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epoxy woven composite under combined hygrothermal conditions
International Journal of Fatigue 22 (2000) 809–820 www.elsevier.com/locate/ijfatigue
Durability of a graphite/epoxy woven composite under combined hygrothermal conditions Sneha R. Patel *, Scott W. Case Materials Response Group, Department of Engineering Science and Mechanics, Virginia Polytechnic Institute and State University, Blacksburg, VA 24061, USA Received 1 November 1999; received in revised form 14 February 2000; accepted 1 May 2000
Abstract The objective of this effort was to evaluate and model the effects of moisture, temperature and combined hygrothermal conditions on the strength and life of a graphite/epoxy woven composite material system. The material system specified is a candidate for use in an advanced subsonic aircraft engine, and the imposed environmental conditions were considered to be representative of engine service conditions. Fatigue and residual strength data showed that initial and residual tensile properties and fatigue life of the material were only minimally affected by any of the imposed environmental conditions for the fatigue stress levels considered in this study. Based on this data, it was shown a residual-strength-based life prediction approach could be used to model strength and life with reasonable results. Fatigue damage progression and accumulation were found to be dependent on testing environment, suggesting that more adverse effects of environment on strength and life may be manifested for other types of loading (i.e., offaxis loading). 2000 Elsevier Science Ltd. All rights reserved. Keywords: Durability; Environment; Polymer; Composite; Residual strength
1. Introduction For the past several decades, the aircraft industry has been one of the primary drivers of composite materials research. Composite systems offer an advantage over traditional aircraft metals because they tend to exhibit higher strength/weight and stiffness/weight ratios, thus making the aircraft lighter and improving performance. Woven composite material systems are particularly useful for such applications because they offer ease in manufacturing as well as increased impact resistance. The woven graphite/epoxy focused on in this study is currently being considered for use in the engine of an advanced subsonic aircraft. An idealization of the service conditions the engine in question is expected to undergo is given by the mission profile in Fig. 1. The mission profile depicts an aircraft that would be in flight approximately 90 min per day, during which time the engine
* Corresponding author. Tel.: +1-540-231-7493; fax: +1-540-2319187. E-mail address: [email protected] (S.R. Patel).
Fig. 1.
Mission profile for aircraft engine.
would be operating at 120°C. The remainder of the time, the aircraft would be in storage at 30°C and 85% relative humidity. Service loads were not specified for the study. Previous investigators have considered the effects of similarly combined moisture and thermal cycling on the properties of graphite/epoxy materials. For instance, Shyprykevich and Wolter [1] found that combined hygrothermal aging had little effect on the tensile strength
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of a graphite/epoxy composite material system. Degradation of compressive strength was found to occur, however, and correlate with moisture content. A more recent study by Obst et al. [2] focused on the effects of the temperature, moisture and aging conditions described in Fig. 1 on the microcracking behavior experienced by a woven graphite/epoxy composite during fatigue. The authors found that hygrothermally cycled material and material tested at elevated temperature had higher crack densities close to failure than unaged material tested at room temperature. Since both hygrothermal cycling and temperature were found to have an effect on the fatigue damage accumulation, it may be postulated that fatigue life is also affected. Smith and Weitsman [3] conducted a study on the effect of moisture on the fatigue behavior of a graphite/epoxy system and found that previously saturated specimens fatigued in water had lower fatigue lives than previously saturated specimens fatigued in air and unsaturated specimens fatigued in air. Case [4] studied the effects of elevated temperature on the strength and stiffness degradation of IM7/K3B notched and unnotched composites. A change in matrix properties as a result of elevated temperature (177°C) was evident from the decrease in transverse strength and stiffness properties from room-temperature properties. Moreover, in the unnotched specimens, a greater reduction in residual strength was observed as a function of fatigue cycles. Such studies provide a strong indication that environmental conditions play a role in the fatigue behavior of composite material systems. However, the literature contains little data concerning the effects of combining mechanical fatigue and changing environment on the fatigue life of a graphite/epoxy composite material. Thus, the goal of the present study was to evaluate the durability and damage tolerance of the specified material in fatigue loading under individual and combined service environmental conditions. In addition, since material life prediction tools that account for service conditions can play a valuable role in materials selection and design, the application of a residual-strength-based life prediction approach developed by Reifsnider and colleagues [5–7] was considered.
2. Experimental methods The material used in this study consisted of PR500 epoxy reinforced with woven (five-harness satin, [0/90]4s) AS-4 carbon fiber. The composite panels were manufactured by DOW-UT using a resin-transfer molding process. The thickness and fiber volume fraction of each panel were nominally 2.8 mm and 0.55, respectively. Panels were cut and ground into coupons 20.3 cm long by 2.54 cm wide. One of the main objectives of this effort was to use
a residual-strength-based approach for life prediction of the specified material, so the testing program was designed to obtain residual strength data. Since relating damage state to residual strength was also of interest, several intermediate steps were taken to track damage progression. Finally, as the change in remaining strength and damage state with different environmental conditions was of particular concern, the tests were conducted under each of four conditions: 1. room temperature (to provide baseline behavior); 2. elevated temperature (120°C, engine operating condition); 3. wet [saturated and then tested at 85% relative humidity (RH) at 30°C, storage condition]; and 4. hygrothermal cycling — alternation between temperature and humidity condition during fatigue. Specimens to be tested under the humidity condition were stored in a humidity chamber so that they could achieve saturation. For each different environmental condition, the following steps were taken to track damage progression and measure remaining strength: 1. quasi-static tests to determine initial tensile properties; 2. non-destructive evaluation (NDE) to determine prefatigue damage state of material (discussed in more detail below); 3. fatigue tests to establish S–N curves and estimate life; 4. fatigue cycling (not to failure) to induce damage state; 5. NDE to assess fatigue-induced damage state of material; and 6. quasi-static tests on fatigued specimens to determine residual properties. Several techniques of NDE were utilized to track damage state, each contributing different types of information. Edge replication was used primarily to study matrix cracking in the 90° tows and fiber/matrix debonding. Edge replicas were taken with the specimen loaded to 16 kN. This load was found to open up cracks without further damaging the specimen. Penetrant-enhanced Xradiography was used to observe delamination and transverse and longitudinal cracks. During fatigue testing, load/stroke data were recorded to monitor dynamic stiffness loss. The above listed procedures were performed on all test specimens. In some cases, scanning electron microscopy (SEM) was used to examine failure surfaces and acoustic emission was used during quasi-static loading to monitor damage events. All mechanical tests were performed in load control on an MTS servohydraulic load frame. Tension–tension fatigue testing was conducted with an R ratio of 0.1 and a frequency of 10 Hz. Since no service loads were speci-
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fied for the application, fatigue loads of 85% UTS, 75% UTS and 70% UTS, where UTS is the ultimate tensile strangth, were chosen as testing conditions. These loads were chosen mainly because they fit best with the time frame allotted for the study. For the “environmental” testing, the grips of the load frame were enclosed in an environmental chamber, so that the entire specimen was exposed to the prescribed condition. Specimens that were tested at elevated temperature were allowed 10 min to equilibrate once the chamber achieved the prescribed condition. During the “wet” fatigue testing, saturated specimens were kept from drying out by fatiguing them in the specified moisture condition and by storing them in the humidity chamber between tests. In the case where hygrothermal cycling was imposed on specimens in conjunction with fatigue loading, a modified version of the hygrothermal cycle shown in Fig. 1 was implemented. This modified version of the hygrothermal cycle was necessary because the time frame associated with the typical mission profile was such that the specimen would fail due to mechanical fatigue before even one full hygrothermal cycle was completed. The modified cycle was based partially on the mission profile and partially on the expected fatigue life of the material at room temperature, and it was imposed as follows: 1. fatigue specimen at 30°C and 85% RH for 144,000 cycles (4 h); 2. remove specimen from testing chamber and place in humidity chamber (unloaded) overnight, allowing oven and grips to heat to 120°C; 3. fatigue at 120°C for 54,000 cycles (90 min); 4. remove specimen from testing chamber and again place in humidity chamber overnight, allowing oven and grips to cool to 30°C; and 5. repeat process until specimen fails or desired number of cycles has been reached. Specimens had to be taken out of the testing chamber between conditions because the size of the grips made it difficult to modify conditions in a reasonably short period of time.
