# Dynamic Fracture and Crack Arrest Toughness Evaluation of High-Performance Steel Used in Highway Bridges

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## Abstract

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## 1. Introduction

_{c}represents the ability of a material to resist crack initiation. Similarly, the arrest toughness, K

_{a}, measures a material’s ability to arrest a crack that has already begun to propagate. As with impact testing, fracture toughness and crack arrest testing are often performed at a range of temperatures representing minimum service temperatures for the desired material application, which is often well below room temperature. For the ferritic steels being examined herein, each of the tests mentioned results in lower shelf behavior in terms of absorbed impact energy or toughness at cold temperatures, characterized by brittle cleavage fracture. At elevated temperatures, ductile tearing is the dominant failure mechanism, resulting in an upper shelf of material behavior. The area in between these two shelves is known as the transition region. Marked by increased scatter and mixed failure modes, failure in this region is often the result of small amounts of ductile tearing that reach a carbide or other inclusion that acts as an initiation site for cleavage fracture.

_{o}. This is defined as the temperature at which the median curve is equal to 100 MPa√m (91 ksi√in).

## 2. Materials and Methods

#### 2.1. Charpy V-Notch Impact Testing

#### 2.2. Fracture Toughness Testing

_{o}, reached a nominal length of one-half of the specimen width, 5 mm (0.197 in), resulting in a fatigue crack length of 0.7 mm (0.028 in). At this point, side grooves were machined on the faces on either side of the notch, centered on the specimen in line with the fatigue crack.

_{c}, which includes both plastic and elastic components.

_{e}, and a plastic component, A

_{p}. The area is divided into the two components based on the reciprocal of the slope of the initial elastic loading, called the compliance, C

_{o}. The plastic component of the J-integral, J

_{p}, is calculated by evaluating the plastic area, A

_{p}. Elastic fracture toughness, K

_{e}, is calculated based on specimen geometry and load at fracture, and it is subsequently converted into the elastic component of the J-integral, J

_{e}. Summing J

_{e}and J

_{p}, the resulting J

_{c}can then be converted into an elastic–plastic equivalent critical fracture toughness, K

_{Jc}, using Equation (2):

#### 2.3. Master Curve Analysis

_{min}is equal to 20 MPa√m (18 ksi√in). The subscript x is the desired specimen thickness to normalize to, and the subscript o is the specimen thickness tested. To convert to 1T fracture toughness, B

_{x}is equal to 25.4 mm (1.0 in).

_{oQ}, using Equation (4):

_{i}is the test temperature in degrees Celsius corresponding to the ith specimen, K

_{Jc(i)}is the 1T fracture toughness in MPa√m corresponding to the ith specimen, and δ

_{i}is 1 if the ith specimen is uncensored and 0 if it is censored. Data can be censored by ductile crack growth that exceeds 0.05b

_{o}, in this case 0.25 mm (0.01 in), or by values of fracture toughness that are greater than the limit, K

_{Jclimit}. This size-dependent limit is intended to guarantee adequate levels of crack-tip constraint and is provided in Equation (5):

_{o}is the initial remaining ligament in millimeters, σ

_{ys}is the material 0.2 percent offset yield stress in MPa, and ν is Poisson’s ratio for the material in question. Here, as well as throughout this article, the yield stress is taken to be the nominal yield stress of each grade, 485 and 690 MPa (70 and 100 ksi), respectively. Censored K

_{Jc}values still contribute to the calculation of the reference temperature but not to the validity of the dataset for dataset size requirements. K

_{Jc}values that exceed K

_{Jclimit}are instead replaced with the limit value for use in Equation (4), while the largest uncensored K

_{Jc}value from the dataset is substituted for crack-growth censored K

_{Jc}values.

_{o}. If the sum of the weighted, valid, uncensored specimens is equal to or greater than 1.0, then the dataset is considered to be valid. The weighting factors for each of the three temperature ranges as prescribed by ASTM E1921 [10] are shown in Table 2. Weighting factors for specimens that are tested colder than the reference temperature are lower due to the assumption that they make a reduced accuracy contribution to calculation of the reference temperature.

