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Effects of catalyst content, anhydride blending and nanofiller addition on anhydride-cured epoxidized linseed oil based thermosets Christopher N. Kuncho, Daniel F. Schmidt, and Emmanuelle Reynaud Ind. Eng. Chem. Res., Just Accepted Manuscript • DOI: 10.1021/acs.iecr.6b03919 • Publication Date (Web): 13 Jan 2017 Downloaded from http://pubs.acs.org on January 14, 2017

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Industrial & Engineering Chemistry Research

Effects of catalyst content, anhydride blending and nanofiller addition on anhydride-cured epoxidized linseed oil based thermosets Christopher N. Kuncho1 Daniel F. Schmidt,*1 and Emmanuelle Reynaud2 1

Department of Plastics Engineering, University of Massachusetts Lowell, One University Avenue, Lowell, Massachusetts 01854, United States

2

Department of Mechanical Engineering, University of Massachusetts Lowell, One University Avenue, Lowell, Massachusetts 01854, United States *

[email protected]

KEYWORDS: bioepoxy, epoxidized linseed oil, anhydride, hardness, dynamic mechanical analysis

ABSTRACT

A series of formulations based on epoxidized linseed oil, ELO cured with various anhydride hardeners were investigated to assess the potential for enhanced elastic properties and thermal transitions through variations in catalyst concentration (1,8-diazabicyclo[5.4.0]undec-7-ene,

ACS Paragon Plus Environment

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DBU), hardener type (nadic methyl anhydride, NMA, methyltetrahydrophthalic anhydride, MTHPA, and their blends with phthalic anhydride, PA) and nanofiller (organically modified hydrotalcite, Perkalite F100S) content. To facilitate rapid screening, Shore D hardness was used as a proxy for Young’s modulus, and the limits of this approach were tested via comparisons with modulus measurements made via dynamic mechanical analysis (DMA). ELO cured with MTHPA at a 1:1 epoxy / anhydride molar ratio with 4.5 phr DBU gave a particularly attractive combination of rigidity, homogeneity and alpha transition temperature rarely seen in thermosets comprised of a 100% bio-based epoxy component. Nanocomposite formation provided a modest increase in modulus with no change in thermal transitions.

INTRODUCTION In 2013 the global market for epoxy systems was valued at 6.64 billion dollars, with 6.9% annual growth forecasted to 2020 to a 10.55 billion dollar market dominated by coating and structural applications1. In tandem, emergent bio-based resins are anticipated to expand their niche in the market, creating competitive and novel opportunities1. In this context, industrial interest in bio-based resins is clear. As far as the scientific literature is concerned, epoxidized vegetable oils in particular have attracted attention both as additives for bisphenol A diglycidyl ether (BADGE) based systems, in coatings2-4 and as the sole basis structural thermosets5-7. While significant research has been conducted on the subject of bio-based epoxy formulations8-9 and a number of systems have been commercialized (e.g. Super SAP, Epicerol, Greenpoxy, etc.), much of the focus has been on materials derived from blends of conventional (petroleum-derived) and bio-based epoxies. In contrast, the work reported here focuses on formulations comprised of a


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100% bio-based epoxy component – in particular, an epoxidized vegetable oil. Investigations of such oils investigated have largely been based on their availability and their degree of unsaturation10. Some common oils of interest are epoxidized soybean oil (ESO)5, 10-15, epoxidized canola oil (ECO)16-17, and epoxidized linseed oil (ELO)7, 18-20. Each of these resins have shown different capacities for achieving fully formed networks; as expected, however, not all epoxidized oils are in fact equal. As natural products, epoxidized vegetable oils (EVOs) possess no absolute chemical structure, but are, in fact, blends of structurally similar triglycerides, which have varying fatty acid chain lengths attached to a parent glycerol molecule21. This blend of molecular weights, functionalities, and thus equivalent weights leads to differing behaviors when reacted with hardeners, as well as a lack of absolute stoichiometry to adhere to with regards to formulation. This can be a disadvantage when compared to traditional BADGE epoxies, which can be made to display less variability. A broader molecular weight distribution is directly correlated to a compromise in physical properties, as well as a broadening of transition temperatures, which in turn affects end use capabilities for structural applications22. As a result, the conventional wisdom used in formulating traditional epoxy systems (especially with respect to hardener and filler content) may not directly translate to bio-based resins. This highlights the need for a method of screening a large number of formulations rapidly in order to develop a better understanding of potentially complicated structure-properties relations in these bio-based systems while at the same time eliminating low performance formulations early in the process18. Indeed, most physical and thermal properties testing requires careful casting and forming of significant numbers of samples in a highly reproducible fashion, and those methods that do not, such as differential scanning


