artificial retardation of fatigue crack growth by the infiltration of cracks by foreign materials

Fatigue & Fracture of Engineering Materials & Structures 1998; 21: 835–846

ARTIFICIAL RETARDATION OF FATIGUE CRACK GROWTH BY THE

INFILTRATION OF CRACKS BY FOREIGN MATERIALS

C. S. S, K. C. H and R. Z. LDepartment of Mechanical Engineering, National Taiwan University, No. 1, Sec. 4, Roosevelt Road, Taipei, 10617, Taiwan,

Republic of China

Received in final form February 1998

Abstract—The effects of inducing artificial crack closure into fatigue cracks in AISI 304 stainless steel byinfiltrating foreign materials have been investigated. The foreign materials used include pure epoxy resinand resin mixed with 0.3 mm and 4 mm TiO2 , 4 mm Fe, as well as 18 mm AISI 316L stainless steel. In allthe cases studied, different degrees of crack growth retardation have been achieved. When the particlesize was small enough or when the prop-opening load for infiltration was large enough, crack arrestoccurred. Crack retardation and arrest were mainly caused by the infiltrated material rather than thepropping load. A rigid-wedge model was found to have limited value in predicting the possible outcomeof an infiltration. On the other hand, the degree of crack closure immediately on resumption of a testafter infiltration could tell whether the treatment was going to be successful or not.

Keywords—Crack infiltration materials; Fatigue crack closure; Crack arrest; Fatigue damage repair,Rigid-wedge model.

INTRODUCTION

Research in fatigue has largely been concentrated on understanding the mechanisms and pre-dicting the progress of failures. Relatively little work has been done on repairing fatigue damage.Early studies on the possibility of fatigue damage repair have been trying to revert such damages.They tried to remove the fatigue-damaged surface material [1–4] or revert damage throughintermittent heat treatment [5–9]. Despite the varying degrees of success reported in some ofthese works, the difficulty to treat full scale components and the lack of consistent performancerendered these early attempts futile for practical engineering applications. Above all, these attemptswere based on the infinite life design concept, which would invariably break down when pre-existing flaws occurred.

A number of more practical methods have been employed for cracked components. Theseinclude: (i) grinding removal [10]; (ii ) weld overlay [10]; and (iii) stop drilling [11–15]. Apromising approach identified in a recent comparison among some possible fatigue damage repairmethods is the introduction of artificial crack closure by infiltrating foreign material [15] into afatigue crack. Other preliminary works [16–19] also showed that this is a feasible repair method,and worth further studies to improve its reproducibility and effectiveness. In the current work, theeffectiveness of a number of different infiltrants has been compared. A better infiltrating methodwhich can lead to crack arrest and give reproducible results has been developed.

FATIGUE CRACK GROWTH RETARDATION BY INFILTRATION

It is well known that a tensile overload can cause fatigue crack growth retardation [e.g. 20].This beneficial effect has been made use of as a by-product in the proof-testing of pressure vesselsto increase their fatigue life [21]. The deliberate use of tensile overload to prolong fatigue life mayrisk causing a momentary stretch growth or even catastrophic fracture. It would be useful if we

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836 C. S. S et al.

can induce the mechanisms that cause overload retardation by some means other than applyinga tensile overload directly. The mechanisms underlying retardation may be varied [2,22–27], yetpremature crack closure stands out to be one of the most significant mechanisms [27–29]. In arecent study of the overload effect in AISI 304 stainless steel, plasticity-induced crack closure hasbeen identified to be the major mechanism for crack growth retardation [30]. Residual plasticdeformation leaves behind extra material on the crack flanks, causing premature crack closureand reduction in the effective crack driving force.

A number of preliminary works indicated that foreign materials may be introduced into thecrack flanks to achieve a similar purpose. Kitagawa et al. [16] showed that the ingress of straingauge adhesive was able to arrest the growth of a crack at low stress intensity level. Vecchio et al.[17] introduced a needle tip into the wake of the crack tip and reduced the growth rate slightly.More systematic works showed that epoxy resin [15,18,31], alumina powder [15], electro-de-posited nickel [19,32] and solder [19] can achieve retardation to different extents.

In our preliminary work [15,18], vacuum suction was employed to help transport the infiltrantinto the crack. On breaking open the tested specimens, it was found that the degree of infiltrantpenetration varied across the thickness. The retardation effect was therefore not consistentlyreproducible.

