Structural Integrity and Failure

**Chapter 1**

**Abstract**

fatigue crack growth rate.

safety of welded joints were also developed.

damages lead to catastrophic fractures.

enough time, reach a critical size and cause breakage.

**1. Introduction**

**3**

Treatment Analysis of Welding

The largest number of welded structures in operating conditions is exposed to variable loads, which is why the share of fatigue fracture in the failure of welded structures is higher than others. The essence of construction with fracture safety is that the structure can withstand the designed load in the designed time. If a crack is detected during operation, it is possible to predict the development of damage during the service life as well as the load-bearing capacity of the structure depending on the development of damage. The paper describes a new system for monitoring fatigue crack growth, which is based on the change in the resistance of the measuring foil during crack growth. The system is compatible with the basic settings of the ASTM E647–86 standard, which refers to the determination of the

**Keywords:** welded joint, fatigue crack, fatigue threshold, crack growth rate

Mass application of welded structures began with the development of welding procedures on the one hand and the development of steels with suitable properties on the other. Along with welding processes, in parallel, methods for assessing the

Construction materials and welded joints can contain defects and microcracks that are the beginnings of fractures. Exploitation conditions can lead to cracking even if there are no defects in the material, e.g., at places of stress concentration caused by the design of the structure. Under the influence of unfavorable exploitation factors, such as fatigue and corrosion, cracks can grow steadily, and after

Fatigue is the phenomenon of gradual destruction of a material due to the long-term action of a periodically changing load. Damage to structures, caused by material fatigue, represents 50 ÷ 90% of all damage to structures in exploitation [1]. The significance of fatigue damage is obvious because a large number of such

The traditional, well-known, S-N approach is based on the experimental determination of the dependence of the stress amplitude from the number of cycles to fracture. This standard method is built into many standards and regulations and is widely used in the design of welded and other structures. In this test, as a rule, only the number of changes of the load to fracture under the action of a constant range load

Structure in the Presence of a

Crack Type Defects

*Mersida Manjgo and Meri Burzic*

### **Chapter 1**

## Treatment Analysis of Welding Structure in the Presence of a Crack Type Defects

*Mersida Manjgo and Meri Burzic*

### **Abstract**

The largest number of welded structures in operating conditions is exposed to variable loads, which is why the share of fatigue fracture in the failure of welded structures is higher than others. The essence of construction with fracture safety is that the structure can withstand the designed load in the designed time. If a crack is detected during operation, it is possible to predict the development of damage during the service life as well as the load-bearing capacity of the structure depending on the development of damage. The paper describes a new system for monitoring fatigue crack growth, which is based on the change in the resistance of the measuring foil during crack growth. The system is compatible with the basic settings of the ASTM E647–86 standard, which refers to the determination of the fatigue crack growth rate.

**Keywords:** welded joint, fatigue crack, fatigue threshold, crack growth rate

### **1. Introduction**

Mass application of welded structures began with the development of welding procedures on the one hand and the development of steels with suitable properties on the other. Along with welding processes, in parallel, methods for assessing the safety of welded joints were also developed.

Construction materials and welded joints can contain defects and microcracks that are the beginnings of fractures. Exploitation conditions can lead to cracking even if there are no defects in the material, e.g., at places of stress concentration caused by the design of the structure. Under the influence of unfavorable exploitation factors, such as fatigue and corrosion, cracks can grow steadily, and after enough time, reach a critical size and cause breakage.

Fatigue is the phenomenon of gradual destruction of a material due to the long-term action of a periodically changing load. Damage to structures, caused by material fatigue, represents 50 ÷ 90% of all damage to structures in exploitation [1]. The significance of fatigue damage is obvious because a large number of such damages lead to catastrophic fractures.

The traditional, well-known, S-N approach is based on the experimental determination of the dependence of the stress amplitude from the number of cycles to fracture.

This standard method is built into many standards and regulations and is widely used in the design of welded and other structures. In this test, as a rule, only the number of changes of the load to fracture under the action of a constant range load

### *Structural Integrity and Failure*

is determined, and the standard only requires information on the magnitude of the stress at which crack and fracture initiation does not occur after a certain number of cycles (usually between 10<sup>6</sup> and 10<sup>8</sup> cycles).

In the presence of cracks, the question arises of its development under the action of a variable load.

Fatigue crack growth is a very complex process that depends on a number of variables [2]:


The method of fracture mechanics is based on linearly elastic fracture mechanics and originates from the Paris' law of 1962, and is still applied, although the impact of large plastic deformations around the crack tip has not been fully taken into account. The constants that occur such as "C" i "m" in Paris' law da/dN=C ΔKm, must be determined separately for each material and the specific test conditions. These data are essential for three types of fatigue analysis:

**Figure 1.**

**Figure 2.**

**Figure 3.**

**5**

*Alternating load scheme. R = 1.*

*Dynamic specimen according to ASTM E466.*

*S-N diagram of speciemn taken out of butt welded joint and tested at room temperature.*

*Treatment Analysis of Welding Structure in the Presence of a Crack Type Defects*

*DOI: http://dx.doi.org/10.5772/intechopen.94832*


### **2. Determination of dynamic characteristics of welded joint**

Metal fatigue is defined as the process of cumulative damage under the action of variable load, which is manifested by the appearance of fatigue cracks and fractures. The fatigue strength of welded joints is determined by testing the specimens at variable load until a crack or fracture occurs.

The test was performed on a high-frequency AMSLER pulsator. The highfrequency pulsator can achieve a sinusoidal alternating load in the range from 100 kN to +100 kN. In order to more fully assess the behavior of the material under the action of variable load, and having in mind the dimensions of the specimen, the most critical case of the action of variable load was made, namely alternating variable load tension - pressure (R = 1), **Figure 1**.

It is clear that the strength at high cyclic fatigue depends on the properties of the constituents of the welded joint. Therefore, data are needed for BM and WM, but also for HAZ, which makes testing of welded joints in high-cyclic fatigue complex and expensive. The aim of the test is to determine the points in the S-N diagram (construction of the Wehler curve) and to determine the permanent dynamic strength *S <sup>f</sup>* . The test procedure as well as the specimen are defined according to ASTM E466 [3]. The appearance of the test tube with variable load is shown in **Figure 2**.

The determination of the maximum dynamic stress at which no crack-type error is initiated in smooth construction forms is shown graphically in the form of Weller curves (S-N diagrams) in **Figure 3**. for butt-welded joint specimen and **Figure 4** for specimen removed from BM.

*Treatment Analysis of Welding Structure in the Presence of a Crack Type Defects DOI: http://dx.doi.org/10.5772/intechopen.94832*

**Figure 1.** *Alternating load scheme. R = 1.*

is determined, and the standard only requires information on the magnitude of the stress at which crack and fracture initiation does not occur after a certain number of

In the presence of cracks, the question arises of its development under the action

Fatigue crack growth is a very complex process that depends on a number of

• the intensity of the effective stress field at the crack tip defined by the K-factor;

The method of fracture mechanics is based on linearly elastic fracture mechanics and originates from the Paris' law of 1962, and is still applied, although the impact of large plastic deformations around the crack tip has not been fully taken into account. The constants that occur such as "C" i "m" in Paris' law da/dN=C ΔKm, must be determined separately for each material and the specific test conditions.

Metal fatigue is defined as the process of cumulative damage under the action of variable load, which is manifested by the appearance of fatigue cracks and fractures. The fatigue strength of welded joints is determined by testing the specimens at

The test was performed on a high-frequency AMSLER pulsator. The highfrequency pulsator can achieve a sinusoidal alternating load in the range from 100 kN to +100 kN. In order to more fully assess the behavior of the material under the action of variable load, and having in mind the dimensions of the specimen, the most critical case of the action of variable load was made, namely alternating

It is clear that the strength at high cyclic fatigue depends on the properties of the constituents of the welded joint. Therefore, data are needed for BM and WM, but also for HAZ, which makes testing of welded joints in high-cyclic fatigue complex and expensive. The aim of the test is to determine the points in the S-N diagram (construction of the Wehler curve) and to determine the permanent dynamic strength *S <sup>f</sup>* . The test procedure as well as the specimen are defined according to ASTM E466 [3].

The determination of the maximum dynamic stress at which no crack-type error is initiated in smooth construction forms is shown graphically in the form of Weller curves (S-N diagrams) in **Figure 3**. for butt-welded joint specimen and **Figure 4** for

• environment (aggressiveness, temperature, humidity),

These data are essential for three types of fatigue analysis:

• to accurately determine fatigue crack behavior

• to estimate the life of the structure

variable load until a crack or fracture occurs.

variable load tension - pressure (R = 1), **Figure 1**.

• to calculate fatigue damage

specimen removed from BM.

**4**

• mechanical and metallurgical characteristics of the material

**2. Determination of dynamic characteristics of welded joint**

The appearance of the test tube with variable load is shown in **Figure 2**.

cycles (usually between 10<sup>6</sup> and 10<sup>8</sup> cycles).

of a variable load.

*Structural Integrity and Failure*

• type and form of load;

variables [2]:

**Figure 2.** *Dynamic specimen according to ASTM E466.*

**Figure 3.** *S-N diagram of speciemn taken out of butt welded joint and tested at room temperature.*

intensity factor, crack opening, and contour J-integral. Paris' crack growth law, which determines the dependence of the load and the corresponding range of stress intensity factors, with the crack growth rate per cycle, is generally accepted today [4].

*Treatment Analysis of Welding Structure in the Presence of a Crack Type Defects*

The need to introduce fracture mechanics into the study of fatigue behavior

which were exposed to loads with a constant amplitude *Δ*P, **Figure 5** [5].

Many data on fatigue crack growth were obtained by examining CT specimen,

In **Figure 5** a typical form of crack length dependence on the number of load cycles *a-*N for three load range levels ΔP = const where ΔP1 < ΔP2 < ΔP3 is shown. It is noticeable that with the increasing number of cycles N and crack length "a", the crack growth rate defined by the slope of the tangent increases steadily. Also, with the increase in the load range ΔP, there is a faster increase of the speed gradient. In other words, the crack, for example, of length a1 in **Figure 1**, grows

Numerous theoretically and empirically defined dependencies in the form *da/dN = f(P, a)* can be found in the literature, which emphasizes the importance of load and cracks length. The first to define the range of stress intensity factor *ΔK = f(σ, a)* in the form of fatigue crack growth rate as a basic parameter were Paris

The crack growth rates *da/dN* as a function of ΔK are determined from the corresponding curve a-N, graphically, or numerically. The experimental results presented on the double-logarithmic scale usually have a characteristic S-shape,

It is noticeable that the crack propagation is initially accelerated (area I), then passes into the phase of stable growth (area II), to finally pass into the phase of

*Crack growth dependence a = f(N) for three levels of load ranges ΔP = const where ΔP1 < ΔP2 < ΔP3.*

<sup>1</sup>*=*<sup>2</sup> <sup>¼</sup> *<sup>Y</sup>*Δ*σ π*ð Þ *<sup>a</sup>*

<sup>1</sup>*=*<sup>2</sup> (1)

**3.1 Fatigue crack growth rate da/dN i ΔKth: Paris' law**

*DOI: http://dx.doi.org/10.5772/intechopen.94832*

arose from the analysis of crack growth under cyclic loading.

faster at the load amplitude ΔP3 than at the load ΔP2 or ΔP1.

*ΔK* ¼ *Kmax* � *Kmin* ¼ *Y*ð Þ *σmax* � *σmin* ð Þ *πa*

and co-workers [5].

**Figure 5.**

**7**

schematically shown in **Figure 6**.

critical crack expansion (area III).

### **Figure 4.**

*S-N diagram of speciemn extracted from the new BM and tested at room and operating temperatures.*

To construct one Weller curve and determine the permanent dynamic strength, it is necessary to test the specimen at 6 to 7 different load levels. According to the ASTM E 466 standard, three specimen were tested for each load level, which is a total of 21 specimen. Therefore, this test is extremely expensive and justified when design data are required, primarily from the aspect of fatigue and fracture mechanics. So when designing parts exposed to long-term variable load in the total design life of the structure.

The traditional, well-known, S-N approach is based on the experimental determination of the dependence of the stress amplitude from the number of cycles to fracture.

### **3. Fatigue analysis from fracture mechanics angle**

The origin of the fault may be related to design and construction, technology, and producing the structure, control, and testing. The continuance of construction then depends only on the possibility and conditions of crack growth from the initial failure. The essence of construction with fracture safety is that the structure can withstand the designed load in the designed time. If a crack is detected during operation, it's possible to predict the development of damage during the service life as well as the load-bearing capacity of the structure depending on the development of damage [4].

The most important characteristics for the operational safety of structures are the ones that describe the appearance and growth of cracks under the influence of variable load. A generally accepted characteristic, in this case, is fatigue strength. Accordingly, the design of structural parts based on possible material fatigue is based on the use of fatigue strength, and empirical recommendations, derived from the analysis of parts failure in operation and extensive testing [4].

The appearance of a fatigue crack requires that the behavior of the material around the crack tip is considered based on the micromechanical aspect. The initiation of cracks and their growth condition that the micromechanical aspect of the behavior of the material becomes important for the assessment of the operational safety of structural parts. The existence of a singularity in the form of a crack tip indicates the application of fracture mechanics and its parameters, such as stress

intensity factor, crack opening, and contour J-integral. Paris' crack growth law, which determines the dependence of the load and the corresponding range of stress intensity factors, with the crack growth rate per cycle, is generally accepted today [4].

### **3.1 Fatigue crack growth rate da/dN i ΔKth: Paris' law**

The need to introduce fracture mechanics into the study of fatigue behavior arose from the analysis of crack growth under cyclic loading.

Many data on fatigue crack growth were obtained by examining CT specimen, which were exposed to loads with a constant amplitude *Δ*P, **Figure 5** [5].

In **Figure 5** a typical form of crack length dependence on the number of load cycles *a-*N for three load range levels ΔP = const where ΔP1 < ΔP2 < ΔP3 is shown.

It is noticeable that with the increasing number of cycles N and crack length "a", the crack growth rate defined by the slope of the tangent increases steadily. Also, with the increase in the load range ΔP, there is a faster increase of the speed gradient. In other words, the crack, for example, of length a1 in **Figure 1**, grows faster at the load amplitude ΔP3 than at the load ΔP2 or ΔP1.

Numerous theoretically and empirically defined dependencies in the form *da/dN = f(P, a)* can be found in the literature, which emphasizes the importance of load and cracks length. The first to define the range of stress intensity factor *ΔK = f(σ, a)* in the form of fatigue crack growth rate as a basic parameter were Paris and co-workers [5].

$$
\Delta K = K\_{\text{max}} - K\_{\text{min}} = Y(\sigma\_{\text{max}} - \sigma\_{\text{min}}) \left(\pi a\right)^{1/2} = Y \Delta \sigma \left(\pi a\right)^{1/2} \tag{1}
$$

The crack growth rates *da/dN* as a function of ΔK are determined from the corresponding curve a-N, graphically, or numerically. The experimental results presented on the double-logarithmic scale usually have a characteristic S-shape, schematically shown in **Figure 6**.

It is noticeable that the crack propagation is initially accelerated (area I), then passes into the phase of stable growth (area II), to finally pass into the phase of critical crack expansion (area III).

**Figure 5.** *Crack growth dependence a = f(N) for three levels of load ranges ΔP = const where ΔP1 < ΔP2 < ΔP3.*

To construct one Weller curve and determine the permanent dynamic strength, it is necessary to test the specimen at 6 to 7 different load levels. According to the ASTM E 466 standard, three specimen were tested for each load level, which is a total of 21 specimen. Therefore, this test is extremely expensive and justified when design data are required, primarily from the aspect of fatigue and fracture mechanics. So when designing parts exposed to long-term variable load in the total design

*S-N diagram of speciemn extracted from the new BM and tested at room and operating temperatures.*

The traditional, well-known, S-N approach is based on the experimental determination of the dependence of the stress amplitude from the number of cycles to

The origin of the fault may be related to design and construction, technology, and producing the structure, control, and testing. The continuance of construction then depends only on the possibility and conditions of crack growth from the initial failure. The essence of construction with fracture safety is that the structure can withstand the designed load in the designed time. If a crack is detected during operation, it's possible to predict the development of damage during the service life as well as the load-bearing capacity of the structure depending on the development

The most important characteristics for the operational safety of structures are the ones that describe the appearance and growth of cracks under the influence of variable load. A generally accepted characteristic, in this case, is fatigue strength. Accordingly, the design of structural parts based on possible material fatigue is based on the use of fatigue strength, and empirical recommendations, derived from

The appearance of a fatigue crack requires that the behavior of the material around the crack tip is considered based on the micromechanical aspect. The initiation of cracks and their growth condition that the micromechanical aspect of the behavior of the material becomes important for the assessment of the operational safety of structural parts. The existence of a singularity in the form of a crack tip indicates the application of fracture mechanics and its parameters, such as stress

the analysis of parts failure in operation and extensive testing [4].

**3. Fatigue analysis from fracture mechanics angle**

life of the structure.

*Structural Integrity and Failure*

fracture.

**Figure 4.**

of damage [4].

**6**

*da*

*Treatment Analysis of Welding Structure in the Presence of a Crack Type Defects*

*da*

*dN* <sup>¼</sup> *<sup>C</sup>* � ð Þ <sup>Δ</sup>*<sup>K</sup>* � <sup>Δ</sup>*Kth*

**3.2 Determination of fatigue crack growth parameters**

based on a simple physical model:

*DOI: http://dx.doi.org/10.5772/intechopen.94832*

valid in the case of type i loads.

Np, i.e., Nu = Ni + Np, **Figure 7** [5].

stress intensity range at the crack tip:

**Figure 7.**

**9**

*da*

*dN* <sup>¼</sup> *<sup>C</sup>* � <sup>Δ</sup>*K<sup>m</sup>*

Klesnil and Lucas modified the Paris law, taking into account the crack propagation threshold, and thus obtained the crack growth equation valid in areas I and II:

*dN* <sup>¼</sup> *<sup>C</sup>* � <sup>Δ</sup>*K<sup>m</sup>* � <sup>Δ</sup>*K<sup>m</sup>*

and which, unlike the previous equations, which were obtained empirically, is

McEvily developed an expression that is valid for the whole crack growth curve,

From all these Eqs. (4), (5), (6) and (7) integration can obtain the time required for crack growth from any arbitrary to a critical length. Also, all the above terms are

The basic progress that fracture mechanics has made in the sphere of material fatigue is in the analytical breakdown of the fatigue fracture phenomenon into the period of creation, in which the fatigue crack occurs, and the period of growth or expansion that follows and in which the resulting crack increases to a critical size at which a sudden fracture occurs. Thus, the total number of cycles, Nu, after which a fracture occurs, is divided by the number of cycles required for the fatigue crack to form, Ni, and the number of cycles for it to increase to the critical value for fracture,

Analysis of the stress state and deformation at the top of a rising fatigue crack by linear elastic fracture mechanics (LEFM) led to the formulation of the Paris equation for all metals and alloys, which relates the fatigue crack growth rate to the

Although the Paris cracks growth equation is not valid in the whole range, between low velocities near the fatigue threshold ΔKth, and high velocities KIc, the large linear midpoint of the curve covered by the Paris relation proved to be by far the most important from a practical point of view and fatigue crack growth.

*The share of the initiation period Ni and the fatigue crack growth period Np in the total fatigue life Nu.*

<sup>2</sup> <sup>1</sup> <sup>þ</sup>

ð Þ� <sup>1</sup> � *<sup>R</sup> KIc* � <sup>Δ</sup>*<sup>K</sup>* (5)

*th* (6)

*KIc* � *<sup>K</sup>*max (7)

Δ*K*

**Figure 6.**

*The principle form of change of growth rate da/dN = f(ΔK) for R = and directions of displacement of S-curve for relations R*6¼*0.*

From the point of view of the crack growth mechanism and different intensities of influencing factors, three areas can be observed on this curve [6–8].

In area I, the crack propagation velocity tends to zero as the range of stress intensity factors approaches the crack propagation threshold:

$$
\Delta K\_{th} = K\_{th,\max} - K\_{th,\min} = \Delta \sigma \sqrt{\pi \cdot a\_{th}} \cdot Y \tag{2}
$$

where: ath- the length of the initial crack.

In area II the crack grows linearly in the log–log diagram, so it can be described by the equation:

$$\frac{da}{dN} = \mathbf{C} \cdot \Delta K^m \tag{3}$$

where C and m are the material constants determined experimentally. This law is known as Paris' law.

In area III, the crack grows rapidly as the range of stress intensity factors approaches *ΔK*c:

$$
\Delta K\_c = K\_{lc} - K\_{c,\text{min}} = \Delta \sigma \sqrt{\pi \cdot a\_c} \cdot Y \tag{4}
$$

where ac - critical crack length, KIc - fracture toughness.

As Paris' law is valid only in area II, attempts were made to find equations that would describe crack growth in other areas of growth as well. One such is the Forman equation describing crack growth in areas II and III:

*Treatment Analysis of Welding Structure in the Presence of a Crack Type Defects DOI: http://dx.doi.org/10.5772/intechopen.94832*

$$\frac{da}{dN} = \frac{C \cdot \Delta K^m}{(1 - R) \cdot K\_{lc} - \Delta K} \tag{5}$$

Klesnil and Lucas modified the Paris law, taking into account the crack propagation threshold, and thus obtained the crack growth equation valid in areas I and II:

$$\frac{da}{dN} = \mathbf{C} \cdot \left(\Delta K^m - \Delta K\_{th}^m\right) \tag{6}$$

McEvily developed an expression that is valid for the whole crack growth curve, and which, unlike the previous equations, which were obtained empirically, is based on a simple physical model:

$$\frac{da}{dN} = C \cdot (\Delta K - \Delta K\_{th})^2 \left(1 + \frac{\Delta K}{K\_{Ic} - K\_{\text{max}}}\right) \tag{7}$$

From all these Eqs. (4), (5), (6) and (7) integration can obtain the time required for crack growth from any arbitrary to a critical length. Also, all the above terms are valid in the case of type i loads.

### **3.2 Determination of fatigue crack growth parameters**

The basic progress that fracture mechanics has made in the sphere of material fatigue is in the analytical breakdown of the fatigue fracture phenomenon into the period of creation, in which the fatigue crack occurs, and the period of growth or expansion that follows and in which the resulting crack increases to a critical size at which a sudden fracture occurs. Thus, the total number of cycles, Nu, after which a fracture occurs, is divided by the number of cycles required for the fatigue crack to form, Ni, and the number of cycles for it to increase to the critical value for fracture, Np, i.e., Nu = Ni + Np, **Figure 7** [5].

Analysis of the stress state and deformation at the top of a rising fatigue crack by linear elastic fracture mechanics (LEFM) led to the formulation of the Paris equation for all metals and alloys, which relates the fatigue crack growth rate to the stress intensity range at the crack tip:

Although the Paris cracks growth equation is not valid in the whole range, between low velocities near the fatigue threshold ΔKth, and high velocities KIc, the large linear midpoint of the curve covered by the Paris relation proved to be by far the most important from a practical point of view and fatigue crack growth.

**Figure 7.** *The share of the initiation period Ni and the fatigue crack growth period Np in the total fatigue life Nu.*

From the point of view of the crack growth mechanism and different intensities

*The principle form of change of growth rate da/dN = f(ΔK) for R = and directions of displacement of S-curve for*

*π* � *ath*

*dN* <sup>¼</sup> *<sup>C</sup>* � <sup>Δ</sup>*K<sup>m</sup>* (3)

*π* � *ac*

<sup>p</sup> � *<sup>Y</sup>* (2)

<sup>p</sup> � *<sup>Y</sup>* (4)

In area I, the crack propagation velocity tends to zero as the range of stress

<sup>Δ</sup>*Kth* <sup>¼</sup> *Kth*, max � *Kth*, min <sup>¼</sup> <sup>Δ</sup>*<sup>σ</sup>* ffiffiffiffiffiffiffiffiffiffiffiffi

In area II the crack grows linearly in the log–log diagram, so it can be described

of influencing factors, three areas can be observed on this curve [6–8].

*da*

where C and m are the material constants determined experimentally.

In area III, the crack grows rapidly as the range of stress intensity factors

<sup>Δ</sup>*Kc* <sup>¼</sup> *KIc* � *Kc*, min <sup>¼</sup> <sup>Δ</sup>*<sup>σ</sup>* ffiffiffiffiffiffiffiffiffiffi

would describe crack growth in other areas of growth as well. One such is the

As Paris' law is valid only in area II, attempts were made to find equations that

where ac - critical crack length, KIc - fracture toughness.

Forman equation describing crack growth in areas II and III:

intensity factors approaches the crack propagation threshold:

where: ath- the length of the initial crack.

This law is known as Paris' law.

by the equation:

**Figure 6.**

*relations R*6¼*0.*

*Structural Integrity and Failure*

approaches *ΔK*c:

**8**

### *Structural Integrity and Failure*

The application of the Paris equation has proved particularly fruitful in the field of fatigue of structures made of high and very-high strength materials.

The ASTM E647 standard [9] prescribes the measurement of the fatigue crack growth rate da/dN, which develops from an existing crack, and the calculation of the stress intensity factor range, *Δ*K. This means that the test tube should have a tiring crack. There are two important limitations in the ASTM E647 standard: the growth rate must be greater than 10–8 m/cycle to avoid the fatigue threshold area, *Δ*Kth, and the load should be of constant amplitude [4].

Steel of quality A-387 Gr was used to determine the dependence of the fatigue crack growth rate per cycle da/dN and the range of stress intensity factors *Δ*K. 91 15 mm thick [10]. The chemical composition and mechanical properties of the base material are given in **Tables 1** and **2**.

The welded joint is made with two welding processes and two additional materials.

• Root welding - TIG welding, additional material is wire marked BOEHLER C 9 MV-IG, diameter 2,4 mm (international designation W CrMo 91 according to EN ISO 21952-A).

The scheme of the measuring foil and the method of registering crack growth is

Determining the dependence of the fatigue crack growth rate on the cycle da/dN and the range of stress intensity factors ΔK is reduced to determining the coefficient **C** and the exponent **m** in the Paris equation. The fatigue crack growth rate should be attributed to the current crack length, and, to the range of stress intensity factor, ΔK, which depends on the specimen geometry and crack length, and to the variable

<sup>W</sup><sup>3</sup> <sup>p</sup> � f að Þ *<sup>=</sup>*<sup>W</sup> (8)

The appearance of the prepared specimen for determining the fatigue crack

Determining the stress intensity factor range uses the formula

*Treatment Analysis of Welding Structure in the Presence of a Crack Type Defects*

<sup>Δ</sup><sup>K</sup> <sup>¼</sup> <sup>Δ</sup><sup>F</sup> � <sup>L</sup> B ffiffiffiffiffiffiffi

shown in **Figure 9**.

*Modern system for dynamic tests [10].*

**Figure 8.**

force range, ΔF=Fg - Fd.

a - crack length.

**Figure 9.**

**11**

growth parameters is given in **Figure 10**.

*DOI: http://dx.doi.org/10.5772/intechopen.94832*

where: L - range of supports, mm; B - specimen thickness, mm;

W - width (height) of the specimen, mm, and.

*The scheme of the measuring foil and the method of registering crack growth.*

• Filling - REL welding, additional material is an electrode marked BOEHLER FOX C9 MV, diameter 3,50 and 4,00 mm (international designation E CrMo 91 B 4 2 H5 according to EN ISO 3580-A).

Determination of fatigue crack growth rate da/dN and fatigue threshold ΔKth was performed on standard Charpy tubes by the method of bending the tube at three points on a resonant high-frequency pulsator.

The test was performed at the same minimum and maximum load ratio R = 1. The achieved frequency ranged from 175 to 195 Hz depending on whether the crack passed through the base metal, weld metal, or the heat-affected zone, and on the magnitude of the load. The medium load and its amplitude were registered with accuracy 3 Ncm.

Prior to the test, the specimen were mechanically prepared and measuring tapes - foils were glued to the prepared tubes, with which the crack growth was monitored. RMF A-5 measuring foils with a measuring length of 5 mm were used for testing. In order to be able to monitor the crack growth using a measuring foil, the FRACTOMAT crack growth detection device was used, **Figure 8**.


### **Table 1.**

*Chemical composition of the tested batch of steel SA 387 Gr. 91.*


### **Table 2.**

*Mechanical properties of the base material.*

*Treatment Analysis of Welding Structure in the Presence of a Crack Type Defects DOI: http://dx.doi.org/10.5772/intechopen.94832*

The application of the Paris equation has proved particularly fruitful in the field of

The ASTM E647 standard [9] prescribes the measurement of the fatigue crack growth rate da/dN, which develops from an existing crack, and the calculation of the stress intensity factor range, *Δ*K. This means that the test tube should have a tiring crack. There are two important limitations in the ASTM E647 standard: the growth rate must be greater than 10–8 m/cycle to avoid the fatigue threshold area,

Steel of quality A-387 Gr was used to determine the dependence of the fatigue crack growth rate per cycle da/dN and the range of stress intensity factors *Δ*K. 91 15 mm thick [10]. The chemical composition and mechanical properties of the base

• Root welding - TIG welding, additional material is wire marked BOEHLER C 9 MV-IG, diameter 2,4 mm (international designation W CrMo 91 according to

• Filling - REL welding, additional material is an electrode marked BOEHLER FOX C9 MV, diameter 3,50 and 4,00 mm (international designation E CrMo 91

Determination of fatigue crack growth rate da/dN and fatigue threshold ΔKth was performed on standard Charpy tubes by the method of bending the tube at

The test was performed at the same minimum and maximum load ratio R = 1. The achieved frequency ranged from 175 to 195 Hz depending on whether the crack passed through the base metal, weld metal, or the heat-affected zone, and on the magnitude of the load. The medium load and its amplitude were registered with

Prior to the test, the specimen were mechanically prepared and measuring tapes - foils were glued to the prepared tubes, with which the crack growth was monitored. RMF A-5 measuring foils with a measuring length of 5 mm were used for testing. In order to be able to monitor the crack growth using a measuring foil,

**Chemical composition, mas. % C Si Mn P S Cr Mo Ni V Nb Cu** 0.129 0.277 0.443 0.001 0.001 8.25 0.874 0.01 0.198 0.056 0.068

445 580–760 18 40

**Elongation A [%]**

**Impact energy Kv [J] +20°C**

the FRACTOMAT crack growth detection device was used, **Figure 8**.

The welded joint is made with two welding processes and two additional

fatigue of structures made of high and very-high strength materials.

*Δ*Kth, and the load should be of constant amplitude [4].

B 4 2 H5 according to EN ISO 3580-A).

three points on a resonant high-frequency pulsator.

*Chemical composition of the tested batch of steel SA 387 Gr. 91.*

**Tensile strength Rm [MPa]**

material are given in **Tables 1** and **2**.

EN ISO 21952-A).

*Structural Integrity and Failure*

accuracy 3 Ncm.

**Table 1.**

**Table 2.**

**10**

**Yield strength Rp02 [MPa]**

*Mechanical properties of the base material.*

materials.

The scheme of the measuring foil and the method of registering crack growth is shown in **Figure 9**.

The appearance of the prepared specimen for determining the fatigue crack growth parameters is given in **Figure 10**.

Determining the dependence of the fatigue crack growth rate on the cycle da/dN and the range of stress intensity factors ΔK is reduced to determining the coefficient **C** and the exponent **m** in the Paris equation. The fatigue crack growth rate should be attributed to the current crack length, and, to the range of stress intensity factor, ΔK, which depends on the specimen geometry and crack length, and to the variable force range, ΔF=Fg - Fd.

