Effects of molecular weight and thermal history on scratch behavior of polypropylene thin sheets

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Tribology International 41 (2008) 425–433 www.elsevier.com/locate/triboint

Effects of molecular weight and thermal history on scratch behavior of polypropylene thin sheets E. Moghbellia, R.L. Browninga, W.-J. Booa, S.F. Hahnb, L.J.E. Feickc, H.-J. Suea, a

Polymer Technology Center, Department of Mechanical Engineering, Texas A&M University, 3123 TAMU, College Station, TX 77843-3123, USA b The Dow Chemical Company, 2301 N. Brazosport Blvd., Freeport, TX 77541, USA c Advanced Composites, Inc., 1062 4th Avenue, Sidney, OH 45362, USA Received 18 December 2006; received in revised form 31 August 2007; accepted 26 September 2007 Available online 5 November 2007

Abstract The effects of molecular weight (Mw) and thermal history on the scratch behavior of polypropylene (PP) thin sheets and a commercial thermoplastic olefin (TPO) have been investigated. Scratch parameters like the critical load for onset of scratch visibility and scratch coefficient of friction (SCOF) were utilized in this evaluation. The results suggest that scratch performance is improved when the Mw and surface crystallinity of PP are high. Correlation between scratch resistance and surface crystallinity of PP is established and discussed. Approaches for the preparation of scratch-resistant TPOs are also addressed. Published by Elsevier Ltd. Keywords: Polypropylene; Thermoplastic olefins; Scratch behavior; Skin-core morphology; Surface crystallinity

1. Introduction Scratch behavior of metallic and ceramic materials has been widely explored since the 1950s [1,2]. Now, owing to the widespread use of plastics in durable goods applications, and especially due to their soft surface nature, polymer scratch behavior has drawn significant attention in recent years [3–11]. Previously, attempts to systematically quantify the scratch resistance of polymers have been problematic due to the lack of standardized test methods and appropriate evaluation tools. The scratch resistance of a material had previously been defined as the ability of a material to withstand abrasion with another body [12,13]. Most recently, Sue et al. have developed a test methodology (ASTM D7027-05) for evaluating scratch resistance of polymers and coatings [4–7]. The standard methodology outlines conditions for a linear load increase test and a set of well-defined scratch evaluation parameters, such as scratch visibility, depth, and width. Using this

Corresponding author. Tel.: +1 979 845 5021; fax: +1 979 845 3081.

E-mail address: [email protected] (H.-J. Sue). 0301-679X/$ - see front matter Published by Elsevier Ltd. doi:10.1016/j.triboint.2007.09.008

methodology, meaningful and highly repeatable results have been attained. Through a detailed examination of polymer scratch behavior, it has been shown that the scratch behavior of plastics is complex and depends on various parameters [4], such as scratch load and speed [4,14–16], coefficient of friction [3,14–17], geometry and number of scratch tips [3,15,16], and types and amounts of fillers and additives incorporated [8,10,18]. In addition, material properties will also certainly have an impact on the scratch behavior of plastics. In previous studies, rudimentary investigations and discussions showed how elastic modulus, yield stress, tensile strength, coefficient of friction, and viscoelastic recovery would affect scratch behavior of polymers [3,8,9]. Even though the experiments yielded meaningful results, more careful and systematic experimental studies of model systems are still needed. Another set of parameters that has a significant effect on material properties, and therefore must be considered, are processing and molding conditions. In injection molding and extrusion of semi-crystalline polymers, the melt cooling rate and mold temperature will definitely play an important role in determining the physical and mechanical properties of the molded article. Generally, plastics are

