Stress Corrosion Cracking in a Dissimilar Metal Butt Weld in a 2 inch Nozzle. Master of Engineering in Mechanical Engineering

Transcription

1 Stress Corrosion Cracking in a Dissimilar Metal Butt Weld in a 2 inch Nozzle by Thomas E. Demers An Engineering Project Submitted to the Graduate Faculty of Rensselaer Polytechnic Institute in Fulfillment of the Requirements for the Degree of Master of Engineering in Mechanical Engineering Approved: Ernesto Gutierrez-Miravete, Project Adviser Rensselaer Polytechnic Institute Hartford, CT August, 2013

2 Table of Contents 1.0 Introduction/Background and Purpose Input and Methodology Stress Intensity Factor (K) Solution Stress Input Crack Growth Methodology and Growth Laws Limiting Flaw Depth Methodology Results and Discussion Calculation of Stress Intensity Factor K I Calculation of Stress Corrosion Cracking Growth Stainless Steel Growth Rate Alloy 82/182 Growth Rate at 617 F Alloy 82/182 Growth Rate at 650 F Summary and Conclusion References... 45

3 List of Tables

4 List of Figures

5 List of Symbols

6 Acronyms

7 Keywords

8 Acknowledgements I would like to thank Professor Ernesto Guiterrez-Miravete and my Current and Former Westinghouse colleagues David Ayres, Warren Bamford, and Reddy Ganta for their advice on this project and subject matter. I would like to thank Ya T. Wu for his work developing the residual stresses used as input to this report, and his assistance in providing additional information and clarification when necessary. I would also like to thank Westinghouse Electric Company for their financial support of my engineering masters degree.

9 Abstract

10 1.0 Introduction/Background and Purpose Alloy 82/182 dissimilar metal (DM) butt welds have been used in nuclear power plant reactor coolant systems (RCS) to join stainless steel piping to low alloy or carbon steel components (see Figure 1-1). These welds are typically between RCS nozzles and RCS piping. Alloy 82/182, a nickel-based alloy also referred to as Alloy 600, has been found to be susceptible to pressurized water stress corrosion cracking (PWSCC) as documented in several industry documents [3, 9, 11]. This is also mentioned in the input residual stress calculation [1]. PWSCC is also referred to as stress corrosion cracking (SCC). Per [3], PWSCC in Alloy 82 and 182 welds involves an intergranular cracking mechanism, which is referred to as intergranular stress corrosion cracking (IGSCC). Figure 1-1: Typical Dissimilar Metal Nozzle Butt Weld [1] IGSCC refers to stress corrosion cracking that occurs along the grain boundaries within a material. This is referred to as active path dissolution in [7], which involves accelerated corrosion along a path of higher than normal corrosion susceptibility. While IGSCC is the mechanism experienced in Alloy 82/182, it is important to realize that the cracking is not occurring due to sensitization, which is typically discussed for IGSCC in stainless steels [3]. According to [8], SCC is cracking due to conjoint corrosion and straining of a metal due to residual and applied stresses. Furthermore, as described in [7], SCC depends on the simultaneous achievement of three requirements: A susceptible material An environment that causes SCC for that material Sufficient tensile stress to induce SCC All three criteria are satisfied. Alloy 82/182 has been shown to be susceptible to SCC, and its susceptibility is documented in multiple industry documents, including [3]. The high pressure and corrosive environment of a pressurized water reactor causes SCC for the material, specifically on the inside diameter surface of the weld. Finally, the weld experiences sufficient tensile stress to induce SCC. In this case the high tensile stresses are due to a combination of highly tensile weld residual stresses on the inside surface with an applied stress. This matches the description from [8]. As mentioned in [8], the effect of the applied stress is mainly to open up the cracks, thereby allowing easier diffusion of corrosion products away from the crack tip, which in turn allows the crack tip to corrode faster.

