Effect of orientation and loading rate on the toughness transition curve of a ship steel

Transcription

1 Effect of orientation and loading rate on the toughness transition curve of a ship steel X. Wu% J. Morrison', & D.A. Hull" ''DREA Dockyard Laboratory (Pacific), Canada. 'Engineering Material Research, Canada. Abstract Toughness measurements are needed to assess the risk of brittle fracture in warship hulls under operational conditions. This demands laboratory testing at minimum service temperatures and loading rates equivalent to hull impact events. However plate thickness is insufficient to permit valid Kjc determinations at the required strength and toughness levels. The master curve approach specified in ASTM Standard El921 provides an alternative means of quantifying the cleavage fracture resistance in the brittle-to-ductile transition temperature regime. Limited replicate testing yields a reference temperature T<,. It also defines a material-specific fracture toughness master curve, and a lower bound failure probability, which can be applied to material selection and structural integrity analyses. The determination of TO and master curves for a 400 MPa yield strength steel plate is described. Variations in specimen orientation and test loading rate have been investigated, and comparisons are made between the reference temperature results and those from other standardized fracture tests, ft is concluded that the master curve approach is a useful way of quantifying hull toughness for warship fracture control. Introduction The hulls of naval vessels are designed to provide an effective and durable platform for the operation of a variety of devices, including weapons systems [1]. They must have maximum operational availability in severe environments, and be able to withstand the dynamic loads expected in combat. They must also, if necessary, continue to function while damaged. Although much of the early work in fracture mechanics and fracture control plans was carried out in naval laboratories, structural integrity was ensured for the most part using empirical methods [2, 3]. These relied on semi-quantitative correlation between material properties and structural performance to ensure adequate toughness [4]. This

2 94 Damage and Fracture Mechanics VI approach has been very successful, and warships have rarely experienced serious cracking problems. However, this has been achieved by adopting very conservative assumptions about their structural robustness in the presence of flaws. Future warships will have to achieve greater performance and structural efficiency, both by weight reduction and by designing for higher operating stresses. At the same time pressure to minimize maintenance budgets and maximize operational availability will mean sometimes delaying high cost repairs unless ship safety or survivability is in question. Increasing sophistication in both fracture testing and structural modelling will therefore be needed to demonstrate adequate brittle fracture resistance and damage tolerance in a more rigorous fashion than has previously been possible. Economic considerations for the most part also preclude the use of large-scale tests for structural integrity assessment, so that laboratory measurements using small specimens can be expected to remain the primary source of toughness data. In recent construction programs the material specifications for the steels used in Canadian warships have relied on tensile strength and Charpy impact energy, primarily for quality control. These steels are both stronger and tougher than those used in commercial shipping. The data obtained from traditional impact tests cannot provide a quantitative description of the fracture behaviour of the welded structure. In addition, except at very low temperatures, plate thickness (typically less than 15 mm) is insufficient to permit valid K^ determination using test methods such as those in ASTM Standards E399 and El820. An alternative method of quantifying cleavage fracture resistance in the brittle-to-ductile transition temperature regime is to use the master curve approach specified in ASTM Standard El921. This procedure combines a statistical approach to failure probability with an elastic-plastic analysis of cleavage instability [5]. Limited replicate testing of precracked specimens yields a median K,c at the onset of cleavage, from which can be calculated a reference temperature TO, corresponding to a median Kjc of 100 MPaVm for a IT (B=25mm) specimen. Ferritic steels tend to follow a common transition curve shape [6]. The reference temperature data can be used to generate a fracture toughness master curve for the transition region, together with any desired lower bound failure probability curve. Since the minimum specimen thickness required for ASTM El921 is much less than for similar fracture mechanics-based tests, the master curve approach has generated considerable interest, particularly in the nuclear power industry [7], and the technique offers some interesting possibilities for both naval material specification and ship fracture control. The Dockyard Laboratory (Pacific) is investigating how master curve technology could be used to characterize the transition fracture behaviour of structural steels and welds used in the hulls of Canadian frigates. Reference temperatures and master curves have been generated for one such steel in both the T-L and L-T orientations, and compared with data from Charpy, dynamic tear, and crack tip opening displacement tests. ASTM standard El921 currently includes only quasi-static loading rates, but in this work testing has been carried out at impact velocities up to 1 m/s in order to provide an appropriate simulation of service conditions.