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poor heat conductors which contributes to the formation of skin-core morphology during molding, where a gradient of molecular orientation, density, and crystallinity develops from the skin toward the core of the molded part. In other words, the skin of the polymer has a higher cooling rate in comparison with the inner core region, which results in a lower crystallinity near the surface. It is noted that the thickness of the skin depends strongly on the mold temperature, cooling rate, molecular weight (Mw), and use of nucleation agent(s) [19,20]. In applications where high crystallinity is favored, the relatively low crystallinity of the skin region can lead to various undesirable effects, such as decreased barrier properties, lowered chemical resistance, etc. Many approaches are known to be effective in controlling skin-core morphology as examined elsewhere [21,22]. For example, a higher overall cooling rate and a higher Mw will result in an increase of skin thickness. The skin and the core are expected to have different physical and mechanical properties [21]. Thereby, a variation in skin-core morphology is expected to strongly affect scratch behavior of polymers. Polypropylene (PP) is one of the most utilized commodity plastics today because of its good combination of properties, recyclability, and low cost. However, some disadvantages may limit its use, namely, its poor lowtemperature impact strength, surface crazing upon repeated flexing, poor scratch resistance, etc. [21]. Furthermore, the mechanical properties of PP are known to be greatly influenced by processing conditions [21,23,24], especially in the case of injection-molded pieces [24–26] or after post-processing thermal treatments such as annealing [27,28]. This makes injection-molded PP an ideal material for investigating the influence of processing conditions on the scratch behavior of plastics. In this study, the effects of skin-core morphology differences on scratch behavior are explored for two model PP systems having two different Mw. Various scratch parameters for evaluation will be examined and discussed. The findings will also be correlated with the scratch behavior of an experimental thermoplastic olefin (TPO) system where different injection molding mold temperatures were varied. The usefulness of varying Mw and mold temperatures to prepare scratch resistant PP will be discussed. 2. Experimental 2.1. Materials Specimens of PP with weight-average Mw of 416,000 and 305,000 g/mol were supplied by the Dow Chemical Company. Melt flow index (MFI) was measured based on ASTM D1238 (230 1C, 2.16 kg weight) and were evaluated to be 1.5 g/10 min for the high Mw system and 4.5 g/10 min for the low Mw system. Gel permeation chromatography (GPC) analysis was performed at 145 1C in trichlorobenzene and the Mw obtained with respect to polystyrene standards. The

tacticity was measured by 13C NMR analysis and found to be 98.5% based on the triad distribution. Owing to the fact that both high and low Mw PP specimens were subjected to a quenching and annealing treatment upon receipt, the systems are so-named: HQ and LQ (quenched high- and low-Mw PP) and HA and LA (annealed high- and low-Mw PP). The dimensions of the injection-molded sheets were measured as 124 mm  124 mm  1 mm. In addition, an experimental TPO system, containing 78 wt% polypropylene/ethylene–propylene–rubber (PP/EPR), 20 wt% talc filler and 2 wt% carbon black pigment, was obtained from Advanced Composites to demonstrate the importance of cooling rate on a commercial TPO system. The TPO was injection-molded into 80 mm  160 mm  3 mm plaques where the temperature of the mold wall was held at 26.7, 37.7, and 48.9 1C. 2.2. Heat treatment After receipt of the samples, heat treatment was performed in two regimes: rapid-cooling (quenching) and slow-cooling (annealing). This was achieved by heating the PP sheets to 140 1C between two 25.4 mm thick steel plates in a Napco E5851 vacuum oven and holding at this temperature for 30 min. The quenched samples were immediately cooled to 0 1C by ice/water bath immersion. The annealed samples were gradually cooled to room temperature at an approximate cooling rate of 0.5 1C/min. Temperatures were checked using an electronic thermocouple. The heating schedule is shown graphically in Fig. 1. 2.3. Scratch testing A recently developed standardized test method, ASTM D7027-05, was used for the scratch testing of the samples. A constant scratch speed of 100 mm/s with a linearly progressive load range of 1–50 N was used for a scratch 175 0.5 hr 150 Temperature (°C)

426

125 Heating 100

Annealed

75

Quenched

50 25 0 0

100

200 Time (min)

300

400

Fig. 1. Graphical representation of schedule for heat treatment of model PP systems.