_{o}

_{,scrn}, is calculated. This is repeated until T

_{o}

_{,scrn}for a given iteration is less than 0.5 °C (0.9 °F) warmer than the previous T

_{o}

_{,scrn}. The dataset is then considered homogeneous if it satisfies the inequality shown in Equation (6):

_{oQ}. To determine β, an equivalent value of the median toughness, K

_{Jc(med)}

^{eq}, is calculated using Equation (7). β has a value of 18.0 °C (32.4 °F) if K

_{Jc(med)}

^{eq}is greater than 83 MPa√m (75 ksi√in), 18.8 °C (33.8 °F) if K

_{Jc(med)}

^{eq}is between 83 and 66 MPa√m (75 and 60 ksi√in), and 20.1 °C (36.2 °F) otherwise.

_{oIN}, is used in place of the provisional reference temperature, T

_{oQ}. This is useful for structural analysis and fitness-for-service applications but is overly conservative for the purposes of material characterization and was thus not used in the analysis contained herein. Instead, all inhomogeneous datasets, regardless of size, are reported using the screened reference temperature, T

_{o}

_{,scrn}, in place of T

_{oQ}. Bimodal and multimodal analysis is possible for datasets with 20 or more specimens but is not discussed in detail as it is not applicable to the data analyzed.

_{Jc(med)}, is constructed at a range of temperatures using Equation (8):

_{o}are in degrees Celsius, and K

_{Jc(med)}is in MPa√m. This curve represents a cumulative probability of failure of 50% based on the probabilistic distribution of flaws throughout a material thickness. Tolerance bounds corresponding to 5% and 95% cumulative probability of failure can also be constructed through the use of Equation (9):

#### 2.4. Crack Arrest Testing

_{o}, is nominally 28 mm (1.1 in) past the load line, approximately 2.5 mm (0.1 in) from base metal. Side grooves, centered on the slot with an included angle of 45 degrees, are employed and are similar to those used in SE(B) specimens. Each groove removes the outer eighth of the specimen; therefore, specimen net thickness, B

_{N}, is equal to 0.75B. Weld deposit, notch, and side-groove details are also shown in Figure 3.

_{n}

_{,max}, calculated using Equation (10):

_{o}/W) is a specimen geometry function that is defined in Equation (11), n is the cycle number, and all other variables are as previously defined.

_{o}

_{,limit}, was reached. A schematic representation of the crack arrest cyclic loading protocol is presented in Figure 5.

_{Ia}, is calculated based on specimen geometry; incremental CMOD values recorded during testing, (δ

_{p})

_{1}and (δ

_{p})

_{n−}

_{1}; the CMOD value at crack initiation, δ

_{o}; and the CMOD value at crack arrest, δ

_{a}. However, if the CMOD limit was reached prior to fracture initiation, the specimen was re-machined to remove material in the plastic zone that formed around the initial notch. The test was then carried out again at a lower temperature in an attempt to initiate fracture. Crack arrest testing proved to be much more difficult than impact and fracture toughness testing. Of the 55 crack arrest specimens fabricated, only 41 fractured, and only 10 of these produced valid K

_{Ia}values.

#### 2.5. Impact Energy–Fracture Toughness Correlations

_{dyn}, is given in Equation (13):

_{dyn}is the dynamic fracture toughness in MPa√m, E is the modulus of elasticity in GPa, and CVN is the impact energy in Joules. The effects of loading rate are then addressed by shifting K

_{dyn}values colder to an equivalent static fracture toughness, K

_{st}. This is accomplished using Equation (14):

_{shift}is the temperature shift from dynamic to static toughness in degrees Celsius. This correlation is limited to CVN values in the lower half of the temperature transition curve. Once shifted, the master curve method can be applied to the resulting K

_{st}values in the same manner as K

_{Jc}values, and a reference temperature estimate, T

_{o}

_{,est}, can be obtained.

_{o}

_{,est}. BS 7910 Method J.2.2 [5] utilizes the test temperatures that correspond to 27 and 40 J (20 and 30 ft-lbf), T

_{27}

_{J}and T

_{40}

_{J}, respectively, to estimate reference temperature through the application of Equations (15) and (16):

_{K}is a constant typically taken to be 25 °C (45 °F) that accounts for scatter in CVN results. If estimates from both equations are available, the reference temperature estimate is the average of the two. Similarly, API 579 “Fitness-For-Service” Annex 9F.4.3.2 [6] estimates reference temperature based on the test temperature corresponding to 28 J (20 ft-lbf), T

_{28}

_{J}. If insufficient data are available to form a full temperature–transition curve, T

_{28}

_{J}is approximated using the following procedure.

