General model for comparative tensile mechanical properties of composites fabricated from fly ash and virgin/recycled high-density polyethylene.
Fly ash, a waste product from coal-fired power plants, represents a large-scale problem for waste disposal because as much as ~750 million tons are produced per annum but only ~16% of this amount is utilized, the remainder of which goes to landfill and ponds [1-4]. Although fly ash has found technical applications, such as adsorbents and fillers, its main applications are in the geotechnical and construction fields, largely as an additive to concrete. It also has been studied as a filler in a range of virgin polymers [5-7] but there has been limited use of fly ash in recycled plastics, which represent an important potential resource. Relevant to the present work is the publication by Kumar et al. , who published a brief literature survey of fly ash additions to high-density polyethylene. The following list provides a summary of the reports of polymers that have been compounded with fly ash to form composites.
Virgin high-density polyethylene HDPE [9-13] Recycled high-density polyethylene RHDPE [14-17] Virgin low-density polyethylene LDPE [12, 18-23] Recycled low-density polyethylene RLDPE  Virgin polypropylene PP [12, 23, 25-29] Virgin polyphenylene oxide PPO  Virgin polyoxymethylene POM  Virgin polyetherketone PEEK  Virgin polyvinyl alcohol PVA [33-36] Virgin polyvinyl chloride PVC  Recycled polyethylene terephthalate RPET [38-41] Virgin acrylonitrile butadiene styrene ABS  Recycled rubber --  Virgin phenolic resin --  Virgin formamide and urea plasticized FUPTPS  thermophastic starch Virgin glycerol-plasticized GPTPS  thermoplastic starch Virgin epoxy -- [46-49] The number of studies of polymers reinforced with ceno- spheres (hollow fly ash) is much smaller. Virgin high-density polyethylene HDPE [50, 51] Recycled high-density polyethylene RHDPE  Virgin low-density polyethylene LDPE  Virgin polycarbonate PC  Virgin polyester -- [55, 56]
Fly ash generally is spherical in shape, with sizes ranging typically from 10 to 90 [micro]m [1-4]. The major fraction of fly ash consists of solid spheres although there is a small proportion of hollow cenospheres [19, 25]. Chemically and mineralogically, fly ash is principally a vitreous aluminosilicate, containing various levels of alkalies, alkaline earths, transition metals, and some precipitated crystalline mullite (3[Al.sub.2][O.sub.3] x 2Si[O.sub.2]), the proportions of which depend on the coal composition and processing [57, 58]. While the main application of fly ash is as a pozzolanic additive to concrete, the plastics industry also is capable of playing an important role in exploiting this underutilized raw material, which has the potential to impact significantly on environmental issues [33-36].
The manufacturing of polymer matrix composites reinforced with particulate fillers is widespread. However, common fillers, such as calcium carbonate and silica [5-7], are relatively expensive, so their potential replacement by a lightweight and abundant waste material, such as fly ash, represents an attractive commercial application. The literature [9-56] shows that such additions generally lead to selected progressive improvements in a range of properties relative to the matrix polymer.
HDPE has important applications in many areas, including packaging, containers, bags, consumer hardgoods, pipes, electrical insulation, prosthetics, and biomedical implants [59, 60]. Owing to the environmental and economic effects of the disposal of such materials, the utilization of RHDPE is an attractive commercial option. The literature reveals that there have been several attempts to improve the mechanical properties of HDPE and RHDPE by the addition of fly ash. Of these, three are directly comparable to the present work [9, 16, 50], although one  presents relative data (without standard deviations), so only the trends can be compared. These comparative studies are of note because they present basic mechanical properties data that, when combined with the data for the present work, can be used to synthesize a generalized model for the load-extension behavior of fly ash-HDPE and fly ash-RHDPE composites. Consequently, the present work reports the results of a more focussed examination of the comparative mechanical and micro-structural data for specimens with low-level additions of up to 10 wt% fly ash to RHDPE in order to supplement the published data.
The comparative data for the mechanical properties as a function of fly ash addition level ([less than or equal to] 10 wt%) to recycled high-density polyethylene (RHDPE) reveal maxima at the initial fly ash addition level of 2.5 wt% for the tensile elastic modulus (+25%) and tensile strength (+10%); a slight general increase in the yield stress (+6%); and significant general decreases in the yield strain (-61%), elongation at break (-92%), and Charpy impact strength (-55%). This behavior is a function of five dependent variables: dispersant volume, dispersant particle size, intrinsic flaw size (viz., dispersant size), generated flaw size (viz., void size), and interfacial bond strength and associated load transfer. The resultant trends are consistent with a generalized model of the load-displacement curves as a function of fly ash addition level. This model is delineated by three regimes characterized by ductile deformation, where dispersion strengthening and stress concentration are the dominant factors; crazing, where debonding and cavitation occur; and brittle failure, where the fibrils in the ligaments fail.
The RHDPE pellets (~6.0 mm diameter x ~3.5 mm height) were supplied by DONMAR Industries, Sydney, NSW, Australia and the fly ash was supplied by Cement Australia, Brisbane, QLD, Australia. Although the basic characteristics of the RHDPE were not provided or assessed, neat HDPE exhibits properties as given in Table 1 . The effects of the temperature and stress applied during recycling on polyethylene are well known [62-64]. That is, stress application during heating alters the polymer chain length, degree of branching, crystallisation, and other factors such that the stiffness, strength, and yield stress generally decrease while yield strain and elongation at break tend increase.