Determining the stress intensity factor range uses the formula

$$
\Delta \mathbf{K} = \frac{\Delta \mathbf{F} \cdot \mathbf{L}}{\mathbf{B} \sqrt{\mathbf{W}^3}} \cdot \mathbf{f} (\mathbf{a}/\mathbf{W}) \tag{8}
$$

where: L - range of supports, mm;

B - specimen thickness, mm;

W - width (height) of the specimen, mm, and.

a - crack length.

**Figure 9.** *The scheme of the measuring foil and the method of registering crack growth.*

**Figure 10.** *The appearance of the prepared specimen for parameter testing [10].*

The geometric term f (a/W) is given by the expression:

*Dependence diagram da/dN - ΔK for speciemens with fatigue crack tip in WM [10].*

*Treatment Analysis of Welding Structure in the Presence of a Crack Type Defects*

*DOI: http://dx.doi.org/10.5772/intechopen.94832*

<sup>W</sup> <sup>1</sup> � <sup>a</sup> W

2 1 <sup>þ</sup> <sup>2</sup> <sup>a</sup> W � � <sup>1</sup> � <sup>a</sup>

Three groups of specimens depending on the location of the crack tip were

Based on the test flow and the obtained dependences of crack length **a** - number of cycles **N**, the fatigue crack growth rate da/dN is calculated. Depending on the applied variable load expressed through the change of the range of voltage intensity

� � 2, 15 � 3, 93 <sup>a</sup>

� �<sup>2</sup> h i � �

W

<sup>W</sup> <sup>þ</sup> 2, 7 <sup>a</sup>

� �<sup>3</sup>*=*<sup>2</sup> (9)

W

f að Þ¼ *=*W

**Figure 12.**

examined, namely:

**13**

<sup>3</sup> � ffiffiffiffi a W <sup>p</sup> � 1, 99 � <sup>a</sup>

Group I - specimens with a crack tip in BM, Group II - specimens with a crack tip in WM and. Group III - specimens with a crack tip in HAZ.

factor, ΔK, log da/dN - log(ΔK) curves are drawn.

**Figure 11.** *Dependence diagram da/dN - ΔK for specimens with fatigue crack tip in BM [10].*

*Treatment Analysis of Welding Structure in the Presence of a Crack Type Defects DOI: http://dx.doi.org/10.5772/intechopen.94832*

**Figure 12.** *Dependence diagram da/dN - ΔK for speciemens with fatigue crack tip in WM [10].*

The geometric term f (a/W) is given by the expression:

$$\mathbf{f}(\mathbf{a}/\mathbf{W}) = \frac{\mathbf{3} \cdot \sqrt{\frac{\mathbf{3}}{\mathbf{W}}} \cdot \left[\mathbf{1}, \mathbf{9} \mathbf{9} - \frac{\mathbf{a}}{\mathbf{W}} \left(\mathbf{1} - \frac{\mathbf{a}}{\mathbf{W}}\right) \left(2, \mathbf{15} - \mathbf{3}, \mathbf{93} \frac{\mathbf{a}}{\mathbf{W}} + 2, \mathbf{7} \left(\frac{\mathbf{a}}{\mathbf{W}}\right)^2\right)\right]}{2 \left(\mathbf{1} + 2 \frac{\mathbf{a}}{\mathbf{W}}\right) \left(\mathbf{1} - \frac{\mathbf{a}}{\mathbf{W}}\right)^{3/2}} \tag{9}$$

Three groups of specimens depending on the location of the crack tip were examined, namely:

Group I - specimens with a crack tip in BM,

Group II - specimens with a crack tip in WM and.

Group III - specimens with a crack tip in HAZ.

Based on the test flow and the obtained dependences of crack length **a** - number of cycles **N**, the fatigue crack growth rate da/dN is calculated. Depending on the applied variable load expressed through the change of the range of voltage intensity factor, ΔK, log da/dN - log(ΔK) curves are drawn.

**Figure 10.**

*Structural Integrity and Failure*

**Figure 11.**

**12**

*The appearance of the prepared specimen for parameter testing [10].*

*Dependence diagram da/dN - ΔK for specimens with fatigue crack tip in BM [10].*

**4. Conclusions**

**Test tube label**

**Table 4.**

**Table 5.**

**Test tube label**

**Test temperature, °C**

*DOI: http://dx.doi.org/10.5772/intechopen.94832*

*Fatigue crack growth parameters for notched specimens in WM.*

*Fatigue crack growth parameters for notched specimens in HAZ.*

**Test temperature, °C**

**Fatigue threshold ΔKth, MPa m1/2**

*Treatment Analysis of Welding Structure in the Presence of a Crack Type Defects*

**Fatigue threshold ΔKth, MPa m1/2**

**ZUT-1** <sup>20</sup> 6,6 3,05 <sup>10</sup><sup>10</sup> 4,01 3,12 <sup>10</sup><sup>6</sup> **ZUT-2** 6,8 3,07 <sup>10</sup><sup>10</sup> 4,04 3,37 <sup>10</sup><sup>6</sup> **ZUT-3** 6,5 2,85 <sup>10</sup><sup>10</sup> 4,09 3,51 <sup>10</sup><sup>6</sup>

**<sup>M</sup>Š-1** <sup>20</sup> 7,2 3,88 <sup>10</sup><sup>10</sup> 3,62 2,56 <sup>10</sup><sup>6</sup> **<sup>M</sup>Š-2** 7,1 4,05 <sup>10</sup><sup>10</sup> 3,71 2,07 <sup>10</sup><sup>6</sup> **<sup>M</sup>Š-3** 7,4 3,93 <sup>10</sup><sup>10</sup> 3,80 2,48 <sup>10</sup><sup>6</sup>

**Coefficient C**

**Coefficient C**

**Exponentm da/dN, μm/cikl, pri**

**Exponentm da/dN, m/cikl, pri**

**ΔK = 10 MPa m1/2**

**ΔK = 10 MPa m1/2**

cracking.

**15**

ily errors that occur in exploitation.

structure elements and time as a whole.

safety of structures must be controlled.

Practice has shown that structures that are exposed to the effects of variable load during operation are most prone to accidents and fractures. The causes are primar-

• The reason for the application of fracture mechanics is based on the fact that in the presence of errors, which inevitably occur as a result of imperfections in production processes and/or in operation, there is a loss of load-bearing

• The reduction of load-bearing capacity in order to achieve the reliability and

• The obtained values of the fatigue threshold ΔKth represent important data on the quality of the tested materials from the aspect of its micromechanical

determination, the magnitudes and periods of action of the variable load are known for the formation and then for the growth of the tired crack until

properties, ie on the behavior in the presence of cracks. After its

**Figure 13.** *Dependence diagram da/dN - Δ for specimens with a fatigue crack tip in HAZ [10].*

Characteristic diagrams of fatigue crack growth rates, da/dN - change of stress intensity factor range, ΔK, for specimens with fatigue crack tip in BM, **Figure 11**., for specimens with fatigue crack tip in WM, **Figure 12**., and for specimens with a fatigue crack tip in HAZ, **Figure 13**.

The obtained values of the parameters of the Paris equation, coefficient C and exponent m, fatigue threshold ΔKth, and fatigue crack growth rate, da/dN, at the value of ΔK = 10 MPa m1/2, are given in **Table 3** for specimens with a notch in BM, **Table 4** for specimens with a notch in WM, and **Table 5** for speciemns with a notch in HAZ.


**Table 3.** *Fatigue crack growth parameters for notched specimens in BM.*


*Treatment Analysis of Welding Structure in the Presence of a Crack Type Defects DOI: http://dx.doi.org/10.5772/intechopen.94832*

**Table 4.**

*Fatigue crack growth parameters for notched specimens in WM.*


**Table 5.**

Characteristic diagrams of fatigue crack growth rates, da/dN - change of stress intensity factor range, ΔK, for specimens with fatigue crack tip in BM, **Figure 11**., for specimens with fatigue crack tip in WM, **Figure 12**., and for specimens with a

*Dependence diagram da/dN - Δ for specimens with a fatigue crack tip in HAZ [10].*

The obtained values of the parameters of the Paris equation, coefficient C and exponent m, fatigue threshold ΔKth, and fatigue crack growth rate, da/dN, at the value of ΔK = 10 MPa m1/2, are given in **Table 3** for specimens with a notch in BM, **Table 4** for specimens with a notch in WM, and **Table 5** for speciemns with a notch

> **Coefficient C**

**Exponentm da/dN, μm/cikl,**

**pri ΔK = 10 MPa m1/2**

**Fatigue threshold ΔKth, MPa m1/2**

**OM-1** <sup>20</sup> 6,9 2,98 <sup>10</sup><sup>10</sup> 3,62 1,24 <sup>10</sup><sup>6</sup> **OM-2** 6,8 3,07 <sup>10</sup><sup>10</sup> 3,58 1,17 <sup>10</sup><sup>6</sup> **OM-3** 7,1 2,85 <sup>10</sup><sup>10</sup> 3,59 1,11 <sup>10</sup><sup>6</sup>

fatigue crack tip in HAZ, **Figure 13**.

*Structural Integrity and Failure*

**Test temperature, °C**

*Fatigue crack growth parameters for notched specimens in BM.*

in HAZ.

**Table 3.**

**14**

**Test tube label**

**Figure 13.**

*Fatigue crack growth parameters for notched specimens in HAZ.*

### **4. Conclusions**

Practice has shown that structures that are exposed to the effects of variable load during operation are most prone to accidents and fractures. The causes are primarily errors that occur in exploitation.


*Structural Integrity and Failure*

### **Author details**

Mersida Manjgo<sup>1</sup> \* and Meri Burzic<sup>2</sup>

1 Džemal Bijedić University, Faculty of Mechanical Engineering, Mostar, Bosnia and Herzegovina

2 University of Belgrade, Innovation Center, Faculty of Mechanical Engineering, Belgrade, Serbia

**References**

[1] Fuchs H. O., Stephens R. I. Metal Fatigue in Engineering. A Wiley-Intersciene Publication. New York: John Wiley and Sons; 1980.

*DOI: http://dx.doi.org/10.5772/intechopen.94832*

*Treatment Analysis of Welding Structure in the Presence of a Crack Type Defects*

[2] Podrug S. Mehanika loma. Split: Fakultet elektrotehnike, strojarstva i

[3] ASTM E466–82. Standard Practice for Conducting Constant Amplitude Axial Fatigue Test of Metalic Materials. Annual Book of ASTM Standard; 1986.

brodogradnje; 2009.

[4] M. Manjgo, Kriterijumi

disertacija, 2008

Podgorica

Verlag; 1970.

Mechanics; 2003.

Crack Growth

**17**

prihvatljivosti prslina u zavarenim spojevima posuda pod pritiskom od mikrolegiranih čelika, Doktorska

[5] V. Ćulafić, Uvod u mehaniku loma,

Bruchmechanik. Munchen: Carl Hanser

[7] Ekberg A. Fatigue crack propagation.

[8] Ritchie R. O. Mechanisms of fatiguecrack propagation in ductile and brittle solids. Netherlands: International Journal of Fracture 100: 55–83, 1999.

[9] ASTM 647 Standard Test Method for Constant-Load-Amplitude Fatigue

[10] M. Manjgo, M. Burzić, Optimal welding parameters of SA 387 Gr. 91

thick steel plates in corrosive environment, Elobarat po projektu Eureka, Beograd-Mostar, 2016–2019.

[6] Heckel K. Einfuhrung in die technische Anwendung der

Gothenborg: Chalmers Solid

\*Address all correspondence to: mersida.manjgo@unmo.ba

© 2020 The Author(s). Licensee IntechOpen. This chapter is distributed under the terms of the Creative Commons Attribution License (http://creativecommons.org/licenses/ by/3.0), which permits unrestricted use, distribution, and reproduction in any medium, provided the original work is properly cited.

*Treatment Analysis of Welding Structure in the Presence of a Crack Type Defects DOI: http://dx.doi.org/10.5772/intechopen.94832*

### **References**

[1] Fuchs H. O., Stephens R. I. Metal Fatigue in Engineering. A Wiley-Intersciene Publication. New York: John Wiley and Sons; 1980.

[2] Podrug S. Mehanika loma. Split: Fakultet elektrotehnike, strojarstva i brodogradnje; 2009.

[3] ASTM E466–82. Standard Practice for Conducting Constant Amplitude Axial Fatigue Test of Metalic Materials. Annual Book of ASTM Standard; 1986.

[4] M. Manjgo, Kriterijumi prihvatljivosti prslina u zavarenim spojevima posuda pod pritiskom od mikrolegiranih čelika, Doktorska disertacija, 2008

[5] V. Ćulafić, Uvod u mehaniku loma, Podgorica

[6] Heckel K. Einfuhrung in die technische Anwendung der Bruchmechanik. Munchen: Carl Hanser Verlag; 1970.

[7] Ekberg A. Fatigue crack propagation. Gothenborg: Chalmers Solid Mechanics; 2003.

[8] Ritchie R. O. Mechanisms of fatiguecrack propagation in ductile and brittle solids. Netherlands: International Journal of Fracture 100: 55–83, 1999.

[9] ASTM 647 Standard Test Method for Constant-Load-Amplitude Fatigue Crack Growth

[10] M. Manjgo, M. Burzić, Optimal welding parameters of SA 387 Gr. 91 thick steel plates in corrosive environment, Elobarat po projektu Eureka, Beograd-Mostar, 2016–2019.

**Author details**

Mersida Manjgo<sup>1</sup>

Belgrade, Serbia

**16**

Bosnia and Herzegovina

*Structural Integrity and Failure*

\* and Meri Burzic<sup>2</sup>

\*Address all correspondence to: mersida.manjgo@unmo.ba

provided the original work is properly cited.

1 Džemal Bijedić University, Faculty of Mechanical Engineering, Mostar,

2 University of Belgrade, Innovation Center, Faculty of Mechanical Engineering,

© 2020 The Author(s). Licensee IntechOpen. This chapter is distributed under the terms of the Creative Commons Attribution License (http://creativecommons.org/licenses/ by/3.0), which permits unrestricted use, distribution, and reproduction in any medium,

**Chapter 2**

**Abstract**

**1. Introduction**

**19**

Building

Structural Modeling and Dynamic

Increasing energy demand urge the developing countries to consider different types of energy sources. Owing the fact that the energy production capacity of renewable energy sources is lower than a nuclear power plant, developed countries like US, France, Japan, Russia and China lead to construct nuclear power plants. These countries compensate 80% of their energy need from nuclear power plants. Further, they periodically conduct tests in order to assess the safety of the existing nuclear power plants by applying impact type loads to the structures. In this study, a sample third-generation nuclear reactor building has been selected to assess its seismic behavior and to observe the crack propagations of the prestressed outer containment. First, a 3D model has been set up using ABAQUS finite element program. Afterwards, modal analysis is conducted to determine the mode shapes. Nonlinear dynamic time history analyses are then followed using an artificial strong ground motion which is compatible with the mean design spectrum of the previously selected ground motions that are scaled to Eurocode 8 Soil type B design spectrum. Results of the conducted nonlinear dynamic analyses are considered in

**Keywords:** nuclear reactor building, prestressed containment, nonlinear dynamic

Large growth of energy consumption is an essential issue that shall be considered, especially for developing countries. Nuclear power industry seems to take a key role over the continuously increasing energy demand. First known nuclear power station for the generations of electricity was built for commercial use in Russia in 1954, with an output of 5 MW(e). Afterwards, USA constructed a pressurized light water reactor (PWR) with 60 MW(e) which first produced electricity. Many countries over the world followed this first attempt and a rapid growth of nuclear energy production started. Though renewable energy seems to take role in the approaching era, it is estimated that the percentage of the generated electricity from these sources will not be deemed to nuclear power plants. It can be forecasted that up to 2050, nuclear energy and nuclear power plants (NPP) will play a major

Nowadays, supporters of nuclear power admit that this is indeed a technology

which is more expensive than its counterparts. Throughout its development,

Analysis of a Nuclear Reactor

*Evrim Oyguc, Abdul Hayır and Resat Oyguc*

terms of stress distributions and crack propagations.

role in producing electricity worldwide.

time history analysis, ABAQUS 3D finite element model

### **Chapter 2**

## Structural Modeling and Dynamic Analysis of a Nuclear Reactor Building

*Evrim Oyguc, Abdul Hayır and Resat Oyguc*

### **Abstract**

Increasing energy demand urge the developing countries to consider different types of energy sources. Owing the fact that the energy production capacity of renewable energy sources is lower than a nuclear power plant, developed countries like US, France, Japan, Russia and China lead to construct nuclear power plants. These countries compensate 80% of their energy need from nuclear power plants. Further, they periodically conduct tests in order to assess the safety of the existing nuclear power plants by applying impact type loads to the structures. In this study, a sample third-generation nuclear reactor building has been selected to assess its seismic behavior and to observe the crack propagations of the prestressed outer containment. First, a 3D model has been set up using ABAQUS finite element program. Afterwards, modal analysis is conducted to determine the mode shapes. Nonlinear dynamic time history analyses are then followed using an artificial strong ground motion which is compatible with the mean design spectrum of the previously selected ground motions that are scaled to Eurocode 8 Soil type B design spectrum. Results of the conducted nonlinear dynamic analyses are considered in terms of stress distributions and crack propagations.

**Keywords:** nuclear reactor building, prestressed containment, nonlinear dynamic time history analysis, ABAQUS 3D finite element model

### **1. Introduction**

Large growth of energy consumption is an essential issue that shall be considered, especially for developing countries. Nuclear power industry seems to take a key role over the continuously increasing energy demand. First known nuclear power station for the generations of electricity was built for commercial use in Russia in 1954, with an output of 5 MW(e). Afterwards, USA constructed a pressurized light water reactor (PWR) with 60 MW(e) which first produced electricity. Many countries over the world followed this first attempt and a rapid growth of nuclear energy production started. Though renewable energy seems to take role in the approaching era, it is estimated that the percentage of the generated electricity from these sources will not be deemed to nuclear power plants. It can be forecasted that up to 2050, nuclear energy and nuclear power plants (NPP) will play a major role in producing electricity worldwide.

Nowadays, supporters of nuclear power admit that this is indeed a technology which is more expensive than its counterparts. Throughout its development,

starting from the 1980s, the sector improved the design of reactors. Furthermore, the scientists working in this era aimed to overcome human error or equipment malfunctions. These resulted more robust, fuel efficient and advanced reactor designs where the cost of the construction is also reduced.

been set up using ABAQUS finite element program. Afterwards, modal analysis is conducted to determine the mode shapes and followed by nonlinear dynamic time history analysis using previously selected ground motions to determine the stress

*Structural Modeling and Dynamic Analysis of a Nuclear Reactor Building*

**2. Details of the 3D analytical model of case-studied nuclear reactor**

Both the Sandia National Laboratories and United States Nuclear Regulatory Commission (USNRC) conducted test programs concerning about the seismic performance of prestressed containments under severe accident conditions. The experiment model was 1:4 prestressed containment and it was conducted in two phases. Determination of the ultimate load capacity and seismic response were carried in the first phase, whereas beyond design basis response was the aim of second phase. More

A sectional view of the considered test building is illustrated in **Figure 1**. The structure has an overall height of 16.4 m, the inner radius of the dome is 5.37 m and the cylinder wall has a thickness of 0.325 m. The bottom part of the structure is fixed, gallery and buttresses are incorporated into the system. Prestressing force is applied by using horizontal and vertical tendons which are anchored at the buttresses and gallery, respectively. Around the gallery and buttresses steel rebars are

details about this experiment may be found in Hessheimer et al. [19, 20].

used to enhance the capacity of the double-shell containment building.

distribution and crack propagation.

*DOI: http://dx.doi.org/10.5772/intechopen.94956*

**building**

**Figure 1.**

**21**

*Sectional view of the considered test building.*

Next generation nuclear reactors may be classified into two broad categories: evolutionary and revolutionary [1]. While the Generation III and III+ types belong to the former, Generation IV reactors fell into the latter category. Crimello [2] reported that the design of Generation III type nuclear reactors based on minimizing the risk and thus increase safety. Advanced Boiling and Pressurized Water reactors and Enhanced CANDU 6 are all examples to this type. Passive safety features are incorporated into Generation III type reactors to reach the revolutionary reactors Generation III+. Some examples to these reactors may be: AP1000, VVER 1200 and APR 1400.

Generally, an NPP is composed of five principal buildings: the nuclear island, the annex building, the turbine building, the diesel generator building and radwaste buildings. The nuclear island consists of the containment building, the shield building and the auxiliary building. The containment is one of the most important components of an NPP because it serves as the final barrier under postulated accident conditions.

In the last decade, there is a keen interest on modeling the reactor buildings by using simplified 2D lumped-mass stick model [3–7]. The results of the analysis may not be considered as accurate as 3D finite element modeling, although the latter one is more time consuming. In the 3D approach, a containment building is generally modeled using either by shell or 3D brick elements [8–10]. In a recent approach to save time, 3D lumped mass stick models were developed and used in the literature to represent the seismic behavior of containment building [11]. This approach is mostly preferred when the coupling effects between the containment building and auxiliary buildings are considered. Improvements in computational efficiency increased the use of 3D finite element models which are capable of defining high levels of structural detail.

While a possible damage in an infra-structure can be repaired or retrofitted within mean time, a possible damage in an NPP may cause catastrophic damage. The massive damage in an NPP may be observed when an internal accident happens such as loss of coolant accident (LOCA) or when an external event (airplane crash, earthquake, explosions etc.) happens. To overcome these deficiencies the majority of the latest NPP's are constructed with double containment. Further, in advance, the containment structure constitutes an ultimate barrier against the dissemination of fissile products towards the general public [12].

Kwak and Kim [13] highlighted that most of the recent containments are composed of a dome, a wall and a foundation which are laterally prestressed. The International Federation for Structural Concrete (fib) [14] reported that the response of these shell-type concrete structures due to external events should be experimentally and analytically studied to evaluate their safety. There are numerous experiments in the literature which can be regarded as representative experimental studies. Sandia National Laboratories [15] brought foreword the cost and timeconsuming features of the conducted experiments and emphasized that these costly experiments often do not precisely simulate the loading and support conditions of the actual structure.

In this study, seismic behavior of the prestressed outer containment of a thirdgeneration nuclear reactor building is evaluated. Since Turkish Building Earthquake Code (TBEC) [16] does not cover the earthquake resistant design of nuclear structures, international standards such as ASCE 4-16 [17] and ASCE 43-05 [18] have been adopted for the analysis. First, a 3D model of the outer containment vessel has *Structural Modeling and Dynamic Analysis of a Nuclear Reactor Building DOI: http://dx.doi.org/10.5772/intechopen.94956*

been set up using ABAQUS finite element program. Afterwards, modal analysis is conducted to determine the mode shapes and followed by nonlinear dynamic time history analysis using previously selected ground motions to determine the stress distribution and crack propagation.

### **2. Details of the 3D analytical model of case-studied nuclear reactor building**

Both the Sandia National Laboratories and United States Nuclear Regulatory Commission (USNRC) conducted test programs concerning about the seismic performance of prestressed containments under severe accident conditions. The experiment model was 1:4 prestressed containment and it was conducted in two phases. Determination of the ultimate load capacity and seismic response were carried in the first phase, whereas beyond design basis response was the aim of second phase. More details about this experiment may be found in Hessheimer et al. [19, 20].

A sectional view of the considered test building is illustrated in **Figure 1**. The structure has an overall height of 16.4 m, the inner radius of the dome is 5.37 m and the cylinder wall has a thickness of 0.325 m. The bottom part of the structure is fixed, gallery and buttresses are incorporated into the system. Prestressing force is applied by using horizontal and vertical tendons which are anchored at the buttresses and gallery, respectively. Around the gallery and buttresses steel rebars are used to enhance the capacity of the double-shell containment building.

**Figure 1.** *Sectional view of the considered test building.*

starting from the 1980s, the sector improved the design of reactors. Furthermore, the scientists working in this era aimed to overcome human error or equipment malfunctions. These resulted more robust, fuel efficient and advanced reactor

Next generation nuclear reactors may be classified into two broad categories: evolutionary and revolutionary [1]. While the Generation III and III+ types belong to the former, Generation IV reactors fell into the latter category. Crimello [2] reported that the design of Generation III type nuclear reactors based on minimizing the risk and thus increase safety. Advanced Boiling and Pressurized Water reactors and Enhanced CANDU 6 are all examples to this type. Passive safety features are incorporated into Generation III type reactors to reach the revolutionary reactors Generation III+. Some examples to these reactors may be: AP1000,

Generally, an NPP is composed of five principal buildings: the nuclear island, the annex building, the turbine building, the diesel generator building and radwaste buildings. The nuclear island consists of the containment building, the shield building and the auxiliary building. The containment is one of the most important components of an NPP because it serves as the final barrier under postulated

In the last decade, there is a keen interest on modeling the reactor buildings by using simplified 2D lumped-mass stick model [3–7]. The results of the analysis may not be considered as accurate as 3D finite element modeling, although the latter one is more time consuming. In the 3D approach, a containment building is generally modeled using either by shell or 3D brick elements [8–10]. In a recent approach to save time, 3D lumped mass stick models were developed and used in the literature to represent the seismic behavior of containment building [11]. This approach is mostly preferred when the coupling effects between the containment building and auxiliary buildings are considered. Improvements in computational efficiency increased the use of 3D finite element models which are capable of defining high

While a possible damage in an infra-structure can be repaired or retrofitted within mean time, a possible damage in an NPP may cause catastrophic damage. The massive damage in an NPP may be observed when an internal accident happens such as loss of coolant accident (LOCA) or when an external event (airplane crash, earthquake, explosions etc.) happens. To overcome these deficiencies the majority of the latest NPP's are constructed with double containment. Further, in advance, the containment structure constitutes an ultimate barrier against the dissemination

Kwak and Kim [13] highlighted that most of the recent containments are com-

In this study, seismic behavior of the prestressed outer containment of a thirdgeneration nuclear reactor building is evaluated. Since Turkish Building Earthquake Code (TBEC) [16] does not cover the earthquake resistant design of nuclear structures, international standards such as ASCE 4-16 [17] and ASCE 43-05 [18] have been adopted for the analysis. First, a 3D model of the outer containment vessel has

posed of a dome, a wall and a foundation which are laterally prestressed. The International Federation for Structural Concrete (fib) [14] reported that the response of these shell-type concrete structures due to external events should be experimentally and analytically studied to evaluate their safety. There are numerous experiments in the literature which can be regarded as representative experimental studies. Sandia National Laboratories [15] brought foreword the cost and timeconsuming features of the conducted experiments and emphasized that these costly experiments often do not precisely simulate the loading and support conditions of

designs where the cost of the construction is also reduced.

VVER 1200 and APR 1400.

*Structural Integrity and Failure*

accident conditions.

levels of structural detail.

the actual structure.

**20**

of fissile products towards the general public [12].

**3. Constitutive material models**

*DOI: http://dx.doi.org/10.5772/intechopen.94956*

*Structural Modeling and Dynamic Analysis of a Nuclear Reactor Building*

**3.1 Constitutive model for concrete**

and summarized in the followings.

*a. Decomposition of the strain rate.*

where *ε*´ is the total strain rate, *ε*´

0; *Del* <sup>¼</sup> ð Þ <sup>1</sup> � *<sup>d</sup> <sup>D</sup>el*

the plastic part of the strain rate.

*b. Stress–strain relations.*

the Cauchy stress by Eq. (4).

*c. Hardening variables.*

*pl <sup>t</sup>* and ~*εpl*

lated using Eq. (2).

where *Del*

preferred.

*<sup>d</sup>* <sup>¼</sup> *<sup>d</sup> <sup>σ</sup>*, <sup>~</sup>*εpl* .~*<sup>ε</sup>*

in Eq. (5).

**23**

independent model which is defined by Eq. (1).

*<sup>σ</sup>* <sup>¼</sup> ð Þ <sup>1</sup> � *<sup>d</sup> <sup>D</sup>el*

The constitutive relation of concrete material is simulated by concrete damaged plasticity model (CDP) of ABAQUS, which is proposed by Lubliner et al. [21] and by Lee and Fenves [22]. This model is preferred by the scientists owing the fact that the model is capable of defining the concrete behavior under cyclic or dynamic loading properly. Further, the material model aims to properly simulate the effects of residual damage. Both the tensile and compressive damages are taken into account in the CDP model. Main contents of the CDP model have been discussed

This is done by assuming an additional strain rate decomposition for the rate

*el* <sup>þ</sup> *<sup>ε</sup>*´

This is dominated by scalar damaged elasticity value which can easily be calcu-

*pl* <sup>¼</sup> *<sup>D</sup>el* : *<sup>ε</sup>*´ � *<sup>ε</sup>*´

*pl* (1)

*pl* (2)

*pl* is

*el* is the elastic part of the strain rate, and *ε*´

0; and *d*, show the initial stiffness, the degraded elastic

*pl* (3)

*σ* ¼ ð Þ 1 � *d σ*´ (4)

*ε*´ ¼ *ε*´

<sup>0</sup> : *ε*´ � *ε*´

*<sup>σ</sup>def* <sup>¼</sup> *<sup>D</sup>el*

stiffness and the scalar stiffness degradation variable, respectively. The scalar stiffness degradation may be associated with the failure mechanism of the concrete which has a range 0≤ *d*≤ 1. Thus, the effective stress is calculated using Eq. (3), and

<sup>0</sup> : *ε*\_ � *ε*\_

ð Þ 1 � *d* in Eq. (4) is the factor that represents the effective load carrying area. When *d* ¼ 0, reduction in the elastic stiffness is not expected and it is the condition where ´*σ* ¼ *σ*. When *d* 6¼ 0, damage is possible and use of the effective stress value will give more robust results. Hence, in problems dealing with plasticity ´*σ* value is

It is also possible to relate *d* value by a set of hardening variables, ~*εpl*, and *σ*; such,

pression, respectively. The governing equation of the hardening variables is given

*<sup>c</sup>* are referred to define damage states in tension and com-

**Figure 2.** *Illustration of the three-dimensional model of the reactor building.*

The three-dimensional model of the test structure has been modeled using shell elements. Smeared steel layers have been used to simulate both the steel reinforcements and prestressing tendons. The created three-dimensional model has been illustrated in **Figure 2**. To satisfy the experiment conditions the containment is assumed as fix, thus boundary conditions are assigned to the structure at the bottom. In literature two alternatives are valid to define the prestressing force: assigning an initial stress value and altering with the temperature of the structure. Here the former methodology is applied to properly simulate the prestressing force. The abrupt changes around the openings are neglected. Incremental load steps are used to define the internal pressure values.

Yielding of reinforcing steel and prestressed tendons has been selected as the criteria of failure in any location in the containment. The total mass of the structure has been calculated as 2,956,294 kg and the results of the eigenvalue analysis are summarized in **Table 1**.


### **Table 1.**

*Eigenvalue output.*

### **3. Constitutive material models**

### **3.1 Constitutive model for concrete**

The constitutive relation of concrete material is simulated by concrete damaged plasticity model (CDP) of ABAQUS, which is proposed by Lubliner et al. [21] and by Lee and Fenves [22]. This model is preferred by the scientists owing the fact that the model is capable of defining the concrete behavior under cyclic or dynamic loading properly. Further, the material model aims to properly simulate the effects of residual damage. Both the tensile and compressive damages are taken into account in the CDP model. Main contents of the CDP model have been discussed and summarized in the followings.