_{V}, at the test temperature, T, is calculated. If impact energy data are available at multiple temperatures, the temperature at which C

_{V}is nearest to 28 J (20 ft-lbf) is chosen. Next, if direct measurements of lower and upper shelf impact energies, C

_{V-LS}and C

_{V-US}, are not available, they are estimated using Equations (17) and (18):

_{V-LS}and C

_{V-US}are lower and upper shelf impact energy in Joules, and SA is the percent shear area of the CVN fracture surface. If the shear area is not available, C

_{V-US}can be estimated as 200 J (147.6 ft-lbf) for steels with 0.01% sulfur or less, or 100 J (73.8 ft-lbf) for steels with a sulfur content greater than 0.01%.

_{V-US}. The value of C in degrees Celsius is estimated using Equation (19):

_{28}

_{J}, is then determined using Equation (20):

_{28}

_{J}, 1.0 if T is less than T

_{28}

_{J}, and 0.5 if the hyperbolic tangent fitting parameter C is known rather than estimated through Equation (19). The β factor ensures that this extrapolation is conservative due to the significant degree of uncertainty in the C parameter obtained from Equation (19). Finally, the master curve reference temperature estimate, T

_{o}

_{,est}, is obtained through the use of the estimated or known T

_{28}

_{J}and C

_{V-US}in Equation (21):

_{o}is a temperature shift of +18 °C (+32.4 °F), which corresponds to one standard deviation on the reference temperature estimate from CVN data.

#### 2.6. Stress Intensity Rate Adjustment

_{o}

_{,adj}, from a known or estimated static reference temperature, T

_{o}

_{,sta}:

_{o}

_{,adj}and T

_{o}

_{,sta}are both in degrees Celsius, $\dot{K}$ is the stress intensity rate in MPa√m/s, and Γ is defined by Equation (23), where all variables are as previously defined:

## 3. Results and Discussion

#### 3.1. Charpy V-Notch Results

_{27}

_{J}and T

_{40}

_{J}, and the static reference temperature estimates, T

_{o}

_{,est}, from each of the three methods described above for all eight plates. It can be seen that for the HPS 690 W (100 W) plates, all three methods are in close agreement, with only a maximum of approximately 4 °C (7.2 °F) separating the three reference temperature estimates for a given plate. The range of estimates for the HPS 485 W (70 W) is much larger, with a maximum difference of approximately 39 °C (70.2 °F) between the Barsom and Rolfe and BS 7910 estimates for Plate I. This is likely caused by the increased scatter present in the HPS 485 W (70 W) results.

#### 3.2. Fracture Toughness Results

_{Jc}values that are not crack growth or limit censored are represented by circles, while values that are crack growth or limit censored are represented by an X. For datasets that fail the homogeneity screening criterion shown in Equation (6), such as in Figure 8b, screening censored values are shown by a downwards triangle. In all cases, censored K

_{Jc}values are represented by squares.

_{o}are reported, while the screened reference temperature T

_{o}

_{,scrn}is reported for the four inhomogeneous datasets.

#### 3.3. Crack Arrest Results

_{KIaQ}, appear to be reasonable estimates despite having validity factors below 0.5. The reference temperatures range from −10 to 8.6 °C (14 to 47 °F), significantly warmer than the fracture toughness reference temperatures presented above.

_{o}

_{,sta}, T

_{o}

_{,dyn}, and T

_{KIaQ}is provided in Figure 10. Results from the five HPS 485 W (70 W) plates are shown with solid lines and symbols, and the three HPS 690 W (100 W) plates have dashed lines and open symbols.