The chemical composition of the fly ash was determined by X-ray fluorescence (XRF, Philips PW2400); the results are given in Table 2. It is well known that the mineralogy of fly ash typically consists of a vitreous matrix containing various amounts of precipitated mullite (3[Al.sub.2][O.sub.3] x 2Si[O.sub.2]) .
The particle size distribution of the fly ash was determined by laser diffraction (Malvern Mastersizer S, Malvern Instruments, Malvern, Worcestershire, UK). As indicated in Fig. 1, the fly ash showed a bimodal particle size distribution, with a major maximum at ~13 [micro]m (range ~3-50 [micro]m, ~87 vol%) and a minor maximum at ~0.6 [micro]m (range ~0.4-1.3 [micro]m, ~13 vol%).
The RHDPE pellets and fly ash were blended by dry rotational mixing in ~350 mL batches for 24 h using a polyethylene mill (15 cm diameter, 20 cm length) rotated at 150 rpm, without media. The fly ash addition levels were 0-10 wt% in 2.5 wt% increments. Prior to injection molding, sample homogeneity was ensured by visual examination of the RHDPE-fly ash mixtures, which revealed conversion of the green RHDPE pellets to dark grey owing to homogeneous coating of the former by the latter. The resulting samples then were compounded by injection molding (Boy 15S Horizontal Injection Molder, Exton, PA, 15 tons) at 220[degrees]C and 100 MPa for 5 min. After injection molding, sample homogeneity was assessed visually by color consistency, which indicated no obvious inhomogeniety. No evidence of pyrolysis was observed following these procedures, which is as expected since, as shown in Table 1, processing temperatures of HDPE typically are in the range 190[degrees]C-274[degrees]C . The stainless steel combined die produced (1) a variant of the Type I standard dog-bone shape (ASTM D638; dimensions in parenthesis when different) of dimensions length 128 mm (165 [+ or -] 6.5 mm), gauge length 50 mm, gauge width 13 mm, and thickness 4 mm (3.6 [+ or -] 0.4 mm) and (2) a variant of the bar shape (ASTM D6110; dimensions in parenthesis when different) of dimensions length 100 mm (127.75 [+ or -] 1.25 mm), width 10 mm (3.0-12.7 mm), and thickness 4 mm (12.70 [+ or -] 0.15 mm). Examples of the injection molded samples are shown in Fig. 2.
Tensile testing of each dog-bone sample was performed at room temperature using a universal testing machine (Instron 1185, Norwood, MA) at the crosshead speed of 50 mm [min.sup.-1], which is according to ASTM D638 and the speed used for the three comparative studies [9, 16, 50].
The tensile elastic moduli were determined using the standard method (ISO 1209-2) of identification of the linear portion of the load-extension data and the calculation of the corresponding slope. The tensile strength, measured according to ASTM D638, was determined conventionally by identification of the local maximum corresponding to the maximal stress application . The yield stress and yield strain were determined by identification of the endpoints of the linear portion of the load-extension data. The elongation at break was assessed by identification of the point of return to nil stress during testing.
The averages and standard deviations for the preceding tests were determined on the basis of the data for five samples at each fly ash addition level.
Impact Strength Testing
The impact strength testing was done using the standard Charpy impact method (ASTM D6110; dimensions in parenthesis when different) using a tabletop pendulum unit (Toyo Seiki Seisaku-Sho, Tokyo, Japan). The moulded notch in the center of one side of each bar sample was of depth 1.6 mm (2.54 [+ or -] 0.15 mm), radius 0.32 mm (0.25 [+ or -] 0.05 mm), and angle 25[degrees] (45[degrees] [+ or -] 1[degrees]). The pendulum was elevated to 135[degrees] and the calculated striking speed was 2.4 m [s.sup.-1]. The averages and standard deviations were determined on the basis of the data for five samples at each fly ash addition level.
In the case of the tensile testing, samples of dimensions ~13 mm x 25 mm were cut by hand using a hacksaw parallel to the fracture surfaces. In the case of the notched Charpy impact testing specimens that did not exhibit through-specimen failure, this was finalized by hand subsequently. One half of each fractured surface was gold coated and examined using a scanning electron microscope (SEM, Hitachi S-3400x, Tokyo, Japan, 15 kV) in secondary ion emission mode.
RESULTS AND DISCUSSION
Tensile and Impact Strength Properties
Figure 3 shows typical load-extension data as a function of fly ash addition level. Tables 3 and 4 provide a numerical summary of the data for the present work and those of the comparative studies. Table 3 gives the relevant data for the different publications and Table 4 compares the sample/preparation parameters and apparent data trends. These data show that neat RHDPE is the most ductile of the samples, which suggests dispersion strengthening by the fly ash particles . The behavior of the tensile elastic modulus at low levels of addition of fly ash to RHDPE in the present work is mirrored in comparative studies I , II , and III . According to rule-of-mixtures considerations , it would be expected that the stiffness would increase proportionately to the amount of hard ceramic phase added. However, the trend as a function of increasing fly ash addition level suggests that a second relevant factor is likely to be increasing stress concentration from the fly ash particles . As shown in Fig. 3, at the highest fly ash addition level of 10 wt%, the greater deformation associated with the higher amount of the ceramic phase also suggests the possibility of embrittlement from recrystallization , crazing , or significant ligament stretching and decrease in width .