11 Both the DM butt weld and the ID repair weld are modeled in [1] as Alloy 82/182 (see Figure 2-2). When using the term Alloy 82/182, 82 refers to a weld made with a wire and 182 refers to a weld made with a rod. Per [3], cracks in Alloy 182 are found to grow more rapidly due to SCC than those in Alloy 82, but both are susceptible. The evaluations performed herein assume the material is Alloy 182 and use the higher growth rates associated with Alloy 182. An inside diameter (ID) repair weld composed of Alloy 82/182 material, modeled in [1], would result in a high tensile residual stress on the inside surface of the weld. This high tensile stress results in the weld being more susceptible to SCC crack initiation and more rapid growth due to SCC at the surface, and is therefore undesirable. Weld overlays (WOL) made of Alloy 52/152 (also referred to as Alloy 690) are frequently applied preemptively to compress the entire DM weld, including the ID repair [11]. As indicated in [3] and [10], Alloy 52/152 material is very resistant to PWSCC, and no industry guidance on calculating SCC growth in 52/152 welds has yet been developed because the growth is expected to be much slower. Additionally, PWSCC requires that the crack be exposed to the coolant on the inside surface. The Alloy 52/152 WOL is applied to the outside surface of the DM weld and pipe, and is not exposed to the coolant unless there is a through-thickness flaw in the DM weld, which is not expected without first having part-through wall indications. The analysis performed herein uses linear elastic fracture mechanics to calculate stress intensity factors based on handbook solutions [2]. These values for stress intensity factors (K I ), are then used as input to K-dependent growth calculations using guidance from [3].

12 2.0 Input and Methodology 2.1 Stress Intensity Factor (K) Solution This project will study the growth of semi-elliptical, circumferentially oriented flaws in the butt weld of a 2-inch pipe. The growth laws used herein (discussed in Section 2.3) require a stress intensity factor as an input. Stress intensity factors for semi-elliptical surface flaws can be found in API-579 [2]. Specifically, for evaluating stress corrosion cracking (SCC), flaws are postulated on the inside diameter (ID) of the component. This is because flaws on the ID will grow due to SCC, while flaws on the outside diameter and embedded flaws would not grow due to SCC because they would not be exposed to the coolant in the pipe. Section C.5.14 of [2] provides the solution for a Cylinder Surface Crack, Circumferential Direction Semi-Elliptical Shape, Through-Wall Fourth Order Polynomial Stress Distribution with a Net Section Bending Stress. This is applicable to the scenario evaluated herein. Equation 2-1: Mode I Stress Intensity Factor Solution In Equation 2-1: a = the flaw depth t = the wall thickness σ 0 through σ 4 are the coefficients of the fourth order polynomial of stress distribution through the wall thickness. σ 5 and σ 6 are the net section bending stresses about the x-axis and y-axis. For the depth point, only σ 5 is necessary for the calculation. Additionally, only one bending stress is used herein, so the σ 6 term is discarded. The influence coefficients G 0, G 1, G 5 and G 6 for inside and outside surface cracks can be determined using the following equations:

13 where the parameters A ij are provided in Table C.14 of [2] for an inside diameter crack and β = 2φ/π φ is the angle between the surface and the point being evaluated. In this case, φ = π/2 to obtain the stress intensity factor solution at the deepest point. G 2, G 3, and G 4 are calculated using the following equations for the deepest point (φ= π/2): where M 1 through M 3 are calculated as: and Q is calculated as:

14 The ratio a/c is always less than zero for the aspect ratios evaluated herein. p c = crack face pressure (2500 psi in this calculation) Figure 2-1: Component and Crack Geometry

15 2.2 Stress Input Stress input is available in [1] that contains the weld residual stresses. Two sets of stresses are available; one that is applicable after an inside diameter (ID) repair weld prior to application of a weld overlay, and one that is applicable after an ID repair weld with a weld overlay. The stresses are developed using a 2-D axisymmetric model of a 2-inch nozzle (see geometry summarized in Table 2-2). Figure 2-2 depicts the model. Figure 2-2: 2-D Axisymmetric Nozzle Model [1] Stresses are provided in [1] along the centerline of the ID repair weld. The stress path can be seen in Figure 2-3. The inside surface path is not used in this evaluation, since the stresses along this path are not useful in evaluating a circumferentially oriented flaw (a stress profile through the thickness is needed for this evaluation). Figure 2-3: Stress Paths at the DM Weld [1]

