they were submerged in liquid nitrogen for a short time. A Vega Tescan II SEM was used for taking the micrographs after sputtering the fracture surface with gold to prevent electric charges. Material density was measured according to EN ISO 1183 (test temperature 23[degrees]C, ethanol as submersion liquid) in order to check the comparability of the compounded materials and the even distribution of the components. In order to check for interfering effects due to degradation isothermal oxidation induction time (OIT) was measured according to EN ISO 11357-6 with a Mettler Toledo DSC device.
RESULTS & DISCUSSION
Influence of Processing Temperature
In this section of the article, we discuss the influence of the processing temperature in compounding on the properties of the glass fiber reinforce polypropylene. All mixtures show similar densities, regardless of the used temperature profile. The values range between 1,112 and 1,123 g/[cm.sup.3], and the results exhibit low scattering. Hence, we can assume that the specimens exhibit comparable fiber contents and due to the low scattering, are mostly free of porosities or shrink holes, which would interfere with any further measurements and their interpretation.
The tensile test show a strong dependency between tensile modulus and strength versus the temperature profile (Fig. 2). In fact, the tensile strength and modulus are increasing with increasing barrel temperature. We can also observe that the results from the two runs at 200[degrees]C with the same screw at different days yield nearly the same results, only differing a little in the scattering in tensile straight. This shows the reproducibility of the setup. As stated above, we see that with increasing temperature and therefore decreasing melt viscosity, which is also represented by the melt volume flow rate in Fig. 4, the mechanical properties are increasing, likely due to the reduced glass fiber damage through a less viscous melt.
Another interesting fact here is that the 200[degrees]C temperature profiles, where only the barrel zones 4 and 5 are raised or lowered 20[degrees] C exhibit very similar mechanical properties to the profiles, which are completely set to 180[degrees]C and 220[degrees]C, respectively. We suspect this being due to the fact that the glass fibers are dosed into barrel zone 4, and in this and the following screw sections the fiber bundles have to be dispersed into the polymer by the kneading elements there. After that, mainly conveing is carried out in the screw, which is less critical to break the glass fibers, which are then already dispersed in the melt. Therefore, the temperature profiles where only these zones are set to 220[degrees]C (i.e., 200 + 20[degrees]C) and the 180[degrees]C (i.e., the 200-20[degrees]C) are giving very similar results. In addition, the glass fibers also act as a heat sink, therefore it is beneficial for preserving their length to increase the melt temperature, at least at the barrel zones where the fibers are introduced into the melt.
Similar results as for the tensile properties can be found in the Charpy impact test results (Fig. 3). The impact strenght increases with increasing melt temperature and also a correlation of the temperature shift in barrel zone 4 and 5 with the corresponding temperature profiles where all barrel zones are the same temperature can be found. Again, the explanation can be the lower viscosity of the polypropylene, as shown in Fig. 4 as the relation between the melt flow volume (MVR) and the capillary temperature, leading to a preservation of the glass fiber length. In order to check the composites, especially the matrix, in regard to thermal degradation due to the different processing temperatures, we carried out oxygen induction time (OIT) measurements. The OIT values for all the samples lie between 13 and 11 min, with no clear trend to be seen from these data. Therefore, we conclude that the thermal load in the investigated range from 180[degrees]C to 260[degrees]C did not affect the samples differently and therefore has no influence on the results.
Fiber length measurements support the interpretations of our findings so far, as it can be seen in Fig. 5, where the cumulative fiber length distribution is plotted. Fiber lengths are very similar in the lower length region, which is due to the fact that the method applied does not distinguish between length and diameter, and the fibers used in this study exhibit diameter of approximately 15 [micro]m, therefore being found in this region. For higher lengths, that is, in the region between 30 to 400 [micro]m, the distributions show a direct correlation between fiber length and temperature; the higher the melt temperature, the higher the fiber length. For better visibility, this is also shown zoomed in the insert. There is also some scattering present, which can be seen from the fiber length distribution of the barrel temperature setting 200 + 20[degrees]C, which shows a dent at around 700-[micro]m fiber length. This could be due to fiber agglomerates and is not considered here further.