### *a. Decomposition of the strain rate.*

This is done by assuming an additional strain rate decomposition for the rate independent model which is defined by Eq. (1).

$$
\acute{\varepsilon} = \acute{\varepsilon}^{el} + \acute{\varepsilon}^{pl} \tag{1}
$$

where *ε*´ is the total strain rate, *ε*´ *el* is the elastic part of the strain rate, and *ε*´ *pl* is the plastic part of the strain rate.

### *b. Stress–strain relations.*

The three-dimensional model of the test structure has been modeled using shell elements. Smeared steel layers have been used to simulate both the steel reinforcements and prestressing tendons. The created three-dimensional model has been illustrated in **Figure 2**. To satisfy the experiment conditions the containment is assumed as fix, thus boundary conditions are assigned to the structure at the bottom. In literature two alternatives are valid to define the prestressing force: assigning an initial stress value and altering with the temperature of the structure. Here the former methodology is applied to properly simulate the prestressing force. The abrupt changes around the openings are neglected. Incremental load steps are

Yielding of reinforcing steel and prestressed tendons has been selected as the criteria of failure in any location in the containment. The total mass of the structure has been calculated as 2,956,294 kg and the results of the eigenvalue analysis are

 0.36315 0.60262 9.5910E-02 0.52506 0.72461 0.11533 4.6031 2.1455 0.34146 6.6812 2.5848 0.41138 12.831 3.5820 0.57009 18.840 4.3405 0.69081 31.426 5.6059 0.89220 40.622 6.3735 1.0144 43.410 6.5887 1.0486 45.345 6.7339 1.0717

**(rad/time) (cycles/time)**

**Mode No Eigenvalue Frequency**

used to define the internal pressure values.

*Illustration of the three-dimensional model of the reactor building.*

summarized in **Table 1**.

**Table 1.** *Eigenvalue output.*

**22**

**Figure 2.**

*Structural Integrity and Failure*

This is dominated by scalar damaged elasticity value which can easily be calculated using Eq. (2).

$$\sigma = (\mathbf{1} - d) D\_0^{el} : \left( \acute{\varepsilon} - \acute{\varepsilon}^{pl} \right) = D^{el} : \left( \acute{\varepsilon} - \acute{\varepsilon}^{pl} \right) \tag{2}$$

where *Del* 0; *Del* <sup>¼</sup> ð Þ <sup>1</sup> � *<sup>d</sup> <sup>D</sup>el* 0; and *d*, show the initial stiffness, the degraded elastic stiffness and the scalar stiffness degradation variable, respectively. The scalar stiffness degradation may be associated with the failure mechanism of the concrete which has a range 0≤ *d*≤ 1. Thus, the effective stress is calculated using Eq. (3), and the Cauchy stress by Eq. (4).

$$\overline{\sigma}^{d \in \mathfrak{f}} = D\_0^{cl} : \left( \dot{\varepsilon} - \dot{\varepsilon}^{pl} \right) \tag{3}$$

$$
\sigma = (1 - d)\sigma \tag{4}
$$

ð Þ 1 � *d* in Eq. (4) is the factor that represents the effective load carrying area. When *d* ¼ 0, reduction in the elastic stiffness is not expected and it is the condition where ´*σ* ¼ *σ*. When *d* 6¼ 0, damage is possible and use of the effective stress value will give more robust results. Hence, in problems dealing with plasticity ´*σ* value is preferred.

### *c. Hardening variables.*

It is also possible to relate *d* value by a set of hardening variables, ~*εpl*, and *σ*; such, *<sup>d</sup>* <sup>¼</sup> *<sup>d</sup> <sup>σ</sup>*, <sup>~</sup>*εpl* .~*<sup>ε</sup> pl <sup>t</sup>* and ~*εpl <sup>c</sup>* are referred to define damage states in tension and compression, respectively. The governing equation of the hardening variables is given in Eq. (5).

$$
\hat{\epsilon}^{pl} = \begin{bmatrix} \hat{\epsilon}\_t^{pl} \\ \hat{\epsilon}\_c^{pl} \end{bmatrix}; \dot{\hat{\epsilon}}^{pl} = h\left(\overline{\sigma}, \hat{\epsilon}^{pl}\right). \dot{\epsilon}^{pl} \tag{5}
$$

Using the Drucker-Prager hyperbolic function, the *G* parameter may be defined using Eq. (13). Here *ψ* is the dilation angle at high confining pressure; *σt*<sup>0</sup> is the uniaxial tensile stress at failure; and *ϵ* refers to an eccentricity parameter.

<sup>2</sup> <sup>þ</sup> *<sup>q</sup>*<sup>2</sup>

� *ptanψ* (13)

� � (14)

*<sup>t</sup> dt* and ~*εpl*

*ε pl <sup>c</sup>* ¼ �~\_ *ε pl*

� �, 0ð Þ <sup>≤</sup>*dt* <sup>≤</sup><sup>1</sup> (15)

� �, 0ð Þ <sup>≤</sup>*dc* <sup>≤</sup><sup>1</sup> (16)

*<sup>c</sup>* <sup>¼</sup> <sup>Ð</sup>*<sup>t</sup>* 0~\_ *ε pl <sup>c</sup> dt* are

, (*i* ¼ 1, 2, *::*) are other

11, in uniaxial

ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ð Þ *ϵσt*0*tanψ*

This section will be considered for both uniaxial and cyclic loading protocols.

The assumption relies on the fact that the uniaxial stress–strain curves can be

In Eq. (14) the subscripts *t* and *c* refer to tension and compression, respectively;

predefined field variables. Under uniaxial loading conditions the effective plastic

compression (ABAQUS 2020). The response of concrete to uniaxial loading is as illustrated in **Figure 3**. The left illustration is for when the system is under effect

To properly evaluate the degraded response of concrete two independent damage variables which are assumed as functions of the strain values are introduced: *dt* and *dc*. These variables can be calculated following Eqs. (15) and (16). If there is no damage in the material then these variables are assumed to be equal to zero, on the contrary these variables are taken as one when the material is fully damaged.

11, in uniaxial tension and ~\_

*<sup>c</sup>* , ~\_ *ε pl <sup>c</sup>* , *θ*, *fi*

*pl <sup>t</sup>* <sup>¼</sup> <sup>Ð</sup>*<sup>t</sup>* 0~\_ *ε pl*

represented in terms of stress and plastic strain values using Eq. (14).

*G* ¼

*σ<sup>t</sup>* ¼ *σ<sup>t</sup>* ~*ε*

*pl <sup>t</sup>* , ~\_ *ε pl <sup>t</sup>* , *θ*, *fi* � �, *<sup>σ</sup><sup>c</sup>* <sup>¼</sup> *<sup>σ</sup><sup>c</sup>* <sup>~</sup>*εpl*

*<sup>c</sup>* are the equivalent plastic strain rates, ~*ε*

the equivalent plastic strains, *θ* is the temperature, and *fi*

*dt* ¼ *dt* ~*ε*

*dc* <sup>¼</sup> *dc* <sup>~</sup>*εpl*

of compression only and the right one corresponds to tension case.

*pl <sup>t</sup>* , *θ*, *fi*

*<sup>c</sup>* , *θ*, *fi*

Defining *EO* as the initial modulus of the material, the stress values may be

*ε pl <sup>t</sup>* <sup>¼</sup> ~\_ *ε pl*

*f. Damage and stiffness degradation.*

*DOI: http://dx.doi.org/10.5772/intechopen.94956*

• Uniaxial loading

strain rates are given as: ~\_

determined using Eq. (17).

*Response of concrete to uniaxial loading.*

~\_ *ε pl <sup>t</sup>* and ~\_ *ε pl*

**Figure 3.**

**25**

q

*Structural Modeling and Dynamic Analysis of a Nuclear Reactor Building*

These variables control the yield surface, the degradation of the elastic stiffness and the dissipated fracture energy.

### *d. Yield function.*

The yield function, *F σ*, ~*εpl* � �, represents a surface in effective stress space, which determines the states of failure or damage. For the inviscid plastic-damage model, *F σ*, ~*εpl* � �≤0.*F* can also be defined in terms of *σ*, as in Eq. (6).

$$F(\overline{\sigma}, \bar{\varepsilon}^{pl}) = \frac{1}{1 - a} \left( \overline{q} - 3a\overline{p} + \beta \langle \bar{\varepsilon}^{pl} \rangle \langle \hat{\overline{\sigma}}\_{\max} \rangle - \gamma \langle -\hat{\overline{\sigma}}\_{\max} \rangle \right) - \overline{\sigma}\_{\varepsilon} \langle \bar{\varepsilon}\_{\varepsilon}^{pl} \rangle \le 0 \tag{6}$$

*α* and *γ* in Eq. (6) are dimensionless constants. *p* and *q* are the effective hydrostatic pressure and the Von Mises equivalent effective stress values, respectively. Further these parameters can be calculated by Eqs. (7) and (8). The deviatoric stress may then be expressed using Eq. (9). Here *σ* ^*max* refers to the maximum eigenvalue of the effective stress value.

$$\overline{p} = -\frac{1}{3}\overline{\sigma} : I \tag{7}$$

$$
\overline{q} = \sqrt{\frac{3}{2}\overline{\mathbb{S}} : \mathbb{S}}\tag{8}
$$

$$
\overline{S} = \overline{p}I + \overline{\sigma} \tag{9}
$$

When *σ* ^*max* <sup>¼</sup> 0 is the case, Eq. (6) reduces to the Drucker-Prager yield statement. Then the *α* coefficient defined in (6) may be determined using the initial equibiaxial (*σ<sup>b</sup>*0Þ and uniaxial compressive yield stress (*σc*0) as presented in Eq. (10). Experimental results revealed the fact that 0*:*08≤*α* ≤0*:*12 [21].

$$a = \frac{\sigma\_{b0} - \sigma\_{c0}}{2\sigma\_{b0} - \sigma\_{c0}}\tag{10}$$

The *β* ~*εpl* � � is then defined using these *α* parameters as given in Eq. (11). Here *σ<sup>t</sup>* and *σ<sup>c</sup>* are the effective tensile and compressive cohesion stresses, respectively.

$$\beta\left(\tilde{\varepsilon}^{pl}\right) = \frac{\overline{\sigma}\_{\varepsilon}\left(\tilde{\varepsilon}\_{\varepsilon}^{pl}\right)}{\overline{\sigma}\_{t}\left(\tilde{\varepsilon}\_{t}^{pl}\right)}(1-a) - (1+a) \tag{11}$$

*e. Flow rule.*

Defining a flow potential value *G*, plastic flow rule may be expressed using Eq. (12). Here \_ *λ* is a plastic multiplier which always has a positive value. Following the Kuhn-Tucker relations, which are \_ *<sup>λ</sup><sup>F</sup>* <sup>¼</sup> 0; \_ *λ*≥ 0; *F* ≤0, the *G* parameter is defined in the *σ* space.

$$
\dot{\epsilon}^{pl} = \dot{\lambda} \frac{\partial G(\overline{\sigma})}{\partial \overline{\sigma}} \tag{12}
$$

*Structural Modeling and Dynamic Analysis of a Nuclear Reactor Building DOI: http://dx.doi.org/10.5772/intechopen.94956*

Using the Drucker-Prager hyperbolic function, the *G* parameter may be defined using Eq. (13). Here *ψ* is the dilation angle at high confining pressure; *σt*<sup>0</sup> is the uniaxial tensile stress at failure; and *ϵ* refers to an eccentricity parameter.

$$\mathbf{G} = \sqrt{\left(\epsilon \sigma\_{t0} tann\nu\right)^2 + \overline{q}^2} - \overline{p}tann\nu \tag{13}$$

### *f. Damage and stiffness degradation.*

This section will be considered for both uniaxial and cyclic loading protocols.

• Uniaxial loading

<sup>~</sup>*εpl* <sup>¼</sup> <sup>~</sup>*<sup>ε</sup>*

*F σ*, ~*εpl* � �≤0.*F* can also be defined in terms of *σ*, as in Eq. (6).

<sup>1</sup> � *<sup>α</sup> <sup>q</sup>* � <sup>3</sup>*α<sup>p</sup>* <sup>þ</sup> *<sup>β</sup>* <sup>~</sup>*εpl* � � *<sup>σ</sup>*

deviatoric stress may then be expressed using Eq. (9). Here *σ*

maximum eigenvalue of the effective stress value.

and the dissipated fracture energy.

*d. Yield function.*

*Structural Integrity and Failure*

*<sup>F</sup> <sup>σ</sup>*, <sup>~</sup>*εpl* � � <sup>¼</sup> <sup>1</sup>

When *σ*

*e. Flow rule.*

Eq. (12). Here \_

**24**

defined in the *σ* space.

the Kuhn-Tucker relations, which are \_

*pl t* ~*εpl c*

; ~\_ *ε pl*

These variables control the yield surface, the degradation of the elastic stiffness

The yield function, *F σ*, ~*εpl* � �, represents a surface in effective stress space, which determines the states of failure or damage. For the inviscid plastic-damage model,

> ^*max* � � � *<sup>γ</sup>* �*<sup>σ</sup>*

*α* and *γ* in Eq. (6) are dimensionless constants. *p* and *q* are the effective hydrostatic pressure and the Von Mises equivalent effective stress values, respectively. Further these parameters can be calculated by Eqs. (7) and (8). The

> *<sup>p</sup>* ¼ � <sup>1</sup> 3

ment. Then the *α* coefficient defined in (6) may be determined using the initial equibiaxial (*σ<sup>b</sup>*0Þ and uniaxial compressive yield stress (*σc*0) as presented in Eq. (10). Experimental results revealed the fact that 0*:*08≤*α* ≤0*:*12 [21].

> *<sup>α</sup>* <sup>¼</sup> *<sup>σ</sup><sup>b</sup>*<sup>0</sup> � *<sup>σ</sup><sup>c</sup>*<sup>0</sup> 2*σ<sup>b</sup>*<sup>0</sup> � *σ<sup>c</sup>*<sup>0</sup>

and *σ<sup>c</sup>* are the effective tensile and compressive cohesion stresses, respectively.

*c* � �

Defining a flow potential value *G*, plastic flow rule may be expressed using

*<sup>λ</sup><sup>F</sup>* <sup>¼</sup> 0; \_

*λ* is a plastic multiplier which always has a positive value. Following

*σ<sup>t</sup>* ~*ε pl t*

> *ε*\_ *pl* <sup>¼</sup> \_ *λ <sup>∂</sup>G*ð Þ *<sup>σ</sup>*

*<sup>β</sup>* <sup>~</sup>*εpl* � � <sup>¼</sup> *<sup>σ</sup><sup>c</sup>* <sup>~</sup>*εpl*

The *β* ~*εpl* � � is then defined using these *α* parameters as given in Eq. (11). Here *σ<sup>t</sup>*

r

ffiffiffiffiffiffiffiffiffiffiffiffi 3 2 *S* : *S*

^*max* <sup>¼</sup> 0 is the case, Eq. (6) reduces to the Drucker-Prager yield state-

*q* ¼

^*max* � �Þ � *<sup>σ</sup><sup>c</sup>* <sup>~</sup>*εpl*

� �≤ 0 � (6)

<sup>¼</sup> *<sup>h</sup> <sup>σ</sup>*, <sup>~</sup>*εpl* � �*:ε*\_

*pl* (5)

*c*

^*max* refers to the

(8)

(10)

*σ* : *I* (7)

*S* ¼ *pI* þ *σ* (9)

� � ð Þ� <sup>1</sup> � *<sup>α</sup>* ð Þ <sup>1</sup> <sup>þ</sup> *<sup>α</sup>* (11)

*λ*≥ 0; *F* ≤0, the *G* parameter is

*<sup>∂</sup><sup>σ</sup>* (12)

" #

The assumption relies on the fact that the uniaxial stress–strain curves can be represented in terms of stress and plastic strain values using Eq. (14).

$$
\sigma\_t = \sigma\_t(\hat{\boldsymbol{\varepsilon}}\_t^{pl}, \dot{\hat{\boldsymbol{\varepsilon}}}\_t^{pl}, \theta, \boldsymbol{f}\_i), \sigma\_c = \sigma\_c(\hat{\boldsymbol{\varepsilon}}\_t^{pl}, \dot{\hat{\boldsymbol{\varepsilon}}}\_c^{pl}, \theta, \boldsymbol{f}\_i) \tag{14}
$$

In Eq. (14) the subscripts *t* and *c* refer to tension and compression, respectively; ~\_ *ε pl <sup>t</sup>* and ~\_ *ε pl <sup>c</sup>* are the equivalent plastic strain rates, ~*ε pl <sup>t</sup>* <sup>¼</sup> <sup>Ð</sup>*<sup>t</sup>* 0~\_ *ε pl <sup>t</sup> dt* and ~*εpl <sup>c</sup>* <sup>¼</sup> <sup>Ð</sup>*<sup>t</sup>* 0~\_ *ε pl <sup>c</sup> dt* are the equivalent plastic strains, *θ* is the temperature, and *fi* , (*i* ¼ 1, 2, *::*) are other predefined field variables. Under uniaxial loading conditions the effective plastic strain rates are given as: ~\_ *ε pl <sup>t</sup>* <sup>¼</sup> ~\_ *ε pl* 11, in uniaxial tension and ~\_ *ε pl <sup>c</sup>* ¼ �~\_ *ε pl* 11, in uniaxial compression (ABAQUS 2020). The response of concrete to uniaxial loading is as illustrated in **Figure 3**. The left illustration is for when the system is under effect of compression only and the right one corresponds to tension case.

To properly evaluate the degraded response of concrete two independent damage variables which are assumed as functions of the strain values are introduced: *dt* and *dc*. These variables can be calculated following Eqs. (15) and (16). If there is no damage in the material then these variables are assumed to be equal to zero, on the contrary these variables are taken as one when the material is fully damaged.

$$d\_t = d\_t\left(\widetilde{\varepsilon}\_t^{pl}, \theta, f\_i\right), (\mathbf{0} \le d\_t \le \mathbf{1}) \tag{15}$$

$$d\_c = d\_c(\tilde{\varepsilon}\_c^{pl}, \theta, f\_i), (\mathbf{0} \le d\_c \le \mathbf{1}) \tag{16}$$

Defining *EO* as the initial modulus of the material, the stress values may be determined using Eq. (17).

**Figure 3.** *Response of concrete to uniaxial loading.*

$$
\sigma\_t = (\mathbf{1} - d) E\_O \left( \varepsilon\_t - \tilde{\varepsilon}\_t^{pl} \right), \sigma\_\varepsilon = (\mathbf{1} - d) E\_O \left( \varepsilon\_\varepsilon - \tilde{\varepsilon}\_\varepsilon^{pl} \right) \tag{17}
$$

When tension loading protocol is the case, cracks propagate in a direction transverse to the stresses. Further, cracks reduce the load bearing capacity of the material and this increases the *σ* value. On the contrary, when compressive loading is the case, cracks propagate in a direction parallel to the stresses. This minimize the effect of cracks. Naming *σ<sup>t</sup>* and *σ<sup>c</sup>* as the effective uniaxial cohesion stresses values at tension and compression, respectively, Eqs. (18) and (19) may be applicable to determine the size of the yield surface.

$$\overline{\sigma}\_{t} = \frac{\sigma\_{t}}{(1 - d\_{t})} = E\_{0} \left( \varepsilon\_{t} - \widetilde{\varepsilon}\_{t}^{pl} \right), \tag{18}$$

$$\overline{\sigma}\_{\mathfrak{c}} = \frac{\sigma\_{\mathfrak{c}}}{(1 - d\_{\mathfrak{c}})} = E\_0 \left( \varepsilon\_{\mathfrak{c}} - \widetilde{\mathfrak{e}}\_{\mathfrak{c}}^{pl} \right) \tag{19}$$

### • Cyclic loading

The interaction of the cracks makes this type of loading protocol more complex to understand the behavior. Previous experiments revealed the fact that there is indeed some recovery of the stiffness and this is named as "the stiffness recovery effect (SRE)". This is a significant behavior of concrete under cyclic loading, especially when the load changes from tension to compression. To correlate *d* and *E* values, Eq. (20) has been proposed for the CDP model.

$$E = (\mathbf{1} - d)E\_0 \tag{20}$$

properly, equivalent stress *σ*<sup>0</sup> is introduced first. This stress value is indeed a function of strain and temperature, as presented in Eq. (25). Here *θ* is the considered temperature and *εpl* gives the value of the equivalent plastic strain, given in Eq. (26).

*pl* <sup>¼</sup> *<sup>σ</sup>* : *<sup>ε</sup>*´

Rice [23] brought forward the use of this model especially when the Bauschinger effect is forceful. The strain rate decomposition may be calculated from Eq. (27). When Eq. (27) is integrated through the contour of the yield surface Eq. (28) is obtained.

**Parameters Values** Elastic modulus (MPa) 3.6 � <sup>10</sup><sup>4</sup> Ultimate tensile strength (MPa) 2.85 Ultimate compressive strength (MPa) 38.5 Poisson ratio 0.2

Thermal expansion coefficient <sup>1</sup> � <sup>10</sup>�<sup>5</sup> Viscosity parameter 0.005 Dilation angle 36.31 Eccentricity 0.1 *σ<sup>b</sup>*0*=σco* 1.16 *KC* 0.667

) 2400

*σ*<sup>0</sup>*ε*´

*Structural Modeling and Dynamic Analysis of a Nuclear Reactor Building*

*DOI: http://dx.doi.org/10.5772/intechopen.94956*

**Figure 4.**

Density (kg/m<sup>3</sup>

*Parameters of the concrete material.*

**Table 2.**

**27**

*Cyclic loading protocol for ω<sup>t</sup>* ¼ 0 *and ω<sup>c</sup>* ¼ 1*.*

*<sup>f</sup>*ð Þ¼ *<sup>σ</sup> <sup>σ</sup>*<sup>0</sup> *<sup>ε</sup>pl*, *<sup>θ</sup>* (25)

*<sup>d</sup><sup>ε</sup>* <sup>¼</sup> *<sup>d</sup>εel* <sup>þ</sup> *<sup>d</sup>εpl* (27)

*pl* (26)

It is previously shown that *d* may be related with *dt* and *dc*values. When cyclic loading protocol is applied ABAQUS follows the relation given in Eq. (21). In Eq. (21) *st* and *sc* are stress functions which includes the SRE associated with stress reversals.

$$(\mathbf{1} - d) = (\mathbf{1} - \mathfrak{s}\_t d\_c)(\mathbf{1} - \mathfrak{s}\_c d\_t), \mathbf{0} \le \mathfrak{s}\_t, \mathfrak{s}\_c \ge \mathbf{1} \tag{21}$$

Introducing *ω<sup>t</sup>* and *ω<sup>c</sup>* material weight factors responsible from the tensile and compressive stiffness recovery, Eqs. (22) and (23) is proposed. The constant *r* in Eqs. (22) and (23) is determined from Eq. (24).

$$\mathfrak{s}\_{l} = \mathfrak{1} - \alpha\_{l} r \* (\overline{\sigma}\_{11}); \mathbf{0} \le \alpha\_{l} \le \mathbf{1} \tag{22}$$

$$\mathfrak{s}\_{\mathfrak{c}} = \mathfrak{1} - a\_{\mathfrak{c}} (\mathfrak{1} - r \ast (\overline{\sigma}\_{11})); \mathbb{0} \le a\_{\mathfrak{c}} \le \mathfrak{1} \tag{23}$$

$$r\*(\overline{\sigma}\_{11}) = H(\overline{\sigma}\_{11}) = \begin{cases} 1 \text{ if } \overline{\sigma}\_{11} > 0 \\ 0 \text{ if } \overline{\sigma}\_{11} < 0 \end{cases} \tag{24}$$

This observation in concrete may be accepted as a proof of SRE when cracks close as the load goes from tension to compression. When vice versa is valid, then the tensile stiffness is not recovered. This is ensured in ABAQUS by assuming *ω<sup>t</sup>* ¼ 0 and *ω<sup>c</sup>* ¼ 1. Cyclic load behavior is illustrated in **Figure 4**.

### **3.2 Constitutive model for steel**

Isotropic hardening model is used to simulate the elastic–plastic steel behavior. As well as the rebars, prestressed tendons are defined using steel material. This indicates that the yield surface changes in all directions. To define the yield function *f*ð Þ *σ*

*Structural Modeling and Dynamic Analysis of a Nuclear Reactor Building DOI: http://dx.doi.org/10.5772/intechopen.94956*

**Figure 4.** *Cyclic loading protocol for ω<sup>t</sup>* ¼ 0 *and ω<sup>c</sup>* ¼ 1*.*

*σ<sup>t</sup>* ¼ ð Þ 1 � *d EO ε<sup>t</sup>* � ~*ε*

*pl t* 

When tension loading protocol is the case, cracks propagate in a direction transverse to the stresses. Further, cracks reduce the load bearing capacity of the material and this increases the *σ* value. On the contrary, when compressive loading is the case, cracks propagate in a direction parallel to the stresses. This minimize the effect of cracks. Naming *σ<sup>t</sup>* and *σ<sup>c</sup>* as the effective uniaxial cohesion stresses values at tension and compression, respectively, Eqs. (18) and

¼ *E*<sup>0</sup> *ε<sup>t</sup>* � ~*ε*

The interaction of the cracks makes this type of loading protocol more complex to understand the behavior. Previous experiments revealed the fact that there is indeed some recovery of the stiffness and this is named as "the stiffness recovery effect (SRE)". This is a significant behavior of concrete under cyclic loading, especially when the load changes from tension to compression. To correlate *d*

It is previously shown that *d* may be related with *dt* and *dc*values. When cyclic loading protocol is applied ABAQUS follows the relation given in Eq. (21). In Eq. (21) *st* and *sc* are stress functions which includes the SRE associated with

Introducing *ω<sup>t</sup>* and *ω<sup>c</sup>* material weight factors responsible from the tensile and compressive stiffness recovery, Eqs. (22) and (23) is proposed. The constant *r* in

This observation in concrete may be accepted as a proof of SRE when cracks close as the load goes from tension to compression. When vice versa is valid, then the tensile stiffness is not recovered. This is ensured in ABAQUS by assuming

Isotropic hardening model is used to simulate the elastic–plastic steel behavior. As well as the rebars, prestressed tendons are defined using steel material. This indicates that the yield surface changes in all directions. To define the yield function *f*ð Þ *σ*

<sup>¼</sup> *<sup>E</sup>*<sup>0</sup> *<sup>ε</sup><sup>c</sup>* � <sup>~</sup>*εpl*

(19) may be applicable to determine the size of the yield surface.

ð Þ 1 � *dc*

*<sup>σ</sup><sup>t</sup>* <sup>¼</sup> *<sup>σ</sup><sup>t</sup>* ð Þ 1 � *dt*

*<sup>σ</sup><sup>c</sup>* <sup>¼</sup> *<sup>σ</sup><sup>c</sup>*

and *E* values, Eq. (20) has been proposed for the CDP model.

Eqs. (22) and (23) is determined from Eq. (24).

*r* ∗ ð Þ¼ *σ*<sup>11</sup> *H*ð Þ¼ *σ*<sup>11</sup>

*ω<sup>t</sup>* ¼ 0 and *ω<sup>c</sup>* ¼ 1. Cyclic load behavior is illustrated in **Figure 4**.

• Cyclic loading

*Structural Integrity and Failure*

stress reversals.

**3.2 Constitutive model for steel**

**26**

, *<sup>σ</sup><sup>c</sup>* <sup>¼</sup> ð Þ <sup>1</sup> � *<sup>d</sup> EO <sup>ε</sup><sup>c</sup>* � <sup>~</sup>*εpl*

*pl t* 

*c*

*E* ¼ ð Þ 1 � *d E*<sup>0</sup> (20)

ð Þ¼ 1 � *d* ð Þ 1 � *stdc* ð Þ 1 � *scdt* , 0≤ *st*, *sc* ≥1 (21)

*st* ¼ 1 � *ωtr* ∗ ð Þ *σ*<sup>11</sup> ; 0 ≤*ω<sup>t</sup>* ≤ 1 (22)

*sc* ¼ 1 � *ωc*ð Þ 1 � *r* ∗ ð Þ *σ*<sup>11</sup> ; 0 ≤*ω<sup>c</sup>* ≤1 (23)

1 *if σ*<sup>11</sup> >0 0 *if σ*<sup>11</sup> <0

(24)

*c*

(17)

, (18)

(19)

properly, equivalent stress *σ*<sup>0</sup> is introduced first. This stress value is indeed a function of strain and temperature, as presented in Eq. (25). Here *θ* is the considered temperature and *εpl* gives the value of the equivalent plastic strain, given in Eq. (26).

$$f(\sigma) = \sigma^0(\varepsilon^{pl}, \theta) \tag{25}$$

$$
\sigma^0 \acute{\epsilon}^{pl} = \sigma : \acute{\epsilon}^{pl} \tag{26}
$$

Rice [23] brought forward the use of this model especially when the Bauschinger effect is forceful. The strain rate decomposition may be calculated from Eq. (27). When Eq. (27) is integrated through the contour of the yield surface Eq. (28) is obtained.

$$d\varepsilon = d\varepsilon^{cl} + d\varepsilon^{pl} \tag{27}$$


### **Table 2.**

*Parameters of the concrete material.*


**Table 3.**

*Parameters of the steel material.*

$$
\varepsilon = \varepsilon^{el} + \varepsilon^{pl} \tag{28}
$$

probabilistic framework, which provides a robust methodology to integrate hazard curves, component fragility curves and consequence functions and to capture the dispersions in each of these elements for evaluating the performance of a building. Importantly, the fragility curves used in the analysis are defined in terms of struc-

> **B Moderate permanent distortion**

> > SDB-1B ASCE7

> > SDB-2B ASCE7

SDB-3B ASCE 43-05

SDB-4B ASCE 43-05

SDB-5B ASCE 43-05

**C Limited permanent distortion**

> SDB-1C ASCE7

> SDB-2C NA

SDB-3C ASCE 43-05

SDB-4C ASCE 43-05

SDB-5C ASCE 43-05

**D Essentially elastic**

> SDB-1D NA

> SDB-2D NA

SDB-3D ASCE 43-05

SDB-4D ASCE 43-05

SDB-5D ASCE 43-05

tural response parameters.

**1** SDB-1A

**2** SDB-2A

**3** SDB-3A

**4** SDB-4A

**5** SDB-5A

**Table 4.**

**Figure 5.**

**29**

**SDC Limit state**

*Structural Modeling and Dynamic Analysis of a Nuclear Reactor Building*

**A Large permanent distortion (short of collapse)**

*DOI: http://dx.doi.org/10.5772/intechopen.94956*

ASCE7

ASCE7

ASCE 43-05

ASCE 43-05

ASCE 43-05

*Seismic design procedure defined in ASCE 43-05 [18] for SDC 3, 4, and 5.*

*Graded approach defined in ASCE 43-05 [18].*

The elasticity can be then written in terms of two temperature-dependent material parameters. These parameters are generally selected as the bulk modulus, *K*, and the shear modulus, *G*, which are given Eqs. (29) and (30).

$$K = \frac{E}{\Im(1 - 2\nu)}\tag{29}$$

$$G = \frac{E}{2(1+\nu)}\tag{30}$$

The constitutive material parameters used in this study for concrete and steel are as given in **Tables 2** and **3**, respectively.