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calorimetry (DSC), or parallel plate rheology, require expensive equipment and non-negligible measurement times. Following screening, a more in-depth analysis of thermoset properties is required, with special attention paid to thermal and mechanical properties. Thermal transitions are of significant importance when considering materials for structural applications. Due to the nature of most crosslinked materials, these transitions typically occur over a broader range of temperatures when compared to thermoplastics. Traditionally, transition temperatures are captured by (DSC), thermomechanical analysis (TMA), and dynamic mechanical rheological testing (DMRT)23. Mechanical properties also play an important role in determining potential areas of application, with the need to modify such properties driving further modifications in formulation, the use of additives, etc. For example, typical BADGE formulations are rigid and brittle24, and often require toughening in the form of additives25 or in some cases blending with thermoplastics26 to achieve the desired properties and / or failure modes. In the context of such properties investigations, previous investigations into anhydride-cured bio-based epoxies highlight their potential5-6, 27-30. This work examines the effects of hardener blending and the addition of a commercial layered nanofiller on anhydride-cured epoxidized linseed oil (ELO) formulations, as observed through changes in hardness, storage modulus and thermal transition temperature. The first part of the study focuses on optimization of resin composition while at the same time assessing the applicability of known correlations between Shore D hardness and elastic modulus31-33. NMA and MTHPA were tested, either by themselves or in blends, to understand the effects of anhydride selection and blending and identify optimized compositions. In the second part of the study, the effect of the addition of a commercial hydrotalcite nanofiller on the thermomechanical properties of an optimized resin formulation


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was evaluated, given in the context of our group’s prior work observation of enhanced mechanical performance in amine-cured ELO in particular34-36. Such nanofillers can also provide self-extinguishing behavior37-41 and augmented thermal properties42-47; an examination of such properties is outside of the scope of the current submission but will be presented in a separate report. Here, the overall significance of the efforts reported is in clarifying the potential of formulations comprised of a 100% bio-based epoxy component to provide (thermo)mechanical properties appropriate for structural applications.

EXPERIMENTAL Materials The resin used for these experiments is an epoxidized linseed oil (ELO) acquired from American Chemical Service. This material, Epoxol 9-5, was selected based on its higher oxirane content (9.57%) (epoxy equivalent weight, EEW, of 167.18 g/eq) as compared to 6.9% (EEW: 231.88 g/eq) for epoxidized soybean oil (ESO) and 8.6% (EEW: 186.04 g/eq) for bisphenol A diglycidyl ether (BADGE). The epoxy groups are secondary in nature, which is expected to equate to a lower reactivity when compared to resins containing primary oxirane rings, as with BADGE, but should yield stiffer materials than those made with other epoxidized vegetable oils (EVOs) such as ESO, due to its higher concentration of epoxide groups, and thus the potential for a higher crosslink density. The hardeners used for this series of experiments are nadic methyl anhydride (NMA) purchased from Polysciences, methyltetrahydrophthalic anhydride (MTHPA), donated by Huntsman as Aradur 917, and phthalic anhydride (PA) purchased from Alfa Aesar. NMA was