Another indication, that the findings from the different temperature settings are comparable, is found from the SEM micrographs. In all the micrographs shown in Fig. 6, no differences can be seen in terms of microstructure. First of all, we can observe that a proper mixing of the glass fibers was achieved regardless of the used temperature profile. Fiber distribution is considerably uniform for every sample and no fiber bundles can be found. In all the samples, fiber pull-out as well as fiber breakage is present and also matrix residues adhering to the fibers can be seen. In Fig. 6 by 220[degrees]C and 260[degrees]C, a fiber pull-out and a fiber covered with matrix have been marked through arrows in order to point out the presence of both effects. There is no variation of fiber pull-out number with temperature change. This means, that the set processing parameters do not influence the interfacial interactions negatively, regardless of temperature setting chosen.
Influence of Screw Configuration
In this section, the influence of the different screw geometries investigated in this study on the properties of the composites and fiber length of the glass fibers is discussed. As the first parameter, the density of the composites was checked to ensure the comparability of the fiber contents. The density of all the composites lies between 1,112 and 1,121 g/[cm.sup.3], therefore a comparison is valid.
With lower shear intensities applied, the fiber length should be preserved in the process, therefore resulting in higher composite properties. We can observe the same correlation for tensile (Fig. 7) and Charpy impact properties (Fig. 8). The highest elastic and impact properties are yielded by the low shear intenstiy configurations. Both the TSE, due to the short screw length of these elements compared to the other screws, and the 30[degrees] stagger angle kneading block, are the least shear intensive designs, which causes less damage to glass fibers and thus, better final properties of the composites. More shear intense configurations, namely the 90[degrees] and standard configurations, feature lower elastic modulus and tensile strength as well as lower impact properties. It should be noted, that mixing with the standard configuration results in the lowest properties, even though the most shear intensive combination of elements, that is, the A90[degrees] angle kneading block is reduced to half of the length compared to the 90[degrees] configuration. The combination of a 60[degrees] angle forward conveying elements leads to a higher shear effect. This could be due to less conveying effect of the 60[degrees] mixing block before the high shear 90[degrees] block, resulting in higher residence time of the material in the high shear intense section. Another possible explanation could be that there is a combined effect due to the medium shear intensity of the 60[degrees] mixing block achieved through the gaps in between elements combined with the following high shear 90[degrees] block.
Here, the OIT was again used to check the influence of the process on the stability of the polypropylene. The OIT values lie between 12 and 7 min (Table 4). Although no clear trend can be identified, one can see that the screws with nominally lower shear input, that is, the ones with 30[degrees] kneading blocks and TME seem to yield shorter OIT than the other screws used, which should exhibit higher shear. This could be due to the lower shear heating brought into the melt by the screw, which has to be compensated by other mechanisms. This results in higher shear and viscosity due to the lower melt temperature and in succession, in higher depletion of the stabilizer.
In terms of shear intensity, the residual fiber length correlates with the estimated shear loading, with some scattering effects. As shown in Fig. 9, the standard screw exhibits the shortest fiber lengths, followed by the screw with the 90[degrees] kneading blocks and followed by the screws with the 30[degrees] kneading blocks and the TME elements. There is some scattering present, that is, for the two experiments with the 30[degrees] kneading blocks screw; the repetition of the sample yields shorter fiber lengths. This could be due to some measurement influences, as the first measurement of the 30[degrees] configuration exhibits an unusual dent in the region of 500-(tm fiber length and above. This could arise either from a portion of longer fibers or from some measurement uncertainties, like fiber agglomerates, or due to scattering present in the fiber length in the composite itself due to the lower shear input. We did omit the SEM micrographs in this section, because these do not give any additional information.
In order to evaluate the composites in this study in regard to their interfacial shear strength (IFSS) and to see any influences of the processing on the interfacial interaction, we applied a "rule-of-mixtures" model, based on the work from Kelly and Tyson , as shown in Eq. (1). The model is valid for subcritical fibers only, as we expect not to have supercritical ones, and the influence of porosity is not accounted for, because this influence is negligible for injection molded universal test specimen, as these show porosities below 1% due to the high pressure applied in the process.