3. Results and discussion 3.1. Quasi-static results The results from the initial quasi-static testing are given in Table 1, and typical stress–strain curves are shown in Fig. 2. The number in parentheses under “Condition” indicates the number of specimens tested under each condition. Stiffness was measured by determining the slope of the stress–strain curve from 0.1% strain to
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Table 1 Initial tensile properties Condition
Average strength (MPa)
Average stiffness (GPa)
Average Poisson’s ratio
25°C (3) 120°C (6) Saturated (4)
745±29 695±66 693±69
65.3±0.4 63.8±7.0 65.2±1.9
0.049±0.005 0.038±0.017 0.064a
a
Values listed with no error represent single specimen results.
Fig. 2.
Initial stress–strain curves at specified conditions.
0.3% strain, as prescribed by ASTM D3039-95A for non-linear stress–strain curves. The Poisson’s ratio was determined in a similar fashion with the transverse strain versus axial strain curve. Considering the scatter in the data, the results indicate no conclusive effect of temperature, moisture or aging on the material tensile properties. One interesting feature of the stress–strain curves is that they tend to curve slightly upward. This result is in contrast to typical stress–strain curves for woven materials which contain a “knee”, similar to the one observed for cross-ply materials, where the stress–strain curve turns concave downward. The knee generally occurs where the off-axis tows fail [8]. In this particular system, no such knee appears. One explanation for this behavior is found by examining the crack density data given in Table 2 for a single specimen. The data show that significant cracking does not occur until relatively large load (stress) levels are achieved. This is further exhibited by the acoustic emission data (for the same specimen) shown in Fig. 3. The histogram is skewed to the right, again indicating that most of the cracking activity takes place only at the loads close to failure. Transverse cracking was the sole damage mechanism observed in quasi-statically loaded specimens. The fact that the upward trend does not appear to be monotonic can be explained by the fact that the crack densities at each load were not necessarily measured at the exact same location of the specimen (although the general area
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Table 2 Quasi-static crack density data Load (kN)
Crack density (cracks/cm/ply)
E (MPa)
15.58 22.25 26.70 31.15 35.60 40.05 44.50 48.95 53.40 57.85
2.7 2.7 5.3 4.9 2.7 6.5 9.8 11.3 7.8 13.4
62.1 62.1 62.1 62.1 62.1 62.1 62.1 63.5 62.8 60.7 Fig. 4. Maximum fatigue stress versus cycles to failure for four environmental conditions.
Fig. 3.
Acoustic emission data.
was the same). Thus, the anomalies in the crack density numbers may be due to the variability in crack density at different sections of each ply [9].
data are normalized by the average initial stiffness value of 65 GPa. It should be noted that data for moisturesaturated specimens are especially limited due to manufacturing-related abnormalities in many of the specimens that were saturated for testing. The data show that strength degrades very little, if at all, for low-cycle (high stress) fatigue. The change in stiffness is even less than that of the strength, perhaps an artifact of tow straightening. For high-cycle (low stress) fatigue, the maximum strength degradation of 20–30% occurred for specimens fatigued at 70% UTS, the lowest fatigue stress level considered. Considering the scatter in the data, elevated-temperature and wet residual strength data indicate no clear deviation from room-temperature data. 3.3. Fatigue damage mechanisms
3.2. Fatigue and residual strength testing results The fatigue life data for material tested under each of the four specified environmental conditions are shown in Fig. 4. The data show that only minimal differences in fatigue life occurred as a result of exposure to temperature or moisture or the combination of the two (hygrothermal cycling) during fatigue cycling. What little difference appeared manifested itself predominantly at the lowest stress level, corresponding to the largest number of fatigue cycles. For residual strength testing, differing damage states were induced in specimens by fatiguing them at 85%, 75% or 70% of ultimate tensile strength for either 50% or 75% of life (as estimated from the S–N data). The testing was conducted under each of the four environmental conditions listed previously. Results from the residual strength testing are shown in Table 3, where residual strength data are normalized by the average initial strength value of 745 MPa and residual stiffness
Radiography and edge replication were the two main methods used to study damage mechanisms during the fatigue process. For all specimens tested, the major initial damage mechanisms appeared in the form of transverse microcracks and delaminations at the fiber undulation areas (Fig. 5). Such delamination regions were also observed by Fujii et al. [9], who termed them “meta-delaminations”. This initial damage was followed with growth of transverse cracks, which extended across one transverse tow at maximum length, and the introduction of longitudinal microcracks, which did not grow significantly, around the fiber undulation areas (Fig. 6). Finally, fiber/matrix debonding and edge and interply delaminations (Fig. 7) typical of cross-ply laminates preceded catastrophic fiber fracture. This behavior is not unusual — it is well documented that transverse and longitudinal cracks intersecting with each other, fibers and adjacent plies lead to delamination. The delamination inhibits the ability of the composite to distribute
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Table 3 Residual strength and stiffness Condition
% UTS
Fatigue cycles
Number of specimens tested
Normalized strength
Normalized stiffness
25°C
85 85 74 74 70 70 85 85 74 74 70 70 85 85 75 75 70 70 70 75 75 70 70
1155 1733 85,157 127,736 396,185 594,278 1572 2358 62,700 94,050 242,000 363,000 1155 1733 85,157 127,736 369,185 594,278 2,000,000 144,000 198,000 242,000 363,000
2 2 2 2 2 3 2 3 3 3 3 2 1 1 1 1 1 2 1 3 3 3 2
0.98±0.02 1.00±0.03 0.90±0.14 0.95±0.07 0.81±0.04 0.78±0.07 0.99±0.02 0.97±0.04 0.91±0.04 0.86±0.11 0.84±0.06 0.84±0.05 0.90 0.88 0.79 0.89 0.88 0.83±0.10 0.90 1.00±0.04 0.86±0.10 0.91±0.07 0.84±0.00
0.99±0.05 1.01±0.12 0.95±0.04 0.93±0.08 1.06±0.13 0.81±0.17 1.01±0.00 0.97±0.07 0.92±0.20 0.94±0.06 0.91±0.07 0.88±0.03 0.99 0.91 0.79 0.90 0.84±0.02 0.85 0.97±0.05 0.83±0.04 0.92±0.07 -
120°C
30°C, 85% RH
Hygrothermala
a
Hygrothermal cycling in conjunction with fatigue as discussed in Experimental methods.