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selected for its low viscosity at room temperature as well as its ability to reach full cure more rapidly than other anhydrides evaluated for this effort. MTHPA was used because of its potential to form a significantly more rigid network than NMA, as well as its low viscosity at room temperature. Finally, PA was chosen for its rigidity, low equivalent weight and ability to dissolve in the other two anhydrides, which mitigates its main liability of being a solid at room temperature. The catalyst used for this research, 1,8-diazabicyclo[5.4.0]undec-7-ene (DBU), is a bicyclic amidine base shown in previous research to effectively facilitate the epoxy curing reaction with ELO's secondary oxirane groups. This catalyst has also been shown to work well in other epoxidized vegetable oil based systems, most notably materials based on epoxidized soybean oil30. Finally, the hydrotalcite nanofiller utilized in this effort was AkzoNobel’s Perkalite F100S, an Al/Mg layered double hydroxide nanoclay modified with hydrogenated fatty acids composed of equal parts of C16 and C18 chains48-50. This material was selected for preliminary experiments given prior work showing good compatibility and improved mechanical properties when added to amine-cured ELO-based bioepoxies34-36. This nanofiller has an aggregate size of 22 µm or less (D99 = 22 µm), typical layer thickness of ~0.5 nm, lateral dimensions of between 100 and 150 nm and an aspect ratio of approximately 200-30048-50. The nanofiller concentrations studied in this work were chosen in light of the aforementioned aspect ratios so as to avoid substantially exceeding the estimated percolation threshold of 0.42-0.64 vol% inorganic51, thus ensuring that nanofiller agglomeration will be disfavored and dispersion state will depend primarily on thermodynamic compatibility52.


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Sample Preparation

For effective mixing and degassing of all formulations, a FlackTek Inc. DAC 600.1 VAC-P Speed Mixer was used. In order to facilitate degassing, an awl was used to create a small hole in the center of the speed mixer cup lid. An attached Leybold LVO 100 vacuum pump connected to the Speed Mixer through a Labconco CentriVap -50°C cold trap was also used to rapidly apply vacuum as the formulations were mixed. All reagents, except for catalyst, were preheated at 80°C until a uniform temperature was achieved. Preheated resin was added to an empty speed mixer cup (cup size dependent on batch size) and hot hardener was added directly into the resin using plastic transfer pipettes. The sample was then mixed for 1 minute at 1000 rpm. Catalyst (at an optimized level of 4.5 parts per hundred resin by weight, abbreviated phr) was then added directly to the mixture; at elevated temperatures, the viscosity decreased to the point where mixing between hardener and resin occurred freely, and as a result the catalyst caused a reaction/discoloration when it came in contact with the resin. Filler was then added as needed, and the container sealed. Formulations were mixed in the speed mixer over the course of two consecutive 10 minute mixing cycles (the maximum time allowed for a single mixing run) at 1000 rpm. This ensures that the resin, a natural product, is thoroughly mixed with the rest of the components, and that (where applicable) enough mixing has occurred that the nanofiller is effectively dispersed and distributed. In all cases, following mixing, samples were transferred to appropriate tooling and cured in a Lindberg Blue M forced air convection oven.


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Three sample types were made. 20 gram samples were cured in simple disposable 2” aluminum weighing pans, final sample thickness ~1/3”, was done so that easily accessible flat surfaces were available for hardness testing. The pans were easily peeled from the specimens. Small batches were cast into 2.9”×1.9”×1.0” Freshware CB-105RD silicone molds for dynamic mechanical analysis (DMA) samples. These molds were sacrificial, as it was discovered that, even with mold release, the tested systems adhered strongly to the molds. For larger scale sample production, custom open faced aluminum molds with cavity sizes of 5”×5”×0.197” and 18”×13”×0.197” were milled and used with a mold release applied (Henkel Frekote 700-NC) to avoid adhesion. In all cases, uniform thickness was achieved by careful leveling of the molds before pouring using a bubble level and adjusting set screws that constituted the legs of the aluminum mold base. Samples were sectioned into DMA bars (55 mm×13 mm×3 mm) using a Chicago Electric 4” table saw fitted with a diamond tipped cutting wheel and a rip fence fixed at 0.5”. This allowed for uniform sample widths to a precision of 0.01 mm. Specimen width was scored using Fowler Vernier calipers, and cut using the saw’s guide, which allowed for uniform sample lengths to a precision of 0.1 mm.

Optimized curing conditions were chosen based on preliminary studies of hardness variations as a function of process parameters. In particular, NMA-based formulations were cured for 8 hours at 180°C, while MTHPA-based formulations were cured for 1 hour at 90°C and then 8 hours at 180°C in order to suppress the void formation observed in these systems when heated directly to 180°C for curing.


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The experimental design for this research followed a simple one-variable-at-a-time format where NMA and MTHPA served as the two primary hardeners for the study. PA was blended into the hardener of choice (MTHPA or NMA) in 10 wt% increments up to 50 wt% PA vs. total hardener content in a given formulation, with the effective equivalent weights of each hardener blend calculated using a weighted average of the equivalent weights of the individual hardeners comprising the blend. The effect of the concentration of the Perkalite F100S nanofiller was investigated using the MTHPA both neat, and blended with 30 wt% PA as a matrix. The blended system was selected for this experiment due to its increased hardness and homogeneity when compared to the other anhydride formulations examined, while the pure hardener was included as a baseline.