Tensile Elastic Modulus. The data for Table 3 indicate that the neat RHDPE is more ductile than the fly ash-RHDPE specimens, which is consistent with other studies using fly ash [9, 16, 50], wood ash , and other dispersants [74-76]. However, the tensile elastic moduli determined in the present work are nearly an order of magnitude less than those of comparative studies I (virgin HDPE)  and III (recycled HOPE) ; the data for comparative study II (virgin HDPE)  are not numerical. The probable reason for this variation is related to the processing methods for the two relevant comparative studies, which were by extrusion by progressive heating over the range 130[degrees]C-200[degrees]C  and by uniaxial hot pressing at 160[degrees]C , while the samples of the present work were processed by injection molding at 220[degrees]C. Table 1 shows that the normal range of processing temperatures is 190[degrees]C-274[degrees]C. Hence, the differences are likely to have derived from the heating of virgin HDPE slightly above the minimal processing temperature , heating of recycled HDPE well below the minimal processing temperature , and heating of recycled HDPE well above the minimal processing temperature (present work). As mentioned in Section 2.1, the effects of the temperature and stress applied during recycling on polyethylene serve to decrease the stiffness [62-64, 77], so recycling and higher temperatures reduce the tensile elastic moduli.
The trend for the tensile elastic moduli data for the present work indicates an increase upon addition of fly ash, followed by a decrease with increasing fly ask addition level. In contrast, comparative studies I  and II  reveal increasing moduli with increasing fly ash addition level and decreases in the moduli only at the highest levels of addition. The former data suggest that the cause of the increasing trend is dispersion of hard particles in a soft matrix , although this would be moderated by stress concentration from the particles . The latter data probably derive from the generation of porosity from debonding and cavitation during testing [78-85], which are initiated by the high proportions of ceramic dispersants and result in thinner and weaker ligaments formed by the polymer. The decrease in moduli at the highest fly ash addition levels is unlikely to derive from recrystallization or crazing since these increase the stiffness [71, 72]. Debonding and cavitation induced during processing also are possible, although all of the fabrication methods appear to have been designed to eliminate or minimise this possibility. The absence of decrease in comparative study III  is attributed to insufficient fly ash addition level to cause the decrease for this study's particular combination of materials and their parameters.
It is noted that the decreasing trend for the tensile elastic moduli data is apparent at the lowest fly ash addition levels for the present work while this trend becomes apparent only at the highest levels for comparative studies I  and II . This can be explained for the latter study since the bonding between the fly ash surfaces and the polymer matrix was improved with the use of a silane coupling agent. The difference for the former is attributed to the use of virgin HDPE processed at a relatively high maximal temperature (200[degrees]C), which would have enhanced contact between the long polymer chains and the fly ash particles (provided the polymer chain length was not shortened). In the present study, the use of recycled HDPE at a higher processing temperature (220[degrees]C) probably reduced the polymer chain length and hence the extent of contact, which would have facilitated debonding and cavitation, which is confirmed by the SEM fracture surface images of the present work to be discussed subsequently.
Table 4 shows that the present work (~13 [micro]m) and comparative study I (~14 [micro]m)  used fly ash of approximately the same size and exhibited similar tensile strengths while comparative studies II (~60-160 [micro]m)  and III (<106 [micro]m)  used much larger fly ash sizes. It is well known that larger particles cause greater stress concentration than smaller , so it would be expected that the tensile strengths of the samples of the two former studies would be greater than those of the two latter studies. However, this is not the case. This is attributed to the effect of the silane coupling agent to improve bonding used in comparative studies II  and III , which was not the case in the present work or comparative study I . Further, comparative study II , which involved the largest fly ash size, reported the use of a compatibilizer, which also would have improved bonding.
A major difference in the two sets of data lies in the actual values, where the specimens of comparative studies I  and III  are considerably stiffer than those of the present work. These differences may have resulted from differences in the molecular weights of the polyethylenes , polymer chain scission during shredding from recycling , the effects from the extrusion in the form of strain hardening , orientation of the polymer chains during extrusion or hot pressing , the effects of injection molding or hot pressing in the form of disentanglement of the polymer chains [90, 91], and/or reduction of the molecular weight from heating .
Tensile Strength. The toughening mechanism of particulate-reinforced polymers is well established [70, 71, 78-81]. It can be described in terms of three stages:
1. Stress concentration, where the difference in stiffness between the rigid dispersant and ductile polymer results in stress concentration.
2. Debonding, where tensile force application debonds the dispersant-polymer interfaces, resulting in the onset of void creation, ligament formation, and crazing.
3. Cavitation, where continued tension results in shear yielding, thereby extending the void length, ligament length, and degree of crazing, the latter of which both increases strength and stiffness but also acts as a precursor to brittle failure.
Table 3 shows the tensile strengths as a function of fly ash addition level. Again, the standard deviations are such that the statistical significance of the published data are not conclusive but the present work and comparative studies I  and II  support the view that, at low fly ash addition levels, there is a modest increase in strength with fly ash additions, which is as expected owing to the likelihood of dispersion strengthening of this polymer . The present work and comparative study II  show a decrease in tensile strength at higher fly ash addition levels while comparative study III  shows a consistent decrease in tensile strength. These trends are attributed to the stress concentration from the fly ash particles  and debonding and cavitation, which are evident from the SEM fracture surface images of the comparative studies and the tensile elastic modulus data and the SEM fracture surface images of the present work, which are shown subsequently.