3 Materials and Test Procedure Damage and Fracture Mechanics VI 95 The 15 mm thick control-rolled low alloy steel plate had the chemical composition shown in Table 1, which also shows selected tensile, CTOD, and impact properties together with the specified values, where appropriate. Plate yield strength is significantly greater than the specified minimum. The steel is also very tough. In CTOD tests at quasi-static loading rates, cleavage fracture is usually absent at temperatures above -30 C, while Dynamic CTOD tests on similar plates are ductile at -15 C. Table 1. Chemical composition and mechanical properties. cchemical Compos;ition (v\'t%) c Mn Si S P Ni V Ti Al Mechanical P ropertie. T-L L-T Specifiec (L-T) 0.2% Proof^Stress ( VlPa) 450 (T) 447 (L) 350 ITlin. UTS( 'MPa) Elong ation (' >/o) 26 2' 7 22m in Charr_ >y V@ -40 C CO min* DT(g) -20 CCO ! 35 n/a CTOI]@-2( ] C (rrim) n/a NDrr( c) n/a * Additional Charpy requirements are +20 C and -60 C. Full Charpy and Dynamic Tear curves are shown in figure 1. The T-L curves have lower upper shelf energies, and tend to fall off gradually. There is, however, an ambiguity in the plate L-T Charpy values between -30 C and -80 C where the data follow two different trends. One has a gradual transition with high energies. The other has an abrupt drop to a level close to the T-L lower shelf. This behaviour may be related to the mid-thickness splitting which can occur in larger specimens during ductile tearing in CTOD tests at higher temperatures. This type of fracture is associated with texture and microstructural effects produced by the controlled rolling. The generation of master curves from T-L and L-T specimens is intended to clarify the resulting uncertainty about the variability of low temperature toughness with orientation in this plate. In contrast, there was little directionality of either tensile properties or the NDT Temperature. One interpretation of the difference between DT and NDTT is that the latter is related more to crack arrest than crack initiation. Kjc fracture toughness measurements were made on fatigue pre-cracked singleedge notched bend specimens using the ASTM El921 procedure. Full thickness specimens with W/B=2 were used. The test procedure, which is similar to other pre-cracked specimen test methods, is intended to produce a cleavage instability or pop-in. The required test temperature is estimated from the 28J Charpy energy

4 96 Damage and Fracture Mechanics VI cs Temperature ( C) 20 Figure 1: Charpy and dynamic tear test results. temperature (T = Tigj + C, where, for this thickness plate, C is about -30 C). Since the plate had high toughness, the required test temperatures were very low. The irregular Charpy curve resulted in uncertainty about the suitable 28J Charpy temperature for L-T specimens. The test temperature is intended to produce a median measured Kjc of about 100 MPaVm. Provided that the IT Kjc(med) is above 83 MPaVm, a minimum of 6 valid test results is sufficient to permit reference temperature calculation. In this program, for each test condition, 12 specimens were made available. The first few could be used to optimize the test temperature to give an expected K^ed) value within the desired range. All specimens were side-grooved 10% of the thickness on each face after precracking to a/w=0.5. Three different loading rates were employed during these tests, with displacement rates of 0.05 mm/s (quasi-static), 6 mm/s (intermediate) and 1 m/s (dynamic). The highest elastic stress intensity rate was approximately 7x10^ MPaVm/s from the 1 m/s impact. This rate comfortably exceeds that expected in the worst case hull loading scenario resulting from wave slamming [8]. During a test, the actual rate at the onset of cleavage could be significantly lower, depending on the amount of prior plastic deformation. For this reason stress intensity rates are often calculated using time to failure. Individual K,c values must fall below the specified upper limit, which is determined by ligament size and material strength, otherwise a data censoring procedure is invoked. In this test program, the upper limit for a valid Kjc was about 250 MPaVm. Precise values require an estimate of tensile yield stress for the material at test temperature. In practice, tensile properties were measured at the two lowest displacement rates at various temperatures, and a conservative value for the highest rate was then inferred. Any test for Kje which either fails to produce a cleavage instability or exceeds a specified amount of stable crack