Test Methods Hardness (Shore D)

Hardness testing was performed according to ASTM D2240 using a Fowler Shore D durometer; results were reported as the average of five measurements, with the standard deviation calculated from these same measurements. The individual results were then converted to an estimated Young’s modulus using Equation 1 below, as outlined by Mix et al and governed by their assumptions that the material should behave in an elastic fashion in the context of the hardness measurement33.


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Equation 1. The Mix and Giacomin model for estimating Young’s modulus based on Shore D hardness.

This allowed for statistical analysis of the converted values to be performed instead of being left with a converted average, which obfuscated the inherent variability in the measurement.

DMA

Dynamic mechanical rheological testing was performed using a TA Instruments Q800 dynamic mechanical analyzer (DMA). Elastic behavior and glass transition temperature were assessed in flexure (dual cantilever beam) via temperature sweeps from -140°C to 220°C using a maximum displacement of 25 µm and a frequency of 1 Hz. The storage modulus at 25°C was taken as the room temperature Young’s modulus, while information on the alpha transition was derived from the location of the associated peak in the loss modulus curves52. Instrumental error was assessed by repeatedly (n = 20) fixturing and measuring the thermal transitions and elastic moduli of a polystyrene specimen (chosen given roughly similar moduli and main relaxation temperatures), then propagating the relative error measured in this fashion to measurements made on the epoxy resins studied here.

Distortion Temperature Under Load DTUL values were determined using a method outlined in a white paper provided by TA Instruments54 which allowed for this measurement to be made using the Q800 DMA while ensuring parity with ASTM D648.


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RESULTS AND DISCUSSION Effect of process parameters Preliminary studies concerning the effects of curing conditions were carried out at 180°C based on the observation that lower temperatures resulted in a very sluggish cure response while higher temperatures favored degradation. First, samples of ELO / NMA / 5 wt% DBU were cured for 48 hours at 180°C to ensure complete crosslinking and post-curing. Shore D hardness values were then taken on both sides of these completely crosslinked sample to check for even curing and phase separation. Replicates were then cured at 180°C while systematically decreasing the cure time until a significant reduction in the Shore D hardness of the samples was observed. Optimal cure times were defined as the minimum cure times at 180°C necessary to reach the same Shore D hardness observed following 48 hours of curing at 180°C.

Effect of epoxy / anhydride stoichiometry The modification of reaction stoichiometry represents a common means of altering the properties of epoxy / anhydride networks. As a consequence, following optimization of the cure cycle using a 1:1 epoxy / anhydride molar ratio, a small number of experiments were performed to assess the effects of small variations (±10 mol% in 5 mol% increments) in reaction stoichiometry. As it was observed that Shore D hardness was at a maximum at the idealized 1:1 epoxy / anhydride molar ratio, this stoichiometry was defined as optimal and used for all subsequent work.

Effect of catalyst content


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The effect of catalyst content on hardness was studied in both the ELO / MTHPA formulation (Figure 1) and the ELO / NMA formulation (Figure 2). Here, in order to develop a more complete appreciation for the behavior of these samples, Shore D hardness measurements were carried out on the the upper (air-contacting) surface (Sample Top) and lower (mold-contacting) surface (Sample Bottom) of fully cured specimens that had been cast in open-face molds. In addition, the difference between the values measured from the top and bottom of these samples (∆ Hardness) was used to assess sample homogeneity. Regardless of anhydride hardener in use, an initial increase in hardness with DBU content is followed by a more or less constant hardness at the highest catalyst concentrations. While the hardness of the NMA formulation is more variable in general (hence the larger number of concentrations studied) and much more sensitive to DBU content under these curing conditions, the highest hardness values are reached in the MTHPA formulation. The optimized catalyst concentration was defined as the lowest concentration providing maximum hardness and minimum difference in hardness (termed ∆ hardness) between the top and bottom of the sample in both the MTHPA and NMA cured networks. This resulted in selection of 4.5 phr DBU as the optimal catalyst concentration. The non-uniformity of hardness between the top and bottom of the samples, first noted in prior research, is a common occurrence with ELO resins, and likely occurs in other materials as well. It is hypothesized that this is due to localized loss of hardener or lower molecular weight resin components due to volatilization at the matrix/air interface, with support offered in the literature55.