Fig. 5.
Radiograph of elevated-temperature specimen fatigued for 2358 cycles at 633 MPa (85% UTS).
load, so the fibers fail as the stress increases to a critical level [10,11]. One damage mechanism unique to the woven material is microcracking at the warp/weft tow cross-over points, evident in both quasi-static specimens (transverse cracking only) and fatigue specimens (transverse and longitudinal cracking). These microcracks indicate that the stresses are relatively high at the cross-over points even
for a material in which the weave occurs as infrequently as in a five-harness satin weave. In fatigue specimens, microcracks may cause the fibers crossing over and under the resin-rich area between them to debond fairly early in the fatigue process, resulting in the previously described “meta-delaminations”. The extent to which these damage events occurred, and the rate at which they occurred, did vary with testing
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Fig. 6. Radiograph of elevated-temperature specimen fatigued for 62,700 cycles at 552 MPa (75% UTS).
conditions and fatigue stress levels. For example, an edge replica and a radiograph of a typical specimen fatigued at elevated temperature at 85% UTS are shown in Fig. 8. In this case, only transverse cracking and meta-
Fig. 7.
Fig. 8.
delaminations are readily apparent. In addition, the fracture surface is clean and cohesive, similar to that found in a quasi-static failure, as shown in Fig. 9. On the other hand, for a typical specimen fatigued at elevated temperature at 70% UTS (Fig. 10), we see a far greater amount of fiber/matrix debonding and interply delamination. The fracture surface, shown in Fig. 11, is far less cohesive than was found for the specimens tested at high fatigue stress level. In many cases, the specimens did not completely separate into two parts. Instead, enough plies pulled apart within the composite to constitute failure, but a few intact plies still held together. This was true even in many of the residual strength tests for which the specimen strength was relatively low. The increased level of damage accumulated during the life of the composite with decreasing fatigue stress levels is manifested in the dynamic stiffness loss curves acquired during fatigue testing (Fig. 12). Note that the stiffness loss follows the classic “S shape” discussed in [10]. During the course of fatigue testing, dynamic stiffness was monitored by way of the extensometer and stroke data. In many cases, the extensometer data drifted substantially, possibly as a result of the damage pro-
Radiograph of elevated-temperature specimen fatigued for 242,000 cycles at 522 MPa (70% UTS).
Edge replica (top) and radiograph (bottom) for a specimen fatigued at elevated temperature at 633 MPa (85% UTS) for 2358 cycles.
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Fig. 9. Failure fracture surface of a room-temperature quasi-static specimen (left) and an elevated-temperature specimen cycled at 85% UTS (right).
Fig. 10.
Edge replica (top) and radiograph (bottom) for a specimen fatigued at elevated temperature at 522 MPa (70% UTS) for 242,000 cycles.
gressing within the composite and particularly in the surface ply on which the extensometer was mounted. Thus, the stroke data were used to track dynamic stiffness loss. The results show that specimens fatigued at lower fatigue stress levels underwent greater stiffness loss. Such results are in line with residual strength and damage tolerance ideas. Considering the concept that fatigue failure of a specimen will occur when its remaining strength degrades to the value of the applied load, it follows that a specimen fatigued at a lower stress level must reach a higher damage state for the strength to degrade to the point of failure. The increased level of damage corresponds to a greater stiffness loss. Differences in the rate and extent of damage accumulation were also observed for the different environmental conditions. Fig. 13 compares room-temperature and elevated-temperature dynamic stiffness loss curves. The
fact that the stiffness loss is slightly greater for elevatedtemperature specimens than for room-temperature specimens indicates that elevated-temperature specimens suffered a greater extent of damage than did room-temperature specimens. Radiographs such as those shown in Fig. 14 verify that the elevated-temperature specimens delaminated more than room-temperature specimens. Edge replicas also indicate a greater amount of fiber/matrix debonding in the elevated-temperature specimens. Moreover, in comparing typical SEM images of a room-temperature quasi-static fracture surface and an elevated-temperature quasi-static fracture surface (Fig. 15), it is readily apparent that the fibers in the elevated-temperature specimen are debonded cleanly from the resin while the fibers in the room-temperature specimen are still well bonded to the resin.
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Fig. 11.
Failure fracture surface of an elevated-temperature specimen fatigued at 70% UTS.
Fig. 12. Dynamic stiffness loss curves (from stroke data) for specimens tested at 120°C for specified percentages of UTS.
Fig. 13. Comparison of room-temperature and elevated-temperature dynamic stiffness loss curves at specified percentages of UTS.
Saturated specimens also seemed to exhibit a greater amount of fiber/matrix debonding than did room-temperature specimens, suggesting that the fiber/matrix interface may have been degraded by the moisture. An SEM image of a typical quasi-static fracture surface from a saturated specimen is shown in Fig. 16. The fibers at the fracture surface of the moisture-saturated specimen are indeed barer than those of an unsaturated specimen (see Fig. 15). Dynamic stiffness loss curves for the saturated specimens were generally inconsistent and therefore offered no indication as to the rate of damage accumulation. Stiffness loss curves of specimens tested in conjunction with hygrothermal cycling were difficult to evaluate.
Interruptions in fatigue cycling incurred while modifying conditions in the environmental chamber resulted in discontinuities in the stiffness loss curves. However, edge replicas and radiographs showed no significant change in delamination or fiber/matrix debonding processes from that found in room-temperature fatigue specimens. Transverse cracking was also monitored as a function of fatigue cycles. From the data given in Table 4, it can be observed that crack density does increase as a function of cycles (neglecting fatigue stress level) for much of the material lifetime before leveling off. While the crack density of the moisture-saturated specimens appears not to grow with cycles, the limited nature of the data (as discussed above under Fatigue and residual
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Fig. 14.
Fig. 15.
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Radiographs of specimens fatigued at room temperature and elevated temperature.
SEM micrographs of quasi-static fracture surfaces for room-temperature (left) and elevated-temperature (right) specimens.