The conclusions concerning the decreases in tensile strength are supported by the fly ash particle size data:
* Present work and comparative study I : The fine particle sizes (~13 gm and ~14 gm), low volume fractions ([less than or equal to] 4.7 vol%), and absence of silane coupling agent or compatibilizer in the two studies suggest the potential for similar properties, which is the case. This is attributed to the effect of dispersion strengthening moderated by stress concentration. In the present work, the large difference in stiffnesses between fly ash and polymer matrix enhanced debonding and cavitation in all samples. In contrast, the greater stiffness of the polymer in comparative study I  probably reduced debonding and cavitation. At the highest volume fraction (~23 vol%), the distribution density of these finer particles probably was sufficiently high to affect flow and hence facilitate debonding. While this resulted in a decrease in the tensile elastic modulus owing to the generation of voids, their small size resulted in minimal effect on the tensile strength.
* Comparative study II : The coarse particle size (~60-160 gm at the highest volume fractions (26 and 39 vol%) is likely to have facilitated stress concentration. The tensile elastic moduli were not given, so the differences in stiffness cannot be assessed. Because the data for tensile elastic modulus and tensile strength show maxima as a function of fly ash addition level, these data suggest that the increases resulted from dispersion strengthening. The decrease in tensile strength is attributed debonding and cavitation because it is clear that the decrease in the tensile elastic modulus was likely to have derived from the same phenomena and the SEM fracture surface images showed void formation, even though a silane coupling agent and compatibilizer were used to hinder this. The decreases in tensile elastic modulus and tensile strength would have been facilitated by the relatively thin ligament width that resulted from the coarse particle size.
* Comparative study III : The intermediate particle size (<106 [micro]m) at low to medium volume fractions (~4.7-23 vol%) reveals the balanced opposition of dispersion strengthening and stress concentration. That is, the gradual stiffening (except at the highest fly ash addition level) suggests the effect of dispersion strengthening while the gradual decrease in tensile strength suggests the effect of stress concentration. These features were moderated by the effect of the silane coupling agent, which was intended to hinder debonding and cavitation. However, the SEM fracture surface images clearly indicate debonding.
These studies indicate the importance of enhancement of the bonding between the fly ash surfaces and the polymer matrix with the use of a silane coupling agent in comparative study III  and with this as well as an HDPE-g-dibutyl maleate compatibilizer in comparative study II . That is, the coupling agent established a bond between polymer and dispersant [95, 96] while the compatibilizer altered the polymer to increase adhesion [97, 98]. Improvements in the tensile strengths by up to ~25% for composites containing silane-coated fly ash compared to those with uncoated fly ash confirm enhanced bonding between dispersant and matrix with the use of silane . The effect of the silane and compatibilizer is more marked, with increases in tensile strength by up to ~70%, which were attributed by the authors to the reaction between the amine groups in the APTS ([3-aminopropyl]triethoxysilane) coupling agent and the ester groups of the HDPE-gdibutyl maleate compatibilizer .
Because the tensile strengths reported in the present work and in comparative study I  are lower than those reported in comparative study III , it is probable that the enhanced bonding resulting from the silane coupling agent was the cause of this difference. Comparative study IE  confirms the effectiveness of the silane coupling agent as these data directly compare the results for composites containing coated and uncoated fly ash. However, it also is possible that these differences are artefacts of the raw materials and/or processing methods, the reasons for which are mentioned above [86-92].
Elongation at Break. Table 3 shows that the elongations (viz., strains) at break as a function of fly ash addition level are consistent for the present work and comparative studies I  and III ; the same trend was reported in comparative study II . It is noted that the data for comparative study I  were omitted mistakenly from the relevant table in this publication. However, the load-displacement plot for additions of fly ash of particle size ~22 [micro]m could be used for comparison. These data are consistent with the gradual stiffening associated with dispersion strengthening.
Although the value for the elongation at break of ~1300% for the neat polymer reported in comparative study III  is significantly higher than the values reported in the present work and comparative study I , it is within the range for HDPE (~10-1500%) . The maximal processing temperature of 160[degrees]C was significantly lower than that of the present work (220[degrees]C) and comparative study I (200[degrees]C), which reflects the probable retention of longer polymer chain lengths.
Yield Stress. Table 3 shows the yield stresses at low fly ash addition levels. While these data show a very slight increase in the yield stress with increasing fly ash addition, the differences are not statistically significant. Nonetheless, if the trend is indicative of actual behavior, again, it is consistent with the effects of dispersion strengthening.
Yield Strain. Table 3 shows the yield strains at low fly ash addition levels. These data show that the yield strain decreased significantly at the fly ash addition levels of 7.5 and 10 wt%. Because Fig. 3 indicates that the load-extension behavior for the fly ash addition level of 7.5 wt% is essentially identical to those of the lower fly ash addition levels, the decrease in the yield strain at 7.5 wt% fly ash is unexpected. This will be discussed subsequently.
Charpy Impact Strength. Table 3 shows the Charpy impact strength as a function of fly ash addition level. These data are consistent with the preceding mechanical data and reflect the trend for impact resistance indicated in Tables 3 and 4. However, the data are not directly comparable since they consist of impact strength in kJ [m.sup.-2] in the present work and impact resistance in J [m.sup.-1] in comparative study I . Further, the standard deviations of all of these data cause overlapping ranges, thereby bringing into question the statistical significance. Nonetheless, if the data are indicative, then they suggest that the impact properties are affected adversely by fly ash additions resulting from the stress concentration from the fly ash particulates and the low fracture toughness of the vitreous fly ash .
Mechanistic Model. Although none of the comparative studies [9, 16, 50] reported yield stresses or yield strains, the load-extension data were presented graphically in comparative studies I  and III . While the data plots in comparative study I  were not suitable for determination of these parameters graphically, the trends were apparent and consistent with that of the present work; the data plots in comparative study III  could not be differentiated.