16 Figure 2-4 is used to justify use of results taken along the stress path through the center of the ID repair weld. It can be observed that there is not a significant difference in the stress profile across the repair weld width, and that the center of the ID repair weld is where the most severe tensile and residual stresses are experienced. While cracks may grow slightly away from the center of the weld (adjacent to the edge of the weld or in the heat affected zones) using stresses from the center of the weld is appropriate and conservative for evaluating flaws in this region based on this being the location of the most severe tensile stresses. Additionally, a flaw growing along the edge of the DM or ID repair weld may not be oriented perfectly circumferentially (perpendicular to the surface of the pipe), however the ASME code [4] indicates that flaws should be projected into the plane perpendicular to the stresses being evaluated. Here, the concern is a circumferential flaw being driven by axial stresses, thus flaws growing along the edge of the weld or in the heat affected zone can be projected into the circumferential plane, and considered bounded by the evaluation performed herein. This concept is shown in Figure 2-6. Figure 2-5 is provided for additional information. The entire cross section of the original pipe and DM weld is put into compression after a WOL; therefore the location of the stress cut is inconsequential for this condition. Figure 2-4: Axial Residual Stress Contour Plot (psi) at 650 F after ID Repair [1] Figure 2-5: Axial Residual Stress Contour Plot (psi) at 650 F after WOL Repair [1]

17 Figure 2-6: Illustration of Flaw Projection into Circumferential Plane Stress distributions along paths through the center of the ID repair weld are provided at 70 F and 650 F in [1]. This calculation will determine stress corrosion cracking during plant operation; therefore the stress distribution experienced during normal operation is most applicable for studying the growth of a flaw. Normal operating conditions correspond to 650 F and 2,250 psi [1]. These stress distributions can be seen in Figure 2-7 and Figure 2-8 as a function of % through the wall thickness. The flaws evaluated in this analysis are oriented circumferentially, which will be driven by the axial stress. It can be observed in Figure 2-7 that, for an ID Repair, the stress on the inside surface is tensile, and the stress on the outside surface is compressive. This is intuitive, because the repair weld is on the inside surface. As this repair weld cools it will experience tension, and will put the rest of the pipe wall in compression. Figure 2-7 provides stresses from 0% to 100% through the wall thickness. In Figure 2-8, for the WOL, the stress in the DM weld is compressive and the stress in the WOL is tensile. This is intuitive, because the WOL is on the outside surface of the pipe. As the WOL cools it will experience tension, and will put the entire pipe (and DM weld) cross section into compression. Compressive stress will serve to stunt crack growth. Figure 2-8 provides stresses from 0% to 216% of the original wall thickness (percentages above 100% would indicate locations within the weld overlay, which is not susceptible to cracking and not evaluated herein).

18 Figure 2-7: Axial and Hoop Residual Stresses through the Wall Thickness at 650 F after ID Repair Figure 2-8: Axial and Hoop Residual Stresses through the Wall Thickness at 650 F after WOL

19 The stresses that will drive a circumferentially oriented weld are the tensile axial stresses, because they act normal to the surface of the flaw and will cause crack opening. As explained in [1], the stresses already contain the axial stress due to the 2,250 psi internal pressure experienced by the pipe at normal operation. Steady-state thermal stresses due to the temperature gradient are also included, because only the ID of the pipe sees the 650 F normal operating temperature. In addition to the pressure, residual, and thermal stresses contained in the stress distributions from [1], representative stresses due to mechanical loads must be considered. For this analysis, the mechanical loads are summarized in Table 2-1. The loads provided are applicable to normal operating conditions. These stresses are arbitrary and based only on experience. Table 2-1: Mechanical Loading Axial (lb) Bend (in lb) Normal Operation 0 10,000 To determine the stresses, the geometry of the butt weld is needed. This geometry is summarized in Table 2-2. Table 2-2: Butt Weld Geometry Dimensions Butt Weld Inside Radius (R i ) in Outside Radius (R o ) in Thickness (t) in Area (A) In 2 Moment of Inertia (I) In 4 The axial stress is calculated using Equation 2-2. Where: F/A + M c/i Equation 2-2: Combined Mechanical Stress Equation F = axial force for a given loading condition A = cross sectional area M = moment for a given loading condition c = radius where the stress result is desired I = moment of ineria The mechanical stresses calculated for normal operating conditions are provided in Table 2-3. Table 2-3: Normal Operating Condition Mechanical Stresses Stress Units Axial 0 psi Bend (OD) psi Bend (ID) psi