Fig. 16. SEM micrograph of a quasi-static fracture surface of a moisture-saturated specimen.
strength testing results) makes such a conclusion difficult to reach. It should be noted that each data point represents the average result of different specimens, thus accounting for the lack of monotonic behavior. The fatigue crack saturation density is generally higher than that found during quasi-static loading. The difference, also observed by Obst et al. [2], may be explained in several ways. One point that should be made is that the cracks in the fatigue specimens opened up more than the cracks in the quasi-static specimens and were easier to locate. In fatigued specimens, the delaminations may have made it easier for the cracks to open. We may also speculate that the crack saturation density is greater for fatigue specimens than for quasistatic specimens because cracks have more opportunity to grow, particularly as the individual plies start to carry increasing loads. The data also show that the material fatigued in conjunction with hygrothermal cycling maintains a higher crack density than material tested under any of the other
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Table 4 Fatigue crack density data Condition 25°C
120°C
30°C, 85% RH
Hygrothermal
a
% UTS 85 85 75 75 70 70 85 85 75 75 70 70 85 85 75 75 70 70 70 70 70 70 70 70 70 75
Cycles 1155 1733 85,157 127,736 396,185 594,278 1572 2358 62,700 94,050 242,000 363,000 1155 1733 85,157 127,736 369,185 594,278 54,000 198,000 252,000 450,000 648,000 846,000 1,000,000 198,000
Crack density (cracks/ply/cm)a 11.83 11.48 13.41±1.90 13.14±6.11 17.08±2.79 15.87±2.75 9.01±4.64 11.20±3.95 15.73±2.50 15.35±2.42 17.30±1.45 18.33±2.36 17.34 13.66 13.77 11.48 15.13 12.12 6.79±2.67 19.88 22.95 20.98 20.04 20.22 20.95 16.21±2.70
Values listed with no error represent single specimen results.
three conditions. However, the room-temperature, elevated-temperature and moisture data appear to be close together. Thus, one observation that can be made is that the alternation of hot/wet conditions does affect the fatigue behavior of the material in a way not realized by either of the conditions alone. Another observation that can be made is that the difference in the dynamic stiffness loss between specimens tested at room temperature and elevated temperature is mostly due to delamination. This is not surprising since very little stiffness loss was observed to occur with cracking in the quasi-static testing. Moreover, O’Brien [12] linked stiffness reduction with delamination development (in 1982).
4. Life prediction The life prediction methodology used in this study is one originally developed by Reifsnider and Stinchcomb [7] and further developed by Reifsnider and co-workers. Details of the method can be found in [5–7]. According to the present modeling approach, failure of the laminate is assumed to occur when its residual strength degrades to the value of the maximum applied load. In addition, the residual strength is assumed to be a damage metric such that equivalent damage states are assumed to be
represented by equivalent residual strengths. Based on kinetics arguments, Reifsnider [5] developed a damage evolution integral,
冕 t2
⌬Fr⫽⫺ (1⫺Fa)jtj−1 dt,
(1)
t1
by which the change in normalized residual strength (⌬Fr) of a material may be calculated by implementing an appropriate failure criterion, Fa. In the case of mechanical fatigue, the characteristic time, t, can be expressed as n/N where n is the number of fatigue cycles the material has undergone out of a total of N (usually fatigue life) cycles. For t1 equal to zero, initial normalized residual strength, Fr1, equal to one, and constant Fa (constant maximum amplitude fatigue) Fr⫽1⫺(1⫺Fa)tj ,
(2)
where j is an experimentally determined shape parameter. The value of j may depend on such factors as failure mode, environmental conditions, laminate layup, etc. Because the study was meant to examine the behavior of the material under alternating service conditions, it was thought that a j value (and corresponding residual strength curve) would have to be found for each individual environment. One could then use this information to predict the life of the material with changing conditions by adjusting the j value in Eq. (2) as needed. However, since the experimental data reflected no significant and quantifiable effects of environment on the residual strength of the material, a single j value, obtained by fitting room-temperature residual strength data to Eq. (2), was found to be sufficient. In fitting the data, the maximum stress criterion was used for Fa, Fa⫽
sapplied max , Xt
(3)
and t was taken to be n/N, where average values of room-temperature fatigue life data for each stress level were substituted for N and the number of applied fatigue cycles as given in Table 3 were substituted for n. In this way, the value of j was determined to be 1.3 for the material considered in this study. As can be seen from the residual strength data and predicted residual strength curves shown in Fig. 17, reasonable agreement was found between experimental and predicted residual strength values for all of the environments considered. However, elevated-temperature results match the calculated residual strength curves more closely if the curves are calculated using average elevated-temperature lifetimes (N). The j value, however, remains 1.3. This result in shown in Fig. 18.
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Fig. 17. Residual strength data and predictions based on room-temperature data.
Fig. 18.
Residual strength data and predictions for elevated temperature (120°C).
5. Summary and conclusions 1. Fatigue life and residual strength of the woven graphite/epoxy composite material studied are, at best, minimally affected by the elevated temperature, moisture or combined hygrothermal conditions imposed and stress levels considered, despite differences in damage accumulation processes. 2. The major damage mechanism resulting from quasistatic loading is cracking in transverse tows. The cracks extend across a single transverse fiber bundle at maximum length. In addition, the majority of the cracks appear relatively close to failure but do not reduce the axial elastic modulus. 3. Damage progression during fatigue occurred in the following general order: transverse microcracking and
delaminations around the fiber bundle undulation regions, longitudinal microcracking around fiber bundle undulation regions, growth of transverse cracks across the entire height of the transverse fiber bundle, and cross-ply-like growth of edge and interply delamination. The rate at which damage progression occurred generally depended on the maximum fatigue stress amplitude and environment in that: 3.1. damage accumulation was found to be greater with decreasing fatigue stress levels; 3.2. fiber/matrix debonding and delamination were found to be greater for the elevated-temperature and moisture-saturated specimens than for roomtemperature specimens; and 3.3. hygrothermal cycling during fatigue was found to result in greater crack density during fatigue
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than did any of the other three conditions.These results suggest that the fatigue life of this material may be more adversely affected by environment for off-axis loading situations than was captured by the fiber-direction loading case considered in this study. 4. A residual-strength-based life prediction model was used in which an experimentally determined model parameter was obtained from room-temperature data and applied for all conditions with reasonable results.
[3] [4]
[5]
[6]
Acknowledgements The authors wish to acknowledge the financial support of Pratt and Whitney, the National Science Foundation (under grant #CMS-9872331), and the Virginia Space Grant Consortium. In addition, the authors would like to thank NASA Glenn Research Center (particularly Mike Castelli) for performing hygrothermal aging on composites.
[7]
[8] [9]
[10]
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Griesheim GE, Tosi KF. The effect of hygrothermal cycling on the microcracking behavior of fabric laminates. In: Technology transfer in a global community: Proceedings of International SAMPE Technical Conference vol. covena (CA): SAMPE 1996;28:994–1002. Smith LV, Weitsman YJ. The immersed fatigue of response of polymer composites. Int J Fract 1996;82:31–42. Case SW. Mechanics of fiber-controlled behavior in polymeric composite materials. Ph.D. dissertation. Blacksburg (VA): Virginia Polytechnic Institute and State University, 1996. Reifsnider KL. Use of mechanistic life prediction methods for the design of damage-tolerant composite material systems. In: Mitchell MR, Landgraf RW, editors. Advances in fatigue lifetime predictive techniques, vol. 2. Philadelphia (PA): American Society for Testing and Materials, 1993:3–18. Case SW, Reifsnider KL. MRLife 11: a strength and life prediction code for laminated composite materials. Blacksburg (VA): Materials Response Group, Virginia Polytechnic Institute and State University, 1998. Reifsnider KL, Stinchcomb WW. A critical-element model of the residual strength and life of fatigue-loaded composite coupons. In: Hahn HT, editor. Composite materials: fatigue and fracture. Philadelphia (PA): American Society for Testing and Materials, 1986:298–313. Ishikawa T, Chou TW. Stiffness and strength behavior of woven fabric composites. J Mater Sci 1982;17:3211–20. Fujii T, Amijima S, Okubo K. Microscopic fatigue processes in a plain-weave glass-fibre composite. Compos Sci Technol 1993;49(4):327–33. Reifsnider KL. Damage and damage mechanics. In: Reifsnider KL, editor. Fatigue of composites, vol. 4. New York: Elsevier Science Publishers, 1990:11–77. Highsmith AL. The effect of delamination on the behavior of crossply laminates. Mech Compos Mater 1988;AMD-92:123–33. O’Brien TK. Characterization of delamination onset and growth in a composite laminate. In: Reifsnider KL, editor. Damage in composite materials. Philadelphia (PA): American Society for Testing and Materials, 1982:140–67.