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90 80 70

Hardness (Shore D)

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60 50 40 30 20

Sample Top Sample Bottom ∆ Hardness

10 0 1

2

3

4

5

6

DBU Concentration (phr)

Figure 1: Shore D hardness of the upper (air-contacting) surface (Sample Top) and lower (moldcontacting) surface (Sample Bottom) of cast samples of fully cured ELO / MTHPA as a function of DBU content, and the difference between these values (∆ Hardness).


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90 80 70

Hardness (Shore D)

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60 50 40

Sample Top Sample Bottom ∆ Hardness

30 20 10 0 0

2

4

6

8

10

DBU Concentration (phr)

Figure 2: Shore D hardness of the upper (air-contacting) surface (Sample Top) and lower (moldcontacting) surface (Sample Bottom) of cast samples of fully cured ELO / NMA as a function of DBU content, and the difference between these values (∆ Hardness).

Effect of phthalic anhydride (PA) substitution Figure 3 shows the changes in Shore D hardness values as the amount of PA is increased in ELO formulations containing optimal catalyst concentrations. Above 50 wt% PA, material handling became difficult due to the tendency of the anhydride blends to solidify. The 100 wt% PA systems only differ in catalyst loading, using the respective loadings for each parent formulation. MTHPA systems were more consistent, and consistently harder than NMA formulations trending toward the 100 wt% PA systems values, with those containing 10-30 wt%


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PA being the most uniform. In contrast, trends in the NMA system were less clear-cut, with good uniformity in hardness observed at 20 wt% and 40 wt% PA loadings.

95

90

Hardness (Shore D)

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85

80

75

MTHPA Top MTHPA Bottom NMA Top NMA Bottom

70

65 0

20

40

60

80

100

Phthalic Anhydride (wt%)

Figure 3: Hardness of ELO formulations containing MTHPA or NMA as a function of PA concentration.

In order to explore the applicability of Shore D hardness / modulus correlations in our materials system in particular, a plot of DMA storage modulus (E’) at 25°C versus Shore D hardness for all formulations in the ELO/MTHPA/PA blend series is shown in Figure 4, with the line indicated on the figure representing the best fit of the Mix and Giacomin model relating these two quantities, using Poisson’s ratio as the adjustable parameter. This fit is imperfect for several reasons. First, Poisson’s ratio may not be the same in all samples. Second, hardness measurements are local whereas storage modulus measurements involve the entire sample; indeed, data from formulations based on NMA-PA blends were omitted from the plot due to


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increased inhomogeneity. Third, Shore D hardness values shares with many other hardness values the fact that their precision is degraded both generally and as far as modulus correlations are concerned when making measurements at the extremes of the scale; indeed, the measurement range defined for the Shore D scale is 20-90, vs. the Shore D modulus values measured here, which are in the mid-80s. This effect is further exacerbated by the very limited range of the Shore D values treated via this approach. While these results do not invalidate the Shore D hardness method as a useful screening tool in the context of the work reported here, they do emphasize the limitations associated with using such data to assess Young’s moduli.

1900

1800

1700

E' at 25°C

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1600

1500

1400

ELO/MTHPA/PA Series Data Mix/Giacomin Model

1300 84

85

86

87

88

89

90

Hardness (Shore D, sample bottom)

Figure 4: E’ vs Shore D hardness (sample bottom) with Mix/Giacomin fit for all formulations in the ELO/MTHPA/PA blend series using Poisson’s ratio as the fitting parameter.

In order to further study the predictions made by the Mix & Giacomin model in the context of the current results, Table 1 shows, for all ELO/MTHPA systems, the measured Shore D hardness


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and E’ values at 25°C as well as the hardness values calculated from the E’ values using the aforementioned model and assuming a Poisson’s ratio of ν = 0.3656.

Table 1: Experimental hardness and E’ data for all formulations in the ELO/MTHPA/PA blend series , as well as hardness data predicted from the E’ values using the Mix & Giacomin model with ν = 0. 3656.