_{o}

_{,dyn}and T

_{KIaQ}is similar in magnitude to the difference between T

_{o}

_{,sta}and T

_{o}

_{,dyn}. The grade HPS 485 W (70 W) Plate D exhibits a greater change from fracture to arrest toughness than the grade HPS 690 W (100 W) Plates E and F. From these limited results, a preliminary correlation between T

_{o}

_{,sta}and T

_{KIaQ}is presented in Equation (24):

_{arrest}is an experimental temperature shift in degrees Celsius, and σ

_{ys}is the material 0.2 percent offset yield stress in MPa. This correlation incorporates the observed inverse relationship between yield stress and the difference between fracture and arrest toughness. The values of 207 and 27,600 are constants chosen to provide the best fit to the data available. Resulting values of ΔT

_{arrest}are 99 °C (179 °F) for HPS 485 W (70 W) and 57 °C (103 °F) for HPS 690 W (100 W). This closely matches the experimentally obtained differences between T

_{o}

_{,sta}and T

_{KIaQ}of 98 °C (176 °F) for Plate D, 46 °C (83 °F) for Plate E, and 64 °C (115 °F) for Plate F.

#### 3.4. Reference Temperature Rate Adjustment

_{o}

_{,adj}, from a known static reference temperature, T

_{o}

_{,sta}, using Equation (22), the stress intensity rate $\dot{K}$ was taken as the mean of the individual stress intensity rates for each of the eight plates. Mean stress intensity rates ranged from 2720 to 3650 MPa√m/s (2470 to 3320 ksi√in/s) and are summarized in Table 5. The difference between T

_{o}

_{,sta}and T

_{o}

_{,adj}can be considered to be a theoretical rate-dependent temperature shift, hereafter referred to as ΔT

_{theoretical}. This temperature shift, as well as the temperature shift prescribed by the second stage of the Barsom and Rolfe method shown in Equation (14), T

_{shift}, and the experimental temperature shift between quasi-static and dynamic refence temperatures contained in Table 4, ΔT

_{experimental}, are shown in Table 5. It should be noted that the large ΔT

_{experimental}for Plates D and H is in part due to the dynamic reference temperatures for these plates being screened to the warmer T

_{o}

_{,scrn}while the quasi-static reference temperatures are not. Because both quasi-static and dynamic reference temperatures for Plate I are screened, the difference is more similar to Plates A and J.

_{shift}is consistently greater in magnitude than ΔT

_{theoretical}by approximately 10 °C (18 °F) for HPS 690 W (100 W) and by 18 °C (32 °F) for HPS 485 W (70 W). The Barsom and Rolfe temperature shift is reasonably accurate for all plates except for Plates D and H, likely due to the use of screened reference temperatures mentioned above. The average difference between T

_{shift}and ΔT

_{experimental}, not including Plates D and H, is 1.9 °C (3.4 °F), with a maximum error of 11 °C (20 °F). The theoretical temperature shift from Equation (22) closely matches the experimental temperature shift for Plate C but underestimates the difference between quasi-static and dynamic reference temperature for all other plates. In this case, an underestimation of temperature shift results in an estimated dynamic reference temperature, T

_{o}

_{,adj}, that is colder than the experimentally determined dynamic reference temperature, T

_{o}

_{,dyn}. Because colder reference temperatures correspond to increased material fracture toughness, this error results in an unconservative estimate of dynamic fracture toughness.

#### 3.5. Correlation Evaluations

_{dyn}estimates obtained using Equation (13) and neglecting the temperature shift of Equation (14). Dynamic reference temperature estimates for the BS 7910 and API 579 correlations adjust static estimates based on experimental stress intensity rate using the rate adjustment shown in Equation (22).

## 4. Conclusions

_{experimental}and T

_{shift}for all plates, excluding D and H, of 1.9 °C (3.4 °F).

_{Ia}values of the 55 specimens fabricated and tested. Due to the small dataset sizes, valid arrest toughness reference temperatures were not obtained. The estimated crack arrest reference temperatures were between −10 and 9 °C (14 and 47 °F). A correlation between quasi-static and arrest reference temperature in the form of a temperature shift is presented. This correlation is based on the observed inverse relationship between yield stress and the difference between fracture initiation and arrest toughness. The proposed correlation may be used by engineers to estimate the crack arrest capability of a material when only fracture initiation information is available.