-Rehman and Thomason [19] deduced theoretically that the strength of the infiltrant willnot affect the extent of retardation. Experimental results on this topic are still limited but some ofthem suggest that a softer infiltrant is less effective in bringing about retardation [18,33]. It is notclear whether this is the result of improper infiltrant penetration as no specific attention was paidto the reproducibility of results in these tests.

In the present study, bottled pressurized nitrogen has been employed to force the infiltrant intothe crack. This method offers a much higher infiltration pressure (~2–4 atm) than the vacuumsuction method (<1 atm). Moreover, the present set-up is much cheaper and more adaptable toin situ applications. Besides, the basic epoxy resin infiltrant has been further reinforced by differentmetallic powders in the hope of evaluating the effect of infiltrant strength on retardation. Althoughcurrently tests were carried out at room temperature, powder-reinforced resins have practicalimplications for higher temperature applications: under such conditions, the resin may have beendecomposed, but the powder may still provide retardation effect. It is therefore worthwhile tostudy the possibility and effect of infiltrating different powder materials and different sized powder-reinforced resin.

EXPERIMENTAL PROCEDURES

Compact tension specimens of width 50 mm and thickness 5 mm were machined from an AISI304 stainless steel sheet stock. Fatigue tests were carried out on an MTS 810 closed-loop servo-hydraulic machine. A sinusoidal waveform with a frequency of 10 Hz was used. Crack length wasmonitored using a travelling microscope to a resolution of 0.01 mm. Manual load shedding wasemployed to maintain a constant baseline DK of 23 MPaEm throughout all the tests. The R ratio(min. load/max. load) was 0.1.

After a crack growth increment of ~5 mm, the growth rate and crack closure response bothsettled down to steady state values. The test was then interrupted and the crack held open at0.95Kmax . An alumina wedge was inserted into the crack mouth until it is hand-tight. The specimenwas then removed from the testing machine and infiltrated with powder-reinforced epoxy resinusing pressurized nitrogen. The detailed infiltration method has been reported in Ref. [34]. The

Fatigue & Fracture of Engineering Materials & Structures, 21, 835–846 © 1998 Blackwell Science Ltd

Artificial retardation of fatigue crack growth 837

resin was allowed to harden for 3 days before the alumina wedge was removed. The test was thenresumed using the same constant stress intensity range of 23 MPaEm.

Powder-reinforced resin was prepared with a resin to powder weight ratio of 2 to 1. Since aparticular powder was available in a very limited range of sizes, a number of different materialshave been used. The powders used to reinforce epoxy resin included 0.3 mm TiO2 , 4 mm TiO2 ,4 mm Fe and 18 mm AISI 316L stainless steel.

In the course of the experiment it was found that the large-diameter powder tended to blockthe resin from getting into a crack and therefore resulted in poor retardation. To counteract this,the crack has to be propped open at 1.2Kmax and even 1.5 Kmax . Wedging open at these loadsalso ensured perfect penetration of the resin reinforced by smaller diameter powders. As a controlfor comparison, tests have also been carried out on specimens which have been propped open forthe same duration at 0.95Kmax , 1.2 Kmax and 1.5 Kmax without any infiltration.

During the fatigue tests, the crack opening load was monitored by a compliance method aidedby an offset procedure similar to that proposed by Kikukawa et al. [35]. The relevant displacementfor compliance information was measured with a back face strain gauge. The gauge output wasalso closely monitored throughout the wedge insertion and infiltration process to ensure that nounwanted overload had been induced.

A low pass filter with a cut-off frequency of 1 Hz was used to reduce electrical noises. As aresult, testing frequency was reduced to 0.05 Hz during closure measurement to avoid signaldistortion. The degree of crack closure, U, was defined as the fraction of the load range for whichthe crack is open:

U=Kmax−KopKmax−Kmin

=DKeffDK

(1)

where Kop is the stress intensity at which the crack started to become fully open and DKeff is theeffective range of stress intensity available for growing a crack.