20 The stress distributions from [1] used to make Figure 2-7 and Figure 2-8 are provided in Table 2-4, along with the mechanical stress distribution for normal operation. Table 2-4: Stress Distributions from [1] and Mechanical Stress Distributions (psi) % x coordinate (in) Radius (in) Residual Before WOL Residual After WOL Mechanical 0% % % % % % % % % % % % % % % % % % % % % % % % As indicated in Section 2.1; the input should be applied as a fourth order polynomial of the stress distribution from [1] and the net section bending should be provided individually. The polynomial representation of the axial stress distribution from Figure 2-4 (which includes the axial stress due to pressure, steady-state thermal, and residual stress due to the inside surface weld repair) is provided in Figure 2-9. The stress distribution after the WOL is provided in Figure It can be observed in Figure 2-10 that after the WOL is applied, the entire butt weld is put into compression. Compressive stress will stop the growth of a crack due to PWSCC. For the calculations contained herein, the benefit of the compressive stress was ignored to provide a worst case conservative scenario following the WOL. If the compressive stresses were used, no crack growth could be predicted using K-dependent growth equations.

21 Figure 2-9: Stress Distribution for Normal Operating Conditions (650 F) (Before WOL) Figure 2-10: Stress Distribution for Normal Operating Conditions (650 F) (After WOL)

22 2.3 Crack Growth Methodology and Growth Laws Crack growth in Alloy 82/182 welds is described in MRP-115 [3], which is a public document available through EPRI. The Alloy 82/182 growth rates from [3] also match those available for Alloy 82/182 in Section C-8511 of Section XI of the ASME Code [4]. Specifically, it provides the following equations: Equation 2-3: Alloy 182 Growth Rate at 617 F More generically: Equation 2-4: Generic Alloy 182 Growth Rate where: a dot = crack growth rate at temperature T in in/h Q g = thermal activation energy for crack growth (31.0 kcal/mole) R = universal gas constant (1.103x10-3 kcal/mole- R) T = absolute operating temperature at location of crack ( R) T ref = absolute reference temperature used to normalize data ( R) α = power-law constant (2.47x10-7 ) f alloy = 1.0 for Alloy 182 and 1/2.6 for Alloy 82 f orient = 1.0 K = crack tip stress intensity factor (ksi in) β = exponent = 1.6 Additionally, for the purpose of comparison, the growth rate for stainless steel available in Section C-8520 of [4] was also used to predict growth. This allows observation of the increased predicted crack growth in Alloy 82/182 welds versus typical stainless steel components. The crack growth is performed in 1 month intervals or 0.1 month intervals, depending on the set of calculations. 1 month intervals were adequate for the calculations using the stainless steel growth rate due to the long periods of growth, however the allowable operating period is significantly reduced when using the Alloy 82/182 growth rate, and a shorter interval was used to provide more data points over the shortened operating period. All crack calculations start with a flaw that is 10% through the wall of the component.

23 2.4 Limiting Flaw Depth Methodology The flaw depth can be limited by two things: 1. The stress intensity factor when compared to fracture toughness. 2. A limit load evaluation, which ensures that the remaining cross section of the component does not experience plastic collapse. Therefore, in this set of calculations, the stress intensity factor is compared to the material fracture toughness. The fracture toughness of Alloy 82/182 is not readily available, but based on literature [3] it is expected to be high compared to other materials. In order to come up with a conservative value for use in these calculations, a fracture toughness of 150 ksi in is obtained from [5]. This corresponds to stainless steel. For these calculations, a safety factor of 3 will be applied to this fracture toughness and a value of 50 ksi in will be used going forward. As can be observed in Figure 3-1 and Figure 3-2, the calculated stress intensity factor does not exceed the fracture toughness in these calculations. Allowable flaw depths are therefore established based on limit load per Article C-5000 of [4]. Limit load evaluations are meant to ensure there is not net section failure. For these evaluations, Table C from [4] is used, which applies to Level B, or upset, conditions. As can be observed in Figure 2-11, the maximum allowable depth is 75% of the total wall thickness. However as the ratio of the applied stress to the flow stress increases, and the flaw length increases, the allowable depth will reduce. In Table C of [4]: l f = flaw length Figure 2-11: Table C of [4] Stress Ratio = (Primary Membrane Stress + Primary Bending Stress)/Flow Stress Flow Stress = (σ u + σ y )/2 Since the Alloy 82/182 weld is between 105 Gr. 2 carbon steel nozzle and a 316 stainless steel safe end, the minimum flow stress of the three materials will be used. The material properties in Table 2-5 are applicable at a temperature of 650 F, and are interpolated from data available in [1].