### **4. Seismic performance assessment of the selected reactor containment building**

Two basic regulations for NPP's may be named as the 10 CFR 50 and 10 CFR 100 which are published by the Nuclear Regulatory Commission (USNRC) [24, 25]. The former one covers the licensing issues and the latter one explains the steps of evaluating a license. In the latter regulation, two basic earthquakes are defined when performing a seismic hazard assessment: safe shutdown earthquake (SSE) and operating basis earthquake (OBE). While SSE is considered to be the maximum earthquake which could occur at the investigation site, OBE is defined as the earthquake which could be expected to occur at the site during the lifetime of the plant. Further in the regulation, it is stated that the maximum vibratory ground acceleration of the OBE must be at least 33% of the maximum vibratory ground acceleration of the SSE [24, 25].

Nowadays, the methodology for risk analysis involves the use of component fragility curves developed using ground-motion parameters. To obtain the fragility curves the capacity and the demand parameters should be evaluated at first instance. It is a well-known fact that failure of both structural and nonstructural components of an NPP are much more involved with the structural response than the ground parameters.

The FEMA P-58-1 Guidelines [26] provide a basis to improve the risk assessment procedure for NPPs. The guideline develops next-generation tools and new approaches for performance assessment of buildings, with a focus on measuring performance in terms of direct economic loss, casualties and downtime. The FEMA P-58 Guidelines [26] present procedures for performance assessment using a

### *Structural Modeling and Dynamic Analysis of a Nuclear Reactor Building DOI: http://dx.doi.org/10.5772/intechopen.94956*

probabilistic framework, which provides a robust methodology to integrate hazard curves, component fragility curves and consequence functions and to capture the dispersions in each of these elements for evaluating the performance of a building. Importantly, the fragility curves used in the analysis are defined in terms of structural response parameters.


### **Table 4.**

*<sup>ε</sup>* <sup>¼</sup> *<sup>ε</sup>el* <sup>þ</sup> *<sup>ε</sup>pl* (28)

3 1ð Þ � <sup>2</sup>*<sup>υ</sup>* (29)

2 1ð Þ <sup>þ</sup> *<sup>υ</sup>* (30)

The elasticity can be then written in terms of two temperature-dependent material parameters. These parameters are generally selected as the bulk modulus,

**Parameters Rebars Tendons** Elastic modulus (MPa) <sup>2</sup> � <sup>10</sup><sup>5</sup> 1.95 � <sup>10</sup><sup>5</sup> Yield strength (MPa) 486.55 1860 Yield strain (m/m) 0.002613 0.008745 Poisson ratio 0.3 0.3

Thermal expansion coefficient <sup>1</sup> � <sup>10</sup>�<sup>5</sup> <sup>1</sup> � <sup>10</sup>�<sup>5</sup>

) 7850 7850

*<sup>K</sup>* <sup>¼</sup> *<sup>E</sup>*

*<sup>G</sup>* <sup>¼</sup> *<sup>E</sup>*

The constitutive material parameters used in this study for concrete and steel are

**4. Seismic performance assessment of the selected reactor containment**

Nowadays, the methodology for risk analysis involves the use of component fragility curves developed using ground-motion parameters. To obtain the fragility curves the capacity and the demand parameters should be evaluated at first instance. It is a well-known fact that failure of both structural and nonstructural components of an NPP are much more involved with the structural response than

The FEMA P-58-1 Guidelines [26] provide a basis to improve the risk assessment

procedure for NPPs. The guideline develops next-generation tools and new approaches for performance assessment of buildings, with a focus on measuring performance in terms of direct economic loss, casualties and downtime. The FEMA P-58 Guidelines [26] present procedures for performance assessment using a

Two basic regulations for NPP's may be named as the 10 CFR 50 and 10 CFR 100 which are published by the Nuclear Regulatory Commission (USNRC) [24, 25]. The former one covers the licensing issues and the latter one explains the steps of evaluating a license. In the latter regulation, two basic earthquakes are defined when performing a seismic hazard assessment: safe shutdown earthquake (SSE) and operating basis earthquake (OBE). While SSE is considered to be the maximum earthquake which could occur at the investigation site, OBE is defined as the earthquake which could be expected to occur at the site during the lifetime of the plant. Further in the regulation, it is stated that the maximum vibratory ground acceleration of the OBE must be at least 33% of the maximum vibratory ground

*K*, and the shear modulus, *G*, which are given Eqs. (29) and (30).

as given in **Tables 2** and **3**, respectively.

acceleration of the SSE [24, 25].

the ground parameters.

**28**

**building**

Density (kg/m3

*Parameters of the steel material.*

*Structural Integrity and Failure*

**Table 3.**

*Graded approach defined in ASCE 43-05 [18].*

### **Figure 5.**

*Seismic design procedure defined in ASCE 43-05 [18] for SDC 3, 4, and 5.*


**Table 5.**

*Structural deformation limits for limit state defined in ASCE 43-05 [18].*

ASCE 43-05 [18] provides seismic design criteria for structures, systems, and components (SSC) that are used in nuclear facilities. This is executed by using a graded approach where the design criteria are proportional with the relative importance to safety, magnitude of the seismic event and other factors. To achieve this 20 Seismic Design Bases (SDB) are defined, where for each Seismic Design Category (SDC) probabilistic target performance goals are used. This graded approach is summarized in **Table 4**. Further in the same guide for each Limit State (LS) an acceptable level of structural response goal is defined. While International Building Code (IBC) [27] may be followed for SDBs defined by SDC1 and 2; for SDBs defined by SDC 3, 4, and 5, ASCE standard should be followed as illustrated in **Figure 5**. Further in the mentioned code, the Design Basis Earthquake (DBE) is defined by Uniform Hazard Response Spectra (UHRS).

For DBE, Limit State A and D are defined as the intensity of the high and low structural damage, respectively. In other words, Limit State D is where complete elastic behavior is dominant. Damage regions between A and D are called the intermediate levels. The deformation limits associated with each Limit State are described in **Table 5**.

To properly determine the seismic demand, ASCE 43-05 [18] allows use of *linear equivalent static analysis*, *linear dynamic analysis*, *complex frequency response methods*, or *nonlinear analysis.* Moreover, when the fundamental mode is dominant then a single step pushover methodology is allowed while conducting nonlinear analysis. These procedures shall be conducted by following the criteria given in FEMA-356 [28] or ATC-40 [29]. This should be clarified that nonlinear analysis is only permitted when beyond design earthquake is considered for low-rise regular NPPs, where higher mode effects are negligible. Otherwise, elastic models are preferred in the seismic design of safety-related structures.

### **5. Results of the conducted analysis**

The 3D finite element model has been subjected to gravity, wind and snow loads in accordance with Eurocode 1 [30], Section 4. First, to determine the fundamental periods and vibration modes of the considered structure, eigenvalue analyses are performed and eigenvectors have been determined. **Figure 6** shows the results of the modal analysis considering first six modes.

As it can be inferred from **Figure 6**, first three modal results reveal the fact that the dome of the reactor building is the most vulnerable section. The displacement value of the dome is calculated approximately 1 mm for the fundamental mode. When considered the higher modes, stress concentrations around operational gaps are observed.

After determining the mode shapes of the outer containment vessel, nonlinear dynamic time history analyses are conducted to assess the seismic behavior of the structure and the stress distribution. For this purpose, PEER database has been used and seven earthquake excitations are first selected in accordance with ATC 58 [31] and scaled to fit Eurocode 8 [32] Soil type B design spectrum. **Figure 7** compares

*Structural Modeling and Dynamic Analysis of a Nuclear Reactor Building*

*DOI: http://dx.doi.org/10.5772/intechopen.94956*

**Figure 6.**

**31**

*Calculated mode shapes and periods.*

*Structural Modeling and Dynamic Analysis of a Nuclear Reactor Building DOI: http://dx.doi.org/10.5772/intechopen.94956*

**Figure 6.** *Calculated mode shapes and periods.*

After determining the mode shapes of the outer containment vessel, nonlinear dynamic time history analyses are conducted to assess the seismic behavior of the structure and the stress distribution. For this purpose, PEER database has been used and seven earthquake excitations are first selected in accordance with ATC 58 [31] and scaled to fit Eurocode 8 [32] Soil type B design spectrum. **Figure 7** compares

ASCE 43-05 [18] provides seismic design criteria for structures, systems, and components (SSC) that are used in nuclear facilities. This is executed by using a graded approach where the design criteria are proportional with the relative importance to safety, magnitude of the seismic event and other factors. To achieve this 20 Seismic Design Bases (SDB) are defined, where for each Seismic Design Category (SDC) probabilistic target performance goals are used. This graded approach is summarized in **Table 4**. Further in the same guide for each Limit State (LS) an acceptable level of structural response goal is defined. While International Building Code (IBC) [27] may be followed for SDBs defined by SDC1 and 2; for SDBs defined by SDC 3, 4, and 5, ASCE standard should be followed as illustrated in **Figure 5**. Further in the mentioned code, the Design Basis Earthquake (DBE) is defined by

**Limit state Structural deformation limit**

**B** Moderate permanent distortion

**C** Limited permanent distortion

**D** Essentially elastic behavior

*Structural deformation limits for limit state defined in ASCE 43-05 [18].*

**A** Large permanent distortion, short of collapse

*Significant damage*

*Minimal damage*

*No damage*

*Generally repairable damage*

For DBE, Limit State A and D are defined as the intensity of the high and low structural damage, respectively. In other words, Limit State D is where complete elastic behavior is dominant. Damage regions between A and D are called the intermediate levels. The deformation limits associated with each Limit State are

To properly determine the seismic demand, ASCE 43-05 [18] allows use of *linear equivalent static analysis*, *linear dynamic analysis*, *complex frequency response methods*, or *nonlinear analysis.* Moreover, when the fundamental mode is dominant then a single step pushover methodology is allowed while conducting nonlinear analysis. These procedures shall be conducted by following the criteria given in FEMA-356 [28] or ATC-40 [29]. This should be clarified that nonlinear analysis is only permitted when beyond design earthquake is considered for low-rise regular NPPs, where higher mode effects are negligible. Otherwise, elastic models are preferred in

The 3D finite element model has been subjected to gravity, wind and snow loads in accordance with Eurocode 1 [30], Section 4. First, to determine the fundamental periods and vibration modes of the considered structure, eigenvalue analyses are performed and eigenvectors have been determined. **Figure 6** shows the results of

As it can be inferred from **Figure 6**, first three modal results reveal the fact that the dome of the reactor building is the most vulnerable section. The displacement value of the dome is calculated approximately 1 mm for the fundamental mode. When considered the higher modes, stress concentrations around operational gaps

Uniform Hazard Response Spectra (UHRS).

the seismic design of safety-related structures.

the modal analysis considering first six modes.

**5. Results of the conducted analysis**

described in **Table 5**.

**Table 5.**

*Structural Integrity and Failure*

are observed.

**30**

**Figure 7.** *Comparison of Eurocode 8 design spectrum with the selected ground motions spectra.*


**Figure 9.**

**Figure 10.**

**33**

*concentration.*

*and (c) total stress concentration.*

*Nonlinear dynamic time history results of the outer containment: (a) total displacement, (b) total deformation*

*Structural Modeling and Dynamic Analysis of a Nuclear Reactor Building*

*DOI: http://dx.doi.org/10.5772/intechopen.94956*

*Nonlinear dynamic time history results of the prestressed tendons: (a) total displacement and (b) total stress*

### **Table 6.**

*Details of the selected strong ground motions.*

Eurocode 8 [32] design spectrum with the spectra of the selected ground motions. The details of the selected strong ground motions are presented in **Table 6**. Following, an artificial time history record which is compatible with the mean spectrum has been generated and used in the nonlinear dynamic analyses. **Figure 8** represents the artificial ground motion.

### *Structural Modeling and Dynamic Analysis of a Nuclear Reactor Building DOI: http://dx.doi.org/10.5772/intechopen.94956*

**Figure 9.**

*Nonlinear dynamic time history results of the outer containment: (a) total displacement, (b) total deformation and (c) total stress concentration.*

### **Figure 10.**

*Nonlinear dynamic time history results of the prestressed tendons: (a) total displacement and (b) total stress concentration.*

Eurocode 8 [32] design spectrum with the spectra of the selected ground motions. The details of the selected strong ground motions are presented in **Table 6**. Following, an artificial time history record which is compatible with the mean spectrum has been generated and used in the nonlinear dynamic analyses. **Figure 8**

represents the artificial ground motion.

**Figure 7.**

**Table 6.**

**Figure 8.**

**32**

*Artificial ground motion.*

*Comparison of Eurocode 8 design spectrum with the selected ground motions spectra.*

3 Loma Prieta (1989) Coyote Lake Dam -

6 Landers (1992) North Palm Springs Fire

7 Chuetsu-oki Japan (2007)

*Structural Integrity and Failure*

*Details of the selected strong ground motions.*

**No Earthquake Station Mw Mechanism Vs30**

Southwest Abutment

Sta #36

4 Chi-Chi Taiwan (1999) CHY010 7.63 Reverse

1 Imperial Valley (1979) Cerro Prieto 6.53 Strike slip 471.53 15.19 — 2 Irpinia Italy (1980) Rionero In Vulture 6.2 Normal 574.88 22.68 22.69

5 Chi-Chi Taiwan (1999) TCU075 6.3 Reverse 573.02 24.34 26.31

**(m/s)**

7.28 Strike slip 367.84 26.95 26.95

6.93 Reverse Oblique

Matsushiro Tokamachi 6.8 Reverse 640.14 18.16 25.03

Oblique

**Rjb (km)**

561.43 19.97 20.34

538.69 19.93 19.96

**Rrup (km)**

Modified Newton–Raphson iterative procedure has been followed during the dynamic analysis. ABAQUS offers several convergence norms. Besides stopping the iteration in case of convergence, the iteration process is also stopped if a specified maximum number of iterations has been reached or if the iteration obviously leads to divergence. The detection of divergence is based on the same norms as the detection of convergence. A preferred way to check the convergence is the energy norm.

As the consequence of dynamic analysis total displacement, total deformation and total stress concentration values are illustrated in **Figure 9**. It should also be highlighted that results of the modal analysis are in good correlation with the dynamic results and the most vulnerable section is assessed as the dome of the structure. In **Figure 9(a)**, total displacement graph has been presented. The maximum displacement value has been calculated as 7.625 mm at the roof. The deformation graph, which is represented in **Figure 9(b)**, is compatible with the stress concentration that is represented in **Figure 9(c)**. Total displacement and stress concentration graphs of the prestressed tendons are given in **Figure 10(a)** and **(b)**, respectively.

### **6. Conclusions**

In this study, seismic behavior of the outer containment of a sample reactor building has been analyzed. First, details of the 3D model have been presented, and material models are described in detail. Following that, results of the eigenvalue analysis are discussed. Then, nonlinear time history analyses are conducted using the artificial record which is compatible with the mean design spectrum of the selected ground motions that are previously scaled to Eurocode 8 Soil Type B design spectrum.

The primary outcome of the study is that the most critical section of a reactor building may be assessed as the dome of the structure. More care should be given while designing the dome. In the recent literature, to overcome the deficiencies of prestressed concrete, use of fiber reinforced concrete is proposed due to the fact that these fibers in plain concrete directly effects both the ultimate capacity and post-cracking behavior of a conventional prestressed containment. It may also be interpreted from these ongoing studies that the structural performance of components of an NPP improves when these fibers are used with plain concrete.

It is believed that the study should be developed to consider inelastic behavior of soil and soil-structure interaction effects should be taken into account in the dynamic analysis.

**Author details**

**35**

Evrim Oyguc\*, Abdul Hayır and Resat Oyguc Istanbul Technical University, Istanbul, Turkey

provided the original work is properly cited.

\*Address all correspondence to: eoyguc@itu.edu.tr

*Structural Modeling and Dynamic Analysis of a Nuclear Reactor Building*

*DOI: http://dx.doi.org/10.5772/intechopen.94956*

© 2020 The Author(s). Licensee IntechOpen. This chapter is distributed under the terms of the Creative Commons Attribution License (http://creativecommons.org/licenses/ by/3.0), which permits unrestricted use, distribution, and reproduction in any medium,

### **Acknowledgements**

This paper is generated from the PhD study of the corresponding author. The author is grateful for the endless support of her supervisor.

### **Conflict of interest**

The authors declare no conflict of interest.

*Structural Modeling and Dynamic Analysis of a Nuclear Reactor Building DOI: http://dx.doi.org/10.5772/intechopen.94956*

### **Author details**

Modified Newton–Raphson iterative procedure has been followed during the dynamic analysis. ABAQUS offers several convergence norms. Besides stopping the iteration in case of convergence, the iteration process is also stopped if a specified maximum number of iterations has been reached or if the iteration obviously leads to divergence. The detection of divergence is based on the same norms as the detection of convergence. A preferred way to check the convergence is the energy norm.

As the consequence of dynamic analysis total displacement, total deformation and

highlighted that results of the modal analysis are in good correlation with the dynamic results and the most vulnerable section is assessed as the dome of the structure. In **Figure 9(a)**, total displacement graph has been presented. The maximum displacement value has been calculated as 7.625 mm at the roof. The deformation graph, which is represented in **Figure 9(b)**, is compatible with the stress concentration that is represented in **Figure 9(c)**. Total displacement and stress concentration graphs of

In this study, seismic behavior of the outer containment of a sample reactor building has been analyzed. First, details of the 3D model have been presented, and material models are described in detail. Following that, results of the eigenvalue analysis are discussed. Then, nonlinear time history analyses are conducted using the artificial record which is compatible with the mean design spectrum of the selected ground motions that are previously scaled to Eurocode 8 Soil Type B design

The primary outcome of the study is that the most critical section of a reactor building may be assessed as the dome of the structure. More care should be given while designing the dome. In the recent literature, to overcome the deficiencies of prestressed concrete, use of fiber reinforced concrete is proposed due to the fact that these fibers in plain concrete directly effects both the ultimate capacity and post-cracking behavior of a conventional prestressed containment. It may also be interpreted from these ongoing studies that the structural performance of compo-

It is believed that the study should be developed to consider inelastic behavior of

This paper is generated from the PhD study of the corresponding author. The

nents of an NPP improves when these fibers are used with plain concrete.

author is grateful for the endless support of her supervisor.

The authors declare no conflict of interest.

soil and soil-structure interaction effects should be taken into account in the

total stress concentration values are illustrated in **Figure 9**. It should also be

the prestressed tendons are given in **Figure 10(a)** and **(b)**, respectively.

**6. Conclusions**

*Structural Integrity and Failure*

spectrum.

dynamic analysis.

**Acknowledgements**

**Conflict of interest**

**34**

Evrim Oyguc\*, Abdul Hayır and Resat Oyguc Istanbul Technical University, Istanbul, Turkey

\*Address all correspondence to: eoyguc@itu.edu.tr

© 2020 The Author(s). Licensee IntechOpen. This chapter is distributed under the terms of the Creative Commons Attribution License (http://creativecommons.org/licenses/ by/3.0), which permits unrestricted use, distribution, and reproduction in any medium, provided the original work is properly cited.

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[32] Eurocode 8. Design of structures for earthquake resistance. Part 1. General Rules. Specific Rules, Seismic Actions and Rules for Buildings. Belgium:

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[30] Eurocode 1. Design of structures for earthquake resistance. Part 1. General Rules. Specific Rules, Seismic Actions and Rules for Buildings. Belgium:

[28] FEMA 356. Prestandard and Commentary for the Seismic

Club Hills. In: IL. 2003

Reston; 2000

[18] ASCE 43-05. Seismic Design Criteria for Structures, Systems, and Components in Nuclear Facilities, American Society of Civil Engineers. 2005. Alexander bell drive, Virginia. In: U.S.A. 1801

[19] Hessheimer MF, Dameron RA. Containment Integrity Research at Sandia National Laboratories – An Overview, Technical Report NUREG/ CR-6906; SAND2006-2274P. Sandia National Laboratories and USNRC,

[20] Hessheimer MF, Klamerus EW, Lambert LD, Rightley GS, Dameron RA. Over Pressurization Test of a 1:4-Scale Prestressed Concrete Containment Vessel Model, Technical Report National/CR-6810; SAND2003-0840p. Washington, USA: Sandia National Laboratories and USNRC; 2006

[21] Lubliner J, Oliver J, Oller S. Onate E. a plastic-damage model for concrete. International Journal of Solids and Structures. 1989;**25**(3):229-326

[22] Lee J, Fenves GL. Plastic-damage model for cyclic loading of concrete structures. Journal of Engineering Mechanics. 1998;**124**(8):892-900

[23] Rice JR. Continuum Mechanics and Thermodynamics of Plasticity in Relation to Microscale Deformation Mechanisms. Constitutive Equations in

Plasticity: Argon, MIT Press, Cambridge, Massachusettes; 1975

[24] USNRC, Individual Plant Examination of External Events, (IPEEE) for Severe Accident Vulnerabilities, Generic Letter No. 88-20, Supplement 4. 1991. US Nuclear

Regulatory Commission.

**37**

[25] USNRC. Procedural and Submittal Guidance for the Individual Plant

Washington. In: USA. 2006

Virginia, U.S.A.

[9] Park J, Park N, Lee S, Park Y, Choi Y. Seismic analysis of the APR1400 nuclear reactor system using a verified beam element model. Nuclear Engineering and Design. 2017;**313**:108-117

[10] Ambrosini D, Codina RH, Curadelli O, Martinez CA. Structural analysis of the CAREM-25 nuclear power plant subjected to the design basis accident and seismic loads. Annals of Nuclear Energy. 2017;**108**:42-56

[11] Tunon-Sanjur L, Orr RS, Tinic S, Ruiz DP. Finite element modeling of the AP1000 nuclear island for seismic analyses at generic soil and rock sites. Nuclear Engineering and Design. 2007; **237**:1474-1485

[12] De Boeck B. A review of containment accidents. Nuclear Engineering Design. 1993;**145**:279-288

[13] Kwak HG, Kim JH. Numerical models for prestressing tendons in containment structures. Nuclear Engineering and Design. 2006;**236**:1061-1080

[14] fib Task Group on Containment Structures. Nuclear containments. In: International Federation for Structural Concrete (Fib). 2001

[15] Sandia National Laboratories. Pretest round Robin analysis of a prestressed concrete containment vessel model. U.S. Nuclear Regulatory Commission (NRC) and nuclear power engineering corporation (NUPEC). In: NUREG/CR-6678. 2000

[16] TBEC TBEC. Ministry of Public Work and Settlement. Turkey: Ankara; 2018

[17] ASCE 4-16. Seismic Analysis of Safety-Related Nuclear Structures.

*Structural Modeling and Dynamic Analysis of a Nuclear Reactor Building DOI: http://dx.doi.org/10.5772/intechopen.94956*

American Society of Civil Engineers. 2017. 1801 Alexander Bell Drive, Virginia, U.S.A.

**References**

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207-225

1421-1442

**235**(13):1335-1348

**237**(12-13):1474-1485

**36**

[1] Kessides IN. The future of the nuclear industry reconsidered: Risks, uncertainties, and continued promise. Energy Policy. 2012;**48**:185-208

*Structural Integrity and Failure*

containment. Nuclear Engineering and

[9] Park J, Park N, Lee S, Park Y, Choi Y. Seismic analysis of the APR1400 nuclear reactor system using a verified beam element model. Nuclear Engineering and Design. 2017;**313**:108-117

Design. 2018;**328**:197-208

[10] Ambrosini D, Codina RH, Curadelli O, Martinez CA. Structural analysis of the CAREM-25 nuclear power plant subjected to the design basis accident and seismic loads. Annals of Nuclear Energy. 2017;**108**:42-56

[11] Tunon-Sanjur L, Orr RS, Tinic S, Ruiz DP. Finite element modeling of the AP1000 nuclear island for seismic analyses at generic soil and rock sites. Nuclear Engineering and Design. 2007;

[13] Kwak HG, Kim JH. Numerical models for prestressing tendons in containment structures. Nuclear Engineering and Design. 2006;**236**:1061-1080

[14] fib Task Group on Containment Structures. Nuclear containments. In: International Federation for Structural

[15] Sandia National Laboratories. Pretest round Robin analysis of a prestressed concrete containment vessel

model. U.S. Nuclear Regulatory

NUREG/CR-6678. 2000

2018

Commission (NRC) and nuclear power engineering corporation (NUPEC). In:

[16] TBEC TBEC. Ministry of Public Work and Settlement. Turkey: Ankara;

[17] ASCE 4-16. Seismic Analysis of Safety-Related Nuclear Structures.

Concrete (Fib). 2001

[12] De Boeck B. A review of containment accidents. Nuclear Engineering Design. 1993;**145**:279-288

**237**:1474-1485

[2] Crimello R. Achievements and prospects for advanced reactor design and fuel cycles. IAEA – Scientific forum. 2004. In: Nuclear Fuel Cycles Issues and

[3] Varma V, Reddy GR, Vaze KK, Kushwaha HS. Simplified approach for

[4] Huang YN, Whittaker AS, Luco N. Seismic performance assessment of base-isolated safety-related nuclear structures. Earthquake Engineering and Structural Dynamics. 2010;**39**(13):

[5] Hirama T, Goto M, Hasegawa T, Kanechika M, Kei T, Mieda T, et al. Seismic proof test of a reinforced concrete containment vessel (RCCV): Part 1: Test model and pressure test. Nuclear Engineering and Design. 2005;

[6] Leonardo TS, Richard SO, Sener T, Diego PR. Finite element modeling of the AP1000 nuclear island for seismic analyses at generic soil and rock sites. Nuclear Engineering and Design. 2007;

[7] Jeremic B, Tafazzoli N, Ancheta T, Orboviæ N, Blahoianu A. Seismic behavior of NPP structures subjected to realistic 3D, inclined seismic motions, in variable layered soil/rock, on surface or

Engineering and Design. 2013;**265**:85-94

[8] Bily P, Kohoutkova A. An estimation of the effect of steel liner on the ultimate bearing capacity of prestressed concrete

embedded foundations. Nuclear

seismic analysis of structures. International Journal of Structural Stability and Dynamics. 2002;**2**(2): [18] ASCE 43-05. Seismic Design Criteria for Structures, Systems, and Components in Nuclear Facilities, American Society of Civil Engineers. 2005. Alexander bell drive, Virginia. In: U.S.A. 1801

[19] Hessheimer MF, Dameron RA. Containment Integrity Research at Sandia National Laboratories – An Overview, Technical Report NUREG/ CR-6906; SAND2006-2274P. Sandia National Laboratories and USNRC, Washington. In: USA. 2006

[20] Hessheimer MF, Klamerus EW, Lambert LD, Rightley GS, Dameron RA. Over Pressurization Test of a 1:4-Scale Prestressed Concrete Containment Vessel Model, Technical Report National/CR-6810; SAND2003-0840p. Washington, USA: Sandia National Laboratories and USNRC; 2006

[21] Lubliner J, Oliver J, Oller S. Onate E. a plastic-damage model for concrete. International Journal of Solids and Structures. 1989;**25**(3):229-326

[22] Lee J, Fenves GL. Plastic-damage model for cyclic loading of concrete structures. Journal of Engineering Mechanics. 1998;**124**(8):892-900

[23] Rice JR. Continuum Mechanics and Thermodynamics of Plasticity in Relation to Microscale Deformation Mechanisms. Constitutive Equations in Plasticity: Argon, MIT Press, Cambridge, Massachusettes; 1975

[24] USNRC, Individual Plant Examination of External Events, (IPEEE) for Severe Accident Vulnerabilities, Generic Letter No. 88-20, Supplement 4. 1991. US Nuclear Regulatory Commission.

[25] USNRC. Procedural and Submittal Guidance for the Individual Plant

Examination of External Events (IPEEE) for Severe Accident Vulnerabilities, NUREG-1407. 1991. US Nuclear Regulatory Commission.

[26] FEMA P-58-1. Seismic Performance Assessment of Buildings Volume 1- Methodology. Applied Technology Council. California: Redwood City; 2012

[27] IBC. International Building Code. International Code Council. Country Club Hills. In: IL. 2003

[28] FEMA 356. Prestandard and Commentary for the Seismic Rehabilitation of Buildings. American Society of Civil Engineers. Virginia: Reston; 2000

[29] ATC 40. Seismic Evaluation and Retrofit of Concrete Buildings Volume 1. Applied Technology Council. California: Redwood City; 1996

[30] Eurocode 1. Design of structures for earthquake resistance. Part 1. General Rules. Specific Rules, Seismic Actions and Rules for Buildings. Belgium: European Committee for Standardization, Bruxel; 2008

[31] ATC 58. Next- Generation Performance-Based Seismic Desigb Guidelines,Program Plan for New and Existing Buildings. FEMA 445. California, US: Redwood City; 2006

[32] Eurocode 8. Design of structures for earthquake resistance. Part 1. General Rules. Specific Rules, Seismic Actions and Rules for Buildings. Belgium: European Committee for Standardization, Bruxel; 2008

**39**

**Chapter 3**

**Abstract**

based on recent scientific studies.

model, corrosion damage

**1. Introduction**

Corrosion Effect on Bond Loss

This chapter is devoted to the effects of steel corrosion on bond relationship between steel and concrete. One of the basic assumptions in design of reinforced concrete members is the perfect steel - concrete bond mechanism, so that strain of reinforcing bar is the same as that of the surrounding concrete and these two different materials act as one. However, corrosion of steel reinforcement consists one of the main durability problems in reinforced concrete members, downgrade the bond behavior and therefore their structural integrity. Corrosion degrades the reinforcement itself, reducing the initial cross-section of the steel bar and its mechanical properties. Furthermore, tensile stresses in surrounding concrete caused due to oxides on the corroded reinforcement, lead to the gradual development of tensile field to the surrounding concrete, with spalling of the cover concrete and loss of bond mechanism as a consequence. In this chapter, an overview of damage of reinforced concrete due to steel corrosion is given, focused on the bond mechanism; factors that play key role in the degree of bonding and, also, proposed models of bond strength loss in correlation with the surface concrete cracking due to corrosion are indicated. To conclude, the ongoing research in this area of interest is presented,

**Keywords:** steel corrosion, bond loss, surface cracking, bond strength, predictive

Reinforced concrete consists the most widely used construction material of the existing building stock, providing high bearing capacity in conjunction with low production cost. Due to weakness of plain concrete to withstand tensile forces, steel reinforcing bars are introduced into concrete to enhance its overall mechanical performance. The key consideration so as to ensure that steel and concrete cooperate is the bond mechanism between them. However, corrosion of steel reinforcement constitutes a major degradation factor, which leads to premature aging of RC structures. Recent reports have indicated huge economic impacts due to corrosion damage, since a significant part of the annual budget in many countries is spent on maintenance, repair, and rehabilitation of RC structures [1, 2]. For instance, results of a study conducted by NACE [1] refers that the cost of corrosion is globally

estimated to be US\$2.5 trillion, approximately 3.4 percent of the global GDP.