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length of 100 mm for the PP, whereas the load range for the experimental TPO systems was only 1–30 N to enhance resolution for observing the critical loads for scratch visibility. The tests were performed at room temperature using a stainless steel spherical scratch tip with a diameter of 1 mm. The scratches were all oriented in the direction of the melt flow to prevent variation in scratch damage due to possible orientation effects. A minimum of five tests were performed per sample. 2.4. Scratch damage quantification After scratch tests were performed, the samples were scanned using an Epson Perfection 4870 Photo PC scanner at a true pixel resolution of 3200 dpi in 8-bit grayscale mode. The samples were inherently a milky white color, so a black background was placed behind the sample during scanning to alleviate any internal reflection. The scanned images showing the scratch on each sample were evaluated using a grayscale threshold function in ImageJ software. A detailed description of the digital image analysis with ImageJ can be found in Ref. [10]. In a linearly progressive load scratch test, the critical load where a transition in damage takes place is found as follows: Lv ðF e  F i Þ þ F i (1) L where Fx is the applied normal load (N) at any point in the scratch, x, L is the total scratch length (mm); v is the length of the visible portion of the scratch obtained through Image J measurement (mm); and Fi and Fe are the initial and end load (N), respectively. For clarification, the total scratch length minus the visible portion of the scratch will give the distance to any point in the scratch from the starting point (i.e., Lv ¼ x). At least five samples were given digital treatment to obtain average values and standard deviation.

Fx ¼

2.5. Scratch coefficient of friction The scratch coefficient of friction (SCOF) is defined as the ratio of the tangential force to the normal force at each loading point. Using digital sensors, the scratch testing instrument utilized in this study is capable of determining the actual applied normal load and tangential force in situ at a user-set sampling rate. 2.6. Optical microscopy In order to determine the differences in the skin-core morphology of the different PP systems, thin cross-sections of the scratch were examined. All cross-sections were taken at the location of the scratch which had a normal load equivalent to 37.5 N. Thin sections were taken with an Ultracut-E microtome using a diamond knife at room temperature. The thin section was placed on a glass slide with a drop of immersion oil followed by a cover slip. The

427

TRANSMITTED LIGHT Sample surface

40 μm

Valley of scratch

Glass slide Fig. 2. Schematic illustrating mounting of PP specimens for transmitted light optical microscopy observation of sub surface scratch damage along the longitudinal direction.

thin sections were then analyzed in transmission mode under crossed polars using an Olympus BX60 optical microscope with image capturing software to observe the plastic deformation patterns of the model PP systems. Longitudinal sections along the scratch direction were also examined to determine the sub-surface damage for HA (best case) and LQ (worst case). The samples were carefully cut in half using an Isomets diamond saw equipped with a diamond-tipped blade. The surface corresponding to the center-line axis of the longitudinal cut was wet-polished with a Struers Abramin disk polisher using a succession of FEPA P1200, P2400, and P4000 grit papers (Struers, Buehler) and then affixed to a glass slide with epoxy. The glass slide was then mounted in the diamond saw to cut the sample as close to the glass slide as possible. After this, the slide-mounted specimen was wet-polished to achieve a thickness of around 40 mm to allow for transmitted light observation under crossed polars using the Olympus BX60. A schematic of the mounting setup for OM observation is shown in Fig. 2. The Olympus BX60 was also used in bright field reflected mode to observe damage transitions on the sample surface. 2.7. Differential scanning calorimetry Differential scanning calorimetry (DSC) was performed using a Mettler Toledo DSC821 from 25 to 190 1C using a heating rate of 5 1C/min with nitrogen as the purge gas (80 mL/min). In order to examine the crystallinity of the skin region, about 10 mg of the surface layer (o100 mm in depth) were carefully removed from the samples for DSC analysis. The integration was performed over a range of 150–180 1C and heat of fusion was considered to be 185 J/g for 100% crystalline PP [23]. Percentage crystallinity was taken to be the ratio of the heat of fusion from the DSC scan to the heat of fusion for fully crystalline PP times 100%. 3. Results and discussion 3.1. Heat treatment effects on internal morphology and crystallinity To appreciate the impact of the heat treatment on the model PP systems of this study, the cross-sectional