RESULTS AND DISCUSSION

Baseline fatigue crack propagation data

The fatigue crack growth rate for the 304 stainless steel employed may be related to the stressintensity range DK through Paris’ law:

da

dN=C(DK)m (2)

where C=7.394×10−10, m=3.593, da/dN is in mm/cycle and DK in MPaEm.Elber [26] suggested to take premature crack closure into account. In this case, the crack

growth rate may be correlated through DKeff :

da

dN=Ce (DKeff )me (3)

with Ce=3.597×10−8, me=2.657 and DKeff is as defined in Eq. (1).Different infiltration conditions and infiltrants caused crack growth retardation to a different

extent. The most straightforward parameter to quantify the beneficial effect of a retardation is theamount of life extension expressed in the number of cycles. The amount of life extension iscomputed by the difference between the actual number of cycles elapsed for crack growth to returnto steady state and the cycles elapsed if the crack had grown at the constant baseline rate

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838 C. S. S et al.

throughout. The severity of maximum retardation can be described by a retardation ratio (minimumgrowth rate/baseline growth rate). The overall retardation-affected crack growth increment alsoaffects the effectiveness of a life extension. These three parameters will be used to quantify thefollowing retardation tests.

Retardation caused by prop-opening load

During infiltration, the fatigue crack was propped open using 0.95–1.5 times the maximum stressintensity for about 3 days. Being near to or even higher than the maximum cyclic load, theseopening loads may cause a certain extent of crack tip blunting and may also induce overloadretardation effect. To assess the possible degree of retardation so caused, a number of specimenswas tested after being subjected to the propping loads alone.

The baseline crack propagation rate at an applied constant DK of 23 MPaEm was~6×10−5 mm/cycle. Propping open at 0.95Kmax for 3 days did not cause any observable changein the fatigue crack growth behaviour. Propping open the specimen at 1.2 Kmax and 1.5 Kmaxcaused an immediate acceleration to ~1.5×10−4 mm/cycle on resumption of the baseline cyclicloading. Thereafter, growth rate steeply decreased to ~67 and ~33% of the baseline rate,respectively. This was followed by gradual acceleration back to the steady baseline state. Figure 1shows the typical disturbances on growth behaviour by these two propping loads. The averageretardation affected zones were 3 and 6.5 mm, respectively. The amount of life extension was ~10and ~68 kilocycles for the 1.2Kmax and 1.5 Kmax cases, respectively. Broken lines in Fig. 1 showthe growth rate calculated using Eq. (3) and the measured level of crack closure. These predictedrates agree well with the measured rates throughout the whole growth transients.

Retardation eVect by infiltrating at 0.95Kmax

Figure 2 shows the effect of infiltrating epoxy resin on crack growth behaviour on three differentspecimens. On resumption of cyclic loading after infiltration, growth rate retardation occurredimmediately. In CT1 and CT3, immediate deceleration by over an order of magnitude was observed.Growth rate then gradually increased from this minimum and took crack growth increments ofmore than 5 mm for the cracks to gradually accelerate back to steady state values. In CT2, growth

Fig. 1. Disturbances on crack growth after being propped Fig. 2. Typical growth behaviour after infiltrating pureopen at 1.2 Kmax and 1.5 Kmax . (Arrows show points of epoxy resin at 0.95 Kmax .

overload.)



Fig. 3. Typical crack closure development after infiltrating pure epoxy resin at 0.95Kmax .

rate only decreased by 80%. Recovery to steady state occurred much faster than the other twocases. Figure 3 shows the development of crack closure in these three specimens. The relativeseverity of a decrease in crack opening level and the growth increments required to regain thesteady state condition closely matched the development of the growth rate transients in therespective tests. Broken lines in Fig. 2 represent the predicted growth rate based on Eq. (3) andthe measured crack opening levels. In all three cases, the agreement between the observedand predicted rates is reasonable.

Similar tests have been repeated on six specimens. The respective maximum retardation ratio,minimum U value reached, crack growth increment affected by the infiltration and the amount oflife extension obtained are listed in Table 1. Comparison among these parameters suggests that

Table 1. Retardation behaviour under an infiltration load of 0.95 KmaxAffected growth Life Wedge Wedge

Retardation Min. increment extension proximity d thickness 2hInfiltrant Specimen ratio U (mm) (cycles) (mm) (mm)

Epoxy resin CT1 0.03 0.15 11.25 255 726 0.6 10.2CT2 0.21 0.43 4.36 73 300* 1.1 7.2CT3 0.05 0.22 6.26 301 301 0.4 6.4CT4 0.03 0.17 6.89 288 334 1.6 14.0CT5 0.04 0.18 7.71 557 054 2.1 14.2CT6 0.04 0.15 10.54 410 886 1.7 15.0

Epoxy resin TIA1 0.24 0.35 5.98 69 626* 1.5 9.0+0.3 mm TiO2 TIA2 0.03 0.16 8.61 492 702 0.9 12.0