It is a common knowledge that corrosion process has initially slow rate in nature, since it takes a long period of more than about 10 or 15 years until the aggressive environmental agents diffuse into concrete and reach the steel reinforcement so

between Steel and Concrete

*Charis Apostolopoulos and Konstantinos Koulouris*

### **Chapter 3**

## Corrosion Effect on Bond Loss between Steel and Concrete

*Charis Apostolopoulos and Konstantinos Koulouris*

### **Abstract**

This chapter is devoted to the effects of steel corrosion on bond relationship between steel and concrete. One of the basic assumptions in design of reinforced concrete members is the perfect steel - concrete bond mechanism, so that strain of reinforcing bar is the same as that of the surrounding concrete and these two different materials act as one. However, corrosion of steel reinforcement consists one of the main durability problems in reinforced concrete members, downgrade the bond behavior and therefore their structural integrity. Corrosion degrades the reinforcement itself, reducing the initial cross-section of the steel bar and its mechanical properties. Furthermore, tensile stresses in surrounding concrete caused due to oxides on the corroded reinforcement, lead to the gradual development of tensile field to the surrounding concrete, with spalling of the cover concrete and loss of bond mechanism as a consequence. In this chapter, an overview of damage of reinforced concrete due to steel corrosion is given, focused on the bond mechanism; factors that play key role in the degree of bonding and, also, proposed models of bond strength loss in correlation with the surface concrete cracking due to corrosion are indicated. To conclude, the ongoing research in this area of interest is presented, based on recent scientific studies.

**Keywords:** steel corrosion, bond loss, surface cracking, bond strength, predictive model, corrosion damage

### **1. Introduction**

Reinforced concrete consists the most widely used construction material of the existing building stock, providing high bearing capacity in conjunction with low production cost. Due to weakness of plain concrete to withstand tensile forces, steel reinforcing bars are introduced into concrete to enhance its overall mechanical performance. The key consideration so as to ensure that steel and concrete cooperate is the bond mechanism between them. However, corrosion of steel reinforcement constitutes a major degradation factor, which leads to premature aging of RC structures. Recent reports have indicated huge economic impacts due to corrosion damage, since a significant part of the annual budget in many countries is spent on maintenance, repair, and rehabilitation of RC structures [1, 2]. For instance, results of a study conducted by NACE [1] refers that the cost of corrosion is globally estimated to be US\$2.5 trillion, approximately 3.4 percent of the global GDP.

It is a common knowledge that corrosion process has initially slow rate in nature, since it takes a long period of more than about 10 or 15 years until the aggressive environmental agents diffuse into concrete and reach the steel reinforcement so

as to create the electrochemical cell of corrosion. Concrete cover thickness and a thin passive layer on steel surface, due to high alkalinity of concrete (pH at about 12.5) protect the reinforcement, delaying the penetration and diffusion of corrosive factors. Nevertheless, when aggressive environmental factors, as chloride ions, reach a critical concentration rate, accompanied by reduction of pH below 9, steel reinforcement depassivates and corrosion initiates [3]. Several scientific studies have been conducted, investigating the rate of chlorides' diffusion through the porrosive concrete, aimed at the establishment of a critical value (threshold) of chloride content in order to predict the onset of corrosion [4–6]. The majority of researchers model chloride transport in concrete using the Fick's second law of diffusion, neglecting the chloride interaction with the solid phase [7]. However, there are many uncertainties since many factors influence the rate of chloride penetration into concrete, such as porosity and cracks of concrete, temperature, moisture and salinity of corrosive environment. Hence, due to misinterpretations and various results in literature to date, there is no broadly accepted by the scientific community method of estimating and modeling the onset of corrosion by means of the critical chloride content. Due to the abovementioned, modern international regulations on the design of concrete structures, as BSI EN 206–1 [8], based on long term service life of reinforced concrete, proposed minimum values of cover thickness and concrete classes, depending on the environmental exposure conditions, in order to ensure high protection level of reinforcement against corrosion.

As aggressive agents penetrate and act in limited exposed to corrosion areas rather than the entire length of the reinforcing bars, the corrosion effect is mainly characterized by non-uniformity along the steel bars and is detected by pits on their surface. During corrosion process, steel tends to return to its initial ore form resulting in mass loss and its conversion to iron oxides (rust) on steel surface. Consequences of corrosion damage on steel reinforcement are the reduction of the initial cross-section, resulting in increase of applied mechanical stresses and stress concentration due to pit development, as well as the degradation of its mechanical properties. During the last decades, an effort has been made so as to estimate and quantify the corrosion effect on steel reinforcing bars. The non-uniform distribution of corrosion damage on the steel cross section agreed to be one of the primary cause of mechanical properties' degradation, which has been studied by many researchers [9–13]. Sun [12] and Andisheh et al. [13] tested bare reinforcing bars under monotonic loads, depicting the significant material degradation. Extending the research upon steel corrosion in concrete, experimental studies on embedded reinforcing bars [14–16] indicated more severe corrosion damage, accompanied by narrow pits, which leads to further reduction of mechanical response. Nonetheless, experimental results of both bare and embedded corroded steel reinforcement mainly showed ductility drop rather than reduction of effective stress. To this effect, several studies targeted on the relationship between the degree of corrosion of steel reinforcement and the bearing capacity of corresponding reinforced concrete structures [17–19]. Recently, Kashani et al. [20] presented a state of the art review up to date concerning the current knowledge upon residual capacity of corroded RC elements.

The iron oxides developed due to corrosion phenomenon on steel bars' surface occupy 4 to 6 times greater volume of the mass lost, generating tensile stresses in surrounding concrete with subsequent concrete cracking and spalling of the cover concrete. Hence, corrosion impairs the interface between steel and concrete and therefore affects the bond between them [21]. Bond is the imperative mechanism to denote the transfer of forces between reinforcement and surrounding concrete [22], which is mainly influenced by chemical adhesion, friction and mechanical interlock due to the presence of ribs.

**41**

*Corrosion Effect on Bond Loss between Steel and Concrete*

Besides corrosion phenomenon, the aspects of steel – concrete adhesion depend to a high extent on numerous parameters related to both steel and concrete, namely steel bar geometry, concrete strength and confinement due to transverse reinforcement and concrete cover thickness [23]. The influence of compressive concrete strength on bond behavior of RC specimens has been studied by Abosrra et al. [24] and Zandi and Coronelli [25]. Yalciner et al. [26] conducted an experimental study on bond strength loss due to corrosion taken into account both the compressive strength of concrete fc and the ratio of concrete cover thickness to nominal steel diameter c/D. Testing RC specimens without stirrups, Maslehuddin et al. [27] indicated that although a slight improvement of bond strength is demonstrated in low corrosion levels, sharp degradation pf bond strength is recorded as corrosion increases. Recently, experimental studies by Zandi et al. [28], Lin et al. [29] and Apostolopoulos and Koulouris [30] investigated both the significant role of stirrups spacing and corrosion carrying out eccentric pull out tests on RC elements with

Steel reinforcement is the determinant factor of bearing capacity of reinforced concrete elements. However, in case of structures located in coastal regions (or marine environment), where high chloride contents are indicated, steel reinforcement degrades due to chloride-induced corrosion and subsequently leads to

It is a common knowledge that steel reinforcement is initially protected by concrete cover and a passive layer on its surface. In particular, concrete cover thickness acts as a physical barrier between steel and corrosive environment, delaying the penetration and diffusion of corrosive factors through pores of concrete. At the same time, high alkalinity of concrete due to cement (pH ~ 12.5) results in protection of steel, forming a thin passive layer of ferric oxides on its surface. Chlorides reach the surface of concrete, enter the pore system either by diffusion (in stationary pore water), or by capillary suction of the surface water in which they are dissolved (or by combination of both transport mechanisms) [31]. It is assumed that there is an initiation period, until chloride ions reach the reinforcement, during which substances as water, chloride ions diffuse into concrete and reach the certain concentration necessary to trigger corrosion of the steel reinforcement [32]. This process has slow rate in nature, since it takes a long period of more than about 10 or 15 years, until the aggressive environmental agents reach the steel reinforcement, accompanied by reduction of concrete pH below 9, depassivate it and then corrosion initiates, **Figure 1**. When the passive protection breaks down then onset of steel corrosion (oxidation) takes place and gradually rust occurs on its surface. Corrosion is an electrochemical phenomenon, in which the existence of an anode, a cathode, an electron pathway and electrolyte (ionic pathway) is required. The electrochemical reactions that occur during the corrosion process are presented

<sup>2</sup> Fe Fe 2e → ++ − (1)

) released in the above anodic reaction, there must be

another reaction (cathodic reaction) in order to ensure the electrical neutrality on

steel surface. This cathodic reaction, Eq. (2), consumes water and oxygen.

*DOI: http://dx.doi.org/10.5772/intechopen.94166*

usual design values of concrete cover.

**2. Corrosion of steel reinforcement**

as follows, in Eq. (1) and (2), [33]: The anodic reaction (oxidation),

Since two electrons (2e<sup>−</sup>

durability problems of the entire RC structures.

*Corrosion Effect on Bond Loss between Steel and Concrete DOI: http://dx.doi.org/10.5772/intechopen.94166*

*Structural Integrity and Failure*

as to create the electrochemical cell of corrosion. Concrete cover thickness and a thin passive layer on steel surface, due to high alkalinity of concrete (pH at about 12.5) protect the reinforcement, delaying the penetration and diffusion of corrosive factors. Nevertheless, when aggressive environmental factors, as chloride ions, reach a critical concentration rate, accompanied by reduction of pH below 9, steel reinforcement depassivates and corrosion initiates [3]. Several scientific studies have been conducted, investigating the rate of chlorides' diffusion through the porrosive concrete, aimed at the establishment of a critical value (threshold) of chloride content in order to predict the onset of corrosion [4–6]. The majority of researchers model chloride transport in concrete using the Fick's second law of diffusion, neglecting the chloride interaction with the solid phase [7]. However, there are many uncertainties since many factors influence the rate of chloride penetration into concrete, such as porosity and cracks of concrete, temperature, moisture and salinity of corrosive environment. Hence, due to misinterpretations and various results in literature to date, there is no broadly accepted by the scientific community method of estimating and modeling the onset of corrosion by means of the critical chloride content. Due to the abovementioned, modern international regulations on the design of concrete structures, as BSI EN 206–1 [8], based on long term service life of reinforced concrete, proposed minimum values of cover thickness and concrete classes, depending on the environmental exposure conditions, in order to

ensure high protection level of reinforcement against corrosion.

As aggressive agents penetrate and act in limited exposed to corrosion areas rather than the entire length of the reinforcing bars, the corrosion effect is mainly characterized by non-uniformity along the steel bars and is detected by pits on their surface. During corrosion process, steel tends to return to its initial ore form resulting in mass loss and its conversion to iron oxides (rust) on steel surface. Consequences of corrosion damage on steel reinforcement are the reduction of the initial cross-section, resulting in increase of applied mechanical stresses and stress concentration due to pit development, as well as the degradation of its mechanical properties. During the last decades, an effort has been made so as to estimate and quantify the corrosion effect on steel reinforcing bars. The non-uniform distribution of corrosion damage on the steel cross section agreed to be one of the primary cause of mechanical properties' degradation, which has been studied by many researchers [9–13]. Sun [12] and Andisheh et al. [13] tested bare reinforcing bars under monotonic loads, depicting the significant material degradation. Extending the research upon steel corrosion in concrete, experimental studies on embedded reinforcing bars [14–16] indicated more severe corrosion damage, accompanied by narrow pits, which leads to further reduction of mechanical response. Nonetheless, experimental results of both bare and embedded corroded steel reinforcement mainly showed ductility drop rather than reduction of effective stress. To this effect, several studies targeted on the relationship between the degree of corrosion of steel reinforcement and the bearing capacity of corresponding reinforced concrete structures [17–19]. Recently, Kashani et al. [20] presented a state of the art review up to date concerning the current knowledge upon residual capacity of

The iron oxides developed due to corrosion phenomenon on steel bars' surface occupy 4 to 6 times greater volume of the mass lost, generating tensile stresses in surrounding concrete with subsequent concrete cracking and spalling of the cover concrete. Hence, corrosion impairs the interface between steel and concrete and therefore affects the bond between them [21]. Bond is the imperative mechanism to denote the transfer of forces between reinforcement and surrounding concrete [22], which is mainly influenced by chemical adhesion, friction and mechanical interlock

**40**

corroded RC elements.

due to the presence of ribs.

Besides corrosion phenomenon, the aspects of steel – concrete adhesion depend to a high extent on numerous parameters related to both steel and concrete, namely steel bar geometry, concrete strength and confinement due to transverse reinforcement and concrete cover thickness [23]. The influence of compressive concrete strength on bond behavior of RC specimens has been studied by Abosrra et al. [24] and Zandi and Coronelli [25]. Yalciner et al. [26] conducted an experimental study on bond strength loss due to corrosion taken into account both the compressive strength of concrete fc and the ratio of concrete cover thickness to nominal steel diameter c/D. Testing RC specimens without stirrups, Maslehuddin et al. [27] indicated that although a slight improvement of bond strength is demonstrated in low corrosion levels, sharp degradation pf bond strength is recorded as corrosion increases. Recently, experimental studies by Zandi et al. [28], Lin et al. [29] and Apostolopoulos and Koulouris [30] investigated both the significant role of stirrups spacing and corrosion carrying out eccentric pull out tests on RC elements with usual design values of concrete cover.

### **2. Corrosion of steel reinforcement**

Steel reinforcement is the determinant factor of bearing capacity of reinforced concrete elements. However, in case of structures located in coastal regions (or marine environment), where high chloride contents are indicated, steel reinforcement degrades due to chloride-induced corrosion and subsequently leads to durability problems of the entire RC structures.

It is a common knowledge that steel reinforcement is initially protected by concrete cover and a passive layer on its surface. In particular, concrete cover thickness acts as a physical barrier between steel and corrosive environment, delaying the penetration and diffusion of corrosive factors through pores of concrete. At the same time, high alkalinity of concrete due to cement (pH ~ 12.5) results in protection of steel, forming a thin passive layer of ferric oxides on its surface. Chlorides reach the surface of concrete, enter the pore system either by diffusion (in stationary pore water), or by capillary suction of the surface water in which they are dissolved (or by combination of both transport mechanisms) [31]. It is assumed that there is an initiation period, until chloride ions reach the reinforcement, during which substances as water, chloride ions diffuse into concrete and reach the certain concentration necessary to trigger corrosion of the steel reinforcement [32]. This process has slow rate in nature, since it takes a long period of more than about 10 or 15 years, until the aggressive environmental agents reach the steel reinforcement, accompanied by reduction of concrete pH below 9, depassivate it and then corrosion initiates, **Figure 1**. When the passive protection breaks down then onset of steel corrosion (oxidation) takes place and gradually rust occurs on its surface.

Corrosion is an electrochemical phenomenon, in which the existence of an anode, a cathode, an electron pathway and electrolyte (ionic pathway) is required. The electrochemical reactions that occur during the corrosion process are presented as follows, in Eq. (1) and (2), [33]:

The anodic reaction (oxidation),

$$\text{Fe} \rightarrow \text{Fe}^{2+} + 2\text{e}^- \tag{1}$$

Since two electrons (2e<sup>−</sup> ) released in the above anodic reaction, there must be another reaction (cathodic reaction) in order to ensure the electrical neutrality on steel surface. This cathodic reaction, Eq. (2), consumes water and oxygen.

**Figure 1.** *Stages of corrosion in RC structures.*

**Figure 2.**

*Corrosion products (rust) cause tensile stresses in concrete leading to cracking and spalling of concrete cover.*

The cathodic reaction,

$$\text{2e}^- + \text{H}\_2\text{O} + \frac{1}{2}\text{O}\_2 \rightarrow 2\text{OH}^- \tag{2}$$

The above two chemical equations contain the basic reactions at the first stages of corrosion. Then, as corrosion propagates, hydration reactions are followed, Eq. (3) - Eq. (5), so as to form hydrated ferric oxide (yield red rust) Fe2O3∙H2O:

$$\mathsf{Fe}^{2\*} + 2\mathsf{O}\mathsf{H}^{-} \to \mathsf{Fe}\mathsf{(O\mathsf{H})}\_{2} \tag{3}$$

where is Fe(OH)2 ferrous hydroxide,

$$\mathsf{4Fe}\mathsf{[O\#]}\_{2} + \mathsf{O}\_{2} + 2\mathsf{H}\_{2}\mathsf{O} \to \mathsf{4Fe}\{\mathsf{O}\#\}\_{3} \tag{4}$$

**43**

**Figure 3.**

*via inducing current (right).*

*Corrosion Effect on Bond Loss between Steel and Concrete*

where 4Fe(OH)3 is ferric hydroxide and

where is Fe2O3·H2O hydrated ferric oxide (rust).

quent reduction of structural capacity of RC members.

studies is impressed current density technique.

3 23 2 2 2Fe(OH) Fe O H O 2H O → ⋅+ (5)

As shown in **Figure 2**, oxides (rust), which are formed due to corrosion on steel

surface, occupy 2 to 6 times greater volume of the attacking mass [15], causing tensile stresses in surrounding concrete and, thereafter, leading to gradual concrete cracking development and spalling of the cover concrete. Hence, corrosion phenomenon affects significantly the steel reinforcement, reducing its initial crosssection, degrades its mechanical properties and bond between steel and concrete.

The capacity assessment of corroded reinforced concrete elements consists an engineering task of major importance, since the effects of steel corrosion have become more apparent. Nevertheless, the current international regulations and standards do not determine degradation rules of RC elements; thus, there is need to develop codes, taking into account the deterioration of materials and the subse-

In order to study and quantify the consequences of corrosion, laboratory methods have been developed to simulate and accelerate the natural process. One of the most widespread accelerated corrosion techniques, used for the goals of many

In this accelerated corrosion technique, the reinforcing steel bars (to be corroded) and stainless-steel bars are connected to the positive and the negative pole of a power supply, respectively, and are immersed in tanks, which are filled by a sodium chloride (NaCl) solution in content of 5% (by weight of water). In this way, an electric circuit is generated, since the reinforcing steel bars act as anode of circuit, the stainless-steel bars act as cathode and the NaCl solution is the electrolyte, which allows ions to flow into the circuit. Direct electric current is induced to the reinforcing steel bars through the power supply in order to accelerate the electrochemical reaction of corrosion [34]. This technique is in accordance with ASTM standards, although all the individual parameters needed to create a standard corrosive environment have not

*Simplified procedure of accelerated electro-corrosion (left) - accelerated corrosion experiments on RC specimens* 

*DOI: http://dx.doi.org/10.5772/intechopen.94166*

**3. Laboratory testing**

where 4Fe(OH)3 is ferric hydroxide and

$$2\mathsf{Fe}\langle\mathsf{O}\mathsf{H}\rangle\_{3} \rightarrow \mathsf{Fe}\_{2}\mathsf{O}\_{3}\cdot\mathsf{H}\_{2}\mathsf{O} + 2\mathsf{H}\_{2}\mathsf{O}\tag{5}$$

where is Fe2O3·H2O hydrated ferric oxide (rust).

As shown in **Figure 2**, oxides (rust), which are formed due to corrosion on steel surface, occupy 2 to 6 times greater volume of the attacking mass [15], causing tensile stresses in surrounding concrete and, thereafter, leading to gradual concrete cracking development and spalling of the cover concrete. Hence, corrosion phenomenon affects significantly the steel reinforcement, reducing its initial crosssection, degrades its mechanical properties and bond between steel and concrete.

### **3. Laboratory testing**

*Structural Integrity and Failure*

**Figure 1.**

**Figure 2.**

*Stages of corrosion in RC structures.*

The cathodic reaction,

2 2 2e H O O 2OH 1

*Corrosion products (rust) cause tensile stresses in concrete leading to cracking and spalling of concrete cover.*

Eq. (5), so as to form hydrated ferric oxide (yield red rust) Fe2O3∙H2O:

2

where is Fe(OH)2 ferrous hydroxide,

The above two chemical equations contain the basic reactions at the first stages of corrosion. Then, as corrosion propagates, hydration reactions are followed, Eq. (3) -

<sup>2</sup> − − ++ → (2)

<sup>2</sup> Fe 2OH Fe(OH) + − + → (3)

22 2 <sup>3</sup> 4Fe(OH) O 2H O 4Fe(OH) ++ → (4)

**42**

The capacity assessment of corroded reinforced concrete elements consists an engineering task of major importance, since the effects of steel corrosion have become more apparent. Nevertheless, the current international regulations and standards do not determine degradation rules of RC elements; thus, there is need to develop codes, taking into account the deterioration of materials and the subsequent reduction of structural capacity of RC members.

In order to study and quantify the consequences of corrosion, laboratory methods have been developed to simulate and accelerate the natural process. One of the most widespread accelerated corrosion techniques, used for the goals of many studies is impressed current density technique.

In this accelerated corrosion technique, the reinforcing steel bars (to be corroded) and stainless-steel bars are connected to the positive and the negative pole of a power supply, respectively, and are immersed in tanks, which are filled by a sodium chloride (NaCl) solution in content of 5% (by weight of water). In this way, an electric circuit is generated, since the reinforcing steel bars act as anode of circuit, the stainless-steel bars act as cathode and the NaCl solution is the electrolyte, which allows ions to flow into the circuit. Direct electric current is induced to the reinforcing steel bars through the power supply in order to accelerate the electrochemical reaction of corrosion [34]. This technique is in accordance with ASTM standards, although all the individual parameters needed to create a standard corrosive environment have not

### **Figure 3.**

*Simplified procedure of accelerated electro-corrosion (left) - accelerated corrosion experiments on RC specimens via inducing current (right).*

been determined yet. Nevertheless, the main advantage of this method is the ability to control the rate of corrosion, which usually varies due to changes in the resistivity, oxygen concentration, and temperature.

In **Figure 3** an indicative automatic system (found in Laboratory of Technology & Strength of Materials of University of Patras) is illustrated, which has been used for the performance of electrochemical corrosion tests. The specific system enables implementation of different corrosive conditions, in terms of impressed current density and ponding cyclic corrosion (wet/dry duration).

### **4. Differential aeration corrosion on steel reinforcement**

In case of steel reinforcing bars, embedded in concrete, the corrosion damage is depicted in finite areas along their length, recording significant reduction of initial cross-section, as shown in **Figure 4**. Recent studies have indicated that the existence of voids and pores in concrete allows not only the penetration and diffusion of corrosive agents but also the higher oxygen concentration on steel bars' surface, consisting a favorable condition for differential aeration corrosion. The anodic dissolution rate depends solely upon the potential difference across the electrolytemetal surface.

In order to further investigate and simulate the consequences of differential aeration corrosion, Apostolopoulos et al. [35] conducted accelerated corrosion experiments on bare reinforcing steel bars, taking into account different exposure to corrosion lengths. The results demonstrated that corrosion damage depends on the exposed to corrosion length, as short samples record higher mass loss percentages and more intense pitting, for the same corrosion duration. In particular, specimens with the short exposed to corrosion length demonstrated about 4 times greater percentage mass loss in contrast to specimens with the long exposed to corrosion length, for 300 h of accelerated corrosion time [35], which is due to Differential aeration corrosion.

In case of RC specimens of experimental study with weak concrete cover thickness, where bond forces are tested, differential aeration corrosion phenomena are detected on ribs. In particular, during the phase of concrete hardening, micro-cracks are recorded in the area of the edge of ribs, which lead to accelerate the penetration of aggressive agents to the steel reinforcement, starting from the damage from the ribs (as closer to the outer surface). Thus, ribs are more vulnerable to corrosion, resulting mainly in degradation of steel - concrete bond mechanism which leads to mechanical interlock's loss and rise of slippage. In that manner, corrosion primarily affects bond behavior of RC elements rather than their mechanical properties, even in low corrosion level. **Figure 5** illustrates intense pits on surface due to corrosion, which subsequently causes loss of ribs, especially on the outer half of the steel bar's circumference, adjacent to the external surface of RC specimens.

**45**

*Corrosion Effect on Bond Loss between Steel and Concrete*

*DOI: http://dx.doi.org/10.5772/intechopen.94166*

**5. Corrosion effect on bond strength**

central issue in this chapter.

**Figure 5.**

*circumference (down).*

reinforcement are taken into account.

surface at the ultimate limit state.

Corrosion factor is responsible for reduction of the cross-sectional area of steel reinforcement, degrading the mechanical properties [9–14], and causing cracks and spalling of the cover concrete, which leads to deterioration of steel - concrete interface and, subsequently, to bond loss between them. In order to assess the structural integrity of corroded RC members, many researchers have studied and quantified the consequences of mechanical performance of steel reinforcement due to corrosion and proposed degradation material laws [15–16]. However, as it is abovementioned, bond behavior is primarily affected by corrosion, as the transfer of forces between reinforcement and surrounding concrete weakens. Thus, the reference to degradation of mechanical properties of steel reinforcement is not a

*Corrosion damage on steel surface. Inner half of steel bar's circumference (up) - outer half of steel bar's* 

Regarding the interface characteristics in RC elements, Model Code 2010 [36] recommends the calculation of ultimate bond strength fbd in the non-corroded condition, as presented in Eq. (6), in which concrete and steel class, steel bar diameter and the contribution of confinement via cover thickness and transverse

*bd* = + ⋅ +⋅ (

where fb,0 is the basic bond strength, which depends on concrete and steel quality,

α2 and α3 are coefficients, which represent the influence of passive confinement from concrete cover thickness and from transverse reinforcement, respectively, in

ptr is the mean compression stress perpendicular to the potential splitting failure

Prolonged exposure of reinforced concrete structures to a corrosive environment

causes significant degradation problems in the steel-concrete bond mechanism, allowing relative slip to develop between steel and concrete and reducing the bond strength between them. Consequently, the bond loss effect due to the environmental action prevents the development of full bearing capacity of reinforced concrete elements until they behave as unreinforced members. From the abovementioned, it is obvious that bond mechanism is a main criterion in design of RC members. However, there is a gap in international codes, regarding the quantification of bond strength loss due to corrosion. In particular, even though range of values are proposed by Model Code 2010 for the estimation of reduced bond strength,

) *<sup>b</sup>*, *tr f fp* **23 0 2** (6)

α α

bar geometry and the casting position of steel bar during concreting,

excess of their respective permissible minima, and.

**Figure 4.** *Local reduction of cross-section due to corrosion.*

### **Figure 5.**

*Structural Integrity and Failure*

metal surface.

Differential aeration corrosion.

external surface of RC specimens.

*Local reduction of cross-section due to corrosion.*

oxygen concentration, and temperature.

density and ponding cyclic corrosion (wet/dry duration).

**4. Differential aeration corrosion on steel reinforcement**

been determined yet. Nevertheless, the main advantage of this method is the ability to control the rate of corrosion, which usually varies due to changes in the resistivity,

In **Figure 3** an indicative automatic system (found in Laboratory of Technology & Strength of Materials of University of Patras) is illustrated, which has been used for the performance of electrochemical corrosion tests. The specific system enables implementation of different corrosive conditions, in terms of impressed current

In case of steel reinforcing bars, embedded in concrete, the corrosion damage is depicted in finite areas along their length, recording significant reduction of initial cross-section, as shown in **Figure 4**. Recent studies have indicated that the existence of voids and pores in concrete allows not only the penetration and diffusion of corrosive agents but also the higher oxygen concentration on steel bars' surface, consisting a favorable condition for differential aeration corrosion. The anodic dissolution rate depends solely upon the potential difference across the electrolyte-

In order to further investigate and simulate the consequences of differential aeration corrosion, Apostolopoulos et al. [35] conducted accelerated corrosion experiments on bare reinforcing steel bars, taking into account different exposure to corrosion lengths. The results demonstrated that corrosion damage depends on the exposed to corrosion length, as short samples record higher mass loss percentages and more intense pitting, for the same corrosion duration. In particular, specimens with the short exposed to corrosion length demonstrated about 4 times greater percentage mass loss in contrast to specimens with the long exposed to corrosion length, for 300 h of accelerated corrosion time [35], which is due to

In case of RC specimens of experimental study with weak concrete cover thickness, where bond forces are tested, differential aeration corrosion phenomena are detected on ribs. In particular, during the phase of concrete hardening, micro-cracks are recorded in the area of the edge of ribs, which lead to accelerate the penetration of aggressive agents to the steel reinforcement, starting from the damage from the ribs (as closer to the outer surface). Thus, ribs are more vulnerable to corrosion, resulting mainly in degradation of steel - concrete bond mechanism which leads to mechanical interlock's loss and rise of slippage. In that manner, corrosion primarily affects bond behavior of RC elements rather than their mechanical properties, even in low corrosion level. **Figure 5** illustrates intense pits on surface due to corrosion, which subsequently causes loss of ribs, especially on the outer half of the steel bar's circumference, adjacent to the

**44**

**Figure 4.**

*Corrosion damage on steel surface. Inner half of steel bar's circumference (up) - outer half of steel bar's circumference (down).*

### **5. Corrosion effect on bond strength**

Corrosion factor is responsible for reduction of the cross-sectional area of steel reinforcement, degrading the mechanical properties [9–14], and causing cracks and spalling of the cover concrete, which leads to deterioration of steel - concrete interface and, subsequently, to bond loss between them. In order to assess the structural integrity of corroded RC members, many researchers have studied and quantified the consequences of mechanical performance of steel reinforcement due to corrosion and proposed degradation material laws [15–16]. However, as it is abovementioned, bond behavior is primarily affected by corrosion, as the transfer of forces between reinforcement and surrounding concrete weakens. Thus, the reference to degradation of mechanical properties of steel reinforcement is not a central issue in this chapter.

Regarding the interface characteristics in RC elements, Model Code 2010 [36] recommends the calculation of ultimate bond strength fbd in the non-corroded condition, as presented in Eq. (6), in which concrete and steel class, steel bar diameter and the contribution of confinement via cover thickness and transverse reinforcement are taken into account.

$$f\_{\mathsf{b}\mathsf{t}} = (\alpha\_{\mathsf{2}} + \alpha\_{\mathsf{3}}) \cdot f\_{\mathsf{b},\mathsf{0}} + \mathsf{2} \cdot \mathsf{p}\_{\mathsf{t}} \tag{6}$$

where fb,0 is the basic bond strength, which depends on concrete and steel quality, bar geometry and the casting position of steel bar during concreting,

α2 and α3 are coefficients, which represent the influence of passive confinement from concrete cover thickness and from transverse reinforcement, respectively, in excess of their respective permissible minima, and.

ptr is the mean compression stress perpendicular to the potential splitting failure surface at the ultimate limit state.

Prolonged exposure of reinforced concrete structures to a corrosive environment causes significant degradation problems in the steel-concrete bond mechanism, allowing relative slip to develop between steel and concrete and reducing the bond strength between them. Consequently, the bond loss effect due to the environmental action prevents the development of full bearing capacity of reinforced concrete elements until they behave as unreinforced members. From the abovementioned, it is obvious that bond mechanism is a main criterion in design of RC members. However, there is a gap in international codes, regarding the quantification of bond strength loss due to corrosion. In particular, even though range of values are proposed by Model Code 2010 for the estimation of reduced bond strength,

### *Structural Integrity and Failure*

considering corrosion penetration and surface crack, nevertheless the influence of stirrups spacing as well as the impact of non-uniform type of corrosion damage on steel bar's surface through pits, which is the most common in practice, are not determined yet. Due to this, there is area of research so as to establish degradation models, including the effect of both density of stirrups and local reduction of cross-section in maintenance of bond strength of corroded RC elements. Moreover, the term corrosion penetration, which is found in many regulatory texts, refers to a uniform circumference loss of the circular cross section due to mass loss, case of uniform corrosion which is practically non-existent in real RC structures exposed to chloride induced corrosion.