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morphology of the model PP systems were analyzed via transmitted OM (Fig. 3). Qualitatively speaking, it can be seen that in both Mw cases the skin layer thickness for the annealed samples are significantly less than those of the quenched samples. This is explained by the fact that with increased cooling rate, the opportunity for polymer chains to form crystalline lamellae at the surface decreases in turn. Therefore, by this convention, the skin crystallinity of the quenched specimens should be much lower than that of the annealed specimens. The DSC results shown in Fig. 4 help to substantiate this claim. The skin layer crystallinity values obtained from the DSC thermographs of the four PP systems are quantitatively reported in Table 1. It can be observed that the skin crystallinity of the annealed systems is approximately 5–10% higher than those of the quenched systems, regardless of the Mw.

Fig. 3. Transmitted light optical micrographs of thin cross-sections of model PP systems. The bright top layer in all micrographs corresponds to the skin layer.

Heat Flow (mW) Exo Up

0 -2

HA

-4 HQ

-6 -8

Table 1 Percentage crystallinity of surface layer of model PP systems Model PP system

Crystallinity of surface layer (%)

HA HQ LA LQ

42.8 36.7 45.3 38.7

The effects of post-processing cooling rates on various mechanical properties of semi-crystalline plastics have been investigated in previous studies [22,24–36]. It has been shown that variation in cooling rates can result in variations in the skin thickness observed in the skin-core morphology of the plastic. This variation will definitely have a major influence on the mechanical properties of the samples due to the lower crystallinity of the skin region. However, increasing the skin layer crystallinity will likely result in an increase of the yield stress of the surface layer [29,30]. The variation of Mw has also been shown to have a significant effect on various mechanical properties of semicrystalline polymers [31–37]. Increasing Mw has been shown to be effective in enhancing mechanical properties such as yield stress, elongation at break, and tensile strength of the polymer [31]. However, the increase in Mw for a given system will also lead to reduced molecular mobility for crystallization, as reflected in the MFI values. The skin thicknesses with regard to Mw can be qualitatively compared in Fig. 3. It is apparent that the skin thickness for high Mw PP is higher than that of the low Mw PP for both quenched and annealed cases. The higher mobility of the polymer chains for the low Mw (higher MFI) sample can be a reason causing the decrease in skin thickness. Furthermore, similar to the comparison between different cooling rates, a lower skin thickness here also reflects a higher crystallinity of the skin region. Again, this is attributed to the high MFI of the low Mw system. As shown in Fig. 4 and Table 1, the thermal analysis shows that, at similar cooling rates, the crystallinity in the skin region of high Mw is lower than the low Mw systems. However, in this case, the mechanical properties of the skin layer are affected both by crystallinity and by Mw. The enhancement of mechanical properties due to Mw increase is likely to improve the PP ductility and strength [31], which have been shown to be critical for improving scratch resistance [10].

LA

3.2. Scratch behavior of heat-treated model PP

-10 -12

LQ

-14 -16 -18 30

80

130 Temperature (°C)

Fig. 4. DSC thermographs for model PP systems.

180

The post-processed images of the scratched PP samples obtained with the optical scanner are shown in Fig. 5. Comparing between the annealed and quenched PP systems (HA vs. HQ and LA vs. LQ), it can be stated that the onset of scratch visibility is significantly delayed for the annealed systems. Furthermore, the comparison between the low Mw and high Mw systems (HA vs. LA and HQ vs.

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35.0 N

HA 28.5 N

HQ 27.5 N

LA 18.9 N

LQ 1N

50 N

100 mm

Critical Load for Onset of Scratch Visibility (N)

Fig. 5. Images obtained with an optical scanner (resolution ¼ 3200 dpi) of scratched model PP systems showing onset points of scratch visibility and corresponding critical loads.

40

35.1 N

35

28.8 N

30

26.5 N

HA

25

19.2 N

20

Fz = 36.1 N

Fz = 38.0 N

15 10 5 HQ

0 HA

HQ

LA

LQ

Model System Fz = 22.9 N

Fz = 24.7 N

Fig. 6. Critical load for the onset of scratch visibility for model PP systems.