TIA3 0.02 0.20 8.68 279 337 0.9 11.2TIA4 0.02 0.15 8.07 372 311 1.4 13.4TIA5 0.29 0.41 5.08 49 889* 0.6 6.0TIA6 0.04 0.19 9.41 472 711 1.4 13.4

Epoxy resin TIB1 0.03 0.19 7.80 138 862 0.7 8.2+4 mm TiO2 TIB2 0.09 0.14 4.65 300 713 0.2 6.4

TIB3 0.03 0.16 6.00 115 925 0.9 10.4TIB4 0.04 0.21 5.08 91 726 0.8 8.6

Epoxy resin FE1 0.03 0.18 7.00 190 591 0.9 11.4+4 mm Fe FE2 0.05 0.19 7.02 442 240 0.5 8.2

Epoxy resin SSA1 0.06 0.22 7.18 180 584 1.3 11.0+18 mm 316L SSA2 0.20 0.28 5.35 76 077* 0.8 6.2

*Unsuccessful infiltration attempt.


840 C. S. S et al.

the infiltration of specimen CT2 was not successful, as the minimum growth rate reached and theamount of life extension achieved were far inferior to that of the other tests. Results in these andlater tests show that successful infiltration attempts are marked by a retardation ratio smaller than0.1 and a minimum U value of ~0.2 or lower. By monitoring the U value on resumption of test,one can tell whether the infiltration will be effective to bring about good life extension.

For the remaining five successful cases, although similar minimum growth rates and crackclosure U values were achieved, the amount of life extension varied from 256 to 557 kilocycles.The results show that for successful infiltration that achieved considerable life extension, theamount of life extension cannot be related to the maximum retardation, minimum U value andaffected crack growth increment in a straightforward manner. As an example, CT3 was inferior toCT4 in these three parameters, yet it achieved a better amount of life extension.

The maximum retardation and minimum U value quantify the immediate effect of the infiltratedwedge. The affected growth increment reflects the effect of the wedge. None of these three parameterscontains information about the contribution to crack closure over the entire affected crack length,which probably depends on the morphology of the wedge. On the other hand, the amount of lifeextension reflects the cumulative effect of the infiltrated wedge. The scatter in the amount of lifeextension is probably caused by the variability in the morphology of the infiltrated wedge.

Table 1 also lists the retardation behaviour when infiltration was carried out with epoxy resinmixed with 0.3 mm TiO4 , 4 mm TiO2 , 4 mm Fe and 18 mm AISI 316L stainless steel powders.

Excluding the unsuccessful attempts, addition of 0.3 mm TiO2 to epoxy resin brought about anessentially similar retardation effect as pure resin. Addition of 4 mm TiO2 to epoxy resin onlyachieved a life extension of 92–301 kilocycles. This result is inferior to that of 0.3 mm TiO2 addition.A possible explanation is that the larger particle size hindered the infiltrant from going well intothe crack. Although 4 mm Fe powder has the same nominal size as the 4 mm TiO2 , addition of theFe powder to resin has practically the same retardation effects as pure resin and resin with 0.3 mmTiO2 . This may be explained by the relative density of iron (#7.9) being higher than that of TiO2(#4.3). In a fixed resin to powder weight ratio of 2 : 1, the amount of 4 mm Fe particles is less thanthe corresponding 4 mm TiO2 and so the blocking effect is less pronounced. Finally, with the 18 mmAISI 316L stainless steel powder addition, poor retardation effect was obtained. In fact, duringinfiltration of the 18 mm AISI 316L stainless steel powder-added resin, the infiltrant failed to comeout on the opposite side of the specimen surface in many attempts. This demonstrated that thelarger size 316L powder has a serious blocking effect on the resin carrier.

Figure 4 shows some typical fracture surfaces. The portions which have resin appear darker.The prevalence of the darker colour suggests that the current infiltration method can fill throughthe thickness effectively. Variation in colour within the infiltrated area probably indicated that thefilling was not uniform. This is especially so in the powder-added resin as segregation of powderseems to have occurred near the crack mouth region. Dark colour was seen on both halves of thefracture surfaces, suggesting that fracture of the wedges occurred and resin was left on both of themating surfaces.