24 Table 2-5: Yield Strength, Ultimate Strength, and Flow Stress at 650 F Alloy 82/ Gr. 2 Carbon Steel 316 Stainless σ y (psi) 46,975 26,675 29,775 σ u (psi) 125,750 90, ,500 σ f (psi) 86,363 58,338 69,638 The secondary stress is not included in the calculation of the stress ratio for limit load evaluations. This means that the stress ratio is based purely on the applied external loads, and will remain constant. The only primary stress to consider for the stress ratio is the bending stress of ksi, which is provided in Table 2-3. The ASME code [4] prescribes a safety factor of 2.3 to these loads, which increases the stress to ksi. This results in a stress ratio of ksi/ ksi, or The data from Table C of [4], interpolated for a stress ratio of 0.402, is used directly in the calculations and included in plots in Section 3.2.1, 3.2.2, and The most limiting allowable flaw depth, as determined by stress intensity factor or limit load calculation per Article C-5000 of [4], is used to establish the final allowable flaw depth for a given flaw and loading.

25 3.0 Results and Discussion 3.1 Calculation of Stress Intensity Factor K I For this calculation, all flaws are assumed to have a starting depth of 10% of the wall thickness. This means the initial flaw depth is in. The selection of 10% is arbitrary, but this is a common starting point for flaw evaluations, and a conservatively large estimate of the minimum detectable flaw during an inspection. The K-solution equations have an upper-bound of 80% of the wall thickness. Therefore, K I is provided in units of ksi in for a range of flaw depths from 10% to 80% of the wall thickness ( in to in) in Figure 3-1 and Figure 3-2. Figure 3-1 provides K I due to only mechanical loads. Figure 3-2 provides K I due to the combined mechanical, thermal, and residual stresses. It can be observed that, for both loading conditions, the stress intensity factor remains below the conservative value of 50 ksi in being used as the fracture toughness, and therefore brittle fracture is not a concern. The K-solution for the mechanical loads increases as the flaw grows (see Figure 3-1). This is intuitive because the solution orients the flaw on the side of the pipe loaded in tension due to the mechanical bending moment. The tensile stresses due to the bending moment serve to open up the flaw. As the crack grows, the stress remains positive and the flaw opens up more, increasing the value of K I. This can be understood by using the most rudimentary solution for K I of a flaw in an infinite plate (Equation 3-1), which is taken from [6]. If the stress value (σ) remains constant and the value for the flaw depth (a) increases, the value for K I will continue to increase. K I = σ (πa) Equation 3-1: K I Solution for a Flaw in an Infinite Plate The K-solution for the combined mechanical, thermal, and residual stresses increases as the flaw grows until a depth of around 0.18 inches (just over 50%) (see Figure 3-2). This is intuitive because the flaw starts in the tensile portion of the residual stress at a depth of 10% (see Figure 2-4 for the stress distribution). As the flaw grows, the stress becomes less tensile, which results in the downward turning K I curve. Once the flaw tip enters the compressive portion of the stress distribution, the K I solution starts to decrease. While the K I solution begins to decrease, it remains a positive value for the flaw sizes evaluated herein, because a majority of the stress distribution over the flaw depth remains tensile, even as portions of the flaw (near the crack tip) are penetrating the compressive region.