PA Content 0% 10% 20% 30% 40% 50% 100% 0% 10% 20% 30% 40% 50% 100%

Side Bottom Bottom Bottom Bottom Bottom Bottom Bottom Top Top Top Top Top Top Top

Hardness (Shore D) 87 (+/- 0.00) 85.6 (+/- 1.24) 85.3 (+/- 0.84) 85.7 (+/- 0.45) 86.7 (+/- 0.57) 86.4 (+/- 0.42) 88.2 (+/- 0.57) 85.2 (+/- 0.84) 84.4 (+/- 1.19) 85.1 (+/- 0.65) 85.5 (+/- 0.50) 84.8 (+/- 0.57) 83.7 (+/- 0.84) 87.1 (+/- 0.96)

E' at 25°C Predicted Hardness Difference vs. (MPa) (Shore D) actual hardness (%) 1725 (+/- 53) 86.5 (+/- 2.68) 0.56% 1789 (+/- 55) 86.7 (+/- 2.69) 1.33% 1766 (+/- 55) 86.7 (+/- 2.69) 1.59% 1654 (+/- 51) 86.2 (+/- 2.67) 0.63% 1701 (+/- 53) 86.4 (+/- 2.68) 0.32% 1526 (+/- 47) 85.7 (+/- 2.66) 0.78% 1805 (+/- 56) 86.5 (+/- 2.68) 0.36% 1725 (+/- 53) 86.7 (+/- 2.69) 1.54% 1789 (+/- 55) 86.7 (+/- 2.69) 2.77% 1766 (+/- 55) 86.2 (+/- 2.67) 1.83% 1654 (+/- 51) 86.4 (+/- 2.68) 0.87% 1701 (+/- 53) 85.7 (+/- 2.66) 1.91% 1526 (+/- 47) 86.8 (+/- 2.69) 2.42% 1805 (+/- 56) 86.8 (+/- 2.69) 0.36%

Here it can be seen that Shore D hardness values estimated from DMA E’ measurements at 25°C do not differ from actual Shore D hardness measurements by more than ~3% in the ELO/MTHPA case. While uncertainties in the measurement of the E’ values limit the significance of such correlations when considered in isolation, this emphasizes the potential of the approach for this family of materials. In contrast, a similar treatment of the ELO/NMA/PA blend system (not shown) results in predicted hardness values overestimated by as much as ~22%. This observation is statistically significant, and provides clear and useful evidence of the fact that the elastic properties of the sample surface tend to differ significantly from that of the


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bulk in this family of materials. One explanation for this could be the greater volatility of NMA vs. MTHPA at the high temperatures required for the complete cure of these systems. Volatilization of the hardener could result in reductions in hardness at the sample surface. If these reductions are local in nature, however, the modulus of the overall part as measured via DMA will be only minimally affected, resulting in measured modulus values higher than expected based on the measured hardness values. When these measured modulus values are converted back to hardness, they will tend to be overestimates as a result. We emphasize here that these results should be taken as illustrative of the differences between local and global measurements of elastic properties coupled with limitations in the precision of the relevant measurements, not as evidence of a lack of conformity with the Mix and Giacomin model or the presence of statistical outliers in the Shore D hardness measurements, which we have ruled out using accepted methods57.

The DMA storage modulus data at 25°C for both series of blends is shown in Figure 5. The MTHPA formulation starts at a higher E’ value than the NMA formulation, and both trend toward the same storage modulus as PA content is increased. While it was initially posited that PA, being highly compact, would provide increased moduli in both systems, MTHPA’s superior storage modulus may be explained the combined presence of a methyl group and a double bond, which have been posited to inhibit rotation between conformers and bond distortions in anhydride-cured epoxies, and, in the latter case, to provide reactive sites for additional crosslinking reactions58. As NMA is diluted with PA, the modulus values observed trend towards those of the pure PA network. In the case of MTHPA, however, it appears that the modulus of the blended system may actually fall below that of both the pure MTHPA and pure PA based


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networks. This is best explained by the observation that the dilution of unsaturated anhydrides with sufficient quantities of anhydrides that (like PA) are incapable of double-bond mediated crosslinking tends to result in a modest reduction in modulus58. While the same effect may be operative in the NMA / PA blends as well, the results indicate that it is the addition of the higherperforming PA rather than double-bond mediated crosslinking by the NMA that drives increases in modulus in this system.