Consideration of the data for the present work and the comparative studies [9, 16, 50] provides the basis for a schematic that illustrates the general behavior and considerations relevant to the tensile mechanical properties in terms of load-extension plots. This phenomenological summary is given in Fig. 4. The general features of this model are similar to those underpinning the toughening mechanism of particulate-reinforced polymers [70, 71, 78-81]:
* The neat polymer and composites with low dispersant levels exhibit ductile behavior, which is in the elastic regime.
* At low dispersant levels, the microstructures are dominated by dispersion strengthening, which is moderated by stress concentration.
* At higher dispersant levels, crazing causes the tensile strength to increase and the strain at necking to decrease.
* These microstructures are dominated by the transition from debonding to cavitation, with the consequent increasing formation of voids and ligaments.
* The composites at these higher dispersant levels exhibit plastic behavior and so is in the plastic regime.
* At the highest dispersant levels, crazing increases and ligament width decreases to the point that limited plastic deformation is followed by brittle failure at lower tensile strengths.
* These microstructures are dominated by fibril failure.
* With increasing dispersant levels (not shown), the approach to percolation causes the microstructures to be dominated by the brittle dispersant, which would eliminate plastic deformation and result in true brittle failure behavior.
* With increasing dispersant levels generally, the yield stress increases and the yield strain decreases in the order ductile deformation region [right arrow] crazing region [right arrow] brittle failure region. The trend is a reflection of the dominant effects of dispersion strengthening at lower stress levels.
* With increasing dispersant levels generally, the tensile elastic modulus decreases in the order ductile deformation region [right arrow] brittle failure region [right arrow] crazing region. This trend is a reflection of the dominant effects of void generation and the associated rule-of-mixtures at higher stress levels.
The data for the present work and the comparative studies [9, 16, 50] indicate that the transitions from ductile deformation to crazing to brittle failure are determined principally by the dependent variables (1) dispersant volume, (2) dispersant particle size, (3) intrinsic flaw size (viz., dispersant size), (4) generated flaw size (viz., void size), and (5) strength of dispersant-matrix bond and associated effectiveness of load transfer. The specific features supporting the model and the influence of these variables on it are as follows:
* Present work: For the fine fly ash particle size of ~13 [micro]m (without silane coupling agent), neither the ductile-crazing nor the crazing-brittle transition is observed because the fly ash addition levels are too low.
* Comparative study I: For the fine fly ash particle size of ~14 pm (without silane coupling agent), the ductile-crazing transition occurs in the range ~7.2-10 vol% (15-20 wt%), as revealed by the increase in tensile strength, and the crazing-brittle transition occurs in the range ~ 16-23 vol% (30-40 wt%), as revealed by the decreases in tensile elastic modulus and tensile strength, which result from the generation of voids that form upon debonding and grow upon cavitation.
* Comparative study II: For the coarse fly ash particle size of ~60-160 [micro]m, the ductile-crazing transition is not observed because the fly ash addition levels are too high (13 vol%). However, the crazing-brittle transition occurs in the range 26-39 vol%, depending on the level of compatibilizer, owing to the generation of large voids. The decrease in tensile strength is more marked in the absence of compatibilizer (the effect is not as obvious in the data for the tensile elastic modulus), emphasizing the importance of effective interfacial bonding in the suppression of void generation.
* Comparative study III: For the intermediate fly ash particle size of <106 [micro]m (with silane coupling agent), neither the ductile-crazing nor the crazing-brittle transition is observed. Instead, increasing fly ash addition levels yield clear trends of gradual increase in tensile elastic modulus, deriving from dispersion strengthening, and gradual decrease in tensile strength, deriving from stress concentration from the intrinsic flaws associated with the intermediate fly ash particle size. In this case, the behavior is consistently brittle owing to efficient load transfer between silane-coated fly ash and RHDPE matrix, as will be explained subsequently.
In regard to the prior comment about the decrease in yield strain at the two lowest fly ash addition levels shown in Fig. 3 and Table 3, it is possible that the significant decrease in the yield strain at 7.5 wt% fly ash addition is an indicator of the onset of debonding, to be followed by cavitation. However, this is uncertain since the plots for the yield stresses and yield strains can be seen to be very sensitive to small variations while the data for the variations in the plots for the tensile stresses and strains at necking are more readily apparent. This difference is highlighted in the data distributions shown Table 3. Hence, the data for the yield stresses and strains are considered less indicative than those for the tensile stresses and strains at necking.
It is noted that the data for 10 wt% fly ash addition in Fig. 3 and Table 3 suggest that this addition level has a relatively greater effect on the properties than those of the lower levels. That is, the load-extension data for Fig. 3 and the notched Charpy impact strength data for Table 3 indicate increased brittleness (i.e., decreased fracture toughness) at the highest level of fly ash addition. This observation corresponds to the data for comparative study I  for the higher fly ash addition levels of 15-40 wt%. These load-extension data correspond in that the specimens with [greater than or equal to] 20 wt% fly ash additions exhibit crazing behavior while those [less than or equal to] 15 wt% show ductile behavior. These features are consistent with those summarized in the schematic model of Fig. 4.
A point of difference is that Table 3 shows that the tensile strength of the 10 wt% fly ash addition exhibits no apparent increase over the values of the lower-level additions while the data for comparative study I , which are interpolated from the load-displacement curves, clearly show dispersion strengthening and a ductile-crazing transition in the range 15-20 wt% fly ash addition. The data for the present work do not show the same behavior because the addition levels were too low to initiate the ductile-crazing transition.