LQ) displays a delay in the onset of scratch visibility for the high Mw system. These observations are quantitatively reported in Fig. 6. The onset of scratch visibility in a ductile plastic occurs at the point in the scratch where the applied normal load and resulting tangential force induce ductile drawing of the material along the scratch path, just under the scratch tip. Once the material under the tip has been pulled to its ultimate limit, the tip ‘‘irons’’ over the drawn material and begins the drawing process again. This results in periodic, parabolic-shaped features termed as ‘‘fish scale’’. As the applied normal load continues to increase, there comes a point where the surface material is drawn to its limit. After this point, the scratch tip ruptures the material and penetrates through the sub-surface. As the tip continues to move in the scratch direction, it begins to displace the material around it. In this study, this mechanism will be termed ‘‘ploughing’’. These damage features were observed in all systems and can be viewed in Fig. 7. More detail regarding this feature will be discussed later. The SCOF of the annealed and quenched PP systems are shown and compared in Figs. 8 and 9. In Fig. 8, it can be seen that in both high and low Mw systems, the SCOF for the annealed systems, for the most part, is lower than that

LA

Fz = 27.6 N

Fz = 29.6 N

LQ

Fz = 20.0 N Onset of Visibility

Fz = 22.7 N Onset of Ploughing

Fig. 7. Reflected optical micrographs of scratch damage transitions observed in model PP systems. Scratch direction is from left to right.

of the quenched systems, indicating that at similar normal loads the tangential force (i.e., frictional force) corresponding to the annealed systems is lower compared to the quenched systems. Fig. 9 shows the SCOF based on a comparison of Mw. It is evident that in both quenched and annealed cases, the low Mw systems show a higher SCOF than the high Mw systems. When Mw is increased, the chain entanglement

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1 0.8

0.8

0.7

0.7

0.6 0.5 0.4

0.6 0.5 0.4

0.3

0.3

0.2

0.2

0.1

0.1 0

0 0

20

40 60 Scratch Length (mm)

80

1

0

100

0.7

0.7

SCOF (Fx/Fz)

0.8

0.5 0.4

40 60 Scratch Distance (mm)

80

100

HQ LQ

0.9

0.8 0.6

20

1

LA LQ

0.9

SCOF (Fx/Fz)

HA LA

0.9

SCOF (Fx/Fz)

SCOF (Fx/Fz)

1

HA HQ

0.9

0.6 0.5 0.4

0.3

0.3

0.2

0.2

0.1

0.1 0

0 0

20

40 60 Scratch Distance (mm)

80

100

0

20

40 60 Scratch Distance (mm)

80

100

Fig. 8. Comparisons of SCOF curves for model PP systems under various heat treatment conditions.

Fig. 9. Comparisons of SCOF curves for model PP systems due to Mw variations.

density also increases. The presence of these entanglements can enhance the elastic recovery of the material, thus imparting resistance to plastic deformation which, in turn, decreases the surface scratch coefficient of friction. The results indicate that both the critical load for onset of scratch visibility and SCOF show consistent trends between systems. The trend suggests that lower SCOF will result in a higher critical load, and thus delayed onset for scratch visibility. In Figs. 8 and 9, it is also observed that the difference in SCOF between the systems decreases at higher loads and, eventually, the SCOF of all systems merge. Hence, it can be interpreted that beyond a certain applied normal load, there comes a point where the scratch behavior is no longer dependent on the surface layer. This could suggest a transition to the sub-surface ploughing damage mechanism. The appearance of the scratch damage mechanisms that relate to the fish scale and ploughing transitions described earlier along the scratch direction for all four systems are shown using reflected OM in Fig. 7. The locations of the transitions and the corresponding applied normal loads for the transition are shown. It can be seen that, independent of the PP systems, the scratch damage pattern after the ploughing transition appear to be qualitatively the same for