[[sigma].sub.c] = [[eta].sub.0][[tau].sub.c]l/d [V.sub.f] + [[sigma]'.sub.m](1-[V.sub.f]) (1)
where [sigma] is the ultimate tensile strength (MPa) with indices c for composite and m for matrix, l and d are the fiber length and diameter ([micro]m), respectively, [[eta].sub.0] is a fiber orientation factor, approximated with 0,75 , [V.sub.f] is the fiber volume fraction, and [[tau].sub.c] is the interfacial shear strength (IFSS in MPa).
Provided that there are no porosities, which can be assumed from the injection molded specimens as explained above, the fiber volume fraction can be calculated, using Eq. (2):
[V.sub.f] = [[rho].sub.c] -[[rho].sub.m]/ [[rho].sub.f] - [[rho].sub.m] (2)
where [[rho].sub.c], [[rho].sub.m], and [[rho].sub.f] are the densities of the composite, matrix and fiber, respectively. Inserting [[rho].sub.c] = 1,118 g/[cm.sup.3] (average measured value), [[rho].sub.m] = 0,91 g/[cm.sup.3] (extracted from the material's data sheet), and [[rho].sub.f] = 2,6 g/[cm.sup.3]  results in a fiber volume fraction [V.sub.f] of 0,123.
Equation (1) can be simplified in the linear form y = ax+b to yield following equations:
a = [[eta].sub.0] x [[tau].sub.c] x 1/d x [V.sub.f] (3)
b = [[sigma]'.sub.m] x (1-[V.sub.f]) (4)
Extracting [[sigma].sub.m]', which is the reduced tensile strength of the matrix material at the same elongation as the tensile strength of the composite, which is 3%, from the stress strain curve of the neat polymer, this results in a value of 28,93 MPa, which can be inserted in Eq. (4) to yield the intercept at the y-axis of the linear fit in Fig. 10.
The tensile strength of the composites increases with increasing fiber length (Fig. 10), regardless if the data from the different barrel temperatures or the different screw configurations are used. The comparison of the data from the different barrel temperatures is valid, because we do not see any significant differences between the different OIT data (Fig. 8), which would indicate a difference due to oxidative degradation and also the SEM micrographs show very comparable microstructures. Also the values from the different screw configurations match the linear fit with the calculated starting point d quite well and achieve a good coefficient of determination.
Using all known values in Eq. (3), the IFSS for the composites can be estimated at about 22 MPa. We are aware that this is not an absolute value, because small differences in the input parameters, such as fiber diameter d and orientation factor [[eta].sub.0] result in high differences in the calculated data for [[tau].sub.c]. These data were not determined individually for every composite in this study and are therefore not absolute values but can be used as reliable estimations. Also, the calculated value is in good accordance with the literature .
In this work, we investigated the influence of barrel temperature and screw geometry on the properties of short glass fiber reinforced polypropylene composites. We varied the barrel temperature to influence the viscosity of the polypropylene during processing and the screw configuration to alter the shear input during the process. In conclusion, we found that increasing the barrel temperature (within the limits of the polymer stability) shows a positive effect on the composite properties, due to the reduced polymer viscosity and therefore the lower shear brought into the material. Also the reduction of the shear input by varying the screw geometry shows comparable influences. Both strategies to preserve the fiber length do not negatively influence the interfacial adhesions as was shown by calculating the interfacial shear strength via a "rule-of-mixtures" approach. Comparing both strategies, that is, increasing the melt temperature to reduce shear or alter screw geometry, the former is easier to apply in an existing setup, being therefore more beneficial, while the latter exhibits advantages if the thermal stability of the matrix or additives do not allow for such high temperatures. On the basis of this study, we think that a combination of both, an optimized screw design together with a reasonably high processing temperature for lower matrix viscosities should provide the best results.
The authors want to acknowledge the effort of Patrick Insamer, who was carrying out part of the lab work. We are grateful for funding through "AdInComp - Thermoplastic composites with innovative performance profile due to enhanced interface" in the scheme "European fund for regional development (EFRE)" from the European Union and the Upper Austrian government and funding through the project "ReProMu," funded by the Upper Austrian government.
[1.] J.L. Thomason and M.A. Vlug, Composites: Part A, 27 A, All (1996).
[2.] J.L. Thomason, M.A. Vlug, G. Schipper, and H.G.L.T. Krikort, Compos.: Part A, 27 A, 1075 (1996).