It also is possible that the fly ash particles themselves could fracture since typical failure stresses of glass are low and in the range ~5 MPa  to <50 MPa , although none of the fracture surfaces in the present work or the comparative studies I [9, 16, 50] indicate this, even with impact testing. Therefore, the interfacial bond strength, even with coupling agent and compatibilizer, is less than the failure stresses of both the fly ashes and the thin-walled cenospheres.
The results of the present work and those of comparative study I , which contrast with those of comparative studies II  and III , also are likely to have been affected by the former two studies' low extent of interfacial bonding between the fly ash and RHDPE and consequent poor load transfer. This is confirmed by the data for comparative study III , in which the load-extension data for silane-coated and uncoated fly ash in RHDPE are contrasted. When the fly ash is coated with the silane coupling agent, the failure is brittle but, when the fly ash is uncoated, the failure is ductile. Thus, as mentioned above in regard to the effects of the variables in the case of comparative study III , the brittle ceramic carries the load with effective interfacial bonding (effective load transfer) while the ductile polymer carries the load with ineffective interfacial bonding (poor load transfer).
Figures 5 and 6 show the tensile fracture and notched Charpy impact fracture surfaces, respectively, of the neat RHDPE. The tensile fracture surface shown in Fig. 5 reveals Wallner lines, suggesting that fracture initiated at the bottom left of the image . As the crack progressed, rapid failure occurred. It also shows evidence of lamellar extension with practically no void formation across the entire microstructure of crystalline and amorphous components owing to the relatively low strain rate, which allows the RHDPE grains sufficient time to transform slowly into a fibrillary texture by cold drawing. The notched Charpy impact fracture surface shown in Fig. 6 indicates a fairly flat surface across the entire microstructure owing to the high strain rate during impact fracture. It clearly illustrates that the laminar texture has transitioned into a structure of small voids and short fibrillation, which is suggestive of brittle failure even in the absence of fly ash.
Figures 7-10, which illustrate tensile fracture surfaces of fly ash additions in the range 2.5-10 wt%, respectively, indicate extensive ductile fibrillary deformation of the RHDPE matrix generated at the low strain rate used.
Figures 11-14 illustrate the ribbed surfaces of the notched Charpy impact fracture surfaces of samples with fly ash additions in the same range, respectively, generated at the high strain rate used. These show relatively poor interfacial bonding between the fly ash particulates and the RHDPE matrix, which is evident from the presence of voids in the fracture surfaces resulting from pull-out of the hard, strong, and stiff fly ash particulates.
Of note is Fig. 14, which shows the notched Charpy impact fracture surface of the specimen with 10 wt% fly ash addition. In this microstructure, tendrils of matrix where the reinforcement has pulled away can be seen; such tendrils are not visible in the specimens containing lower fly ash additions.
Brittle fracture surfaces of polymers may display hackles, which consist of divergent lines resembling river patterns radiating outward from the fracture origin; they are indicative of plastic deformation [101, 102]. However, no hackles were observed on the tensile fracture surfaces, confirming that RHDPE is not a naturally brittle polymer. Hence, the relatively rough topography of the tensile fracture surfaces supports the view that failure initiation occurred by void generation at the weakly bonded fly ash-RHDPE interfaces.
The present work synthesises data for low-level additions of fly ash to RHDPE and data from three comparative studies [9, 16, 50] that report higher-level additions of fly ash to HDPE and RHDPE. This synthesis provides the basis for a generalised model for the failure behavior of composites comprising hard brittle spherical dispersants (fly ash) in soft ductile polymeric matrices (HDPE and RHDPE) in terms of load-extension plots. The plots are characterised by three behavioral regions (ductile deformation, crazing, and brittle failure) that are distinguished by two transitions (ductile-crazing and crazing-brittle). The locations of these regions and transitions are a function of five dependent variables: (1) dispersant volume, (2) dispersant particle size, (3) intrinsic flaw size (viz., dispersant size), (4) generated flaw size (viz., void size), and (5) interfacial bond strength and associated load transfer. The present work provides potential design parameters for polymer matrices reinforced with waste fly ash in that consideration of the effects of these variables on the resultant load-extension characteristics can be used to predict failure modes and associated trends in tensile elastic modulus, tensile strength, elongation at break, yield stress, and yield strain as well as indications of fracture toughness through comparative trends in brittleness.
The authors thank DONMAR Industries for providing the recycled high density polyethylene and Cement Australia for providing the fly ash. The authors also acknowledge access to the characterization facilities of the Australian Microscopy and Microanalysis Research Facilities (AMMRF) node at UNSW Australia.
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Imrana I. Kabir, Charles C. Sorrell, Mykanth R. Mada, Sagar T. Cholake, Sri Bandyopadhyay
School of Materials Science and Engineering, UNSW Australia, Sydney, NSW 2052, Australia
Correspondence to: I.I. Kabir; e-mail: email@example.com
Published online in Wiley Online Library (wileyonlinelibrary.com).