all systems independent of their skin-core morphology. As expected, it is also observed that the trend for the location of the fish scale transition for the systems follows that for their critical loads for the onset of scratch visibility. Furthermore, comparison of the scratch transitions between the systems leads to the conclusion that in the case of either high Mw or annealed systems, the fish scale pattern is more evident on the scratch path. The main reason behind the more pronounced fish scale pattern can be the higher ductility of the HA system as compared to the LQ, as has been shown for PP systems in previous studies [24,38–40]. In order to investigate the sub-surface damage caused by the ploughing scratch damage mechanism, longitudinal sections along the scratch direction have been examined for the two extreme cases, as shown in Figs. 10 and 11. It can be observed that in both cases, the fish scale and ploughing transitions are evidenced by the similar types of sub-surface damage caused by the normal and tangential force acting upon that region of the material. However, the damage pattern is more evident and dramatic for HA when compared with LQ. This fact is due to the higher applied normal load required for HA to reach the transitions. That is, for LQ, the critical load for scratch visibility and thus the

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Onset of fish-scale

Onset of ploughing

Fz ≈ 36.1 N

Fz ≈ 38.0 N

431

Fig. 10. Transmitted optical micrographs of sub-surface scratch damage of the high molecular weight annealed model PP system (HA).

Fz ≈ 20.0 N

Fz ≈ 22.7 N

Onset of fish-scale

Onset of ploughing

Fig. 11. Transmitted optical micrographs of sub surface scratch damage of the low molecular weight quenched model PP system (LQ).

~12.4 N 26.7 °C ~16.3 N

37.7 °C

~21.2 N

48.9 °C

1N

100 mm

30 N

Fig. 12. Images obtained with an optical scanner (resolution ¼ 3200 dpi) of scratched TPO specimens injection-molded with different mold wall temperatures showing onset points of scratch visibility and corresponding critical loads.

fish scale transition occurs at a normal load of 18.5 N, while this transition occurs at 35.1 N for HA, which is almost twice the normal load for LQ. Furthermore, the longitudinal section along the scratch direction of the scratched samples can also illustrate the type of damage and fracture mechanisms caused by the scratch process. The apparent microcrack-like feature evident at the edge of each fish scale is believed to be a result of tensile drawing followed by compressive ironing of the PP on the surface. Fundamental work in attempts to elucidate these claims and study the link of traditional mechanical mechanisms to scratch behavior is currently underway.

3.3. Extension to a commercial TPO system Understanding the fundamental scratch behavior discussed above using a model PP is essential. In the case of industrially relevant TPO systems, which contain rubber, fillers, and other additives, it becomes difficult to unambiguously show how material/processing parameters influence the scratch behavior. To illustrate the relevance of the above fundamental study on neat PP to a more complex system, an experimental TPO system will be considered to study if mold wall temperature, which affects the cooling rate of injection-molded polymers, can greatly affect the scratch behavior of TPOs.

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Applied Normal Load (N)

30 25 21.2 N

20 16.3 N

15

12.4 N

10 5 0 26.7 °C

37.7 °C Mold Wall Temperature

by the heat insulating nature of these materials. In this study, quenching and annealing were performed on thin sheets of PP to investigate how the processing conditions affect scratch behavior. The effect of Mw has been investigated as well. The critical load for the onset of scratch visibility and SCOF were utilized to evaluate scratch performance of PP and a commercially available TPO. The results indicate that post-processing as well as Mw has a significant effect on the scratch behavior. The present study shows that scratch resistance of both PP and TPOs can be easily improved by controlling either cooling rate or Mw, or both.

48.9 °C

Fig. 13. Critical load for the onset of scratch visibility for TPO specimens injection-molded with different mold wall temperatures.