2.0 1.8 1.6 1.4

E' (GPa)

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60


1.2 1.0 0.8 0.6

MTHPA NMA

0.4 0.2 0.0 0

20

40

60

80

100


Figure 5: E’ at 25°C of ELO formulations containing MTHPA or NMA as a function of PA concentration.

Figure 6 illustrates variations in alpha transition temperature (Tα), as assessed by the location of the peak in the loss modulus data from the relevant DMA ramps for each system. Tα for the MTHPA formulations was the highest of any tested before the addition of phthalic anhydride. As PA was added to the MTHPA based system Tα was suppressed slightly, up to approximately


19


15°C. The neat NMA system exhibited the lowest Tα of any formulations tested, ~30°C, with steady increases observed as the percentage of PA was increased up to 30 wt%, beyond which a plateau was reached.

100 90 80

Temperature (°C)

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

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70 60 50 40 30

MTHPA NMA

20 10 0 0

20

40

60

80

100


Figure 6: Tα (from the temperature at which the peak in E” is observed) for ELO formulations containing MTHPA or NMA as a function of PA concentration.

The neat PA formulations exhibited Tα values above those noted in both the blended systems and the unblended parent formulations. Tα values for the blended NMA and MTHPA formulations appear to converge in the range of 30-50 wt% PA. Here the trend in the properties of the blended NMA system seems much more consistent with the performance of the neat PA system than does the trend in the blended MTHPA system. One explanation for these observations relates to relative differences in anhydride reactivity. With such differences, the


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formation of a homogeneous, “copolymeric” network displaying properties intermediate between the two “homopolymeric” networks would be favored. Large differences in reactivity, on the other hand, would be expected to give two more or less homopolymeric networks, with such a situation favoring macroscale phase-separation as molecular weight increases prior to the gel point. This could compromise network connectivity and homogeneity, thus reducing E’ as well as the temperature at which the onset of the alpha transition is observed. The reality of the situation appears to be intermediate between these two extremes. DMA data for representative formulations (Figures S1 and S2, Supporting Information) show broad main relaxation peaks (typical of thermosets in general and vegetable oil based materials in particular), with no evidence of the coexistence of distinct phases with significantly different properties at up to 30 wt% PA but the observation of a clear shoulder on the alpha transition peak at 50 wt% PA (albeit without the observation of macroscale phase-separation in any sample in this range). This may be rationalized in light of the argument that anhydride reactivity in an epoxy / anhydride system should track with the dissociation constant of the diacid derived from the anhydride58. Predictions of the dissociation constants of NMA, MTHPA and PA made using the ARChem SPARC physicochemical properties calculator1 indicate that, while the diacids formed from NMA and the MTHPA isomers typical of anhydrides used for epoxy curing59 have very similar pKa1 and pKa2 values (3.88-3.94 and 5.64-5.70, respectively), that the analogous values for the diacid of PA are distinctly different (3.05 and 5.57, respectively). These lower values in the case of the diacid formed from PA would imply that the PA is likely more reactive than either NMA or MTHPA, especially in the case of the first epoxy / anhydride reaction. This difference in reactivity helps to justify an increasing degree of microphase separation as PA

1

See http://archemcalc.com/


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content is increased, but it cannot account for the opposing trends in Tα observed in the blended NMA and MTHPA systems. This is once more explained by modulation of the potential for double-bond mediated crosslinking, which has been shown to impact heat distortion temperature more significantly than stiffness58. As in the case of the modulus data, it would appear that the Tα values of the NMA-based blends are much less affected by dilution, implying that double-bond mediated crosslinking may be less important with NMA than with MTHPA in the systems studied here. In sum, then, while the trends in Tα are more complex than those observed in the room temperature storage modulus, both may be better understood in light of prior work on epoxy / anhydride thermosets.

Effect of Nanofiller

In addition to anhydride blends, the effect of nanofiller on properties was also investigated. Nanocomposite formation was carried out using the highest unblended system (ELO/MTHPA) as well as the highest performing, most homogeneous and most readily processed blend formulation (ELO/MTHPA with 30 wt% PA).