[3.] J.L. Thomason and M.A. Vlug, Compos.: Part A, 28 A, 277 (1997).
[4.] R.D.S.G. Campilho, Natural Fibre Composites, CRC Press, Boca Raton (2016).
[5.] S.K. Ramamoorthy, M. Skrifvars, and A. Persson, Polym. Rev., 55, 107 (2015).
[6.] E.H. Wallace, Pneumatic tire, US Patent, 2782830A (1957)
[7.] M. Ho, H. Wang, J. Lee, C. Ho, K. Lau, J. Leng, and D. Hui, Compos.: Part B, 43, 3549 (2012).
[8.] J.L. Thomason, Compos.: Part A, 33, 1641 (2002).
[9.] S.Y. Fu and B. Lauke, Compos. Sci. Teehnol, 56, 1179 (1996).
[10.] B. Yang, J. Leng, B. He, H. Liu, Y. Zhang, and Z. Duan, J. Reinf. Plast. Comp., 31, 1103 (2012).
[11.] B. Fisa, Polym. Compos., 6, 232 (1985).
[12.] A.M.M. El-Sabbagh, L. Steuernagel, D. Meiners, and G. Ziegmann, J. Appl. Polym. Sci., 131, 40435 (2014).
[13.] L.V. Turkovich and L. Erwin, Polym. Eng. Sci., 23, 743 (1983).
[14.] V.B. Gupta, R.K. Mittal, P.K. Sharma, G. Mennig, and J. Wolters, Polym. Compos., 10, 8 (1989).
[15.] J. Ville, F. Inceoglu, N. Ghamri, J.L. Pradel, A. Durin, R. Valette, and B. Vergnes, Int. Polym. Process., 28, 49 (2013).
[16.] K. Schon and J.L. White, Polym. Eng. Sci., 38, 1757 (1999).
[17.] J.M. Lunt and J.B. Shortall, Plast. Rubber Process., 4, 108 (1979).
[18.] S.H. Bumm, J.L. White, and A.I. Isayev, Polym. Compos., 33, 2147 (2012).
[19.] F. Berzin, B. Vergnes, and J. Beaugrand, Compos.: Part A, 59, 30 (2013).
[20.] U. Yilmazer and M. Cansever, Polym. Compos., 23, 61 (2002).
[21.] F. Inceoglu, J. Ville, N. Ghamri, J.L. Pradel, A. Durin, R. Valette, and B. Vergnes, Polym Compos., 32, 1842 (2011).
[22.] J. Beaugrand and F. Berzin, J. Appl. Polym. Sci., 128, 1227 (2013).
[23.] G. Ozkoc, F. Bayram, and E. Bayramli, Polym. Compos., 26, 745 (2005).
[24.] S.Y. Fu, B. Lauke, E. Mader, C.Y. Yue, and X. Hu, Compos.: Part A, 31, 1117 (2000).
[25.] D. Dospisil, J. Kubat, M. Plesek, and P. Saha, Int. Polym. Process., 4, 303 (1994).
[26.] A. Kelly and W.R. Tyson, J. Mech. Phy. Solid., 13, 329 (1965).
[27.] B. Burgstaller, Int. J. Mater. Prod. Tech., 36, 11 (2009).
[28.] G.W. Ehrenstein, Faserverbundkunststoffe, Carl Hanser Verlag, Munich (2006).
[29.] J.L. Thomason, Comp. Sci. Tech., 62, 1455 (2002).
Blanca Maria Lekube (iD), (1) Bianca Purgleitner, (1) Karoly Renner, (2,3) Christoph Burgstaller (1)
(1) Transfercenter fur Kunststofftechnik GmbH 4600 Wels, Franz-Fritsch-Strasse 11, Austria
(2) Laboratory of Plastics and Rubber Technology, Department of Physical Chemistry and Materials Science, Budapest University of Technology and Economics H-1521 Budapest, P.O. Box 91, Hungary
(3) Institute of Materials and Environmental Chemistry, Research Centre For Natural Sciences, Hungarian Academy of Sciences H-1519 Budapest, P.O. Box 286, Hungary
Correspondence to: B. Lekube; e-mail: firstname.lastname@example.org
This paper contains original research and has not been published earlier in any journal or presented anywhere.