5 Damage and Fracture Mechanics VI 97 extension is discarded. There is also a lower limit Kjc based on a 2% lower tolerance bound, which in these tests was in the range MPaVm. Encountering a value below this limit (an "outlier") triggers a requirement to test additional specimens. The Kjc tests were carried out in either a servo-hydraulic machine or, for the highest loading rate, an instrumented drop tower. A low temperature chamber was used in the servo-hydraulic system, whereas the drop tower tests relied on rapid specimen transfer from a cold bath. Average test temperature was within ±2 C of that desired, with a gradient across the notch of up to 2 C. Table 2 shows test temperatures for the quasi-static tests estimated from Charpy, together with the actual temperatures used for each series. Load line and crack mouth opening displacements were both measured during the tests. In the drop tower, CMOD proved to be more precise than LLD, and was used for estimating the high rate J values. Tensile yield strengths are also shown in Table 2. Table 2. Test matrix and temperatures. Orientation L-T T-L Displacement Rate (mm/s) Test Temperature ( C) 28J Charpy Actual Estimate % Proof Stress (MPa) * * * inferred The underlying rationale for the master curve is a weakest-link model. A maximum likelihood statistical framework is provided in the test method to determine the 100 MPaVm mean toughness reference temperature for a IT specimen. From this the corresponding master curve can be determined, together with any desired tolerance bound. An alternative method using data from different temperatures has recently been added to the standard. Expressions are provided in the standard to adjust individual Kjc values to those equivalent to a IT specimen, and from these to calculate a Weibull scale factor KQ. A median Kjc in MPaVm can be calculated from the equation: and the reference temperature in C determined from the expression: In 70

6 98 Damage and Fracture Mechanics VI where T is the test temperature. The latter equation defines the transition master curve, and a set of coefficients is provided for any required tolerance bound. Slightly different procedures are available for censoring any Kjc values which exceed the specified limit determined largely by specimen size. Results Damage & Fracture Mechanics VI, C.A. Brebbia, A.P.S. Selvadurai, (Editors) Sample quasi-static and dynamic load-displacement curves are shown in figure 2. All of the tests produced the desired cleavage instability before the stable crack extension limit of 5% of remaining ligament, i.e mm. However, in any given batch of specimens a wide range of fracture behaviour was observed, from virtually linear-elastic to near limit load elastic-plastic. For each data set Table 3 summarizes the reference temperature calculations. A high proportion of valid tests was obtained, with typically 8-10 data points per batch at the test temperature. In all cases median toughness was high enough so that, in principle, the minimum of 6 valid specimens was sufficient. In only one instance (L-T at 372 MPaVm/s) did outliers below the lower toughness limit necessitate additional testing (14 specimens). Comparison of tables 2 and 3 shows that in a few instances inexperience resulted in the selected test temperature being well above the reference temperature. The preliminary tests at lower temperatures gave Kjc values which in retrospect turned out to have very low probabilities, and implied (wrongly) that a higher temperature would be more suitable. In these cases, the median toughness will be much higher than 100 MPaVm, although this should have no impact on the calculated reference temperatures or master curves. Figure 3 shows the transition toughness master curves, including both the medians and 5% probability lower bounds. In these curves an upper limit cut-off has been used based on specimen Kjc capacity. Above this limit constraint loss may result in a specimen size dependent TO measurement CMOD (mm) Figure 2: Typical load-cmod curves at different loading rates. 0.5

7 Damage and Fracture Mechanics VI 99 Table 3. Reference temperature results. Orientation L-T T-L Loading Rate (time to fracture) (MPaVm/s) x10" x10" Ixic(med) (0.6T) K (MPaVm) Kjc(med) (IT) Ko To ( C) , Temperature "C 250 -, oo 50 _ Temperature ( C) Figure 3: Master curves with 5% tolerance bound (dashed lines) and TO.