On this basis, many researchers have studied the effect of corrosion on bond between steel and concrete [24–30]. As Maslehuddin et al. [27] reported, bond strength of RC specimens without transverse reinforcement increases slightly in low corrosion degree; however, sharp bond loss takes place with the propagation of corrosion. These findings are in aggrement with the study of Auyeung [37], which demonstrated up to 80% bond strength loss due to only 2% reduction of steel cross-section. A more comprehensive study is presented by Lundgren [38] concerning the contribution of stirrups to the bond behavior of corroded RC elements. In real RC structures the estimation of corrosion penetration and mass loss of steel bars consists a difficult task, since steel bars are embedded in concrete and corrosion damage is non-uniform on steel's surface. However, cracks due to steel corrosion on concrete surface are visible and their width can be easily measured, as shown in **Figure 6**. For this reason, recent scientific studies tend to quantify the corrosion damage of the embedded steel reinforcement through the surface cracking width and, subsequently, estimate the local bond loss [10, 21, 24–30, 37–44]. An empirical correlation between the loss of steel bar's diameter and the average corrosion penetration has been proposed by Torres-Acosta et al. [39]. Moreover, studies of Andrade et al. [40] and Tahersamsi [10] link the surface cracking width

**Figure 6.** *Surface concrete cracking due to steel corrosion (left) -measuring of crack width on concrete (right).*

**47**

**Figure 7.**

*(left) and 40 mm (right).*

*Corrosion Effect on Bond Loss between Steel and Concrete*

with the corrosion damage of steel bar and the loss of bond strength. Recent experimental study of Lin et al. [29] investigated the influence of concrete cover thickness and stirrups on the occurrence of surface cracks and on the subsequent bond strength loss. Gathering various experimental data, a predictive model of bond strength loss as a function of surface concrete cracking has been suggested by

Based on the abovementioned, a broad and ongoing experimental research on corroded RC elements was conducted by Apostolopoulos and Koulouris [30], studying the influence of stirrups spacing (density) and concrete cover thickness on bond behavior of corroded RC specimens, in correlation with the surface cracking width due to corrosion. The results depict close correlation of bond strength with

The depassivation of protective layer on steel reinforcement leads to onset of corrosion, the propagation of which develops various range of surface concrete cracking along the reinforcing bars. As illustrated, firstly in the following **Figure 6** (Rigth) and thereafter in **Figure 7**, the surface cracking width varies depending on

In the case of specimens with concrete cover of 25 mm, the values of average surface crack width followed a common path up to a low corrosion level of 3% (**Figure 7** Left). It is noteworthy that the specimens with dense stirrups (Φ8/60 mm) demonstrated initially higher values of cracking width, since corrosion potential was higher due to the high percentage of steel reinforcement; however, as corrosion degree increases the confinement provided by dense stirrups limited the progressive development of surface cracking. On the other hand, the specimens without stirrups) recorded initially limited range cracking, due to the low percentage of steel, whereas in higher corrosion levels the absence of confine-

Similar results were recorded in the case of specimens with concrete cover of 40 mm (**Figure 7** Right). More specific, specimens without stirrups and specimens with dense and quite dense stirrups (Φ8/60 mm and Φ8/120 mm respectively), for mass loss equal to 3%, depicted similar range of surface crack width. In contrast, specimens with stirrups Φ8/240 mm, for the same percentage of mass loss, recorded sudden and remarkable high range of surface crack. In this group pf specimens, namely with concrete cover thickness equal to 40 mm, the confined cross section is reduced in line to the specimens with concrete cover thickness equal to 25 mm. Hence, poor confinement level in conjuction with corrosion initiation

*Average crack width on concrete surface in function of percentage mass loss of steel bar. Cover thickness 25 mm* 

corrosion level, stirrups spacing and concrete cover thickness.

ment lead to rapid growth rate of cracking width.

impacts the uncontrolled propagation of surface cracking.

*DOI: http://dx.doi.org/10.5772/intechopen.94166*

surface concrete cracking width.

Zhou et al. [41].

### *Corrosion Effect on Bond Loss between Steel and Concrete DOI: http://dx.doi.org/10.5772/intechopen.94166*

*Structural Integrity and Failure*

chloride induced corrosion.

considering corrosion penetration and surface crack, nevertheless the influence of stirrups spacing as well as the impact of non-uniform type of corrosion damage on steel bar's surface through pits, which is the most common in practice, are not determined yet. Due to this, there is area of research so as to establish degradation models, including the effect of both density of stirrups and local reduction of cross-section in maintenance of bond strength of corroded RC elements. Moreover, the term corrosion penetration, which is found in many regulatory texts, refers to a uniform circumference loss of the circular cross section due to mass loss, case of uniform corrosion which is practically non-existent in real RC structures exposed to

On this basis, many researchers have studied the effect of corrosion on bond between steel and concrete [24–30]. As Maslehuddin et al. [27] reported, bond strength of RC specimens without transverse reinforcement increases slightly in low corrosion degree; however, sharp bond loss takes place with the propagation of corrosion. These findings are in aggrement with the study of Auyeung [37], which demonstrated up to 80% bond strength loss due to only 2% reduction of steel cross-section. A more comprehensive study is presented by Lundgren [38] concerning the contribution of stirrups to the bond behavior of corroded RC elements. In real RC structures the estimation of corrosion penetration and mass loss of steel bars consists a difficult task, since steel bars are embedded in concrete and corrosion damage is non-uniform on steel's surface. However, cracks due to steel corrosion on concrete surface are visible and their width can be easily measured, as shown in **Figure 6**. For this reason, recent scientific studies tend to quantify the corrosion damage of the embedded steel reinforcement through the surface cracking width and, subsequently, estimate the local bond loss [10, 21, 24–30, 37–44]. An empirical correlation between the loss of steel bar's diameter and the average corrosion penetration has been proposed by Torres-Acosta et al. [39]. Moreover, studies of Andrade et al. [40] and Tahersamsi [10] link the surface cracking width

*Surface concrete cracking due to steel corrosion (left) -measuring of crack width on concrete (right).*

**46**

**Figure 6.**

with the corrosion damage of steel bar and the loss of bond strength. Recent experimental study of Lin et al. [29] investigated the influence of concrete cover thickness and stirrups on the occurrence of surface cracks and on the subsequent bond strength loss. Gathering various experimental data, a predictive model of bond strength loss as a function of surface concrete cracking has been suggested by Zhou et al. [41].

Based on the abovementioned, a broad and ongoing experimental research on corroded RC elements was conducted by Apostolopoulos and Koulouris [30], studying the influence of stirrups spacing (density) and concrete cover thickness on bond behavior of corroded RC specimens, in correlation with the surface cracking width due to corrosion. The results depict close correlation of bond strength with surface concrete cracking width.

The depassivation of protective layer on steel reinforcement leads to onset of corrosion, the propagation of which develops various range of surface concrete cracking along the reinforcing bars. As illustrated, firstly in the following **Figure 6** (Rigth) and thereafter in **Figure 7**, the surface cracking width varies depending on corrosion level, stirrups spacing and concrete cover thickness.

In the case of specimens with concrete cover of 25 mm, the values of average surface crack width followed a common path up to a low corrosion level of 3% (**Figure 7** Left). It is noteworthy that the specimens with dense stirrups (Φ8/60 mm) demonstrated initially higher values of cracking width, since corrosion potential was higher due to the high percentage of steel reinforcement; however, as corrosion degree increases the confinement provided by dense stirrups limited the progressive development of surface cracking. On the other hand, the specimens without stirrups) recorded initially limited range cracking, due to the low percentage of steel, whereas in higher corrosion levels the absence of confinement lead to rapid growth rate of cracking width.

Similar results were recorded in the case of specimens with concrete cover of 40 mm (**Figure 7** Right). More specific, specimens without stirrups and specimens with dense and quite dense stirrups (Φ8/60 mm and Φ8/120 mm respectively), for mass loss equal to 3%, depicted similar range of surface crack width. In contrast, specimens with stirrups Φ8/240 mm, for the same percentage of mass loss, recorded sudden and remarkable high range of surface crack. In this group pf specimens, namely with concrete cover thickness equal to 40 mm, the confined cross section is reduced in line to the specimens with concrete cover thickness equal to 25 mm. Hence, poor confinement level in conjuction with corrosion initiation impacts the uncontrolled propagation of surface cracking.

### **Figure 7.**

*Average crack width on concrete surface in function of percentage mass loss of steel bar. Cover thickness 25 mm (left) and 40 mm (right).*

### *Structural Integrity and Failure*

Nevertheless, with the evolution of corrosion, specimens with dense connectors, namely Φ8/60 mm, noted a significant decrement of surface cracking development. This particularly notable decrease, recorded in specimens with dense stirrups, was applied to both categories of specimens, one with concrete cover thickness of 25 mm and the other with concrete cover of 40 mm. More precisely, an average crack width equal to 1 mm has been recorded, for mass loss between 8.5% and 9%, respectively.

It is obvious that surface crack width is the outcome of corrosion damage of steel reinforcement; surface cracking is directly linked to various parameters beginning with the existence and amount of transverse reinforcement and cover concrete thickness. For both groups of specimens with different concrete cover (25 mm and 40 mm, respectively), the presence of dense stirrups (Φ8/60 mm) is preceded with a remarkable limitation of the surface cracking evolution to a width threshold of 0.90 mm, corresponding value to the abovementioned average mass loss of 8.5%–9.0%.

In order to investigate the bond behavior, pull out tests of uncorroded and corroded RC specimens conducted, the results of which confirmed that, in both cases of concrete cover thickness of 25 mm and 40 mm, the increase of the average range of surface cracking brought a dramatic decrease of bond strength between concrete and steel bar, **Figure 8**. Moreover, obtained by non-linear regression analysis, exponential predictive models of bond strength loss due to corrosion of steel reinforcement were given, derived from the correlation of bond strength loss of corroded specimens and surface cracking of concrete. The functions of predictive models are as follows, Eq. (7):

$$\frac{\mathcal{L}\_b^{c\mu\prime}}{\mathcal{L}\_b^{\text{asucov}}} = \mathcal{e}^{-A \cdot \varepsilon\_\nu} \tag{7}$$

**49**

respsectively.

**Table 1.**

*Corrosion Effect on Bond Loss between Steel and Concrete*

R2

R2

Among specimens of the same concrete cover, it is clear that densification of transverse reinforcement slows down the progression of bond loss. In particular, specimens with concrete cover equal to 25 mm and dense stirrups Φ8/60 mm, the bond strength performance, even though its initial increase up to a threshold of 0.60 mm crack width, remained stable as in the case of non-corroded specimens. Moreover, among specimens of similar range of cracking, specimens without stirrups recorded a decline of bond strength performance equal to 57%, whereas specimens with wide stirrups (Φ8/240 mm) and quite dense stirrups (Φ8/120 mm)

**Cover (mm) No Stirrups** Φ**8/240** Φ**8/120** Φ**8/60** 25 A 1.435 0.736 0.274 0.117

40 A 1.257 0.724 0.499 0.260

(%) 96.2 96.5 97.7 45.9

(%) 97.5 96.0 91.9 80.1

recorded a decrease of bond strength performance equal to 35% and 15%,

*Parameters (by regression analysis) for the exponential predictive model of bond strength loss.*

Thus, loss of bond strength seems to be inevitable.

Almusallam et al. [43] and Rodriguez et al. [44].

to-be-tested in comparative studies.

this benefits, thereby delaying bond strength degradation.

The wide stirrups spacing (Φ8/240 mm), in case of 25 mm cover thickness, degrades bond strength performance, leading to a residual bond strength equal to 40% of the non-corroded bond strength, corresponding to a crack width of 1.45 mm, whereas the absence of stirrups (specimens without stirrups) recorded bond loss equal to 16% of non-corroded value. The abovementioned outcomes transposed to former practices, as in the existing building stock, the use of transverse reinforcement accounts for four pieces per linear meter, i.e. Φ8/250 mm.

In the case of cover thickness of 40 mm, the absence of stirrups deteriorates rapidly bond strength contrary to the dense fitting of stirrups (Φ8/60 mm) where

The bond strength between steel and concrete demonstrates a denoting drop when increasing the range of surface cracking; it follows from the assessment of both cases of concrete cover that as the range of surface cracking raises, the threshold of bond strength performance reduces. These results come in good agreement with results of former studies, to name Lin et al. [29], Fischer and Ozbolt [42],

Given the tendency to approach the issue of bond strength between concrete and steel, exponential predictive models were developed in order to link the bond strength loss of corroded specimens to the average width of concrete surface

cracking. The exponential model introduces an adequate approach of bond strength loss and comes to agreement with previous studies [43, 44] as corrosion evolves. Hence, so as to assess the bond loss, besides the traditional method of chlorides' measurement, surface cracking measurement occurs as an emerging methodology. Notwithstanding the efforts of the scientific community to correlate experimental results of current literature, the issue of dispersion is pertinent due to several parameters, to cover thickness, concrete class, nominal diameter of steel reinforcement, presence of stirrups. It is also noteworthy that, in existing experimental literature, exponential models are proposed as predictive models of bond behavior of corroded RC elements, regardless of the differences denoted in all parameters

*DOI: http://dx.doi.org/10.5772/intechopen.94166*

where *A* is a parameter that depends on the concrete cover (c) and the amount of transverse reinforcement (stirrups spacing or absence of stirrups). The values of parameter A for each predictive model, and the corresponding values of the R<sup>2</sup> coefficient, are presented in the following **Table 1**. As shown by the predictive curves, stirrups spacing is main influencing factor of bond strength degradation due to corrosion. Hence, there is need to determine specific models of bond loss in order to enhance the current technical codes.

### **Figure 8.**

*Predictive models of bond strength loss correlated to the average surface crack width. Cover thickness 25 mm (left) and 40 mm (right).*


*Corrosion Effect on Bond Loss between Steel and Concrete DOI: http://dx.doi.org/10.5772/intechopen.94166*

### **Table 1.**

*Structural Integrity and Failure*

respectively.

8.5%–9.0%.

tive models are as follows, Eq. (7):

enhance the current technical codes.

Nevertheless, with the evolution of corrosion, specimens with dense connectors, namely Φ8/60 mm, noted a significant decrement of surface cracking development. This particularly notable decrease, recorded in specimens with dense stirrups, was applied to both categories of specimens, one with concrete cover thickness of 25 mm and the other with concrete cover of 40 mm. More precisely, an average crack width equal to 1 mm has been recorded, for mass loss between 8.5% and 9%,

It is obvious that surface crack width is the outcome of corrosion damage of steel reinforcement; surface cracking is directly linked to various parameters beginning with the existence and amount of transverse reinforcement and cover concrete thickness. For both groups of specimens with different concrete cover (25 mm and 40 mm, respectively), the presence of dense stirrups (Φ8/60 mm) is preceded with a remarkable limitation of the surface cracking evolution to a width threshold of 0.90 mm, corresponding value to the abovementioned average mass loss of

In order to investigate the bond behavior, pull out tests of uncorroded and corroded RC specimens conducted, the results of which confirmed that, in both cases of concrete cover thickness of 25 mm and 40 mm, the increase of the average range of surface cracking brought a dramatic decrease of bond strength between concrete and steel bar, **Figure 8**. Moreover, obtained by non-linear regression analysis, exponential predictive models of bond strength loss due to corrosion of steel reinforcement were given, derived from the correlation of bond strength loss of corroded specimens and surface cracking of concrete. The functions of predic-

> *cor b A c uncor b c*

*c*

*w*

− ⋅ = (7)

coef-

*e*

where *A* is a parameter that depends on the concrete cover (c) and the amount of transverse reinforcement (stirrups spacing or absence of stirrups). The values of parameter A for each predictive model, and the corresponding values of the R<sup>2</sup>

ficient, are presented in the following **Table 1**. As shown by the predictive curves, stirrups spacing is main influencing factor of bond strength degradation due to corrosion. Hence, there is need to determine specific models of bond loss in order to

*Predictive models of bond strength loss correlated to the average surface crack width. Cover thickness 25 mm* 

**48**

**Figure 8.**

*(left) and 40 mm (right).*

*Parameters (by regression analysis) for the exponential predictive model of bond strength loss.*

Among specimens of the same concrete cover, it is clear that densification of transverse reinforcement slows down the progression of bond loss. In particular, specimens with concrete cover equal to 25 mm and dense stirrups Φ8/60 mm, the bond strength performance, even though its initial increase up to a threshold of 0.60 mm crack width, remained stable as in the case of non-corroded specimens. Moreover, among specimens of similar range of cracking, specimens without stirrups recorded a decline of bond strength performance equal to 57%, whereas specimens with wide stirrups (Φ8/240 mm) and quite dense stirrups (Φ8/120 mm) recorded a decrease of bond strength performance equal to 35% and 15%, respsectively.

The wide stirrups spacing (Φ8/240 mm), in case of 25 mm cover thickness, degrades bond strength performance, leading to a residual bond strength equal to 40% of the non-corroded bond strength, corresponding to a crack width of 1.45 mm, whereas the absence of stirrups (specimens without stirrups) recorded bond loss equal to 16% of non-corroded value. The abovementioned outcomes transposed to former practices, as in the existing building stock, the use of transverse reinforcement accounts for four pieces per linear meter, i.e. Φ8/250 mm. Thus, loss of bond strength seems to be inevitable.

In the case of cover thickness of 40 mm, the absence of stirrups deteriorates rapidly bond strength contrary to the dense fitting of stirrups (Φ8/60 mm) where this benefits, thereby delaying bond strength degradation.

The bond strength between steel and concrete demonstrates a denoting drop when increasing the range of surface cracking; it follows from the assessment of both cases of concrete cover that as the range of surface cracking raises, the threshold of bond strength performance reduces. These results come in good agreement with results of former studies, to name Lin et al. [29], Fischer and Ozbolt [42], Almusallam et al. [43] and Rodriguez et al. [44].

Given the tendency to approach the issue of bond strength between concrete and steel, exponential predictive models were developed in order to link the bond strength loss of corroded specimens to the average width of concrete surface cracking. The exponential model introduces an adequate approach of bond strength loss and comes to agreement with previous studies [43, 44] as corrosion evolves. Hence, so as to assess the bond loss, besides the traditional method of chlorides' measurement, surface cracking measurement occurs as an emerging methodology. Notwithstanding the efforts of the scientific community to correlate experimental results of current literature, the issue of dispersion is pertinent due to several parameters, to cover thickness, concrete class, nominal diameter of steel reinforcement, presence of stirrups. It is also noteworthy that, in existing experimental literature, exponential models are proposed as predictive models of bond behavior of corroded RC elements, regardless of the differences denoted in all parameters to-be-tested in comparative studies.

Extending the investigation of bond behavior of corroded RC members and focusing on values of maximum pull-out force and, subsequently, bond strength, and not on bond loss, the role of stirrups spacing is more highlighted, where bonding between steel and concrete degrades due to corrosion and the developed maximum pull-out force drops. The usage of dense transverse reinforcement contributes to bond behavior, not only reducing the bond loss rate, but leading to greater values of maximum pull-out force due to confinement, **Figure 9**.

In an effort to estimate that influence of stirrups spacing on non-corroded condition, the presence of wide stirrups (Φ8/240 mm) present maximum bond strength equal to 7.04 MPa, whereas quite dense (Φ8/120 mm) and dense stirrups (Φ8/60 mm) equal to 9.10 MPa and 9.53 MPa, respectively. The percentages attributing the increase of bond strength against specimens without stirrups are 35.9%, 75.6% and 84%, respectively. It is noteworthy that, quite dense stirrups spacing (Φ8/120 mm) result in sufficient bond strength levels, whereas further densification leads to minor increase of bond strength, nevertheless delays the bond loss. Extrapolating the abovementioned on real RC structures, stirrups' densification above a certain threshold, is considered inappropriate since it could lead on the one hand to substantial increase of costs and on the other hand to rapid rise of corrosion rate, due to potential increase.

Moreover, bond stress - relative slip curves are exported from pull-out tests for each corrosion level and for each category of stirrups spacing, respectively. Typical curves of uncorroded and corroded specimens are shown in **Figure 10**.

The harmful influence of corrosion phenomenon on bond behavior between steel and concrete can be seen examining the bond stress - slip curves of **Figure 10**, and in particular group of specimens without stirrups. Uncorroded specimens - where no surface cracking had been initially observed - showed a quasi linear relationship between bond stresses and relative slip till the point of bond strength's development. During this phase, cracks were occurred, parallel to the axis of steel bar, and gradually were developed due to the radial stresses, which are transferred from steel to concrete. After the development of bond strength, sharp decline of bond stresses and complete bond loss followed. In the case of corroded specimens (without

**51**

**Figure 10.**

*Corrosion Effect on Bond Loss between Steel and Concrete*

stirrups) intense surface concrete cracking was recorded, due to steel corrosion, even in low mass loss levels. As a result of this cracking, bond behavior degrades dramatically, as confirmed by the corresponding curves of **Figure 10**, since both the steel-concrete interface is damaged by the corrosion oxides and the confinement

On the other hand, as shown by the groups of specimens with stirrups, surface cracking and degree of bond loss is strongly correlated with stirrups spacing. In particular, specimens with quite wide stirrups (Φ8/240 mm) showed that transverse reinforcement has a positive impact on bond behavior in uncorroded condition, as greater values of bond stress indicated. However, when corrosion occurs, accompanied by surface concrete cracking, bond resistance degradates as reflected in responding curves, where bond strength reduces significantly, and subsequently

Greater contribution of stirrups to bond behavior was noticed on specimens with quite dense stirrups spacing (Φ8 /120 mm), where even higher bond strength values are indicated, and while bond behavior degrades due to corrosion damage nevertheless does not demonstrate a massive drop mainly due to confinement. In addition, in the case of wide stirrups (Φ8 / 60 mm), full use of bond behavior, even after corrosion damage, occurs, with residual bond stress recorded after the peak of bond strength. It is also noteworthy that specimens of this category had finally ended due to failure under tension. Thus, that densification of stirrups make full use of bearing capacity of steel reinforcing bars and subsequently of RC elements. From the aforementioned, it becomes obvious that the use of dense stirrups (Φ8/60 mm) leads to full bonding between steel and concrete, in uncorroded and corroded conditions respectively, marking both high bond strength and high residual bond

To conclude, existence of stirrups contributes to bond behavior, and subsequently, greater density of stirrups affects both the reduction of bond loss' degradation rate due to steel corrosion, as well as the increasing of bond stresses values due to greater confinement. The use of wide stirrups (Φ8 / 240 mm) enhances the

level is deteriorated by the cracks in concrete cover.

*Bond stress - slip curves obtained from pull out tests.*

low values of residual bond stress are recorded.

stress after the ultimate pull out force.

*DOI: http://dx.doi.org/10.5772/intechopen.94166*

**Figure 9.** *Bond strength values in function with average crack width.*

*Corrosion Effect on Bond Loss between Steel and Concrete DOI: http://dx.doi.org/10.5772/intechopen.94166*

*Structural Integrity and Failure*

rate, due to potential increase.

Extending the investigation of bond behavior of corroded RC members and focusing on values of maximum pull-out force and, subsequently, bond strength, and not on bond loss, the role of stirrups spacing is more highlighted, where bonding between steel and concrete degrades due to corrosion and the developed maximum pull-out force drops. The usage of dense transverse reinforcement contributes to bond behavior, not only reducing the bond loss rate, but leading to greater values of

In an effort to estimate that influence of stirrups spacing on non-corroded condition, the presence of wide stirrups (Φ8/240 mm) present maximum bond strength equal to 7.04 MPa, whereas quite dense (Φ8/120 mm) and dense stirrups (Φ8/60 mm) equal to 9.10 MPa and 9.53 MPa, respectively. The percentages attributing the increase of bond strength against specimens without stirrups are 35.9%, 75.6% and 84%, respectively. It is noteworthy that, quite dense stirrups spacing (Φ8/120 mm) result in sufficient bond strength levels, whereas further densification leads to minor increase of bond strength, nevertheless delays the bond loss. Extrapolating the abovementioned on real RC structures, stirrups' densification above a certain threshold, is considered inappropriate since it could lead on the one hand to substantial increase of costs and on the other hand to rapid rise of corrosion

Moreover, bond stress - relative slip curves are exported from pull-out tests for each corrosion level and for each category of stirrups spacing, respectively. Typical

The harmful influence of corrosion phenomenon on bond behavior between steel and concrete can be seen examining the bond stress - slip curves of **Figure 10**, and in particular group of specimens without stirrups. Uncorroded specimens - where no surface cracking had been initially observed - showed a quasi linear relationship between bond stresses and relative slip till the point of bond strength's development. During this phase, cracks were occurred, parallel to the axis of steel bar, and gradually were developed due to the radial stresses, which are transferred from steel to concrete. After the development of bond strength, sharp decline of bond stresses and complete bond loss followed. In the case of corroded specimens (without

curves of uncorroded and corroded specimens are shown in **Figure 10**.

maximum pull-out force due to confinement, **Figure 9**.

**50**

**Figure 9.**

*Bond strength values in function with average crack width.*

**Figure 10.** *Bond stress - slip curves obtained from pull out tests.*

stirrups) intense surface concrete cracking was recorded, due to steel corrosion, even in low mass loss levels. As a result of this cracking, bond behavior degrades dramatically, as confirmed by the corresponding curves of **Figure 10**, since both the steel-concrete interface is damaged by the corrosion oxides and the confinement level is deteriorated by the cracks in concrete cover.

On the other hand, as shown by the groups of specimens with stirrups, surface cracking and degree of bond loss is strongly correlated with stirrups spacing. In particular, specimens with quite wide stirrups (Φ8/240 mm) showed that transverse reinforcement has a positive impact on bond behavior in uncorroded condition, as greater values of bond stress indicated. However, when corrosion occurs, accompanied by surface concrete cracking, bond resistance degradates as reflected in responding curves, where bond strength reduces significantly, and subsequently low values of residual bond stress are recorded.

Greater contribution of stirrups to bond behavior was noticed on specimens with quite dense stirrups spacing (Φ8 /120 mm), where even higher bond strength values are indicated, and while bond behavior degrades due to corrosion damage nevertheless does not demonstrate a massive drop mainly due to confinement. In addition, in the case of wide stirrups (Φ8 / 60 mm), full use of bond behavior, even after corrosion damage, occurs, with residual bond stress recorded after the peak of bond strength. It is also noteworthy that specimens of this category had finally ended due to failure under tension. Thus, that densification of stirrups make full use of bearing capacity of steel reinforcing bars and subsequently of RC elements. From the aforementioned, it becomes obvious that the use of dense stirrups (Φ8/60 mm) leads to full bonding between steel and concrete, in uncorroded and corroded conditions respectively, marking both high bond strength and high residual bond stress after the ultimate pull out force.

To conclude, existence of stirrups contributes to bond behavior, and subsequently, greater density of stirrups affects both the reduction of bond loss' degradation rate due to steel corrosion, as well as the increasing of bond stresses values due to greater confinement. The use of wide stirrups (Φ8 / 240 mm) enhances the

bond strength, in uncorroded conditions, about 35.9% in comparison with group of specimens without stirrups. However, when corrosion occurs, the subsequent surface cracking degrades significantly the confinement and the bond between steel and concrete, about 60% reduction against of uncorroded specimens with Φ8 /240 mm. Specimens with quite dense stirrups (Φ8 /120 mm) indicated higher bond strength values, and while bond behavior degrades due to corrosion damage nevertheless does not demonstrate a massive drop mainly due to confinement, about 32%. Finally, Dense stirrups spacing, specimens with Φ8/60 mm, ensure high level of bond behavior, either in terms of bond strength or of residual bond stress, both in uncorroded and corroded conditions. Furthermore, stirrups spacing of 60 mm results in full anchorage of steel reinforcing bars and make full use of their bearing capacity.

### **6. Conclusions**

The present chapter presents an extensive and ongoing experimental research, which was conducted in Laboratory of Technology and Strength of Materials, in University of Patras, and comes to agreement with corresponding results of other scientific studies. The effect of steel corrosion on bond loss between steel and concrete was deeply investigated, including influencing parameters such as concrete cover thickness, density of stirrups and surface concrete cracking. The results of this research, the following outcomes were obtained:


**53**

**Author details**

Patras, Greece

Charis Apostolopoulos and Konstantinos Koulouris\*

\*Address all correspondence to: kkoulouris@upnet.gr

provided the original work is properly cited.

Department of Mechanical Engineering and Aeronautics, University of Patras,

© 2020 The Author(s). Licensee IntechOpen. This chapter is distributed under the terms of the Creative Commons Attribution License (http://creativecommons.org/licenses/ by/3.0), which permits unrestricted use, distribution, and reproduction in any medium,

*Corrosion Effect on Bond Loss between Steel and Concrete*

*DOI: http://dx.doi.org/10.5772/intechopen.94166*

*Corrosion Effect on Bond Loss between Steel and Concrete DOI: http://dx.doi.org/10.5772/intechopen.94166*

*Structural Integrity and Failure*

bearing capacity.

**6. Conclusions**

bond strength, in uncorroded conditions, about 35.9% in comparison with group of specimens without stirrups. However, when corrosion occurs, the subsequent surface cracking degrades significantly the confinement and the bond between steel and concrete, about 60% reduction against of uncorroded specimens with Φ8 /240 mm. Specimens with quite dense stirrups (Φ8 /120 mm) indicated higher bond strength values, and while bond behavior degrades due to corrosion damage nevertheless does not demonstrate a massive drop mainly due to confinement, about 32%. Finally, Dense stirrups spacing, specimens with Φ8/60 mm, ensure high level of bond behavior, either in terms of bond strength or of residual bond stress, both in uncorroded and corroded conditions. Furthermore, stirrups spacing of 60 mm results in full anchorage of steel reinforcing bars and make full use of their

The present chapter presents an extensive and ongoing experimental research, which was conducted in Laboratory of Technology and Strength of Materials, in University of Patras, and comes to agreement with corresponding results of other scientific studies. The effect of steel corrosion on bond loss between steel and concrete was deeply investigated, including influencing parameters such as concrete cover thickness, density of stirrups and surface concrete cracking. The results of

• The width of surface cracking on concrete due to corrosion of steel reinforcement is closely related both to the cover thickness and to the amount of stirrups

• Existence of stirrups contributes to bond behavior, and subsequently, greater density of stirrups affects both the reduction of bond loss' degradation rate due to steel corrosion, as well as the increasing of bond stresses values due to

• Based on the fact that presence of dense connectors, Φ8/60 mm, was accompanied by a clear limitation of the surface cracking development, it appears that the densification of stirrups (through the confinement) contributes positively

• In conclusion, there is a need for further improvement and strengthening of existing technical codes, introducing predictive models of bond loss in function with corrosion damage, cover thickness, surface concrete cracking

• The development of surface cracking in concrete is associated with an

to maintaining bond between steel reinforcement and concrete.

this research, the following outcomes were obtained:

exponential reduction of bonding forces.

or their absence in RC element.

greater confinement.

and density of stirrups.

**52**

### **Author details**

Charis Apostolopoulos and Konstantinos Koulouris\* Department of Mechanical Engineering and Aeronautics, University of Patras, Patras, Greece

\*Address all correspondence to: kkoulouris@upnet.gr

© 2020 The Author(s). Licensee IntechOpen. This chapter is distributed under the terms of the Creative Commons Attribution License (http://creativecommons.org/licenses/ by/3.0), which permits unrestricted use, distribution, and reproduction in any medium, provided the original work is properly cited.

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doi.org/10.1155/2019/3438743

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in Concrete." (2000).

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[23] Li CQ, Zheng JJ. Propagation of reinforcement corrosion in concrete and its effects on structural deterioration. Magazine of Concrete Research.