Durability of a graphite/epoxy woven composite under combined hygrothermal conditions Sneha R. Patel *, Scott W. Case Materials Response Group, Department of Engineering Science and Mechanics, Virginia Polytechnic Institute and State University, Blacksburg, VA 24061, USA Received 1 November 1999; received in revised form 14 February 2000; accepted 1 May 2000
Abstract The objective of this effort was to evaluate and model the effects of moisture, temperature and combined hygrothermal conditions on the strength and life of a graphite/epoxy woven composite material system. The material system specified is a candidate for use in an advanced subsonic aircraft engine, and the imposed environmental conditions were considered to be representative of engine service conditions. Fatigue and residual strength data showed that initial and residual tensile properties and fatigue life of the material were only minimally affected by any of the imposed environmental conditions for the fatigue stress levels considered in this study. Based on this data, it was shown a residual-strength-based life prediction approach could be used to model strength and life with reasonable results. Fatigue damage progression and accumulation were found to be dependent on testing environment, suggesting that more adverse effects of environment on strength and life may be manifested for other types of loading (i.e., offaxis loading). 2000 Elsevier Science Ltd. All rights reserved. Keywords: Durability; Environment; Polymer; Composite; Residual strength
1. Introduction For the past several decades, the aircraft industry has been one of the primary drivers of composite materials research. Composite systems offer an advantage over traditional aircraft metals because they tend to exhibit higher strength/weight and stiffness/weight ratios, thus making the aircraft lighter and improving performance. Woven composite material systems are particularly useful for such applications because they offer ease in manufacturing as well as increased impact resistance. The woven graphite/epoxy focused on in this study is currently being considered for use in the engine of an advanced subsonic aircraft. An idealization of the service conditions the engine in question is expected to undergo is given by the mission profile in Fig. 1. The mission profile depicts an aircraft that would be in flight approximately 90 min per day, during which time the engine
* Corresponding author. Tel.: +1-540-231-7493; fax: +1-540-2319187. E-mail address: [email protected] (S.R. Patel).
Fig. 1.
Mission profile for aircraft engine.
would be operating at 120°C. The remainder of the time, the aircraft would be in storage at 30°C and 85% relative humidity. Service loads were not specified for the study. Previous investigators have considered the effects of similarly combined moisture and thermal cycling on the properties of graphite/epoxy materials. For instance, Shyprykevich and Wolter [1] found that combined hygrothermal aging had little effect on the tensile strength
0142-1123/00/$ - see front matter 2000 Elsevier Science Ltd. All rights reserved. PII: S 0 1 4 2 - 1 1 2 3 ( 0 0 ) 0 0 0 4 1 - 4
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of a graphite/epoxy composite material system. Degradation of compressive strength was found to occur, however, and correlate with moisture content. A more recent study by Obst et al. [2] focused on the effects of the temperature, moisture and aging conditions described in Fig. 1 on the microcracking behavior experienced by a woven graphite/epoxy composite during fatigue. The authors found that hygrothermally cycled material and material tested at elevated temperature had higher crack densities close to failure than unaged material tested at room temperature. Since both hygrothermal cycling and temperature were found to have an effect on the fatigue damage accumulation, it may be postulated that fatigue life is also affected. Smith and Weitsman [3] conducted a study on the effect of moisture on the fatigue behavior of a graphite/epoxy system and found that previously saturated specimens fatigued in water had lower fatigue lives than previously saturated specimens fatigued in air and unsaturated specimens fatigued in air. Case [4] studied the effects of elevated temperature on the strength and stiffness degradation of IM7/K3B notched and unnotched composites. A change in matrix properties as a result of elevated temperature (177°C) was evident from the decrease in transverse strength and stiffness properties from room-temperature properties. Moreover, in the unnotched specimens, a greater reduction in residual strength was observed as a function of fatigue cycles. Such studies provide a strong indication that environmental conditions play a role in the fatigue behavior of composite material systems. However, the literature contains little data concerning the effects of combining mechanical fatigue and changing environment on the fatigue life of a graphite/epoxy composite material. Thus, the goal of the present study was to evaluate the durability and damage tolerance of the specified material in fatigue loading under individual and combined service environmental conditions. In addition, since material life prediction tools that account for service conditions can play a valuable role in materials selection and design, the application of a residual-strength-based life prediction approach developed by Reifsnider and colleagues [5–7] was considered.