TABLE 1. Typical ranges of selected properties of high-density polyethylene (HDPE) . Type of Property Unit property Physical True density kg [m.sup.-3] Water absorption % Linear mold shrinkage cm [cm.sup.-1] Mechanical Tensile modulus MPa Flexural modulus MPa Ultimate tensile MPa strength Tensile yield strength MPa Flexural yield strength MPa Elongation at break % Elongation at yield % Notched Charpy impact J [cm.sup.-2] strength Rockwell D hardness -- Shore D hardness -- Electrical Electrical resistivity [ohms] x cm Sheet resistance [ohms] Dielectric constant -- Dielectric breakdown kV-mm strength Dissipation factor -- Thermal Melting point [degrees]C Vicat softening point [degrees]C Processing temperature [degrees]c Maximal service [degrees]c temperature Heat capacity J [g.sup.-1] [[degrees]C.sup.-1] Thermal conductivity W [m.sup.-1] [K.sup.-1] at 20[degrees]C Coefficient of thermal [[degrees]C.sup.-1] expansion at 20[degrees]C Optical Visible light % transmission Haze % Gloss % Type of Property Range property Physical True density 918-1400 Water absorption 0.01-1.50 Linear mold shrinkage 0.003-0.020 Mechanical Tensile modulus 180-1600 Flexural modulus 179-1700 Ultimate tensile 10-50 strength Tensile yield strength 2.4-31.7 Flexural yield strength 14-25 Elongation at break 10-1500 Elongation at yield 6.9-15.0 Notched Charpy impact 0.38-11 strength Rockwell D hardness 60-65 Shore D hardness 55-69 Electrical Electrical resistivity [10.sup.6]-[10.sup.16] Sheet resistance [10.sup.6]-[10.sup.15] Dielectric constant 1-3 Dielectric breakdown 19-150 strength Dissipation factor [10.sup.-4]-[10.sup.-2] Thermal Melting point 110-135 Vicat softening point 67-131 Processing temperature 190-274 Maximal service 41-120 temperature Heat capacity 2.2 Thermal conductivity 0.29-0.50 at 20[degrees]C Coefficient of thermal 22-200 x [10.sup.-6] expansion at 20[degrees]C Optical Visible light 80 transmission Haze 6 Gloss 85 TABLE 2. Chemical analysis by XRF of fly ash. Oxide Weight% Si[O.sub.2] 66.13 [Al.sub.2][O.sub.3] 30.79 Ti[O.sub.2] 1.29 [Fe.sub.2][O.sub.3] 0.58 [K.sub.2]O 0.29 MgO 0.13 S[O.sub.3] 0.10 [Na.sub.2]0 0.06 [P.sub.2][O.sub.5] 0.06 CaO 0.05 MnO BLD Total 99.38 TABLE 3. Specific data for comparative studies (fly ash dispersant). Fly ash addition level (wt%) Parameter Reference 0 2.5 Tensile elastic Present 73 [+ or -] 6 91 [+ or -] 3 modulus work (MPa) (a) 9 490 [+ or -] 25 -- 16 560 -- Tensile strength Present 20 [+ or -] 5 22 [+ or -] 5 (MPa) a work 9 23 [+ or -] 1 -- 16 32 -- Yield stress Present 17 [+ or -] 1 17 [+ or -] 1 (MPa) a work 9 -- -- 16 -- -- Yield strain Present 0.51 [+ or -] 0.01 0.51 [+ or -] 0.01 (%) (a) work 9 -- -- 16 -- -- Elongation at Present 276 [+ or -] 112 200 [+ or -] 132 break (%) (a) work 9 (b) 695 [+ or -] 52 -- 16 1300 -- Impact strength Present 986 [+ or -] 5 586 [+ or -] 6 (kJ [m.sup.-2]) work Impact resistance 9 48 [+ or -] 9 -- (J [m.sup.-1]) Impact data 16 -- -- Fly ash addition level (wt%) Parameter Reference 5.0 7.5 Tensile elastic Present 83 [+ or -] 6 88 [+ or -] 5 modulus work (MPa) (a) 9 500 [+ or -] 24 -- 16 -- -- Tensile strength Present 21 [+ or -] 1 22 [+ or -] 3 (MPa) a work 9 23 [+ or -] 1 -- 16 -- -- Yield stress Present 17 [+ or -] 1 18 [+ or -] 1 (MPa) a work 9 -- -- 16 -- -- Yield strain Present 0.49 [+ or -] 0.01 0.27 [+ or -] 0.01 (%) (a) work 9 -- -- 16 -- -- Elongation at Present 120 [+ or -] 65 30 [+ or -] 14 break (%) (a) work 9 (b) 450 [+ or -] 67 -- 16 -- -- Impact strength Present 574 [+ or -] 6 580 [+ or -] 6 (kJ [m.sup.-2]) work Impact resistance 9 39 [+ or -] 5 -- (J [m.sup.-1]) Impact data 16 -- -- Fly ash addition level (wt%) Parameter Reference 10.0 15.0 Tensile elastic Present 88 [+ or -] 5 -- modulus work (MPa) (a) 9 520 [+ or -] 16 575 [+ or -] 46 16 635 -- Tensile strength Present 19 [+ or -] 1 -- (MPa) a work 9 23 [+ or -] 1 25 [+ or -] 1 16 31 -- Yield stress Present 18 [+ or -] 1 -- (MPa) a work 9 -- -- 16 -- -- Yield strain Present 0.20 [+ or -] 0.01 -- (%) (a) work 9 -- -- 16 -- -- Elongation at Present 22 [+ or -] 10 -- break (%) (a) work 9 (b) 295 [+ or -] 166 81 [+ or -] 4 16 47 -- Impact strength Present 466 [+ or -] 7 -- (kJ [m.sup.-2]) work Impact resistance 9 33 [+ or -] 3 32 [+ or -] 4 (J [m.sup.-1]) Impact data 16 -- -- Fly ash addition level (wt%) Parameter Reference 20.0 30.0 Tensile elastic Present -- -- modulus work (MPa) (a) 9 645 [+ or -] 56 875 [+ or -] 7 16 640 820 Tensile strength Present -- -- (MPa) a work 9 28 [+ or -] 1 28 [+ or -] 1 16 28 25 Yield stress Present -- -- (MPa) a work 9 -- -- 16 -- -- Yield strain Present -- -- (%) (a) work 9 -- -- 16 -- -- Elongation at Present -- -- break (%) (a) work 9 (b) 27 [+ or -] 11 21 [+ or -] 5 16 17 18 Impact strength Present -- -- (kJ [m.sup.