Retardation eVect by infiltrating at above Kmax

By propping open the crack at 1.2 Kmax , infiltration became a lot easier. All infiltrants exceptthe 18 mm AISI 316L powder-added resin caused a significant immediate retardation, so that nocrack growth was detected for over 2 000 000 cycles. With a crack length resolution of 0.01 mm,this corresponds to a crack growth rate below 1×10−8 mm/cycle. For the 316L powder-addedresin, the large sized powder again provided considerable hindrance to the resin, and crack growthstill occurred after infiltration. A better retardation ratio of 0.01 and a lower minimum U value of



Fig. 4. Photographic pairs showing resin infiltrated under 0.95 Kmax in: (a) CT1, pure epoxy resin;(b) TIA6, resin+0.3 mm TiO2 powder; and (c) TIB4, resin+4 mm TiO2 powder.

0.09 were achieved (Table 2) in comparison to infiltration at 0.95Kmax . The degree of life extensionamounts to 1.1–2 million cycles.

Finally, by propping open the crack further to 1.5 Kmax for infiltration, the 316L powder-addedresin was then capable of arresting the crack.

All tests mentioned in this section have been repeated at least on two different specimens andthe results are reproducible.

Rigid-wedge model for the artificial infiltration

For the case of oxide debris-induced crack closure, Suresh et al. [36,37] have proposed a rigid-wedge model to quantify the level of crack closure. Figure 5 shows schematically a semi-infinitewedge of constant thickness 2h fitted snugly into a crack. At the time of insertion, the crack lengthwas a0 and the wedge tip was at a distance d behind the crack tip. The crack subsequently grewto the current length a. The stress intensity Kw induced by the wedge is given by [38]:

Kw=hE∞

√2p(a−a0+d)(4)

where E∞=Young’s modulus E for plane stress, E∞=E/(1−n2 ) for plane strain.Researchers -Rehman and Thomason [19,39] argued that due to small deformation and large

plastic constraint, a low strength elastic wedge may be regarded as nearly rigid. Moreover, as thecrack opening angle was small, the infiltrated wedge may be considered to have a constant

Table 2. Retardation behaviour on infiltrating epoxy resin with 18 mm AISI 316L powder

Affected growth Wedge WedgeRetardation Minimum increment Life extension proximity d thickness 2h

Specimen ratio U (mm) (cycles) (mm) (mm)

SSB1 0.01 0.09 8.60 1 125 808 0.9 13.0SSB2 0.01 0.08 8.68 1 994 394 1.6 16.6

Infiltration load: 1.2 Kmax .


842 C. S. S et al.

Fig. 5. Schematic diagram of the rigid-wedge model. Fig. 6. Relation between incremental crack growth from theinfiltrated crack tip and wedge contribution to crack closure

for specimen CT1.

thickness. As a result, the crack closure intensity contributed by the infiltrated wedge may bedescribed by Eq. (4). They used the average of crack tip and crack mouth opening displacementsfor wedge thickness, and estimated d from fractographs. On taking into account this wedgecontribution to crack closure, they found the cracks under investigation should have arrested,contrary to what they have actually observed. At crack lengths sufficiently far away from thewedge tip, crack growth was predicted to occur, but the predicted rates were lower than observed.-Rehman and Thomason attributed the differences between the observed and predicted growthbehaviour to: (i) finite crack length effect; (ii) irregularity in wedge geometry; and (iii) frictionbetween the wedge and crack interfaces.

Instead of trying to estimate the wedge details as above, we tried to deduce d and 2h from themeasured crack closure behaviour as follows: an experimental Kw was first deduced from themeasured U values by deducting the steady state crack opening level without the wedge. (a−a0 )was then plotted against 1/K2w for every test. A typical plot is shown in Fig. 6. Most of the datafall nicely on the straight line:

162.51

K2w= (a−a0 )+0.604 (5)

where Kw is in MPaEm, a is in mm.In all specimens, it was found that for (a−a0 ) smaller than 4 mm, the data points fit well to a

straight line. d and 2h may be deduced from this straight line fit and are listed in Table 1 for each test.Plotting Kw against (a−a0+d) on a log–log scale (Fig. 7) shows that the initial portion of

curve fits well to a power law relationship:

Kw=12.96(a−a0+d)−0.51 (6)

The exponent for (a−a0+d) is very close to the −0.5, as indicated by Eq. (4). For all theinfiltration tests listed in Table 1, the power law exponent ranged from −0.39 to −0.57, with anaverage value of −0.496 and a standard deviation of 0.04. This shows that by substituting



Fig. 7. Relation between incremental crack growth from the Fig. 8. Deduced wedge thickness compared with crackdeduced rigid-wedge tip and the wedge contribution to opening displacement at the corresponding deduced position

crack closure for specimen CT1. of the wedge tip for all specimens.

appropriate d and h, Eq. (4) can describe the infiltration-induced crack closure well when the cracktip is less than ~4 mm from the position where infiltration took place.