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Dizaj EA. Residual capacity of corroded reinforced concrete bridge components: State-of-the-art review. Journal of Bridge Engineering. 2019;**24**(7)

corrosion on the bond between concrete and steel rebar. Cement and Concrete Research. 1997;**27**(12):1811-1815

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2016;**49**:4959-4973. DOI: https://doi. org/10.1617/s11527-016-0836-2

[18] Zandi K. Structrural behavior of deteriorated concrete structures [PhD thesis]. Gothenburg, Sweden: Chalmers University of technology; 2010

[19] Zhu W, François R. Prediction of the residual load-bearing capacity of naturally corroded beams using the variability of tension behaviour of corroded steel bars. Structure and Infrastructure Engineering. 2016;**12**(2):143-158. DOI: 10.1080/15732479.2014.996165

[20] Kashani MM, Maddocks JR, Dizaj EA. Residual capacity of corroded reinforced concrete bridge components: State-of-the-art review. Journal of Bridge Engineering. 2019;**24**(7)

[21] Fu X, Chung DDL. Effect of corrosion on the bond between concrete and steel rebar. Cement and Concrete Research. 1997;**27**(12):1811-1815

[22] Tepfers, R., Z. Achillides, A. Azizinamini, G. Balázs, Agniezka Bigaj-van-Vliet, J. Cabrera, J. Cairns, E. Cosenza, J. D. Uijl, R. Eligehausen, B. Engström, L. Erdélyi, P. Gambarova, J. Jirsa, S. Lane, R. León, J. Magnússon, U. Mayer, S. Mccabe, C. Modena, J. Modniks, T. Mottram, K. Noghabai, Koiji Otsuka, J. Ožbolt, Stavrola J. Pantazopoulou, K. Pilakoutas, G. Plizzari, R. Realfonzo, Jesós Rodriguez, G. Rosati, G. Russo, S. Russo, H. Shima, Christiano Schumm, L. Taerwe, V. Tamuzs, T. Ueda, L. Vandewalle and lisabeth Vintzileou. "Fib Bulletin 10. Bond of Reinforcement in Concrete." (2000).

[23] Li CQ, Zheng JJ. Propagation of reinforcement corrosion in concrete and its effects on structural deterioration. Magazine of Concrete Research. 2005;**57**(5):261-271

[24] Abosrra L, Ashour AF, Youseffi M. Corrosion of steel reinforcement in

concrete of different compressive strengths. Construction and Building Materials. 2011;**25**(10):3915-3925

[25] Zandi K, Coronelli D. Anchorage capacity of corroded reinforcement: Eccentric pull-out tests on beam-end specimens. In: Report No. 2010-06, Department of Civil and Environmental Engineering. Goteborg, Sweden: Chalmers University of Technology; 2010

[26] Yalciner H, Eren O, Sensoy S. An experimental study on the bond strength between reinforcement bars and concrete as a function of concrete cover, strength and corrosion level. Cement and Concrete Research. 2012;**42**(5):643-655

[27] Maslehuddin M, Allam I, Al-Sulaimani G, Al-Mana A, Abduljauwad S. Effect of rusting of reinforcing steel on its mechanical properties and bond with concrete. ACI Materials Journal. 1990;**87**(5):496-502

[28] Zandi K, Coronelli D, Lundgren K. Bond capacity of severely corroded bars with corroded stirrups. Magazine of Concrete Research. December 2011;**63**(12):953-968. DOI: https://doi. org/10.1680/macr.10.00200

[29] Lin H, Zhao Y, Ozbolt J, Hans-Wolf R. Bond strength evaluation of corroded steel bars via the surface crack width induced by reinforcement corrosion. Engineering Structures. 2017;**152**:506-522

[30] Apostolopoulos C, Koulouris K, Apostolopoulos A. Correlation of surface cracks of concrete due to corrosion and bond strength (between steel Bar and concrete). Advances in Civil Engineering;**2019**:12. DOI: https:// doi.org/10.1155/2019/3438743

[31] Ozbolt J, Balabanic G, Periskic G, Kuster M. Modelling the effect of damage on transport processes in

**54**

*Structural Integrity and Failure*

**References**

[1] G. Koch, J. Varney, N. Thompson et al., International Measures of

[2] Hou B, Li X, Ma X, et al. The cost of corrosion in China. NPJ Materials

Degradation. 2017;**1**(1):1-10

1998;**95**(6):675-681

2009;**39**(12):1122-1138

[3] Liu Y et al. Modeling the timeto-corrosion cracking in chloride contaminated reinforced concrete structures. ACI Materials Journal.

[4] Angst U, Elsener B, Larsen CK, Vennesland Ø. Critical chloride content in reinforced concrete—A review. Cement and Concrete Research.

[5] Babaee M, Castel A. Chloride diffusivity, chloride threshold, and corrosion initiation in reinforced alkaliactivated mortars: Role of calcium, alkali, and silicate content. Cement and Concrete Research. 2018;**111**:56-71

[6] Cao Y, Gehlen C, Angst U, Wang L, Wang Z, Yao Y. Critical chloride content in reinforced concrete — An updated review considering Chinese experience.

Cement and Concrete Research.

[7] Richardson MG. Fundamentals of Durable Reinforced Concrete. London:

[8] BSI (British Standards Institution). Concrete. Part 1: Specification, Performance, Production and

Conformity." EN 206-1. London; 2000

[9] Fernandez I, Bairán JM, Marí AR. Mechanical model to evaluate steel reinforcement corrosion effects on σ–ε and fatigue curves. Experimental calibration and validation. Engineering

Structures. 2016;**118**:320-333

2019;**117**:58-68

Spon Press; 2002

Prevention, Application, and Economics of Corrosion Technologies Study, NACE International, Houston, TX, USA, 2016.

[10] Tahershamsi M, Fernandez I, Lundgren K, Zandi K. Investigating correlations between crack width, corrosion level and anchorage

capacity. Structure and Infrastructure Engineering. 2017;**13**(10):1294-1307. DOI: 10.1080/15732479.2016.1263673

[11] Zhang W, Song X, Gu X, Li S. Tensile and fatigue behavior of corroded rebars. Construction and Building

[12] Sun X, Kong H, Wang H, Zhang Z. Evaluation of corrosion characteristics and corrosion effects on the mechanical properties of reinforcing steel bars based on three-dimensional scanning. Corrosion Science. 2018;**142**:284-294

Materials. 2012;**34**:409-417

[13] Andisheh K, Scott A,

2018;**167**:188-202

Palermo A, Clucas D. Influence of chloride corrosion on the effective mechanical properties of steel reinforcement. Structure and Infrastructure Engineering. 2019;**15**(8):1036-1048

[14] Gu X, Guo H, Zhou B, Zhang W, Jiang C. Corrosion non-uniformity of steel bars and reliability of corroded RC beams. Engineering Structures.

[15] Apostolopoulos C, Papadakis VG. Consequences of steel corrosion on the ductility properties of reinforcement bar. Construction and Building Materials. 2008;**22**(12):2316-2324

[16] Apostolopoulos C, Kappatos V. Tensile properties of corroded

[17] Fernandez I, Herrador MF,

embedded steel bars B500c in concrete. International Journal of Structural Integrity. 2013;**4**(2):275-294

Marí AR, et al. Structural effects of steel reinforcement corrosion on statically indeterminate reinforced concrete members. Materials and Structures.

concrete. Construction and Building Materials. 2010;**24**:1638-1648

[32] Apostolopoulos C et al. Chloride-induced corrosion of steel reinforcement – Mechanical performance and pit depth analysis. Construction and Building Materials. 2013;**38**:139-146

[33] Broomfield JP. Corrosion of Steel in Concrete: Understanding, Investigation and Repair. London: E & FN Spon; 1997

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[41] Zhou HJ, Zhou YF, Xu YN, Lin ZY, Xing F. Regression analysis of bond

parameters between corroded rebar and concrete based on reposted test data. International Journal of Corrosion. 2018, 2018:18 Article ID 5309243

**Chapter 4**

**Abstract**

of Works

*M. Rosário Oliveira*

assessment, building works

organizational and technological capacity.

maximize the profitability of the works [1].

**1. Introduction**

**57**

Risk Assessment in the Monitoring

The purpose of this chapter is to present a methodology for developing Control, Measurement and Monitoring Plans. It aims to apply risk-based thinking associated with the works control plan. The failures and rework of the works must not be accepted as inevitable or even as certainties. They must be considered permanent

challenges to their management. The importance of using risk assessment

**Keywords:** Control, Measurement and Monitoring Plan, analysis and risk

companies. It improves with more competitive companies and with better

accurately and subsequently monitor their implementation [2, 3].

Building works productivity is not only improved with more works and more

The work of the Construction is developed within a growing and demanding context where rigor and competence in production management are necessary conditions for the provision of the best service and product, being also essentials to

For building companies, the logical choice to ensure competitive advantage with the rest of industry requires the use of new productivity tools and work production control methods. In many manufacturing industries, production processes have been modified with the implementation of systems that limit the existence of the failures and reworks along the production flow. These industries are confined to factories and can implement efficient monitoring systems that define any process

Defects and rework should not be accepted as inevitable or even as certainties but considered as a permanent challenge to the management of the works, being important to use risk assessment techniques there its planning and control [2].

techniques in the planning and control of the production activities of the works is evident. Control, measurement and monitoring process should provide the assessment of risks and failures, should demonstrate technical compliance of the work, and should improve operational efficiency. Thus, it is important to define a methodology for the preparation of the Control, Measurement and Monitoring Plan (PCMM), to be implemented in the execution of the works, in order to ensure the conformity of the works with its technical and regulatory requirements. It must establish which the trials and control inspections, its acceptance criteria, its purposes, frequencies and responsible and it must also identify and assess its risks.

[42] Fischer C, Oˇzbolt J. An appropriate indicator for bond strength degradation due to reinforcement corrosion. In: Proceedings of the 8th International Conference on Fracture Mechanics of Concrete and Concrete Structures (FraMCoS). Toledo, Spain; March 2013. pp. 1828-1835

[43] Almusallam A, Al-Gahtani A, Aziz A, et al. Effect of reinforcement corrosion on bond strength. Construction and Building Materials. 1996;**10**(2):123-129

[44] Rodriguez J, Ortega LM, Garda AM. Assessment of structural elements with corroded reinforcement. In: Proceedings of International Conference Corrosion and Corrosion Protection of Steel in Concrete. Vol. 1. Sheffield, England: University of Sheffield; July 1994. pp. 172-185

**Chapter 4**

*Structural Integrity and Failure*

Materials. 2010;**24**:1638-1648

[32] Apostolopoulos C et al. Chloride-induced corrosion of steel reinforcement – Mechanical performance and pit depth analysis. Construction and Building Materials.

2013;**38**:139-146

2019.

concrete. Construction and Building

parameters between corroded rebar and concrete based on reposted test data. International Journal of Corrosion. 2018, 2018:18 Article ID 5309243

[42] Fischer C, Oˇzbolt J. An appropriate indicator for bond strength degradation due to reinforcement corrosion. In: Proceedings of the 8th International Conference on Fracture Mechanics of Concrete and Concrete Structures (FraMCoS). Toledo, Spain; March 2013.

[43] Almusallam A, Al-Gahtani A, Aziz A, et al. Effect of reinforcement

Construction and Building Materials.

[44] Rodriguez J, Ortega LM, Garda AM. Assessment of structural elements with corroded reinforcement. In: Proceedings of International Conference Corrosion and Corrosion Protection of Steel in Concrete. Vol. 1. Sheffield, England: University of Sheffield; July 1994.

corrosion on bond strength.

1996;**10**(2):123-129

pp. 172-185

pp. 1828-1835

[33] Broomfield JP. Corrosion of Steel in Concrete: Understanding, Investigation and Repair. London: E & FN Spon; 1997

[34] Sh A. Techniques for inducing accelerated corrosion of steel in concrete. Arabian Journal for Science and Engineering. 2009;**34**(2C):95-104

[35] Apostolopoulos C, Koulouris K, Basdeki M. Damage parameters of rebars in marine environment and fatigue life. In: Rilem SMSS Conference

[36] Model Code EBCEB-FIP. 2010. fib Model Code Concr. Struct. 2010. Lausanne, Switzerland. 2013:152-189

Chung L. Bond behavior of corroded reinforcement bars. ACI Materials

[38] Lundgren K. Effect of corrosion on the bond between steel and concrete: An overview. Magazine of Concrete Research. 2007;**59**(6):447-461

Navarro-Gutierrez S, Teran-Guillen J. Residual flexure capacity of corroded reinforced concrete beams. Engineering Structures. 2007;**29**(6):1145-1152

[40] Andrade C, Cesetti A, Mancini G, Tondolo F. Estimating corrosion attack in reinforced concrete by means of crack opening. Structural Concrete.

[41] Zhou HJ, Zhou YF, Xu YN, Lin ZY, Xing F. Regression analysis of bond

[37] Auyeung Y, Balaguru P,

Journal. 2000;**97**:214-220

[39] Torres-Acosta AA,

2016;**17**(4):533-540

**56**

## Risk Assessment in the Monitoring of Works

*M. Rosário Oliveira*

### **Abstract**

The purpose of this chapter is to present a methodology for developing Control, Measurement and Monitoring Plans. It aims to apply risk-based thinking associated with the works control plan. The failures and rework of the works must not be accepted as inevitable or even as certainties. They must be considered permanent challenges to their management. The importance of using risk assessment techniques in the planning and control of the production activities of the works is evident. Control, measurement and monitoring process should provide the assessment of risks and failures, should demonstrate technical compliance of the work, and should improve operational efficiency. Thus, it is important to define a methodology for the preparation of the Control, Measurement and Monitoring Plan (PCMM), to be implemented in the execution of the works, in order to ensure the conformity of the works with its technical and regulatory requirements. It must establish which the trials and control inspections, its acceptance criteria, its purposes, frequencies and responsible and it must also identify and assess its risks.

**Keywords:** Control, Measurement and Monitoring Plan, analysis and risk assessment, building works

### **1. Introduction**

Building works productivity is not only improved with more works and more companies. It improves with more competitive companies and with better organizational and technological capacity.

The work of the Construction is developed within a growing and demanding context where rigor and competence in production management are necessary conditions for the provision of the best service and product, being also essentials to maximize the profitability of the works [1].

For building companies, the logical choice to ensure competitive advantage with the rest of industry requires the use of new productivity tools and work production control methods. In many manufacturing industries, production processes have been modified with the implementation of systems that limit the existence of the failures and reworks along the production flow. These industries are confined to factories and can implement efficient monitoring systems that define any process accurately and subsequently monitor their implementation [2, 3].

Defects and rework should not be accepted as inevitable or even as certainties but considered as a permanent challenge to the management of the works, being important to use risk assessment techniques there its planning and control [2].

### *Structural Integrity and Failure*

The "Operational Planning and Control" requirement specified in ISO 9001: 2015 indicates that organizations must plan, implement and manage the processes necessary for the supply of the product and service provision (Works) to ensure compliance with customer requirements (Owners) [4, 5].

provides quick ordering of risks at different levels of significance. Adequate scales of the likelihood and consequences criteria and the definition of risk matrix are the inputs essential the risk assessment process. The likelihood criteria scale (Probability) should cover the relevant domain for the case in analysis. The consequence criteria scale (Severity) should cover the range of different types of the consequences to be considered, from the maximum plausible consequence to the smallest plausible consequence to be considered. All scales can have any number

To order the risks, the consequence descriptor (Severity) that best suits the situation is chosen first, then the probability (Probability) of occurrence of these consequences is then defined. The risk level defined in the LCM may be associated

Failure Mode and Effects Analysis (FMEA) is a technique for analyzing the reliability of the products, systems or processes. It is used to identify modes in which components of products, systems or processes may to fail to performance their functions. There are several FMEA applications: design FMEA which is used for components and products; system FMEA which is used for systems and process FMEA which is used for manufacturing processes and procedures and assembly. FMEA is also used in risk assessment and this requires detailed information on the phases of the case under study, to permit a significative analysis its failure modes. To perform a FMEA is fundamental the experience of the evaluators, the knowledge of the history of the failures and the causes, of the decision criteria and/or accep-

Severity, Probability and Detection indices are the inputs for FMEA. Their scales must be adequate to the consequences and the likelihood of the events that combined define the risk matrix. Additionally, the level of risk combined with the failure detection index determines the Risk Priority Number (RPN). The scales these three criteria can have any number of levels, the most common being scales of

To order the risks identifies, the consequence descriptor is chosen first, which best adapts to the situation, then the probability of occurrence of that consequence is defined. With the third detection descriptor, the Risk Priority Number (RPN) is

If we accept that all results of the building works processes are subject to uncertainty, then we can conclude that there is need a risk assessment for each of these. So, we can find the risk assessment in the reception of materials and the control of work in progress, in short, in all critical processes of the works, that it can

According to ISO 9001 standard, production and service provision processes should be implemented under controlled conditions. This determines that the operational processes of the works are implemented a controlled mode, before, during and after its completion, particularly all its critical activities. Among other requirements, this condition includes the implementation of monitoring and measurement activities, in adequate steps, to verify if the criteria of the process or its outputs and

With this aim, it is essential to establish Plans of Control, Measurement and Monitoring (PCMM), assess the risk and define actions for its treatment, and implement the control of operational processes in conformity with the defined criteria. The PCMM is the document that specifies which are the trials and control inspections, the purposes, the acceptance criteria, the frequency, those responsible for the monitoring, and the records of the results obtained, in order to retain the objective evidence to the satisfaction of technical and regulatory requirements of

of the levels, the most common being the scales of 3, 4 or 5 levels [7].

with a decision rule, such as, for example, treating or not treating the risk.

tance of the specific risks, and of the steps of the case under study [9].

defined. RPN is used to prioritize the of risk mitigation actions.

the criteria for acceptance of the product and services were satisfied.

ensure your technical and regulatory conformity.

3, 5 or 10 levels [7, 8, 10–13].

*Risk Assessment in the Monitoring of Works DOI: http://dx.doi.org/10.5772/intechopen.93957*

the Work [2, 14].

**59**

Control, measurement and monitoring process of the works should provide the specific actions to address the risks and opportunities and achieve the objectives specified in their planning [4].

Thus, it is necessary to establish Plans of Control, Measurement and Monitoring Plans (PCMM), assess risk and define actions for its treatment, and implement the control of operational processes in accordance with the defined criteria. Plans of Control, Monitoring and Measurement (PCMM) are required to: i) demonstrate the technical compliance of the Work; ii) continuously improve operational effectiveness [2].

### **2. Risk assessment and plans of control**

Risk assessment is an integral part to the various process of the works, aiming at prevention and its resilience.

To understand the risk assessment is necessary to know the definition adopted for "*risk*" in ISO Guide 73 (*Risk management – Vocabulary - Guidelines for use in standards*) and ISO 31000 (*Risk management - Principles and guidelines*). In according to these standards' "*risk*" is defined as the "*effect of uncertainty on objectives*" [6, 7].

This definition gives the possibility of we considering the risk as a threat or an opportunity. However, it is not gives clue as to how to quantify the risk. For this purpose, we must use the definition of the risk as being the combination of the probability of the occurrence an event (Likelihood or Frequency) and its consequences (Severity), something that is referred to in complementary notes these standards. The risk assessment compares the results of risk analysis with risk criteria (frequency and severity criteria) to determine whether the risk is acceptable or tolerable. It requires the identification and analysis of the events (occurrence or change of a set of circumstances) and to determine the risk level. The risk level is a function of its consequence (or Severity) with its likelihood (or Probability) and measures the magnitude of the risk. The risk valuation criteria are references in respect of which the significance of the risk is assessed [8].

The events that influence the results of processes under analysis can be identified and classified between risks and opportunities. The opportunities are directed to the organization's strategic planning processes and the risks are analyzed by quantifying the probability of occurrence and the severity its effects, to determine its level and the actions of mitigation [4].

Risk assessment may be made at difference degrees of depth and detail, we using various techniques ranging from simple to complex. We must use the risk criteria consistent with the scope of the process under analysis, as well as the technique and the assessment results. Likelihood/Consequence Matrix (LCM) and Failure Mode and Effects Analysis (FMEA) are two of several techniques of great application in risk assessment.

Likelihood/Consequence Matriz (LCM) is a technique that combines the probability of the event under analysis with its effects, to define a qualification of the level of risk. The form of the matrix and the definitions that apply to it depend on the context in which it is used. This technique is used to classify risks and their sources, and to identify your treatments. It can be used in situations where there is not enough data for a detailed analysis or when the situation does not justify the time and effort for a more quantitative analysis. It is relatively easy to use, and it

### *Risk Assessment in the Monitoring of Works DOI: http://dx.doi.org/10.5772/intechopen.93957*

The "Operational Planning and Control" requirement specified in ISO 9001: 2015 indicates that organizations must plan, implement and manage the processes necessary for the supply of the product and service provision (Works) to ensure

Control, measurement and monitoring process of the works should provide the specific actions to address the risks and opportunities and achieve the objectives

Thus, it is necessary to establish Plans of Control, Measurement and Monitoring Plans (PCMM), assess risk and define actions for its treatment, and implement the control of operational processes in accordance with the defined criteria. Plans of Control, Monitoring and Measurement (PCMM) are required to: i) demonstrate the technical compliance of the Work; ii) continuously improve operational

Risk assessment is an integral part to the various process of the works, aiming at

To understand the risk assessment is necessary to know the definition adopted for "*risk*" in ISO Guide 73 (*Risk management – Vocabulary - Guidelines for use in standards*) and ISO 31000 (*Risk management - Principles and guidelines*). In according to these standards' "*risk*" is defined as the "*effect of uncertainty on objectives*" [6, 7]. This definition gives the possibility of we considering the risk as a threat or an opportunity. However, it is not gives clue as to how to quantify the risk. For this purpose, we must use the definition of the risk as being the combination of the probability of the occurrence an event (Likelihood or Frequency) and its consequences (Severity), something that is referred to in complementary notes these standards. The risk assessment compares the results of risk analysis with risk criteria (frequency and severity criteria) to determine whether the risk is acceptable or tolerable. It requires the identification and analysis of the events (occurrence or change of a set of circumstances) and to determine the risk level. The risk level is a function of its consequence (or Severity) with its likelihood (or Probability) and measures the magnitude of the risk. The risk valuation criteria are references in

The events that influence the results of processes under analysis can be identified and classified between risks and opportunities. The opportunities are directed to the organization's strategic planning processes and the risks are analyzed by quantifying the probability of occurrence and the severity its effects, to determine

Risk assessment may be made at difference degrees of depth and detail, we using various techniques ranging from simple to complex. We must use the risk criteria consistent with the scope of the process under analysis, as well as the technique and the assessment results. Likelihood/Consequence Matrix (LCM) and Failure Mode and Effects Analysis (FMEA) are two of several techniques of great application in

Likelihood/Consequence Matriz (LCM) is a technique that combines the probability of the event under analysis with its effects, to define a qualification of the level of risk. The form of the matrix and the definitions that apply to it depend on the context in which it is used. This technique is used to classify risks and their sources, and to identify your treatments. It can be used in situations where there is not enough data for a detailed analysis or when the situation does not justify the time and effort for a more quantitative analysis. It is relatively easy to use, and it

compliance with customer requirements (Owners) [4, 5].

specified in their planning [4].

*Structural Integrity and Failure*

prevention and its resilience.

**2. Risk assessment and plans of control**

respect of which the significance of the risk is assessed [8].

its level and the actions of mitigation [4].

risk assessment.

**58**

effectiveness [2].

provides quick ordering of risks at different levels of significance. Adequate scales of the likelihood and consequences criteria and the definition of risk matrix are the inputs essential the risk assessment process. The likelihood criteria scale (Probability) should cover the relevant domain for the case in analysis. The consequence criteria scale (Severity) should cover the range of different types of the consequences to be considered, from the maximum plausible consequence to the smallest plausible consequence to be considered. All scales can have any number of the levels, the most common being the scales of 3, 4 or 5 levels [7].

To order the risks, the consequence descriptor (Severity) that best suits the situation is chosen first, then the probability (Probability) of occurrence of these consequences is then defined. The risk level defined in the LCM may be associated with a decision rule, such as, for example, treating or not treating the risk.

Failure Mode and Effects Analysis (FMEA) is a technique for analyzing the reliability of the products, systems or processes. It is used to identify modes in which components of products, systems or processes may to fail to performance their functions. There are several FMEA applications: design FMEA which is used for components and products; system FMEA which is used for systems and process FMEA which is used for manufacturing processes and procedures and assembly. FMEA is also used in risk assessment and this requires detailed information on the phases of the case under study, to permit a significative analysis its failure modes. To perform a FMEA is fundamental the experience of the evaluators, the knowledge of the history of the failures and the causes, of the decision criteria and/or acceptance of the specific risks, and of the steps of the case under study [9].

Severity, Probability and Detection indices are the inputs for FMEA. Their scales must be adequate to the consequences and the likelihood of the events that combined define the risk matrix. Additionally, the level of risk combined with the failure detection index determines the Risk Priority Number (RPN). The scales these three criteria can have any number of levels, the most common being scales of 3, 5 or 10 levels [7, 8, 10–13].

To order the risks identifies, the consequence descriptor is chosen first, which best adapts to the situation, then the probability of occurrence of that consequence is defined. With the third detection descriptor, the Risk Priority Number (RPN) is defined. RPN is used to prioritize the of risk mitigation actions.

If we accept that all results of the building works processes are subject to uncertainty, then we can conclude that there is need a risk assessment for each of these. So, we can find the risk assessment in the reception of materials and the control of work in progress, in short, in all critical processes of the works, that it can ensure your technical and regulatory conformity.

According to ISO 9001 standard, production and service provision processes should be implemented under controlled conditions. This determines that the operational processes of the works are implemented a controlled mode, before, during and after its completion, particularly all its critical activities. Among other requirements, this condition includes the implementation of monitoring and measurement activities, in adequate steps, to verify if the criteria of the process or its outputs and the criteria for acceptance of the product and services were satisfied.

With this aim, it is essential to establish Plans of Control, Measurement and Monitoring (PCMM), assess the risk and define actions for its treatment, and implement the control of operational processes in conformity with the defined criteria. The PCMM is the document that specifies which are the trials and control inspections, the purposes, the acceptance criteria, the frequency, those responsible for the monitoring, and the records of the results obtained, in order to retain the objective evidence to the satisfaction of technical and regulatory requirements of the Work [2, 14].

### **3. PCCM form with risk assessment**

This heading, on context of the building works processes, a methodology is proposed for elaboration of the PCMM, following the approach of risk-based thinking. It applies to operational processes considering the most critical work activities.

According to [2] the steps to be followed in the preparation of the PCMM require the definition of the: critical works activities that need to be controlled; inspections and tests to be performed in each critical activity; acceptance, frequency and sampling criteria of the inspections and tests; those responsible for control, measurement and monitoring; records of the results; risk assessment and its effects; and corrective and preventive actions to be implemented. **Figure 1** shows the flowchart of the methodology for preparing the PCMM.

### **3.1 Risk assessment with Likelihood/Consequence Matriz (LCM)**

**Table 1** shows the template PCMM with LCM which takes the form of a matrix of the columns and rows whose contents are explained below.

**Logo**

**61**

**Item**

**Activities**

**Inspection/**

**Purposes**

**Acceptance**

**Frequency/**

**Responsible**

**Re**

**Risks**

**Description**

**Effects**

**Risk**

**Controlled**

**Actions**

**Risks?**

**(13)**

**(6)**

**cords**

**(8)**

**(9)**

**(10)**

**Assessment**

**(11)**

P S RN

**(12)**

*Risk Assessment in the Monitoring of Works DOI: http://dx.doi.org/10.5772/intechopen.93957*

**(7)**

**Sampling**

**(5)**

**criteria**

**(4)**

**(3)**

**Testing**

**(2)**

**(1)**

**(1)**

**PCMM - Plan Control,** 

**Measurement**

 **and Monitoring**

**Process/Work:**

**Revision:**

**Page:**

 **1/1**

**Figure 1.** *Flowchart for preparing PCMM [2].*

### **Logo PCMM - Plan Control, Measurement and Monitoring Revision: Page: 1/1 Process/Work: Item (1) Activities (1) Inspection/ Testing (2) Purposes (3) Acceptance criteria (4) Frequency/ Sampling (5) Responsible (6) Re cords (7) Risks (8) Description (9) Effects (10) Risk Assessment (11) Controlled Risks? (12) Actions (13)** P S RN

### *Risk Assessment in the Monitoring of Works DOI: http://dx.doi.org/10.5772/intechopen.93957*

**3. PCCM form with risk assessment**

*Structural Integrity and Failure*

**Figure 1.**

**60**

*Flowchart for preparing PCMM [2].*

This heading, on context of the building works processes, a methodology is proposed for elaboration of the PCMM, following the approach of risk-based thinking. It applies to operational processes considering the most critical work activities. According to [2] the steps to be followed in the preparation of the PCMM require the definition of the: critical works activities that need to be controlled; inspections and tests to be performed in each critical activity; acceptance,

frequency and sampling criteria of the inspections and tests; those responsible for control, measurement and monitoring; records of the results; risk assessment and its effects; and corrective and preventive actions to be implemented. **Figure 1** shows

**Table 1** shows the template PCMM with LCM which takes the form of a matrix

the flowchart of the methodology for preparing the PCMM.

of the columns and rows whose contents are explained below.

**3.1 Risk assessment with Likelihood/Consequence Matriz (LCM)**


**Table 1.** *PCMM templatewith*

 *LCM [2].* Using the LCM technique for risk assessment in the preparation of the PCMM,

Columns (1) of the PCMM is identified the critical activities of the work that will be monitored. The critical activities are those ensure the technical conformity of the

In the columns (2-3) are used to indicate the kind of the inspections and tests to be applied in quality control, their purpose and that is to be measured to monitor critical activity in analysis. The column (4) is used to propose which normative and regulatory references and criteria for acceptance that are used to analyze the results

At each inspection/test, in the columns (5-7) are indicated their sampling frequency, who is responsible for carrying out the control and which record to use to compile their results. These items are intended to ensure the systematic control,

Risk assessment is carried out at each stage of control, measurement and monitoring. Thus, for each critical activity is identified the events whose outputs may not

The identification of risk and its effects is done on the columns (8-10) of the PCMM. For each critical activity and its control, the Risk is identified, its description is made and its effect is characterized. In this way, the event associated with it will be featured and you can review the respective controls, measurements and monitoring, especially in cases where it is not possible to act on the causes. This characterization will allow the reflection on the consequences of the effect, which will allow to assess its severity using the defined criteria. It is intended to briefly describe the effect of the risk previously identified in order to better identify the

Finally, the risk identified in (8), described in (9) and with the effect identified

**3** .

4

3

2

1

Risk assessment is carried out using two criteria: Probability (P) and Severity (S). According to [2], the score criteria to be used in the estimation of Probability

Then, for each risk or failure mode, its Probability (P) is given it a score. After, the analysis of the consequences and its effects, the same is done for Severity (S). The scale of these scores should be assigned based on our experience with the

The Risk Number (RN) classifies the assessed risk. Thus, if the Probability (P) and Severity (S) scores are multiplied, we obtain the RN in each case. Using **Table 4** found in [2], the RN that we can obtain vary between the minimum value 1 and the

Therefore, high risk is classified when the NB is higher than 9, medium risk is classified when the NB is 9 and the low risk is classified when the NB is lesser than 9.