26 Figure 3-1: Stress Intensity Factor vs. Flaw Depth for Mechanical Loads Only Figure 3-2: Stress Intensity Factor vs. Flaw Depth for Combined Mechanical, Thermal, and Residual Loads

27 3.2 Calculation of Stress Corrosion Cracking Growth For this calculation, all flaws are assumed to have a starting depth of 10% of the wall thickness. This means the initial flaw depth is in. The selection of 10% is arbitrary, but this is a common starting point for flaw evaluations, and a conservatively large estimate of the minimum detectable flaw during an inspection. As mentioned in Section 3.1, K I does not exceed the fracture toughness, and brittle fracture is not expected. Therefore, the allowable depth is set by the ASME code as described in Section 2.4. As shown in Figure 2-11, the allowable end-of-evaluation period flaw depth-to-thickness ratios for circumferential flaws are provided as a function of the stress ratio and the ratio of the flaw length to pipe circumference. The stress ratio is based purely on the mechanical stresses and remains constant as the crack grows. However, as the flaw grows the flaw length to pipe circumference ratio increases, and the allowable flaw depth-to-thickness ratio decreases. To simplify the calculations, an acceptable flaw depth-to-thickness ratio is calculated for each step in the growth calculations, and the corresponding flaw depth is calculated. This can be seen in the individual plots provided for each aspect ratio as a dashed line. The final acceptable flaw depth is where the dashed line intersects the plot of the flaw growth. For aspect ratios of 2 and 4, this is always 75% of the wall thickness (0.258 inches). However for aspect ratios of 10, the allowable flaw depth was 66.3% of the wall thickness (0.228 inches). For all growth rates, the same trends can be observed, however they occur over different periods of time. For the case considering only mechanical loads, the growth has an upwardbending curve. This is because K I is always increasing with larger flaw depth, and the growth laws raise K I to a power greater than 1. For the combined mechanical, thermal, and residual stresses, the growth curves tend to accelerate early on (due to an increased K I for flaws in the tensile region of the residual stress profile), then level off as the flaw enters the compressive region of the residual stress field. This is intuitive, and serves as a sanity check of the results.

28 3.2.1 Stainless Steel Growth Rate Section provides the expected growth of flaws with aspect ratios of 2, 4, and 8 using the stainless steel growth rate from Section C-8520 of [4]. These results are provided simply for comparative purposes to show the accelerated growth in Alloy 82/182 material versus the growth expected in stainless steels. Figure 3-3: Stainless Steel Growth Calculations - Mechanical Only

29 Figure 3-4: Stainless Steel Growth Calculations (AR = 2) - Mechanical Only Figure 3-5: Stainless Steel Growth Calculations (AR = 4) - Mechanical Only

30 Figure 3-6: Stainless Steel Growth Calculations (AR = 8) - Mechanical Only Figure 3-7: Stainless Steel Growth Calculations - Mechanical + Thermal + Residual

31 Figure 3-8: Stainless Steel Growth Calculations (AR = 2) - Mechanical + Thermal + Residual Figure 3-9: Stainless Steel Growth Calculations (AR = 4) - Mechanical + Thermal + Residual

32 Figure 3-10: Stainless Steel Growth Calculations (AR = 8) - Mechanical + Thermal + Residual

33 3.2.2 Alloy 82/182 Growth Rate at 617 F Section provides the expected growth of flaws with aspect ratios of 2, 4, and 8 using the Alloy 82/182 growth rate from [3]. The growth rate calculation, summarized in Section 2.3, is evaluated at 617 F (as originally presented in MRP-115 [3]). While the residual stress reference [1] indicates that 650 F is the normal operating temperature, the actual fluid temperature experienced by the welds is typically much lower. The temperature presented in MRP-115 [3] is considered a best-estimate of the actual fluid temperature that may be seen by the weld during normal operation. Figure 3-11: Alloy 82/182 Growth Calculations at 617 F - Mechanical Only

34 Figure 3-12: Alloy 82/182 Growth Calculations (AR = 2) at 617 F - Mechanical Only Figure 3-13: Alloy 82/182 Growth Calculations (AR = 4) at 617 F - Mechanical Only

35 Figure 3-14: Alloy 82/182 Growth Calculations (AR = 8) at 617 F - Mechanical Only Figure 3-15: Alloy 82/182 Growth Calculations at 617 F - Mechanical + Thermal + Residual

36 Figure 3-16: Alloy 82/182 Growth Calculations (AR = 2) at 617 F - Mechanical + Thermal + Residual Figure 3-17: Alloy 82/182 Growth Calculations (AR = 4) at 617 F - Mechanical + Thermal + Residual