## Author Contributions

## Funding

## Institutional Review Board Statement

## Informed Consent Statement

## Data Availability Statement

## Acknowledgments

## Conflicts of Interest

## Notation

a | Asymptotic minimum of sigmoid fit; |

a_{o} | Initial crack size; |

A | Total area under load-CMOD curve; |

A_{e} | Elastic component of area under load-CMOD curve; |

A_{p} | Plastic component of area under load-CMOD curve; |

b | Asymptotic maximum of sigmoid fit; |

b_{o} | Initial remaining ligament size; |

B | Specimen thickness; |

B_{N} | Net specimen thickness; |

B_{o} | Initial specimen thickness used in thickness adjustment; |

B_{x} | Desired specimen thickness for thickness adjustment; |

c | Location parameter of sigmoid fit; |

C(T) | Compact specimen geometry; |

C_{o} | Compliance, the reciprocal of the initial elastic slope; |

C_{V} | Average CVN impact energy used in API 579 correlation approach; |

CVN(T) | Charpy V-notch impact energy at a given temperature, T; |

C_{V-LS} | Charpy V-notch impact energy at lower shelf behavior; |

C_{V-US} | Charpy V-notch impact energy at upper shelf behavior; |

d | Slope parameter of sigmoid fit; |

E | Modulus of elasticity; |

J_{c} | Critical value of J-integral at failure; |

J_{e} | Elastic component of J-integral calculation; |

J_{p} | Plastic component of J-integral calculation; |

$\dot{K}$ | Stress intensity rate; |

K_{c} | Critical fracture toughness; |

K_{dyn} | Dynamic fracture toughness; |

K_{e} | Elastic fracture toughness; |

K_{Ia} | Crack arrest fracture toughness; |

K_{Jc} | Elastic–plastic equivalent fracture toughness converted from critical J-integral; |

K_{Jc(0.xx)} | Elastic–plastic fracture toughness tolerance limit of xx per cent; |

K_{Jc(i)} | Elastic–plastic fracture toughness of ith specimen; |

K_{Jclimit} | Elastic–plastic fracture toughness limit in master curve analysis; |

K_{Jc(med)} | Median elastic–plastic fracture toughness; |

K_{Jc(med)}^{eq} | Equivalent median toughness used in inhomogeneity screening process |

K_{Jc(o)} | Initial elastic–plastic fracture toughness used in thickness adjustment; |

K_{Jc(x)} | Size adjusted fracture toughness for specimen of thickness B_{N}; |

K_{min} | Absolute minimum fracture toughness equal to 20 MPa√m (18 ksi√in); |

m | Scale parameter of sigmoid fit; |

n | Load cycle count in crack arrest testing; |

r | Number of uncensored data used to determine provisional reference temperature; |

SA | Percent shear area of CVN fracture surface; |

SE(B) | Single edge bend specimen geometry; |

T | Test temperature; |

T_{27}_{J} | Test temperature corresponding to CVN of 27 J; |

T_{28}_{J} | Test temperature corresponding to CVN of 28 J; |

T_{40}_{J} | Test temperature corresponding to CVN of 40 J; |

T_{i} | Test temperature of ith specimen; |

T_{k} | Temperature adjustment based on CVN scatter in BS 7910; |

T_{KIaQ} | Provisional arrest reference temperature; |

T_{o} | Reference temperature where median toughness equals 100 MPa√m (91 ksi√in); |

T_{o,adj} | Rate adjusted reference temperature; |

T_{o,est} | Prediction of static reference temperature correlated from CVN data; |

T_{oQ} | Provisional reference temperature prior to validation; |

T_{o,scrn} | Reference temperature following inhomogeneity screening procedure; |

T_{o,sta} | Quasi-static rate reference temperature; |

T_{shift} | Temperature shift rate adjustment in Barsom and Rolfe Two-Stage correlation; |

W | Specimen width; |

xT | Specimen thickness designation with x in inches; |

β | Sample size uncertainty factor used in inhomogeneity screening process; |

δ_{a} | CMOD corresponding to crack arrest; |

δ_{i} | Validity of ith specimen in master curve analysis, valid = 1, invalid = 0; |

δ_{n,max} | Maximum CMOD limit for a given load cycle in crack arrest testing; |

δ_{o} | CMOD corresponding to crack initiation in crack arrest testing; |

δ_{o,limit} | Overall CMOD limit for crack arrest testing; |

(δ_{p})_{1} | CMOD offset at the end of the first crack arrest load cycle; |

(δ_{p})_{n−1} | CMOD offset at the beginning of the last crack arrest load cycle; |