**Probability Category Description Score**

in according [2] the following guidelines must be used:

obtained in the inspections and in the tests carried out.

measurement and monitoring in each critical activity.

in (10), it is assessed in columns (11) of the PCMM.

(P) and Severity (S) are proposed in **Tables 2** and

High Occurs often.

Medium Probably, it has occurred several times.

Low Probably, it has already occurred.

Remote Probably, but never occurred.

work with your design.

*Risk Assessment in the Monitoring of Works DOI: http://dx.doi.org/10.5772/intechopen.93957*

be what expected.

critical impact on its activity.

activity in question.

maximum value 16.

**Table 2.**

**63**

*Risk probability criteria (P) [2].*

### *Structural Integrity and Failure*

### *Risk Assessment in the Monitoring of Works DOI: http://dx.doi.org/10.5772/intechopen.93957*

Using the LCM technique for risk assessment in the preparation of the PCMM, in according [2] the following guidelines must be used:

Columns (1) of the PCMM is identified the critical activities of the work that will be monitored. The critical activities are those ensure the technical conformity of the work with your design.

In the columns (2-3) are used to indicate the kind of the inspections and tests to be applied in quality control, their purpose and that is to be measured to monitor critical activity in analysis. The column (4) is used to propose which normative and regulatory references and criteria for acceptance that are used to analyze the results obtained in the inspections and in the tests carried out.

At each inspection/test, in the columns (5-7) are indicated their sampling frequency, who is responsible for carrying out the control and which record to use to compile their results. These items are intended to ensure the systematic control, measurement and monitoring in each critical activity.

Risk assessment is carried out at each stage of control, measurement and monitoring. Thus, for each critical activity is identified the events whose outputs may not be what expected.

The identification of risk and its effects is done on the columns (8-10) of the PCMM. For each critical activity and its control, the Risk is identified, its description is made and its effect is characterized. In this way, the event associated with it will be featured and you can review the respective controls, measurements and monitoring, especially in cases where it is not possible to act on the causes. This characterization will allow the reflection on the consequences of the effect, which will allow to assess its severity using the defined criteria. It is intended to briefly describe the effect of the risk previously identified in order to better identify the critical impact on its activity.

Finally, the risk identified in (8), described in (9) and with the effect identified in (10), it is assessed in columns (11) of the PCMM.

Risk assessment is carried out using two criteria: Probability (P) and Severity

(S). According to [2], the score criteria to be used in the estimation of Probability (P) and Severity (S) are proposed in **Tables 2** and **3**.

Then, for each risk or failure mode, its Probability (P) is given it a score. After, the analysis of the consequences and its effects, the same is done for Severity (S).

The scale of these scores should be assigned based on our experience with the activity in question.

The Risk Number (RN) classifies the assessed risk. Thus, if the Probability (P) and Severity (S) scores are multiplied, we obtain the RN in each case. Using **Table 4** found in [2], the RN that we can obtain vary between the minimum value 1 and the maximum value 16.

Therefore, high risk is classified when the NB is higher than 9, medium risk is classified when the NB is 9 and the low risk is classified when the NB is lesser than 9.


**Table 2.** *Risk probability criteria (P) [2].*

**Logo**

**62**

**Item**

**Activities**

**Inspection/**

**Purposes**

**Acceptance**

**Frequency/**

**Responsible**

**Re**

**Risks**

**Description**

**Effects**

**Risk**

**Controlled**

**Actions**

*Structural Integrity and Failure*

**Risks?**

**(13)**

**(6)**

**cords**

**(8)**

**(9)**

**(10)**

**Assessment**

**(11)**

P S RN

**(12)**

**(7)**

**Sampling**

**(5)**

**criteria**

**(4)**

**(3)**

**Testing**

**(2)**

**(1)**

Prepared by:

**Table 1.**

*PCMM template with LCM [2].*

Date: \_/\_/\_

Verified by:

Date: –/–/–

Responsible:

Date:

 –/–/–

**(1)**

**PCMM - Plan Control,** 

**Measurement**

 **and Monitoring**

**Process/Work:**

**Revision:**

**Page:**

 **1/1**


### **Table 3.**

*Risk severity criteria (G) [2].*


**Table 4.**

*Risk number [2].*

With the classification of the RN done, the appreciation of events is began whose outputs are not the desired. Thus, the undesired outputs will be characterized, making it possible to make revision the inspections/tests identified and to adopt another type of control, measurement and monitoring.

In the column (12) the RN high, medium or low is identified and whether adequately is controlled or not.

However, according to the criteria defined in [2], the risk is adequately controlled (Yes) if the NR is lesser than 9 (for medium or low risks) and uncontrolled (No) if NR is higher than 9. For the uncontrolled risks (No) are must indicate the action required to mitigate them.

In the column (13) of the PCMM the actions to mitigate uncontrolled risk are indicated, making it adequately controlled.

### **3.2 Risk assessment with failure and effect modes analysis (FMEA)**

In **Table 5** we see the template of a PCMM with FMEA that takes too the form of a matrix, constituted by a set of columns and rows containing information relating to the same items already explained above.

Using the FMEA technique for risk assessment, in the preparation of the PCMM we can use the following guidelines:

On the PCMM with FMEA form the fields (1) to (10) are the same as those of PCMM with LCM form. Field (11) is used to assess the risk identified in (8), described in (9) and with the effects identified in (10). Here, risk assessment will be carried out according to the three FMEA criteria, described below. The classification of risk depends on the combination of probability (P), gravity (S) and detection (D).

Thus, for each risk or failure mode, we must analysis he consequences of their effects and to estimate its Severity (S). After to analysis its Probability (P) we must assign a it a score. Then, we must assign the Detection Index (D) to order the number priority of risk (NPR).

**Logo**

**65**

**Item**

**Activities**

**Inspection/**

**Purposes**

**Acceptance**

**Frequency/**

**Responsible**

**Re**

**Risks**

**Description**

**Effects**

**Risk**

**Controlled**

**Actions**

**Risks?**

**(13)**

**(12)**

*Risk Assessment in the Monitoring of Works DOI: http://dx.doi.org/10.5772/intechopen.93957*

**(6)**

**cords**

**(8)**

**(9)**

**(10)**

**Assessment**

**(11)**

P S D RPN

**(7)**

**Sampling**

**(5)**

**criteria**

**(4)**

**(3)**

**Testing**

**(2)**

**(1)**

**(1)**

**PCMM - Plan Control,** 

**Measurement**

 **and Monitoring**

**Process/Work:**

**Revision:**

**Page:**

**1/1**

### *Risk Assessment in the Monitoring of Works DOI: http://dx.doi.org/10.5772/intechopen.93957*


With the classification of the RN done, the appreciation of events is began whose

**Probability (P) 1** *12 3 4*

**12 3 4**

**2** *24 6 8* **3** *3 6 9 12* **4** *4 8 12 16*

**Severity Category Description Score** High Requires the re-inspection/rework the whole lot 4 Medium Requires the re-inspection/rework part of the lot 3 Low Requires adjustments in inspection/inspection/testing 2 Negligible Does not require specific actions 1

outputs are not the desired. Thus, the undesired outputs will be characterized, making it possible to make revision the inspections/tests identified and to adopt

**Risk Number Severity (S)**

In the column (12) the RN high, medium or low is identified and whether

However, according to the criteria defined in [2], the risk is adequately controlled (Yes) if the NR is lesser than 9 (for medium or low risks) and uncontrolled (No) if NR is higher than 9. For the uncontrolled risks (No) are must indicate the

In the column (13) of the PCMM the actions to mitigate uncontrolled risk are

In **Table 5** we see the template of a PCMM with FMEA that takes too the form of a matrix, constituted by a set of columns and rows containing information relating

Using the FMEA technique for risk assessment, in the preparation of the PCMM

On the PCMM with FMEA form the fields (1) to (10) are the same as those of

The classification of risk depends on the combination of probability (P), gravity (S)

Thus, for each risk or failure mode, we must analysis he consequences of their effects and to estimate its Severity (S). After to analysis its Probability (P) we must assign a it a score. Then, we must assign the Detection Index (D) to order the

PCMM with LCM form. Field (11) is used to assess the risk identified in (8), described in (9) and with the effects identified in (10). Here, risk assessment will be carried out according to the three FMEA criteria, described below.

**3.2 Risk assessment with failure and effect modes analysis (FMEA)**

another type of control, measurement and monitoring.

adequately is controlled or not.

**Table 3.**

**Table 4.** *Risk number [2].*

*Risk severity criteria (G) [2].*

*Structural Integrity and Failure*

action required to mitigate them.

indicated, making it adequately controlled.

to the same items already explained above.

we can use the following guidelines:

number priority of risk (NPR).

and detection (D).

**64**


### **Table 5.**

*PCMM template with FMEA.*

Therefore, it is necessary to establish the criteria to be used in the estimation of Probability (P), Severity (S) and Detection (D) values. **Table 2** shows the values of Probability (P), **Table 3** shows the values of Severity (S) and **Table 6** shows the

For the assignment of the scale of these scores, the experience we have of the activity under analysis is too important. But we are looking for is the risk priority order. Thus, if the Probability, Severity and Detection scores are multiplied, we will

In accordance with **Table 4** we can classify High Risk as values higher than 9, Medium Risk values of 9 and Low Risk values lower than 9. In each class of risk, we must appreciate the ranking of the priority, namely, which are the most priority

Fields (12) and (13) on the PCMM with FMEA form are the same too as the fields explained above (PCMM with LCM). Thus, according to the criteria defined for the Risk Number, in field (12) we must consider the Risk appropriately controlled if the RN is ≤9 (for Medium or Low Risks), otherwise we should consider it uncontrolled (N). For the uncontrolled risks. Then, we must then indicate the action required to mitigate them. In field (13) we should indicate what actions are necessary to mitigate the most priority uncontrolled Risk, making it adequately

Then, a brief presentation of a PCMM used in a works is made. The objective is to present two practical examples of the use of PCMM with risk assessment, the first under the LCM technique and the second under the FMEA technique, both exam-

This case PCMM were prepared for the critical activities of the building works.

The methodology set out above was followed in preparing of the PCMM with

The use of this procedure allows, on the one hand to plan the inspections and tests to make the critical activities, so as to control, measure and monitor all the work, on the other hand, it helps mitigate the risk in the events with a negative

work" has the highest risk number (RN = 9), and it is classified as medium risk.

**Detection Category Description Score**

's structure.

"Installation False-

3

2

1

However, **Table 7** shows an excerpt of the PCMM with LCM prepared for the

LCM, and risk assessment was made under Likelihood/Consequence Matrix.

It is observed that the planned control for the activity of the

values for Detection (D).

*Risk Assessment in the Monitoring of Works DOI: http://dx.doi.org/10.5772/intechopen.93957*

controlled.

**4. Application examples**

**4.1 PCMM with LCM**

**Table 6.**

**67**

*Risk detection criteria (D).*

obtain the Risk Priority Number (RPN) in each case.

events where the outputs may not be desired.

ples from the methodology above mentioned.

impact on the development of work.

execution of reinforced concrete beams of the building

High High difficulty detection

Moderate Medium difficulty detection

Low Low difficulty detection

### *Structural Integrity and Failure*

### *Risk Assessment in the Monitoring of Works DOI: http://dx.doi.org/10.5772/intechopen.93957*

Therefore, it is necessary to establish the criteria to be used in the estimation of Probability (P), Severity (S) and Detection (D) values. **Table 2** shows the values of Probability (P), **Table 3** shows the values of Severity (S) and **Table 6** shows the values for Detection (D).

For the assignment of the scale of these scores, the experience we have of the activity under analysis is too important. But we are looking for is the risk priority order. Thus, if the Probability, Severity and Detection scores are multiplied, we will obtain the Risk Priority Number (RPN) in each case.

In accordance with **Table 4** we can classify High Risk as values higher than 9, Medium Risk values of 9 and Low Risk values lower than 9. In each class of risk, we must appreciate the ranking of the priority, namely, which are the most priority events where the outputs may not be desired.

Fields (12) and (13) on the PCMM with FMEA form are the same too as the fields explained above (PCMM with LCM). Thus, according to the criteria defined for the Risk Number, in field (12) we must consider the Risk appropriately controlled if the RN is ≤9 (for Medium or Low Risks), otherwise we should consider it uncontrolled (N). For the uncontrolled risks. Then, we must then indicate the action required to mitigate them. In field (13) we should indicate what actions are necessary to mitigate the most priority uncontrolled Risk, making it adequately controlled.

### **4. Application examples**

Then, a brief presentation of a PCMM used in a works is made. The objective is to present two practical examples of the use of PCMM with risk assessment, the first under the LCM technique and the second under the FMEA technique, both examples from the methodology above mentioned.

### **4.1 PCMM with LCM**

This case PCMM were prepared for the critical activities of the building works. However, **Table 7** shows an excerpt of the PCMM with LCM prepared for the execution of reinforced concrete beams of the building's structure.

The methodology set out above was followed in preparing of the PCMM with LCM, and risk assessment was made under Likelihood/Consequence Matrix.

The use of this procedure allows, on the one hand to plan the inspections and tests to make the critical activities, so as to control, measure and monitor all the work, on the other hand, it helps mitigate the risk in the events with a negative impact on the development of work.

It is observed that the planned control for the activity of the "Installation Falsework" has the highest risk number (RN = 9), and it is classified as medium risk.


**Table 6.** *Risk detection criteria (D).*

**Logo**

**66**

**Item**

**Activities**

**Inspection/**

**Purposes**

**Acceptance**

**Frequency/**

**Responsible**

**Re**

**Risks**

**Description**

**Effects**

**Risk**

**Controlled**

**Actions**

*Structural Integrity and Failure*

**Risks?**

**(13)**

**(12)**

**(6)**

**cords**

**(8)**

**(9)**

**(10)**

**Assessment**

**(11)**

P S D RPN

**(7)**

**Sampling**

**(5)**

**criteria**

**(4)**

**(3)**

**Testing**

**(2)**

**(1)**

Prepared by:

Date: \_/\_/\_

> **Table 5.**

*PCMM template with FMEA.*

Verified by:

Date: –/–/–

Responsible:

Date:

–/–/–

**(1)**

**PCMM - Plan Control,** 

**Measurement**

 **and Monitoring**

**Process/Work:**

**Revision:**

**Page:**

**1/1**


**Logo**

**69**

**Item**

**Activities**

**Inspection/**

**Purposes**

**Acceptance**

**Frequency/**

**Responsible**

**Re**

**Risks**

**Description**

**Effects**

**Risk**

**Controlled**

**Actions**

**(9)**

**(10)**

**Assessment**

**Risks?**

**(13)**

**(11)**

**P S RN**

**(12)**

*Risk Assessment in the Monitoring of Works DOI: http://dx.doi.org/10.5772/intechopen.93957*

**(6)**

**cords**

**(8)**

**(7)**

**Sampling**

**(5)**

**criteria**

**(4)**

**(3)**

**Testing**

**(2)**

**(1)**

6

 Placement

Visual

C of

Structural

By

Overseer and

Control

NC of

Deviations in

Pathologies

1 5 5

 Yes

 NA

Placement

Construction

Sheet

Concrete

the final

in the Beam

Placement

quality of the

Beam

Concrete

Engineer

Concrete

…

Prepared by:

*C - Compliance;*

**Table 7.**

*PCMM with LCM for beams.*

 *NC - Not compliance;*

 *NA - Not Applicable.*

Date: \_/\_/\_

Verified by:

Date: –/–/–

Responsible:

Date:

–/–/–

 …

inspection

Concrete

Design

Placement

**(1)**

**PCMM - Plan Control,** 

**Measurement**

**Process/Work:**

 **Execution of Beams**

 **and Monitoring**

**Revision:**

**Page:**

 **1/1**

*Structural Integrity and Failure*


### *Risk Assessment in the Monitoring of Works DOI: http://dx.doi.org/10.5772/intechopen.93957*

**Table 7.**

*PCMM with LCM for beams.*

**Logo**

**68**

**Item**

**Activities**

**Inspection/**

**Purposes**

**Acceptance**

**Frequency/**

**Responsible**

**Re**

**Risks**

**Description**

**Effects**

**Risk**

**Controlled**

**Actions**

*Structural Integrity and Failure*

**(9)**

**(10)**

**Assessment**

**Risks?**

**(13)**

**(11)**

**P S RN**

**(12)**

**(6)**

**cords**

**(8)**

**(7)**

**Sampling**

**(5)**

**criteria**

**(4)**

**(3)**

**Testing**

**(2)**

**(1)**

1

 Receipt

Visual

C with

Purchase

By Delivery

Overseer

 Control

NC with

Variation in

Deviation

2 2 4

 Yes

 NA

Sheet

Purchase

Quantities;

in Activity

Order

Deviations in

3

type

Falsework

inspection

Purchase

Oder

By Lot

Order

Falsework

2

 Receipt

Visual and

C Steel

Project; Steel

By Delivery

Overseer

 Control

NC with

Deviation in

Deviations

1 3 3

 Yes

 NA

Sheet

Specification

Classes and ϕ

in Activity

> of steel

4

specification

By Lot

> for concrete

reinforcement

Reinforcement

Metric

Class and ϕ

inspection

3

 Installation

Visual and

C of

Structural

By Beam

 Overseer

 Control

NC of

Instability of

Deviations

3 3 9

 Yes

 NA

Sheet

Verticality,

the beam

in activities

4 and 6

Planimetry

and Stability

Falsework

Metric

Verticality,

Design

inspection

Planimetry

and

Stability

4

 Installation

Visual and

C of the

Structural

By Beam

 Overseer and

Control

NC of

Deformation

Deviations

2 3 6

 Yes

 NA

and

in Activity

Construction

Sheet

mooring and

positioning

instability of

6

the Beam

Engineer

Reinforcement

Metric

mooring

Design

and

positioning

5

 Receipt

Visual

C with

Purchase

By delivery

 Overseer and

Control

NC of

Deviations in

Activities 7

2 3 6

 Yes

 NA

Construction

Sheet

Purchase

Concrete

stop

Class

Order and

NC Class

Deviations in

Consistency

Consistency

Engineer

Concrete

inspection

Purchase

Order

Structural

and

Order and

Consistency

Class

Design

Testing

Consistency

inspection

**(1)**

**PCMM - Plan Control,** 

**Measurement**

**Process/Work:**

 **Execution of Beams**

 **and Monitoring**

**Revision:**

**Page:**

 **1/1**


**Logo**

**71**

**Item**

**Activities**

**Inspection/**

**Purposes**

**Acceptance**

**Frequency/**

**Responsible**

**Re**

**Risks**

**Description**

**Effects**

**Risk**

**Controlled**

**Actions**

**Risks?**

**(13)**

**Assessment**

**(11)**

**P S D RPN**

**(12)**

*Risk Assessment in the Monitoring of Works DOI: http://dx.doi.org/10.5772/intechopen.93957*

**(9)**

**(10)**

**(6)**

**cords**

**(8)**

**(7)**

**Sampling**

**(5)**

**criteria**

**(4)**

**(3)**

**Testing**

**(2)**

**(1)**

6

 Placement

Visual

C of

Structural

By

Overseer and

Control

NC of

Deviations in

Pathologies

 Yes

 NA

Placement

Construction

Sheet

Concrete

the final

in the Beam

Placement

quality of the

Beam

Concrete

Engineer

Concrete

…

Prepared by:

**Table 8.**

*PCMM with FMEA for beams.*

Date: \_/\_/\_

Verified by:

Date: –/–/–

Responsible:

Date:

–/–/–

 …

inspection

Concrete

Design

Placement

**(1)**

**PCMM - Plan Control,** 

**Measurement**

**Process/Work:**

 **Execution of Beams**

 **and Monitoring**

**Revision:**

**Page:**

 **1/1**

### *Structural Integrity and Failure*


### *Risk Assessment in the Monitoring of Works DOI: http://dx.doi.org/10.5772/intechopen.93957*

**Table8.**

 *PCMM with FMEA for beams.*

**Logo**

**70**

**Item**

**Activities**

**Inspection/**

**Purposes**

**Acceptance**

**Frequency/**

**Responsible**

**Re**

**Risks**

**Description**

**Effects**

**Risk**

**Controlled**

**Actions**

*Structural Integrity and Failure*

**Risks?**

**(13)**

**Assessment**

**(11)**

**P S D RPN**

**(12)**

**(9)**

**(10)**

**(6)**

**cords**

**(8)**

**(7)**

**Sampling**

**(5)**

**criteria**

**(4)**

**(3)**

**Testing**

**(2)**

**(1)**

1

 Receipt

Visual

C with

Purchase

By Delivery

Overseer

 Control

NC with

Variation in

Deviation

 Yes

 NA

Sheet

Purchase

Quantities;

in Activity

Order

Deviations in

3

type

Falsework

inspection

Purchase

Oder

By Lot

Order

Falsework

2

 Receipt

Visual and

C Steel

Project; Steel

By Delivery

Overseer

 Control

NC with

Deviation in

Deviations

 Yes

 NA

Sheet

Specification

Classes and ϕ

in Activity

> of steel

4

specification

By Lot

> for concrete

reinforcement

Reinforcement

Metric

Class and ϕ

inspection

3

 Installation

Visual and

C of

Structural

By Beam

 Overseer

 Control

NC of

Instability of

Deviations

 Yes

 NA

Sheet

Verticality,

the beam

in activities

4 and 6

Planimetry

and Stability

Falsework

Metric

Verticality,

Design

inspection

Planimetry

and

Stability

4

 Installation

Visual and

C of the

Structural

By Beam

 Overseer and

Control

NC of

Deformation

Deviations

 Yes

 NA

and

in Activity

Construction

Sheet

mooring and

positioning

instability of

6

the Beam

Engineer

Reinforcement

Metric

mooring

Design

and

positioning

5

 Receipt

Visual

C with

Purchase

By delivery

 Overseer and

Control

NC of

Deviations in

Activities 7

 Yes

 NA

Construction

Sheet

Purchase

Concrete

stop

Class

Order and

NC Class

Deviations in

Consistency

Consistency

Engineer

Concrete

inspection

Purchase

Order

Structural

and

Order and

Consistency

Class

Design

Testing

Consistency

inspection

**(1)**

**PCMM - Plan Control,** 

**Measurement**

**Process/Work:**

 **Execution of Beams**

 **and Monitoring**

**Revision:**

**Page:**

 **1/1** Given the set decision rule in the methodology, this risk is controlled and does not require mitigation action.

It also was concluded that the use of the proposed methodology for of the PCMM

If we are used to risk assessment for identify from the sources, events and causes

can be a solution to prevent defects and reworks, since it allows you to easily identify which trials and inspections will be implement on the control,

and its possible consequences, then we can determine the level of risk and its acceptance or tolerance. Finally, the risk assessment applied to PCMM helps to

With the examples presented, it was possible to conclude that defects and rework of the works do not have to be accepted as inevitable or even as certainties and should be considered as a permanent challenge to the management of work.

measurement and monitoring critical activities of the works.

*Risk Assessment in the Monitoring of Works DOI: http://dx.doi.org/10.5772/intechopen.93957*

**Author details**

**73**

M. Rosário Oliveira†

† ORCID 0000-0001-8149-7351

provided the original work is properly cited.

indicate which actions are necessary to mitigate uncontrolled risks.

ISEP - School of Engineering Polytechnic of Porto, Porto, Portugal

© 2020 The Author(s). Licensee IntechOpen. This chapter is distributed under the terms of the Creative Commons Attribution License (http://creativecommons.org/licenses/ by/3.0), which permits unrestricted use, distribution, and reproduction in any medium,

\*Address all correspondence to: mro@isep.ipp.pt

It is seen too that the planned control for the activity "Receipt Reinforcement" has the lowest risks number (RN = 3), and it is classified as low risk. With the set decision rule defined in the methodology, this risk is also controlled and does not require any mitigation action.

It is seen that the planned control for the activity of the "Installation Falsework" has the highest risk number (RN = 9), and it is classified as medium risk. Given the set decision rule in the methodology, this risk is controlled and does not require mitigation action.

It is observed that the planned control for the activity "Receipt Reinforcement" has the lowest risks number (RN = 3), and it is classified as low risk. With the set decision rule defined in the methodology, this risk is also controlled and does not require any mitigation action.

### **4.2 PCMM with FMEA**

Keeping the form of the PCMM with LCM shown above, the **Table 8** shows the part of the PCMM with FMEA prepared for the execution of reinforced concrete beams of the building's structure.

The methodology set out above was followed in preparing PCMM with FMEA and risk assessment was made under Failure Mode and Effects Analysis (FMEA).

The use of this procedure also allows, on the one hand to plan the inspections and tests to make the critical activities, so as to control, measure and monitor all the work, on the other hand, it helps prevent the risk in the occurrence of monitoring events with a negative impact on the development of work.

It is seen that the planned control for the activity of the "Installation Falsework" continues to have the highest risk number and the highest risk priority number (RPN = 27). Since it is classified as medium risk, according to the set decision rule defined in the methodology, this risk is controlled and does not require any mitigation action.

It is observed that the planned control for the activity "Receipt Reinforcement" has the lowest risk number and the lowest risk priority number (RPN = 3), it being classified as low risk. With the set decision rule defined in the methodology, this risk is also controlled and does not require any mitigation action.

### **5. Conclusion**

It was concluded the methodology presented for the preparation of Control, Measurement and Monitoring Plans (PCMM), can follow the approach on riskbased thinking and can help to assurance the compliance with technical and regulatory of the works. It was concluded too that measurement and monitoring process can promote the risks and failures assessment and can demonstrate the conformity of the works and can too improve operational efficiency.

The PCMM with risk assessment helps to identification the need for risk mitigation actions in the control of the works. Therefore, the use of risk assessment techniques is important in the planning and control of the works processes.

It was observed that Likelihood/Consequence Matrix (LCM) and Failure Mode and Effects Analysis (FMEA) are two techniques of the risk assessment applicable to PCMM, and introducing the probability, consequences and detection indices, they allow risk classification and priority of their mitigation.

### *Risk Assessment in the Monitoring of Works DOI: http://dx.doi.org/10.5772/intechopen.93957*

Given the set decision rule in the methodology, this risk is controlled and does not

It is seen too that the planned control for the activity "Receipt Reinforcement" has the lowest risks number (RN = 3), and it is classified as low risk. With the set decision rule defined in the methodology, this risk is also controlled and does not

It is seen that the planned control for the activity of the "Installation Falsework" has the highest risk number (RN = 9), and it is classified as medium risk. Given the set decision rule in the methodology, this risk is controlled and does not require

It is observed that the planned control for the activity "Receipt Reinforcement" has the lowest risks number (RN = 3), and it is classified as low risk. With the set decision rule defined in the methodology, this risk is also controlled and does not

Keeping the form of the PCMM with LCM shown above, the **Table 8** shows the part of the PCMM with FMEA prepared for the execution of reinforced concrete

The methodology set out above was followed in preparing PCMM with FMEA and risk assessment was made under Failure Mode and Effects Analysis (FMEA). The use of this procedure also allows, on the one hand to plan the inspections and tests to make the critical activities, so as to control, measure and monitor all the work, on the other hand, it helps prevent the risk in the occurrence of monitoring

It is seen that the planned control for the activity of the "Installation Falsework"

It is observed that the planned control for the activity "Receipt Reinforcement" has the lowest risk number and the lowest risk priority number (RPN = 3), it being classified as low risk. With the set decision rule defined in the methodology, this

It was concluded the methodology presented for the preparation of Control, Measurement and Monitoring Plans (PCMM), can follow the approach on riskbased thinking and can help to assurance the compliance with technical and regulatory of the works. It was concluded too that measurement and monitoring process can promote the risks and failures assessment and can demonstrate the conformity

The PCMM with risk assessment helps to identification the need for risk mitiga-

It was observed that Likelihood/Consequence Matrix (LCM) and Failure Mode and Effects Analysis (FMEA) are two techniques of the risk assessment applicable to PCMM, and introducing the probability, consequences and detection indices,

tion actions in the control of the works. Therefore, the use of risk assessment techniques is important in the planning and control of the works processes.

continues to have the highest risk number and the highest risk priority number (RPN = 27). Since it is classified as medium risk, according to the set decision rule defined in the methodology, this risk is controlled and does not require any mitiga-

events with a negative impact on the development of work.

risk is also controlled and does not require any mitigation action.

of the works and can too improve operational efficiency.

they allow risk classification and priority of their mitigation.

require mitigation action.

*Structural Integrity and Failure*

require any mitigation action.

require any mitigation action.

beams of the building's structure.

**4.2 PCMM with FMEA**

tion action.

**5. Conclusion**

**72**

mitigation action.

It also was concluded that the use of the proposed methodology for of the PCMM can be a solution to prevent defects and reworks, since it allows you to easily identify which trials and inspections will be implement on the control, measurement and monitoring critical activities of the works.

If we are used to risk assessment for identify from the sources, events and causes and its possible consequences, then we can determine the level of risk and its acceptance or tolerance. Finally, the risk assessment applied to PCMM helps to indicate which actions are necessary to mitigate uncontrolled risks.

With the examples presented, it was possible to conclude that defects and rework of the works do not have to be accepted as inevitable or even as certainties and should be considered as a permanent challenge to the management of work.

### **Author details**

M. Rosário Oliveira† ISEP - School of Engineering Polytechnic of Porto, Porto, Portugal

\*Address all correspondence to: mro@isep.ipp.pt

† ORCID 0000-0001-8149-7351

© 2020 The Author(s). Licensee IntechOpen. This chapter is distributed under the terms of the Creative Commons Attribution License (http://creativecommons.org/licenses/ by/3.0), which permits unrestricted use, distribution, and reproduction in any medium, provided the original work is properly cited.

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[4] Oliveira M.R. A Gestão da Qualidade na Construção e a Gestão do Risco. ISEP Moodle, Porto, 2016.

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[6] DNP ISO Guia 73:2011 - Gestão do risco - Vocabulário (ISO Guide 73: 2009). Instituto Português da Qualidade (IPQ), Caparica, 2011.

[7] NP EN 31010:2016 - Gestão do risco - Técnicas de apreciação do risco (ISO/ IEC 31010:2009). Instituto Português da Qualidade (IPQ), Caparica, 2016.

[8] Nuchpho P, Nansaarng S, Pongpullponsak A. Risk Assessment in the Organization by Using FMEA Innovation: A Literature Review. King Mongkut's University Technology Thonburi, Thailand, 2014.

[9] M. Abdelgawad M, Fayek A.R. Risk Management in the Construction Industry Using Combined Fuzzi FMEA and Fuzzy AHP. Journal of Construction Engineering and Management, vol. 136, n° 9, pp. 1028-1036, 2010. DOI: 10.1061/(ASCE)CO.1943-7862. 0000210.

[10] Wang H, Deng X, Zhang Z, Jiang W. A New Failure Mode and Effects Analysis Method. IEEE Acess ISSN 2169-3536*,* vol. 7, p. 12, 2019. DOI: 10.1109/ACCESS.2019.2923064.

[11] Silva S.R.C, Fonseca M, Brito J. Metodologia FMEA e sua Aplicação à Construção de Edificios. LNEC, QIC2006, Lisboa, 2006.

[12] Huang J, You J-X, Liu H-C, Song M-S. Failure mode and effect analysis improvement: A systematic literature. Reliability Engineering and System Safety ISSN 0951-8320*,* vol. 199, p. 12, 2020. DOI: 10.1016/j. ress.2020.106885.

Section 2

Bridge Engineering

**75**

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