37 Figure 3-18: Alloy 82/182 Growth Calculations (AR = 8) at 617 F - Mechanical + Thermal + Residual

38 3.2.3 Alloy 82/182 Growth Rate at 650 F Section provides the expected growth of flaws with aspect ratios of 2, 4, and 8 using the Alloy 82/182 growth rate from [3]. The growth rate calculation, summarized in Section 2.3, is evaluated at 650 F, which is the normal operating temperature described in the weld residual stress calculation [1]. This is meant to be a conservative upper bound for growth during normal operation. In practice, if these results were unacceptable and shown to be overly conservative, the results of the evaluation in Section (for a more realistic temperature) could be used. Figure 3-19: Alloy 82/182 Growth Calculations at 617 F - Mechanical Only

39 Figure 3-20: Alloy 82/182 Growth Calculations (AR = 2) at 617 F - Mechanical Only Figure 3-21: Alloy 82/182 Growth Calculations (AR = 4) at 617 F - Mechanical Only

40 Figure 3-22: Alloy 82/182 Growth Calculations (AR = 8) at 617 F - Mechanical Only Figure 3-23: Alloy 82/182 Growth Calculations at 617 F - Mechanical + Thermal + Residual

41 Figure 3-24: Alloy 82/182 Growth Calculations (AR = 2) at 617 F - Mechanical + Thermal + Residual Figure 3-25: Alloy 82/182 Growth Calculations (AR = 4) at 617 F - Mechanical + Thermal + Residual

42 Figure 3-26: Alloy 82/182 Growth Calculations (AR = 8) at 617 F - Mechanical + Thermal + Residual

43 4.0 Summary and Conclusion Table 4-1 and Table 4-2 provide the time (in months) required for a 10% through-wall semielliptical flaw to grow to the allowable depth. Additionally, the same information is provided in years in Table 4-3 and Table 4-4. A general trend that can be observed is that the time needed for the flaw to grow to the maximum allowable depth decreases as the aspect ratio increases, and that the Alloy 82/182 material grows to the maximum allowable depth much more quickly than a crack would in stainless steel. Table 4-1: Time (in months) to Reach Allowable Flaw Depth after Application of a Weld Overlay Mechanical + Residual (After WOL) Time to Reach Allowable Depth (months) Growth Rate Aspect Ratio Stainless Steel Rate 2, , Alloy 82/182 Rate (617F) Alloy 82/182 Rate (650F) Table 4-2: Time (in months) to Reach Allowable Flaw Depth before Application of a Weld Overlay Mechanical + Residual (Before WOL) Time to Reach Allowable Depth (months) Growth Rate Aspect Ratio Stainless Steel Rate Alloy 82/182 Rate (617F) Alloy 82/182 Rate (650F) Table 4-3: Time (in years) to Reach Allowable Flaw Depth after Application of a Weld Overlay Mechanical + Residual (After WOL) Time to Reach Allowable Depth (years) Growth Rate Aspect Ratio Stainless Steel Rate Alloy 82/182 Rate (617F) Alloy 82/182 Rate (650F) Table 4-4: Time (in years) to Reach Allowable Flaw Depth before Application of a Weld Overlay Mechanical + Residual (Before WOL) Time to Reach Allowable Depth (years) Growth Rate Aspect Ratio Stainless Steel Rate Alloy 82/182 Rate (617F) Alloy 82/182 Rate (650F)