2. Experimental methods The material used in this study consisted of PR500 epoxy reinforced with woven (five-harness satin, [0/90]4s) AS-4 carbon fiber. The composite panels were manufactured by DOW-UT using a resin-transfer molding process. The thickness and fiber volume fraction of each panel were nominally 2.8 mm and 0.55, respectively. Panels were cut and ground into coupons 20.3 cm long by 2.54 cm wide. One of the main objectives of this effort was to use
a residual-strength-based approach for life prediction of the specified material, so the testing program was designed to obtain residual strength data. Since relating damage state to residual strength was also of interest, several intermediate steps were taken to track damage progression. Finally, as the change in remaining strength and damage state with different environmental conditions was of particular concern, the tests were conducted under each of four conditions: 1. room temperature (to provide baseline behavior); 2. elevated temperature (120°C, engine operating condition); 3. wet [saturated and then tested at 85% relative humidity (RH) at 30°C, storage condition]; and 4. hygrothermal cycling — alternation between temperature and humidity condition during fatigue. Specimens to be tested under the humidity condition were stored in a humidity chamber so that they could achieve saturation. For each different environmental condition, the following steps were taken to track damage progression and measure remaining strength: 1. quasi-static tests to determine initial tensile properties; 2. non-destructive evaluation (NDE) to determine prefatigue damage state of material (discussed in more detail below); 3. fatigue tests to establish S–N curves and estimate life; 4. fatigue cycling (not to failure) to induce damage state; 5. NDE to assess fatigue-induced damage state of material; and 6. quasi-static tests on fatigued specimens to determine residual properties. Several techniques of NDE were utilized to track damage state, each contributing different types of information. Edge replication was used primarily to study matrix cracking in the 90° tows and fiber/matrix debonding. Edge replicas were taken with the specimen loaded to 16 kN. This load was found to open up cracks without further damaging the specimen. Penetrant-enhanced Xradiography was used to observe delamination and transverse and longitudinal cracks. During fatigue testing, load/stroke data were recorded to monitor dynamic stiffness loss. The above listed procedures were performed on all test specimens. In some cases, scanning electron microscopy (SEM) was used to examine failure surfaces and acoustic emission was used during quasi-static loading to monitor damage events. All mechanical tests were performed in load control on an MTS servohydraulic load frame. Tension–tension fatigue testing was conducted with an R ratio of 0.1 and a frequency of 10 Hz. Since no service loads were speci-
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fied for the application, fatigue loads of 85% UTS, 75% UTS and 70% UTS, where UTS is the ultimate tensile strangth, were chosen as testing conditions. These loads were chosen mainly because they fit best with the time frame allotted for the study. For the “environmental” testing, the grips of the load frame were enclosed in an environmental chamber, so that the entire specimen was exposed to the prescribed condition. Specimens that were tested at elevated temperature were allowed 10 min to equilibrate once the chamber achieved the prescribed condition. During the “wet” fatigue testing, saturated specimens were kept from drying out by fatiguing them in the specified moisture condition and by storing them in the humidity chamber between tests. In the case where hygrothermal cycling was imposed on specimens in conjunction with fatigue loading, a modified version of the hygrothermal cycle shown in Fig. 1 was implemented. This modified version of the hygrothermal cycle was necessary because the time frame associated with the typical mission profile was such that the specimen would fail due to mechanical fatigue before even one full hygrothermal cycle was completed. The modified cycle was based partially on the mission profile and partially on the expected fatigue life of the material at room temperature, and it was imposed as follows: 1. fatigue specimen at 30°C and 85% RH for 144,000 cycles (4 h); 2. remove specimen from testing chamber and place in humidity chamber (unloaded) overnight, allowing oven and grips to heat to 120°C; 3. fatigue at 120°C for 54,000 cycles (90 min); 4. remove specimen from testing chamber and again place in humidity chamber overnight, allowing oven and grips to cool to 30°C; and 5. repeat process until specimen fails or desired number of cycles has been reached. Specimens had to be taken out of the testing chamber between conditions because the size of the grips made it difficult to modify conditions in a reasonably short period of time.
3. Results and discussion 3.1. Quasi-static results The results from the initial quasi-static testing are given in Table 1, and typical stress–strain curves are shown in Fig. 2. The number in parentheses under “Condition” indicates the number of specimens tested under each condition. Stiffness was measured by determining the slope of the stress–strain curve from 0.1% strain to
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Table 1 Initial tensile properties Condition
Average strength (MPa)
Average stiffness (GPa)
Average Poisson’s ratio
25°C (3) 120°C (6) Saturated (4)
745±29 695±66 693±69
65.3±0.4 63.8±7.0 65.2±1.9
0.049±0.005 0.038±0.017 0.064a
a
Values listed with no error represent single specimen results.
Fig. 2.
Initial stress–strain curves at specified conditions.
0.3% strain, as prescribed by ASTM D3039-95A for non-linear stress–strain curves. The Poisson’s ratio was determined in a similar fashion with the transverse strain versus axial strain curve. Considering the scatter in the data, the results indicate no conclusive effect of temperature, moisture or aging on the material tensile properties. One interesting feature of the stress–strain curves is that they tend to curve slightly upward. This result is in contrast to typical stress–strain curves for woven materials which contain a “knee”, similar to the one observed for cross-ply materials, where the stress–strain curve turns concave downward. The knee generally occurs where the off-axis tows fail [8]. In this particular system, no such knee appears. One explanation for this behavior is found by examining the crack density data given in Table 2 for a single specimen. The data show that significant cracking does not occur until relatively large load (stress) levels are achieved. This is further exhibited by the acoustic emission data (for the same specimen) shown in Fig. 3. The histogram is skewed to the right, again indicating that most of the cracking activity takes place only at the loads close to failure. Transverse cracking was the sole damage mechanism observed in quasi-statically loaded specimens. The fact that the upward trend does not appear to be monotonic can be explained by the fact that the crack densities at each load were not necessarily measured at the exact same location of the specimen (although the general area
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Table 2 Quasi-static crack density data Load (kN)
Crack density (cracks/cm/ply)
E (MPa)
15.58 22.25 26.70 31.15 35.60 40.05 44.50 48.95 53.40 57.85
2.7 2.7 5.3 4.9 2.7 6.5 9.8 11.3 7.8 13.4
62.1 62.1 62.1 62.1 62.1 62.1 62.1 63.5 62.8 60.7 Fig. 4. Maximum fatigue stress versus cycles to failure for four environmental conditions.
Fig. 3.
Acoustic emission data.
was the same). Thus, the anomalies in the crack density numbers may be due to the variability in crack density at different sections of each ply [9].
data are normalized by the average initial stiffness value of 65 GPa. It should be noted that data for moisturesaturated specimens are especially limited due to manufacturing-related abnormalities in many of the specimens that were saturated for testing. The data show that strength degrades very little, if at all, for low-cycle (high stress) fatigue. The change in stiffness is even less than that of the strength, perhaps an artifact of tow straightening. For high-cycle (low stress) fatigue, the maximum strength degradation of 20–30% occurred for specimens fatigued at 70% UTS, the lowest fatigue stress level considered. Considering the scatter in the data, elevated-temperature and wet residual strength data indicate no clear deviation from room-temperature data. 3.3. Fatigue damage mechanisms
3.2. Fatigue and residual strength testing results The fatigue life data for material tested under each of the four specified environmental conditions are shown in Fig. 4. The data show that only minimal differences in fatigue life occurred as a result of exposure to temperature or moisture or the combination of the two (hygrothermal cycling) during fatigue cycling. What little difference appeared manifested itself predominantly at the lowest stress level, corresponding to the largest number of fatigue cycles. For residual strength testing, differing damage states were induced in specimens by fatiguing them at 85%, 75% or 70% of ultimate tensile strength for either 50% or 75% of life (as estimated from the S–N data). The testing was conducted under each of the four environmental conditions listed previously. Results from the residual strength testing are shown in Table 3, where residual strength data are normalized by the average initial strength value of 745 MPa and residual stiffness
Radiography and edge replication were the two main methods used to study damage mechanisms during the fatigue process. For all specimens tested, the major initial damage mechanisms appeared in the form of transverse microcracks and delaminations at the fiber undulation areas (Fig. 5). Such delamination regions were also observed by Fujii et al. [9], who termed them “meta-delaminations”. This initial damage was followed with growth of transverse cracks, which extended across one transverse tow at maximum length, and the introduction of longitudinal microcracks, which did not grow significantly, around the fiber undulation areas (Fig. 6). Finally, fiber/matrix debonding and edge and interply delaminations (Fig. 7) typical of cross-ply laminates preceded catastrophic fiber fracture. This behavior is not unusual — it is well documented that transverse and longitudinal cracks intersecting with each other, fibers and adjacent plies lead to delamination. The delamination inhibits the ability of the composite to distribute
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Table 3 Residual strength and stiffness Condition
% UTS
Fatigue cycles
Number of specimens tested
Normalized strength
Normalized stiffness
25°C
85 85 74 74 70 70 85 85 74 74 70 70 85 85 75 75 70 70 70 75 75 70 70
1155 1733 85,157 127,736 396,185 594,278 1572 2358 62,700 94,050 242,000 363,000 1155 1733 85,157 127,736 369,185 594,278 2,000,000 144,000 198,000 242,000 363,000
2 2 2 2 2 3 2 3 3 3 3 2 1 1 1 1 1 2 1 3 3 3 2
0.98±0.02 1.00±0.03 0.90±0.14 0.95±0.07 0.81±0.04 0.78±0.07 0.99±0.02 0.97±0.04 0.91±0.04 0.86±0.11 0.84±0.06 0.84±0.05 0.90 0.88 0.79 0.89 0.88 0.83±0.10 0.90 1.00±0.04 0.86±0.10 0.91±0.07 0.84±0.00
0.99±0.05 1.01±0.12 0.95±0.04 0.93±0.08 1.06±0.13 0.81±0.17 1.01±0.00 0.97±0.07 0.92±0.20 0.94±0.06 0.91±0.07 0.88±0.03 0.99 0.91 0.79 0.90 0.84±0.02 0.85 0.97±0.05 0.83±0.04 0.92±0.07 -
120°C
30°C, 85% RH
Hygrothermala
a
Hygrothermal cycling in conjunction with fatigue as discussed in Experimental methods.