-2]) work Impact resistance 9 30 [+ or -] 1 28 [+ or -] 3 (J [m.sup.-1]) Impact data 16 -- -- Fly ash addition level (wt%) Parameter Reference 40.0 Tensile elastic Present -- modulus work (MPa) (a) 9 610 [+ or -] 60 16 1010 Tensile strength Present -- (MPa) a work 9 29 [+ or -] 1 16 21 Yield stress Present -- (MPa) a work 9 -- 16 -- Yield strain Present -- (%) (a) work 9 -- 16 -- Elongation at Present -- break (%) (a) work 9 (b) 20 [+ or -] 19 16 10 Impact strength Present -- (kJ [m.sup.-2]) work Impact resistance 9 25 [+ or -] 4 (J [m.sup.-1]) Impact data 16 -- (a) All crosshead speeds 50 mm [min.sup.-1]. (b) The data for elongation at break were omitted inadvertently but the data for the composites using fly ash particle size of ~22 [micro]m (instead of ~14 [micro]m) could be determined graphically, so these have been used. TABLE 4. Sample/preparation parameters and apparent data trends for comparative studies (fly ash and cenosphere dispersants). Reference Sample/preparation Present work parameters Matrix RHDPE Type of dispersant Coal fly ash Silane coupling agent None Compatibilizer None Absolute particle size of ~13 dispersant ([micro]m) Relative particle size of Fine dispersant Addition levels of (wt%) 0, 2.5, 5.0, 7.5, 10 dispersant (d) (vol%) 0, 1.1, 2.3, 3.5, 4.7 Processing instruments Rotational mixer + Injection molder Processing temperature 220 ([degrees]C) Processing pressure (MPa) 100 Apparent data trends Tensile elastic modulus Increase: dispersion/crazing Tensile strength Increase: dispersion/crazing Elongation at break Decrease: dispersion Yield stress Increase: dispersion Yield strain Decrease: Debonding/cavitation Impact strength/ Decrease: resistance stress concentration Reference Comparative Sample/preparation study I  parameters Matrix HDPE Type of dispersant Coal fly ash Silane coupling agent None Compatibilizer None Absolute particle size of ~14 dispersant ([micro]m) Relative particle size of Fine dispersant Addition levels of (wt%) 0, 5, 10, 15, 20, 30, 40 dispersant (d) (vol%) 0, 2.3, 4.7, 7.2, 10, 16, 23 Processing instruments Twin-screw extruder Processing temperature 130-200 ([degrees]C) Processing pressure (MPa) 9-11 Apparent data trends Tensile elastic modulus Increase: dispersion/crazing [right arrow] Decrease: debonding/cavitation Tensile strength Increase: dispersion/crazing Elongation at break Decrease: dispersion Yield stress Not given Yield strain Not given Impact strength/ Decrease: resistance stress concentration Reference Comparative Sample/preparation study II  parameters Matrix HDPE Type of dispersant Cenospheres Silane coupling agent APTSa Compatibilizer HDPE-g-DBMb Absolute particle size of ~60-160c dispersant ([micro]m) Relative particle size of Coarse dispersant Addition levels of (wt%) -- dispersant (d) (vol%) 13, 26, 39 Processing instruments Rheometer + injection molder Processing temperature 160e ([degrees]C) Processing pressure (MPa) 100e Apparent data trends Tensile elastic modulus Increase: dispersion/crazing [right arrow] Decrease: debonding/cavitation: Tensile strength Increase: dispersion/crazing [right arrow] Decrease: debonding/cavitation Elongation at break Decrease: dispersion Yield stress Not given Yield strain Not given Impact strength/ Not given resistance Reference Comparative Sample/preparation study III  parameters Matrix RHDPE Type of dispersant Coal fly ash Silane coupling agent APTSfl] Compatibilizer None Absolute particle size of <106 dispersant ([micro]m) Relative particle size of Intermediate dispersant Addition levels of (wt%) 10, 20, 30, 40 dispersant (d) (vol%) 4.7, 10, 16, 23 Processing instruments Rheometer + Uniaxial hot press Processing temperature 160 ([degrees]C) Processing pressure (MPa) 14 Apparent data trends Tensile elastic modulus Increase: dispersion/crazing Tensile strength Decrease:stress concentration Elongation at break Decrease: sispersion Yield stress Not given Yield strain Not given Impact strength/ Not given resistance (a) (3-Aminopropyl)triethoxysilane. (b) HDPE-g-dibutyl maleate. (c) Particle size not reported; size range estimated from SEM images of fracture surfaces. (d) Conversion of wt% to vol% based on average HDPE true density of 960 kg [m.sup.-3] [591 and fly ash true density of 2170 kg [m.sup.-3] , (e) Private communication, R.R.N. Sailaja.
Please note: Some tables or figures were omitted from this article.
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|Author:||Kabir, Imrana I.; Sorrell, Charles C.; Mada, Mykanth R.; Cholake, Sagar T.; Bandyopadhyay, Sri|
|Publication:||Polymer Engineering and Science|
|Date:||Oct 1, 2016|
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