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ELO/MTHPA ELO/MTHPA + 30% PA

2.0

1.9

E' (GPa)

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60


1.8

1.7

1.6

0.0

0.2

0.4

0.6

0.8

1.0

Inorganic Content (vol%)

Figure 7: E’ as a function of nanofiller concentration (measured as vol% of inorganic platelets) for optimized neat and blended formulations

As shown in Figure 7, DMA storage modulus (E’) at 25°C exhibited an increase of ~10-15% at low nanofiller concentrations in all cases, with a reduction observed at the highest nanofiller content in the neat ELO/MTHPA system in particular. The highest storage modulus, obtained at 0.25 vol% inorganic, is expected to be largely free of agglomerates as it is below the estimated percolation threshold of ~0.42-0.64 vol% inorganic51. Likewise, the moduli observed at higher nanofiller concentrations are entirely in keeping with the expectation of reductions in dispersion quality as the system passes through the estimated percolation threshold. These observations are in keeping with the trends observed in our previous work involving amine-cured ELO nanocomposites34-36, albeit with less significant modulus enhancements and a downturn in


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modulus at lower nanofiller concentrations in this case implying a lower level of compatibility. Small variations in Tα (± 5°C) were observed with no clear trend vs. composition (data not shown), while DTUL was largely unaffected by nanoclay addition. While nanocomposite formation is not the primary focus of this work, these observations are significant in that they clearly show that hardener type strongly impacts the potential for further enhancements in structural performance due to the addition of nanofiller in these systems.

CONCLUSIONS

Blending of liquid anhydride hardeners with phthalic anhydride in anhydride cured epoxidized linseed oil based epoxy systems resulted in changes in hardness, modulus and alpha transition temperature. These effects were most significant in the case of systems cured with nadic methyl anhydride, which experienced increases in all three parameters as a consequence of PA addition. In spite of the fact that systems cured with methyltetrahydrophthalic anhydride experienced little or no gains on PA addition, their properties consistently matched or exceeded those of the analogous nadic methyl anhydride based systems. No significant changes in DTUL were observed in either case. These efforts indicate that methyltetrahydrophthalic anhydride is clearly the higher performing hardener in combination with epoxidized linseed oil, in spite of the fact that it requires stepwise curing in order to enable the formation of uniform, void-free samples. These efforts further demonstrate that the addition of phthalic anhydride to such systems produces only minor improvements in properties, and may in some cases interfere with double-bond mediated secondary crosslinking reactions capable of further enhancing properties through simple dilution of the double bonds.


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In tandem with these efforts, it is shown that the model used in this research to correlate hardness with Young’s modulus is useful for rapid screening and selection of formulations for maximum modulus, and that such correlations can produce accurate results depending on the situation. Limitations of this approach highlighted here include the need to define Poisson’s ratio accurately, the limited precision of the Shore D hardness test in this modulus range, and the fact that indentation hardness testing is by definition much more local (and sensitive to properties variations at the sample surface) vs. modulus measurements. Nevertheless, so long as such limitations are kept in mind, this approach remains a useful tool for the rapid evaluation of new materials. Finally, the addition of nanofiller to the highest performing neat and blend systems resulted in a ~10-15% increase in room temperature storage modulus at lower nanofiller contents, with a possible downturn in modulus at the highest nanofiller content. While this observation may be explained by the possibility that nanofiller dispersion suffers as the percolation threshold is approached, more work is needed to confirm this supposition. Regardless, these results confirm the potential of organically modified hydrotalcite clays as nanofillers compatible with both the chemistry and the curing conditions of epoxidized linseed oil based bioepoxies, in contrast with montmorillonite nanofillers, whose modifiers were posited to undergo significant thermal degradation under similar conditions, while emphasizing the critical relationship between hardener type and the potential for nanofiller-based reinforcement36.

ASSOCIATED CONTENT Supporting Information


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An overlay of DMA data from formulations containing ELO cured with MTHPA, MTHPA with 30 wt% PA, and MTHPA with 50 wt% PA; an overlay of DMA data from formulations containing ELO cured with NMA, NMA with 30 wt% PA, and NMA with 50 wt% PA.

ACKNOWLEDGEMENTS This material is based upon work supported by the National Science Foundation under Grant No. 1230884. The authors would like to extend the greatest of thanks to Dr, Akshay Kokil, Mr. Stephen Driscoll, Mr. Johannes Möller, Mr. Wenhao Liu, Mr. David Rondeau, and Mr. Patrick Casey.


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