Fig. 5.
Radiograph of elevated-temperature specimen fatigued for 2358 cycles at 633 MPa (85% UTS).
load, so the fibers fail as the stress increases to a critical level [10,11]. One damage mechanism unique to the woven material is microcracking at the warp/weft tow cross-over points, evident in both quasi-static specimens (transverse cracking only) and fatigue specimens (transverse and longitudinal cracking). These microcracks indicate that the stresses are relatively high at the cross-over points even
for a material in which the weave occurs as infrequently as in a five-harness satin weave. In fatigue specimens, microcracks may cause the fibers crossing over and under the resin-rich area between them to debond fairly early in the fatigue process, resulting in the previously described “meta-delaminations”. The extent to which these damage events occurred, and the rate at which they occurred, did vary with testing
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Fig. 6. Radiograph of elevated-temperature specimen fatigued for 62,700 cycles at 552 MPa (75% UTS).
conditions and fatigue stress levels. For example, an edge replica and a radiograph of a typical specimen fatigued at elevated temperature at 85% UTS are shown in Fig. 8. In this case, only transverse cracking and meta-
Fig. 7.
Fig. 8.
delaminations are readily apparent. In addition, the fracture surface is clean and cohesive, similar to that found in a quasi-static failure, as shown in Fig. 9. On the other hand, for a typical specimen fatigued at elevated temperature at 70% UTS (Fig. 10), we see a far greater amount of fiber/matrix debonding and interply delamination. The fracture surface, shown in Fig. 11, is far less cohesive than was found for the specimens tested at high fatigue stress level. In many cases, the specimens did not completely separate into two parts. Instead, enough plies pulled apart within the composite to constitute failure, but a few intact plies still held together. This was true even in many of the residual strength tests for which the specimen strength was relatively low. The increased level of damage accumulated during the life of the composite with decreasing fatigue stress levels is manifested in the dynamic stiffness loss curves acquired during fatigue testing (Fig. 12). Note that the stiffness loss follows the classic “S shape” discussed in [10]. During the course of fatigue testing, dynamic stiffness was monitored by way of the extensometer and stroke data. In many cases, the extensometer data drifted substantially, possibly as a result of the damage pro-
Radiograph of elevated-temperature specimen fatigued for 242,000 cycles at 522 MPa (70% UTS).
Edge replica (top) and radiograph (bottom) for a specimen fatigued at elevated temperature at 633 MPa (85% UTS) for 2358 cycles.
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Fig. 9. Failure fracture surface of a room-temperature quasi-static specimen (left) and an elevated-temperature specimen cycled at 85% UTS (right).
Fig. 10.
Edge replica (top) and radiograph (bottom) for a specimen fatigued at elevated temperature at 522 MPa (70% UTS) for 242,000 cycles.
gressing within the composite and particularly in the surface ply on which the extensometer was mounted. Thus, the stroke data were used to track dynamic stiffness loss. The results show that specimens fatigued at lower fatigue stress levels underwent greater stiffness loss. Such results are in line with residual strength and damage tolerance ideas. Considering the concept that fatigue failure of a specimen will occur when its remaining strength degrades to the value of the applied load, it follows that a specimen fatigued at a lower stress level must reach a higher damage state for the strength to degrade to the point of failure. The increased level of damage corresponds to a greater stiffness loss. Differences in the rate and extent of damage accumulation were also observed for the different environmental conditions. Fig. 13 compares room-temperature and elevated-temperature dynamic stiffness loss curves. The
fact that the stiffness loss is slightly greater for elevatedtemperature specimens than for room-temperature specimens indicates that elevated-temperature specimens suffered a greater extent of damage than did room-temperature specimens. Radiographs such as those shown in Fig. 14 verify that the elevated-temperature specimens delaminated more than room-temperature specimens. Edge replicas also indicate a greater amount of fiber/matrix debonding in the elevated-temperature specimens. Moreover, in comparing typical SEM images of a room-temperature quasi-static fracture surface and an elevated-temperature quasi-static fracture surface (Fig. 15), it is readily apparent that the fibers in the elevated-temperature specimen are debonded cleanly from the resin while the fibers in the room-temperature specimen are still well bonded to the resin.
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Fig. 11.
Failure fracture surface of an elevated-temperature specimen fatigued at 70% UTS.
Fig. 12. Dynamic stiffness loss curves (from stroke data) for specimens tested at 120°C for specified percentages of UTS.
Fig. 13. Comparison of room-temperature and elevated-temperature dynamic stiffness loss curves at specified percentages of UTS.
Saturated specimens also seemed to exhibit a greater amount of fiber/matrix debonding than did room-temperature specimens, suggesting that the fiber/matrix interface may have been degraded by the moisture. An SEM image of a typical quasi-static fracture surface from a saturated specimen is shown in Fig. 16. The fibers at the fracture surface of the moisture-saturated specimen are indeed barer than those of an unsaturated specimen (see Fig. 15). Dynamic stiffness loss curves for the saturated specimens were generally inconsistent and therefore offered no indication as to the rate of damage accumulation. Stiffness loss curves of specimens tested in conjunction with hygrothermal cycling were difficult to evaluate.
Interruptions in fatigue cycling incurred while modifying conditions in the environmental chamber resulted in discontinuities in the stiffness loss curves. However, edge replicas and radiographs showed no significant change in delamination or fiber/matrix debonding processes from that found in room-temperature fatigue specimens. Transverse cracking was also monitored as a function of fatigue cycles. From the data given in Table 4, it can be observed that crack density does increase as a function of cycles (neglecting fatigue stress level) for much of the material lifetime before leveling off. While the crack density of the moisture-saturated specimens appears not to grow with cycles, the limited nature of the data (as discussed above under Fatigue and residual
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Fig. 14.