8 100 Damage and Fracture Mechanics VI CJ -20 -, [91 10" 10 10^ 10" Loading Rate (MPam'''"/s) T-L L-T Figure 4: Effect of loading rate on reference temperature. Figure 4 shows a semi-logarithmic plot of reference temperature as a function of the average stress intensity rate based on the time to failure. The reported empirical trend for pressure vessel steels and welds is [9]: = l 7.1 1"C This relation is shown in figure 4. A lower rate dependence was found in the T-L orientation which also has more gradual Charpy and DT transitions. The trends are consistent, but the T-L reference temperatures (and thus master curves) at the two higher loading rates are remarkably similar. Further testing at a rate between these two will clarify whether the rate dependence does indeed level off in this orientation. The dynamic TO values are also both close to the -40 C NDTTs. A semi-empirical analysis of ship failures [8] indicated that a minimum Kjc of 125 MPaVm at a temperature of 0 C and a test rate of 5x10^ MPaVm/s would be sufficient to avoid brittle fracture in surface ships. A maximum wave bending stress of 100 MPa was used to rationalize this criteria. The reference temperature equation indicates that a IT median value of 125 MPaVm would be achieved at a temperature which is 16 C higher than TO, i.e. in the range -20 C to -25 C at the highest loading rate for this ship plate. For 15 mm thickness the corresponding temperatures would be 8 C lower. These are below minimum service temperatures, which are roughly -5 C below the water line, and about -20 C in the superstructure, even if no allowance is made for heat flow from the ship's interior. A K^ of 125 MPaVm at 0 C occurs on the 10% and 5% lower tolerance bounds for IT and 15 mm thickness respectively, so that cleavage fracture is very unlikely in this plate. These low fracture probabilities may not be typical of the corresponding fusion and heat affected zones from butt welded plate which are currently being investigated, nor of lower toughness ship plate.

9 Conclusions Damage and Fracture Mechanics VI 101 The master curve approach has been used to characterize the fracture transition of a low alloy ship plate. This description of cleavage resistance provided useful data to complement upper shelf tearing resistance estimates, structural integrity assessments, and material specification requirements. This particular plate has been shown to be very tough. Work is ongoing to clarify its rate dependence, and the anisotropy of fracture resistance observed in both fracture mechanics and impact tests. References [1] Gates, P.J., Surface Warships, An Introduction to Design Principles, Brassey's Defence Publishers, pp 1-12, [2] Sumpter, J.D.G., Fracture Avoidance in Submarines and Ships, Advances in Marine Structures - 2, Ed. C.S. Smith and R.S. Dow, Elsevier Applied Science, pp 1-22, [3] Vanderveldt, H.H., and Gudas, J.P., Application of Fracture Control Technology in Navy Ships and Submarines, Fracture Mechanics, ed. N. Perrone, H. Liebowitz, D. Mulville, and W. Pilkey,University of Virginia Press, pp 3-15, [4] Sumpter, J.D.G., and Caudrey, Recommended Fracture Toughness for Ship Hull Steel and Weld, Manne SYnfcmras - <3, pp , 1995^ [5] Merkle, J.G., Wallin, K., and McCabe, D.E., Technical Basis for an ASTM Standard on Determining the Reference Temperature TO, for Ferritic Steels in the Transition Range, NUREG/CR-5504, Nuclear Regulatory Commission, [6] Wallin, K., Fracture Toughness Transition Curve Shape for Ferritic Structural Steels, Proc. Fracture of Engineering Materials and Structures, ed. S.E. Teoh and K.H. lee. Elsevier Applied Science, pp 83-88, [7] Kirk, M., Lott, R., Server, W., and Rosinski, S., Initial Reference Temperature and Irradiation Trend Curves for Use with RTjo, a Preliminary Assessment, Proc. /4^ME Pr&s\s%rg KesW #W fzpmg Co/7/g/'e/?cg, Boston, Aug [8] Sumpter, J.D.G, Bird, J., Clarke, J.D., and Caudrey, A.J., Fracture Toughness of Ship Steels, /foya/y^nw/o/? q/wmw/w/7/fzc%\ 131, pp , [9] Joyce, J.A., On the Utilization of High Rate Charpy Test Results and the Master Curve to Obtain Accurate Lower Bound Toughness Predictions in the Ductile-to-Brittle Transition, Small Specimen Test Techniques, ASTM STP 1329, ed. W.R. Corwin, S.T. Rosinski, and E. Van Walle, pp , 1997.