Published online in Wiley Online Library (wileyonlinelibrary.com).
Caption: FIG. 1. Applied screw configurations, (a) Tooth mixing element configuration (b) 90[degrees] configuration (c) 30[degrees] configuration (d) standard configuration (e) zone 1-4 are the same for all configurations. A indicates a neutral conveying zone; F means that the conveying elements follows the same direction as the revolution of the screw. Mixing elements used for different screw configurations (from the original thermo prism drawing sheets), (f) 0[degrees] mixing element (g) 90[degrees] mixing element (h) tooth mixing element (TME) (i) feed screw (FS).
Caption: FIG. 2. Tensile modulus and tensile strength versus applied temperature profile for injection molded short glass fiber polypropylene composites, "r" indicates repetition of an experiment.
Caption: FIG. 3. Charpy impact strength unnotched (aCUe) and notched (aCNe) versus applied compounding temperature profile for injection molded short glass fiber reinforced polypropylene composites, "r" indicates repetition of an experiment.
Caption: FIG. 4. Melt flow volume versus capillary temperature for the unreinforced polypropylene used in this study.
Caption: FIG. 5. Cumulative glass fiber length distribution versus applied compounding temperature profile for injection molded short glass fiber reinforced polypropylene composites, "r" indicates repetition of an experiment.
Caption: FIG. 6. SEM micrographs of cryo-fractured surfaces of the injection molded standard test specimens of the under different temperatures compounded short glass fiber reinforced polypropylene, "r" indicates repetition of an experiment.
Caption: FIG. 7. Tensile modulus and tensile strength versus different compounding screw configurations for injection molded short glass fiber reinforced polypropylene composites, "r" indicates repetition of an experiment.
Caption: FIG. 8. Charpy impact strength unnotched (aCUe) and notched (aCNe) versus different compounding screw configurations for injection molded short glass fiber reinforced polypropylene composites, "r" indicates repetition of an experiment.
Caption: FIG. 9. Cumulative glass fiber length distribution versus different compounding screw configurations for injection molded short glass fiber polypropylene composites, "r" indicates repetition of an experiment.
Caption: FIG. 10. Maximum tensile strength versus residual length weighted fiber length of the injection molded composite specimens produced with different compounding temperatures and screw configurations. Due to very small standard deviations and for clarity purposes the error bars were left out of this figure.
TABLE 1. Properties and proportion of the used materials. Polypropylene Short glass Compatibilizer fiber Type HG313MO HP3299 Orevac CA 100 Proportion [wt%] 67 30 3 [E.sub.Tension] [MPa] 1.500 72,000 880 MFR [g/10min] 30 -- 150-200 Fiber length [mm] 4,5 TABLE 2. Used temperature profiles along the screw zones (set temperatures) in [degrees]C. Experiment Zone 1 2 3 4 5 6 7 Die 180[degrees]C 180 200-20[degrees]C 180 200 200[degrees]C 200 200 + 20[degrees]C 40 170 180 220 200 220[degrees]C 220 240[degrees]C 240 260[degrees]C 260 TABLE 3. Injection molding parameters applied on all produced specimens. Parameter Set values Mass temperature 200[degres]C Tool temperature 40[degrees]C Dosing volume 45 [cm.sup.3] Injection volume 40 [cm.sup.3] Injection speed 0,35 m/s Injection pressure 550 bar Holding pressure 450 bar Holding pressure time 40 s Cycle time 60s TABLE 4. Oxidation induction time (OIT) of the samples produced with different compounding screw configurations. Screw configuration OIT [min] s [min] Std 11,70 1,07 90[degrees] 10,16 0,68 30[degrees] 7,81 0,68 30[degrees]r 8,91 0,40 TME 7,45 1,30
|Printer friendly Cite/link Email Feedback|
|Author:||Lekube, Blanca Maria; Purgleitner, Bianca; Renner, Karoly; Burgstaller, Christoph|
|Publication:||Polymer Engineering and Science|
|Date:||Aug 1, 2019|
|Previous Article:||Influence of Vacuum on the Porosity and Mechanical Properties in Rotational Molding.|
|Next Article:||Properties of Microinjection-Molded Polypropylene/Graphite Composites.|