ΔT_{experimental} | Difference between quasi-static and dynamic reference temperature; |

ΔT_{theoretical} | Theoretical rate-dependent shift applied to reference temperature; |

Γ | Material fitting coefficient used in Wallin rate adjustment; |

σ_{ys} | 0.2% offset yield strength; |

ν | Poisson’s ratio. |

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**Figure 7.**1T fracture toughness results for Plate C at (

**a**) quasi-static and (

**b**) dynamic stress intensity rates.

**Figure 8.**1T fracture toughness results for Plate D at (

**a**) quasi-static and (

**b**) dynamic stress intensity rates.

Plate | Grade MPa (ksi) | Thickness mm (in) | CVN Impact | Charpy-Sized SE(B) | Crack Arrest | |
---|---|---|---|---|---|---|

Static | Dynamic | |||||

A | 485 (70) | 25.4 (1.00) | 24 (24) | 9 (8) | 11 (10) | - |

C | 690 (100) | 19.1 (0.75) | 24 (24) | 8 (8) | 8 (8) | - |

D | 485 (70) | 63.5 (2.50) | 24 (23) | 13 (10) | 17 (17) | 10 (5) |

E | 690 (100) | 38.1 (1.50) | 24 (23) | 15 (12) | 13 (13) | 5 (2) |

F | 690 (100) | 50.8 (2.00) | 24 (24) | 15 (11) | 12 (12) | 3 (3) |

H | 485 (70) | 31.8 (1.25) | 24 (24) | 10 (10) | 8 (8) | - |

I | 485 (70) | 31.8 (1.25) | 24 (24) | 9 (8) | 8 (8) | - |

J | 485 (70) | 38.1 (1.50) | 24 (23) | 8 (8) | 10 (10) | - |

Temperature Range, T–T_{o}°C (°F) | Weighting Factor |
---|---|

+50 to −14 (+90 to −25) | 1/6 |

−15 to −35 (−27 to −63) | 1/7 |

−36 to −50 (−65 to −90) | 1/8 |

Grade MPa (ksi) | Plate | T_{27J}°C (°F) | T_{40J}°C (°F) | T_{o}_{,est}, °C (°F) | ||
---|---|---|---|---|---|---|

B and R | BS 7910 | API 579 | ||||

690 (100) | C | −71 (−96) | −63 (−81) | −62 (−80) | −63 (−81) | −59 (−74) |

E | −81 (−113) | −73 (−100) | −73 (−100) | −73 (−99) | −69 (−93) | |

F | −88 (−127) | −78 (−108) | −77 (−106) | −79 (−110) | −76 (−105) | |

485 (70) | A | −119 (−183) | −110 (−166) | −143 (−225) | −111 (−167) | −131 (−204) |

D | −103 (−153) | −98 (−144) | −105 (−157) | −96 (−141) | −115 (−175) | |

H | −89 (−129) | −86 (−124) | −114 (−174) | −84 (−119) | −102 (−151) | |

I | −91 (−132) | −87 (−124) | −124 (−191) | −85 (−121) | −105 (−156) | |

J | −104 (−156) | −97 (−142) | −125 (−194) | −97 (−142) | −116 (−177) |

Grade MPa (ksi) | Plate | Stress Intensity Rate | Validity | Reference Temperature °C (°F) |
---|---|---|---|---|

690 (100) | C | Quasi-Static | 1.262 | −80 (−111) |

Dynamic | 1.167 | −55 (−66) | ||

E | Quasi-Static | 1.476 | −102 (−151) | |

Dynamic | 1.833 | −56 (−69) | ||

F | Quasi-Static | 1.214 | −95 (−139) | |

Dynamic | 1.619 | −55 (−68) | ||

485 (70) | A | Quasi-Static | 1.000 | −173 (−279) |

Dynamic | 1.500 | −113 (−172) | ||

D | Quasi-Static | 1.000 | −181 (−294) | |

Dynamic | - | −96 (−142) | ||

H | Quasi-Static | 1.000 | −161 (−258) | |

Dynamic | - | −62 (−79) | ||

I | Quasi-Static | - | −161 (−259) | |

Dynamic | - | −98 (−145) | ||

J | Quasi-Static | 1.095 | −134 (−209) | |

Dynamic | 1.196 | −64 (−84) |

Grade MPa (ksi) | Plate | Mean Stress Intensity Rate, $\dot{K}$ MPa√m/s (ksi√in/s) | ΔT_{theoretical}°C (°F) | T_{shift}°C (°F) | ΔT_{experimental}°C (°F) |
---|---|---|---|---|---|