Fig. 12 shows the scanned images (resolution ¼ 3200 dpi) of the scratched surfaces of the TPO model system. Fig. 13 shows the corresponding values of the critical load of onset of scratch visibility for the TPO molded at three mold-wall temperatures of 26.7, 37.7, and 48.9 1C, respectively. From these results, it is readily observed that the TPO exhibits an increase in scratch resistance as the mold wall temperature is increased. This finding is analogous to the trend seen in the behavior of the model PP systems. When the TPO is quenched, or when the mold wall is held at a low temperature, there is a large temperature gradient between the polymer and the mold wall, resulting in rapid cooling. The converse is true for the case of slowly cooled TPO. It is noted that attempts were made to investigate the skin-core morphology of the TPO system using OM and DSC. However, the presence of the talc filler and EPR tended to convolute the results so that the skin-core morphology could not be clearly discerned. The above findings on TPO are in excellent agreement with the model PP systems observed. In other words, a slowly cooled PP or TPO will give rise to greatly improved scratch resistance against visibility. The present study also clearly demonstrates the importance and relevance of a model PP system study to reveal conclusive information for establishing structure–property relationship between PP morphology and its scratch behavior. The implication of the current study signifies that polymer products with improved scratch resistance can be made simply by slight alteration of the processing conditions, whether by postprocessing methods (slow cooling/quenching) or by changing the temperature of the mold wall. This fact will surely be attractive to industries where scratch resistance is of utmost importance. 4. Conclusion Mold temperatures and cooling rates can profoundly affect the mechanical properties of semi-crystalline polymers through the so-called skin-core morphology caused

Acknowledgments The authors would like to thank Dow Chemical and Advanced Composites for the supply of the model PP and TPO samples. Financial support from the Polymer Scratch Behavior Consortium at Texas A&M University is greatly appreciated. References [1] Bowden FP, Tabor D. In the friction and lubrication of solids. Part I. Oxford, UK: Oxford University Press; 1950. [2] Bowden FP, Tabor D. In the friction and lubrication of solids. Part I. Oxford, UK: Oxford University Press; 1964. [3] Xiang C, Sue HJ, Chu J, Coleman B. J Polym Sci, Part B: Polym Phys 2001;39:47. [4] Wong M, Moyse A, Lee F, Sue HJ. J Mater Sci 2004;39:3293. [5] Wong M, Lim GT, Moyse A, Reddy JN, Sue HJ. Wear 2004;256:1214. [6] Browning RL, Lim GT, Moyse A, Sue HJ, Chen H, Earls JD. Surf Coat Technol 2006;201:2970. [7] Browning RL, Jiang H, Moyse A, Sue HJ, Iseki Y, Ohtani K et al. J Mater Sci. 2007, accepted. [8] Xiang C, Sue HJ, Chu J, Masuda K. Polym Eng Sci 2001;41(1):23. [9] Jiang H, Lim GT, Whitcomb JD, Sue HJ. J Polym Sci, Part B: Polym Phys 2007;45:1435. [10] Browning R, Lim GT, Moyse A, Sun L, Sue HJ. Polym Eng Sci 2006;46:601. [11] Wong JSS, Sue HJ, Zeng KY, Li RKY, Mai YW. Acta Mater 2004;52:431. [12] Wu S, Sehanobish K, Christenson C, Newton J. Polym Mat Sci & Eng 1998;79:212. [13] Sadati M, Mohammadi N, Qazvini NT, Tahmasebi N, Koopahi S. Prog Org Coat 2005;53(1):23. [14] Briscoe BJ, Evans PD, Pelillo E, Sinha SK. Wear 1996;200:137. [15] Briscoe BJ, Pelillo E, Sinha SK. Polym Eng Sci 1996;36(24):2996. [16] Briscoe BJ, Evans PD, Biswas SK, Sinha SK. Tribol Int 1996; 29(2):93. [17] Chu J, Rumao L, Coleman B. Polym Eng Sci 1998;38(11):1906. [18] Chu J, Xiang C, Sue HJ, Hollis RD. Polym Eng Sci 2000;40(4):944. [19] Daniels CA. Polymers: Structure and Properties. Lancaster, PA: Technomic Publishing Co.; 1989. [20] Karger-Kocsis J, Cunhaand AM, Fakirov S, editors. Structure development during polymer processing. Dordrecht, The Netherlands: Kluwer Academic Publishers; 1999. [21] Maspoch ML, Gamez-Perez J, Giminez E, Santana O, Gordillo A. J Appl Polym Sci 2004;93:2866. [22] Strebel JJ, Mirabella F, Blythe C, Pham T. Polym Eng Sci 2004; 44(8):1588.

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