Fig. 15.
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Radiographs of specimens fatigued at room temperature and elevated temperature.
SEM micrographs of quasi-static fracture surfaces for room-temperature (left) and elevated-temperature (right) specimens.
Fig. 16. SEM micrograph of a quasi-static fracture surface of a moisture-saturated specimen.
strength testing results) makes such a conclusion difficult to reach. It should be noted that each data point represents the average result of different specimens, thus accounting for the lack of monotonic behavior. The fatigue crack saturation density is generally higher than that found during quasi-static loading. The difference, also observed by Obst et al. [2], may be explained in several ways. One point that should be made is that the cracks in the fatigue specimens opened up more than the cracks in the quasi-static specimens and were easier to locate. In fatigued specimens, the delaminations may have made it easier for the cracks to open. We may also speculate that the crack saturation density is greater for fatigue specimens than for quasistatic specimens because cracks have more opportunity to grow, particularly as the individual plies start to carry increasing loads. The data also show that the material fatigued in conjunction with hygrothermal cycling maintains a higher crack density than material tested under any of the other
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Table 4 Fatigue crack density data Condition 25°C
120°C
30°C, 85% RH
Hygrothermal
a
% UTS 85 85 75 75 70 70 85 85 75 75 70 70 85 85 75 75 70 70 70 70 70 70 70 70 70 75
Cycles 1155 1733 85,157 127,736 396,185 594,278 1572 2358 62,700 94,050 242,000 363,000 1155 1733 85,157 127,736 369,185 594,278 54,000 198,000 252,000 450,000 648,000 846,000 1,000,000 198,000
Crack density (cracks/ply/cm)a 11.83 11.48 13.41±1.90 13.14±6.11 17.08±2.79 15.87±2.75 9.01±4.64 11.20±3.95 15.73±2.50 15.35±2.42 17.30±1.45 18.33±2.36 17.34 13.66 13.77 11.48 15.13 12.12 6.79±2.67 19.88 22.95 20.98 20.04 20.22 20.95 16.21±2.70
Values listed with no error represent single specimen results.
three conditions. However, the room-temperature, elevated-temperature and moisture data appear to be close together. Thus, one observation that can be made is that the alternation of hot/wet conditions does affect the fatigue behavior of the material in a way not realized by either of the conditions alone. Another observation that can be made is that the difference in the dynamic stiffness loss between specimens tested at room temperature and elevated temperature is mostly due to delamination. This is not surprising since very little stiffness loss was observed to occur with cracking in the quasi-static testing. Moreover, O’Brien [12] linked stiffness reduction with delamination development (in 1982).
4. Life prediction The life prediction methodology used in this study is one originally developed by Reifsnider and Stinchcomb [7] and further developed by Reifsnider and co-workers. Details of the method can be found in [5–7]. According to the present modeling approach, failure of the laminate is assumed to occur when its residual strength degrades to the value of the maximum applied load. In addition, the residual strength is assumed to be a damage metric such that equivalent damage states are assumed to be
represented by equivalent residual strengths. Based on kinetics arguments, Reifsnider [5] developed a damage evolution integral,
冕 t2
⌬Fr⫽⫺ (1⫺Fa)jtj−1 dt,
(1)
t1
by which the change in normalized residual strength (⌬Fr) of a material may be calculated by implementing an appropriate failure criterion, Fa. In the case of mechanical fatigue, the characteristic time, t, can be expressed as n/N where n is the number of fatigue cycles the material has undergone out of a total of N (usually fatigue life) cycles. For t1 equal to zero, initial normalized residual strength, Fr1, equal to one, and constant Fa (constant maximum amplitude fatigue) Fr⫽1⫺(1⫺Fa)tj ,
(2)
where j is an experimentally determined shape parameter. The value of j may depend on such factors as failure mode, environmental conditions, laminate layup, etc. Because the study was meant to examine the behavior of the material under alternating service conditions, it was thought that a j value (and corresponding residual strength curve) would have to be found for each individual environment. One could then use this information to predict the life of the material with changing conditions by adjusting the j value in Eq. (2) as needed. However, since the experimental data reflected no significant and quantifiable effects of environment on the residual strength of the material, a single j value, obtained by fitting room-temperature residual strength data to Eq. (2), was found to be sufficient. In fitting the data, the maximum stress criterion was used for Fa, Fa⫽
sapplied max , Xt
(3)
and t was taken to be n/N, where average values of room-temperature fatigue life data for each stress level were substituted for N and the number of applied fatigue cycles as given in Table 3 were substituted for n. In this way, the value of j was determined to be 1.3 for the material considered in this study. As can be seen from the residual strength data and predicted residual strength curves shown in Fig. 17, reasonable agreement was found between experimental and predicted residual strength values for all of the environments considered. However, elevated-temperature results match the calculated residual strength curves more closely if the curves are calculated using average elevated-temperature lifetimes (N). The j value, however, remains 1.3. This result in shown in Fig. 18.
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Fig. 17. Residual strength data and predictions based on room-temperature data.
Fig. 18.
Residual strength data and predictions for elevated temperature (120°C).
5. Summary and conclusions 1. Fatigue life and residual strength of the woven graphite/epoxy composite material studied are, at best, minimally affected by the elevated temperature, moisture or combined hygrothermal conditions imposed and stress levels considered, despite differences in damage accumulation processes. 2. The major damage mechanism resulting from quasistatic loading is cracking in transverse tows. The cracks extend across a single transverse fiber bundle at maximum length. In addition, the majority of the cracks appear relatively close to failure but do not reduce the axial elastic modulus. 3. Damage progression during fatigue occurred in the following general order: transverse microcracking and
delaminations around the fiber bundle undulation regions, longitudinal microcracking around fiber bundle undulation regions, growth of transverse cracks across the entire height of the transverse fiber bundle, and cross-ply-like growth of edge and interply delamination. The rate at which damage progression occurred generally depended on the maximum fatigue stress amplitude and environment in that: 3.1. damage accumulation was found to be greater with decreasing fatigue stress levels; 3.2. fiber/matrix debonding and delamination were found to be greater for the elevated-temperature and moisture-saturated specimens than for roomtemperature specimens; and 3.3. hygrothermal cycling during fatigue was found to result in greater crack density during fatigue
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than did any of the other three conditions.These results suggest that the fatigue life of this material may be more adversely affected by environment for off-axis loading situations than was captured by the fiber-direction loading case considered in this study. 4. A residual-strength-based life prediction model was used in which an experimentally determined model parameter was obtained from room-temperature data and applied for all conditions with reasonable results.
[3] [4]
[5]
[6]
Acknowledgements The authors wish to acknowledge the financial support of Pratt and Whitney, the National Science Foundation (under grant #CMS-9872331), and the Virginia Space Grant Consortium. In addition, the authors would like to thank NASA Glenn Research Center (particularly Mike Castelli) for performing hygrothermal aging on composites.
[7]
[8] [9]
[10]
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