690 (100) | C | 2720 (2470) | 24 (43) | 36 (65) | 25 (45) |

E | 3270 (2980) | 27 (49) | 36 (65) | 46 (82) | |

F | 3230 (2940) | 26 (47) | 36 (65) | 40 (71) | |

485 (70) | A | 3020 (2750) | 43 (78) | 61 (110) | 60 (108) |

D | 2820 (2570) | 42 (75) | 61 (110) | 85 (153) | |

H | 3540 (3220) | 45 (81) | 61 (110) | 99 (179) | |

I | 3350 (3050) | 44 (80) | 61 (110) | 63 (113) | |

J | 3650 (3320) | 44 (79) | 61 (110) | 70 (126) |

Grade MPa (ksi) | Plate | Stress Intensity Rate | T_{o}°C (°F) | T_{o}_{,est}°C (°F) | ||
---|---|---|---|---|---|---|

B and R | BS 7910 | API 579 | ||||

690 (100) | C | Quasi-Static | −80 (−111) | −62 (−80) | −63 (−81) | −59 (−74) |

Dynamic | −55 (−66) | −26 (−15) | −41 (−42) | −37 (−35) | ||

E | Quasi-Static | −102 (−151) | −73 (−100) | −73 (−99) | −69 (−93) | |

Dynamic | −56 (−69) | −37 (−35) | −49 (−57) | −46 (−50) | ||

F | Quasi-Static | −95 (−139) | −77 (−106) | −79 (−110) | −76 (−105) | |

Dynamic | −55 (−68) | −40 (−41) | −54 (−66) | −52 (−62) | ||

485 (70) | A | Quasi-Static | −173 (−279) | −143 (−225) | −111 (−167) | −131 (−204) |

Dynamic | −113 (−172) | −82 (−115) | −71 (−96) | −89 (−128) | ||

D | Quasi-Static | −181 (−294) | −105 (−157) | −96 (−141) | −115 (−175) | |

Dynamic | −96 (−142) | −44 (−47) | −59 (−74) | −75 (−103) | ||

H | Quasi-Static | −161 (−258) | −114 (−174) | −84 (−119) | −102 (−151) | |

Dynamic | −62 (−79) | −53 (−64) | −48 (−54) | −63 (−81) | ||

I | Quasi-Static | −161 (−259) | −124 (−191) | −85 (−121) | −105 (−156) | |

Dynamic | −98 (−145) | −63 (−81) | −49 (−56) | −65 (−85) | ||

J | Quasi-Static | −134 (−209) | −125 (−194) | −97 (−142) | −116 (−177) | |

Dynamic | −64 (−84) | −64 (−84) | −58 (−72) | −75 (−102) |

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## Share and Cite

**MDPI and ACS Style**

Collins, W.N.; Yount, T.D.; Sherman, R.J.; Leon, R.T.; Connor, R.J.
Dynamic Fracture and Crack Arrest Toughness Evaluation of High-Performance Steel Used in Highway Bridges. *Materials* **2023**, *16*, 3402.
https://doi.org/10.3390/ma16093402

**AMA Style**

Collins WN, Yount TD, Sherman RJ, Leon RT, Connor RJ.
Dynamic Fracture and Crack Arrest Toughness Evaluation of High-Performance Steel Used in Highway Bridges. *Materials*. 2023; 16(9):3402.
https://doi.org/10.3390/ma16093402

**Chicago/Turabian Style**

Collins, William N., Tristan D. Yount, Ryan J. Sherman, Roberto T. Leon, and Robert J. Connor.
2023. "Dynamic Fracture and Crack Arrest Toughness Evaluation of High-Performance Steel Used in Highway Bridges" *Materials* 16, no. 9: 3402.
https://doi.org/10.3390/ma16093402