Figure 8 compares the deduced wedge thickness with the crack opening at the corresponding d.Crack opening displacements (COD) were computed using an elastic–plastic finite element analysis,as well as from a Dugdale model for CT specimen proposed by Mall and Newman [40]. CODfrom the Mall–Newman model agrees fairly well with the finite element results if the averagebetween the initial yield (350 MPa) and ultimate tensile (675 MPa) strengths is used as the flowstress in the former.

The crack opening profile evaluated by the finite element method forms an upper bound to allthe deduced wedge thickness. The cases of unsuccessful infiltration (diamond symbols) are associ-ated with small wedge thickness. On the other hand, the deduced thickness values are very closeto the crack opening displacement when the propping load was 1.2 Kmax (star symbols in Fig. 8).The average proximity of infiltration, d, under 0.95Kmax was 1.01 mm, and the average fill-up ratioof the wedge thickness to the COD at the corresponding d was 0.66.

Figure 9 compares the deduced geometry and location of the rigid wedges (shown as hatchedareas) with crack profiles measured near the specimen surface and along the mid-thickness position

Fig. 9. Comparison between the wedge geometry deduced from the rigid-wedge model, and the measuredand calculated wedge profiles. (The vertical arrow indicates the surface crack length during infiltration.)


844 C. S. S et al.

for one of the specimens infiltrated with the 18 mm AISI 316L powder-added resin at 1.2 Kmax .The arrow in the figure indicates the surface crack length at which infiltration has been applied.Fracture surface profiles were measured by a ball-pointed probe traversing along the crack.Combining the profiles from mating fracture surface gives the resultant crack profiles shown inFig. 9. With the average height of the infiltration-free surface taken as the reference zero level, theresultant crack profile reflects the geometry of the infiltrated wedge. The deduced wedge thicknessesare much smaller than the measured wedge heights in general. They only compare well at theposition of the deduced wedge tip. The deduced spacing between the wedge tip to crack tip is alsonot supported by the measured evidence. A total of three specimens have been measured in thisway and the above observations applied to the other two specimens as well. Hence, the usefulnessof the rigid-wedge model to predict the possible outcome of an infiltration is limited, as no objectiveguideline can be drawn about what values d and h should take.

Also shown in Fig. 9 are the crack opening profiles calculated using the Mall–Newman model[40]. The general shape of the measured profiles agrees with the calculated one. Occasional crestsand troughs may rise above or drop below the computed profiles. The troughs may be gas holesor another form of incomplete filing. The occurrence of the crests rising above the crack openingprofile may be explained by the fact that the smallest ball-point probe available has a diameter of1 mm. Such a probe will reveal a small protrusion of reinforcing powder or resin tearing mark,but will be insensitive to the corresponding hole on the mating surface. This may give rise tooverestimation and result in the observed crests rising above the calculated profiles.

CONCLUSIONS

By infiltrating epoxy resin, with various powders added, into a propped-open fatigue crack,crack growth retardation can occur immediately. Infiltration at 0.95Kmax caused life extensionsranging from 76 to 557 kilocycles. When propped open at 1.2 Kmax , crack arrest occurred for allexcept the 18 mm AISI 316L stainless steel powder-added resins. Life extension in the latter casewas between 1125 and 1994 kilocycles. At 1.5 Kmax , even the 316L powder-added resin successfullybrought about crack arrest. Good reproducibility of results has been achieved when the proppingload is 1.2 Kmax or higher.

The difference in retardation behaviour under different infiltration load and reinforcing particlesis believed to be related to how effectively the infiltrant goes into the crack tip, and it is notedthat a smaller amount of life extension occurred if infiltration is hindered by either one or moreof the following: (i) infiltration load is small; (ii ) reinforcing particle size is large; (iii ) the infiltrantcontains too many reinforcing particles.

Although the crack closure behaviour after infiltration can be described by the rigid-wedgemodel when appropriate values of wedge thickness and proximity of wedge to crack tip weresubstituted into Eq. (4), these values were not supported by experimental observation.

Acknowledgement—The authors gratefully acknowledge the financial support provided by the National Science Councilunder the projects NSC-83-0401-E-002-087 and NSC-84-2212-E-002-024.

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