44 Alloy 82/182 versus Stainless Steel A comparison can be made between the time it takes a crack to grow to the maximum allowable depth in stainless steel versus Alloy 82/182. It can be observed in Table 4-1 and Table 4-2 that, for a given aspect ratio, the number of months it takes a 10% through-wall semi-elliptical flaw to grow to the maximum allowable depth in stainless steel is an order of magnitude larger than in Alloy 82/182. This serves to show, and confirm, that crack growth occurs much more rapidly in Alloy 82/182 welds than typical stainless steels used in PWR applications. This is true both before and after the application of a WOL. Using the stainless steel growth rate, the amount of time for a 10% through-thickness flaw to grow to the maximum allowable depth after a WOL is applied is longer than the typical year life [12] of a plant. Crack growth in Alloy 82/182 is on the scale of months, which is closer to the amount of time covered between outages, which occur every 18 to 24 months [13]. This helps to explain why crack growth in Alloy 82/182 is such a concern in the nuclear industry. Crack Growth Before and After WOL Application Even for the worst case scenario evaluated ignoring the beneficial residual stress introduced by a WOL, the time before reaching the maximum allowable flaw depth with the WOL is greater than ten times the amount of time it would take without the WOL. Additionally, it should be noted that the compressive residual stresses introduced by the WOL would serve to prevent SCC from initiating on the inside surface and halt (or significantly slow) the growth of any preexisting flaws. This clearly shows the benefit of applying a WOL. According to [13], outages at nuclear plants occur every 18 to 24 months. According to Table 4-2, the time that it would take a crack in Alloy 82/182 to reach the allowable crack depth ranges from 1.6 months to 9.2 months if a WOL were not applied. This means that if a flaw of 10% is not detected during a particular outage, it could grow to be larger than the allowable depth prior to the next outage. In other words, a crack that could grow from initiation to a size that violates code allowable depths in such a short time that it could not be detected using the typical outage/inspection schedule of a nuclear plant. This implies that failure could occur before a crack is detected. As mentioned in [11], WOL are applied preemptively in the nuclear industry. This means they may be applied even if there is no current indication of cracking due to PWSCC. The results in Table 2-1 help to show why preemptive WOL are used, and how they can help a plant ensure they will detect flaws before they can grow to unacceptable depths. According to Table 2-1, it would take a crack in Alloy 82/182 between 25.8 months to months even when the beneficial compressive residual stress due to WOL is ignored. This means that a 10% flaw may go undiscovered in a given outage, and will still not grow to unacceptable depths before the next outage. This ensures the plant will detect a flaw before pipe failure, and can repair any indications that are found. Considering the actual stress distribution after WOL (shown in Figure 2-10), cracks should not initiate due to PWSCC after a WOL due to the highly compressive stress on the inside surface. Furthermore, pre-existing cracks should not grow when in a compressive stress field. These observations, along with the results in Table 4-1 and Table 4-2, show the benefits of applying a preemptive WOL after performing an ID repair weld.

45 5.0 References 1. Wu, Ya T, Residual Stress Study at the Dissimilar Metal Butt Weld due to the Weld Overlay Repair on 2 inch Nozzle Using ANSYS, Rensselear Polytechnic institute, Hartford, CT, April API 579-1/ASME FFS-1, Fitness-for-Service, Annex C, Compendium of Stress Intensity Factor Solutions, June 5, Materials Reliability Program Crack Growth Rates for Evaluating Primary Water Stress Corrosion Cracking (PWSCC) of Alloy 82, 182, and 132 Welds (MRP-115), EPRI, Palo Alto, CA: ASME Boiler & Pressure Vessel Code, Section XI, BWR-VIP-76NP, Revision 1: BWR Vessel and Internals Project, BWR Core Shroud Inspection and Flaw Evaluation Guidelines. EPRI, Palo Alto, CA: NP. 6. Tada, Hiroshi, Paul C. Paris, and George R. Irwin, The Stress Analysis of Cracks Handbook, Third Edition, The American Society of Mechanical Engineers, New York, NY, Cottis, R. A., Stress Corrosion Cracking, National Physical Laboratory, Guides to Good Practice in Corrosion Control, Wollman, M., Stress Corrosion Cracking (SCC), Clausthal University of Technology, May Materials Reliability Program (MRP) Crack Growth Rates for Evaluating Primary Water Stress Corrosion Cracking (PWSCC) of Thick-Wall Alloy 600 Materials (MRP-55) Revision 1. EPRI. Palo Alto, CA: Materials Reliability Program (MRP), Resistance to Primary Water Stress Corrosion Cracking of Alloys 690, 52, and 152 in Pressurized Water Reactors (MRP-111. EPRI, Palo Alto, CA: Materials Reliability Program (MRP), Technical Basis for Preemptive Weld Overlays for Alloy 82/182 Butt Welds in Pressurized Water Reactors (PWRs) (MRP-169) Revision 1-A. EPRI, Palo Alto, CA: Extending the Operational Life Span of Nuclear Plants, Division of Public Information, International Atomic Energy Agency, February 18, Accessed June 20, < What is an Outage?, Clean Energy Insight: Moving Energy Forward, May, 10, Accessed June 20, <