Pipeline Failure Analysis and Prevention

#### **Chapter 1**

## Failure Analysis of Pipelines in the Oil and Gas Industry

*Mohamed Mohamed Azzam*

#### **Abstract**

The term "failure" can be defined as the inability of a part or assembly to perform its intended function. Despite the significant technological advances, failure incidents frequently occur, thus, causing human and financial consequences. The failure analysis is a crucial engineering tool. It aims to avoid similar cases in the future, thereby preventing accidents, reducing economic losses due to stopping plant production and keeping the environment safe. Furthermore, the failure analysis contributes to redesign, solve manufacturing drawbacks, save money and time, and in some cases, prevents fatality and saves lives. Conversely, failures can also improve engineering practices; indeed, through analyzing failures and implementing preventive measures, significant advances have been obtained in the quality of products and systems. Moreover, a beneficial outcome of failure analysis has been improved codes and specifications governing materials, for instance, API, ASTM, and ASME. In the current chapter, the failure analysis methodology will be discussed in detail with practical examples to know how to perform analysis for any failure cases, particularly in the oil and gas industry.

**Keywords:** failure analysis, pipeline, corrosion, fracture, oil & gas industry

#### **1. Introduction**

The term "failure" can be defined as the inability of a part or assembly to perform its intended function [1]. Based on the simple definition of failure, we can understand that the part of the component is considered failed if it cannot perform its function perfectly for any reason, for instance, a change in dimensions, corrosion, fracture, and so on. Sometimes unspecialized think the part to fail must be broken or fractured, but this is not the case. In other words, each fracture is considered a failure; however, not every failure is considered a fracture. A fracture separates parts into two or more species in response to the applied or residual stresses.

The pipeline can expose to thinning due to erosion-corrosion damage; however, it is still in service; thus, the pipeline can be considered to have failed, although it is still in service without leakage. The thinning mechanism is a failure since the pipeline has lost service life. In other words, the pipeline was designed to serve a specific period; however, the thinning damage has shortened its lifetime, which means the pipeline lost some of its lifetime. The high-pressure gas (HPG) pipeline has been subjected to internal corrosion, which led to a localized metal loss. Thus, the HPG pipeline

has been converted to transfer oil or low-pressure gas due to the fitness for service, which revealed the remaining thickness cannot withstand the high pressure; thus, the pipeline has lost its function to transfer the high-pressure gas.

Material failure can be divided into four types: distortion or plastic deformation, fracture, corrosion, and wear [1]. It is worth noting that two or more physical failures can occur in the same failed part. The root causes of the failure can be divided into three levels; physical roots, human roots, and latent roots. The physical roots can be divided into four categories: design deficiencies, material defects, manufacturing or installation defects, and service life anomalies [2]. The human roots include inadequate inspection and improper equipment installed. The latent roots are the cultural or organizational rules that lead to the human cause; it is not direct roots. The inadequate inspector training is an example of the latent root, where some companies consider the training courses as additional costs that need to be reduced. These companies fail to recognize that this reduction is reflected in the ability of individuals, leading to catastrophic events due to incompetent persons.

The poor design can play a significant role in some failure cases; for example, the pipeline can be designed with a low spot that accumulates the water and causes corrosion. Also, poor design can create a crevice location which accelerates corrosion in the form of crevice corrosion due to the different concentrations of oxygen inside and outside the crevice. Additionally, material selection can be the root cause of some failure cases, using inappropriate material to serve in a harsh environment.

For instance, of inefficient material selection, using 304 austenitic stainless steel in chloride containing environment can lead to severe pitting corrosion or stress corrosion cracking [3]. Moreover, the manufacturing defects in most failure cases play a significant role in the failure. Thus, these defects act as the origin of catastrophic damage. For example, the lack of fusion of manufacturing defects can be a location for crack initiation or a stress concentration.

Furthermore, environmental change can cause premature or unexpected failure before its lifetime. For example, increasing the fluid velocity inside the pipeline accelerates the corrosion rate through erosion damage. Also, if the pipeline is designed according to a specific value (i.e., max limit) of hydrogen sulfide (H2S) and carbon dioxide (CO2), the increase of this limit would lead to anticipated failure mechanisms such as Sulphide stress corrosion cracking and CO2 corrosion, especially in high-pressure gas (HPG) pipelines. In addition, the change in operating conditions may play a crucial role in the failure of pipelines. For example, increasing the operating pressure beyond the design pressure can lead to overloading damage or corrosion by increasing the partial pressure of the corrosive species such as Oxygen (O2), hydrogen sulfide (H2S), and carbon dioxide (CO2) [4]. Also, the inadequate inspection may lead to failure; for example, performing the visual examination without nondestructive examination (NDE) may skip and ignore fine cracks or internal defects that the visual inspection disables to detect. Moreover, the absence of monitoring has a crucial role in some corrosion cases, where different monitoring methods are used to monitor the corrosion rate in the pipeline, such as the corrosion coupon, sand probe, and bio-probe. Without the monitoring method, assessing the operating condition, especially the fluid corrosivity, is not feasible.

Moreover, human error can play a critical role in failure cases through incompetent persons. For example, after conducting the hydro test for repaired tanks or vessels, and during the drainage of the used water, the responsible person can cause collapse due to rapid drain rate or due to the closing of the vent. Also, the non-drain of the water (i.e., Missing to drain) used in the hydro test can cause catastrophic failure, mainly if the pipeline is not used directly after the hydro test and is left for some time.

#### *Failure Analysis of Pipelines in the Oil and Gas Industry DOI: http://dx.doi.org/10.5772/intechopen.108140*

The hydro test water inside the pipeline with stagnation condition will be suitable for SRB colonies to grow and cause corrosion damage, called microbiological induced corrosion (MIC) [5]. Therefore, the water used in the hydro test must be flushed and entirely drained if the pipeline will be used directly after the hydro test. However, suppose the pipeline will not be used directly after the hydro test; in that case, the pipeline must be mothballed through injection of multifunction chemical which contains a mixture of oxygen scavenger, corrosion inhibitor, and biocide.

Indeed, failure analysis aims to determine the causes or factors that have led to an undesired loss of functionality; this considers the failure analysis's direct benefit. Also, failure analysis is an engineering tool for enhancing product quality and failure prevention; this considers the failure analysis's indirect benefit [2]. Furthermore, the failure analysis contributes to redesign, to solve the drawbacks of manufacture, saving money, saving time, and in some cases preventing fatality and saving lives [6].

The failure analysis must be performed based on a scientific base and a standard methodology to identify the damage mechanism and determine the significant root causes; otherwise, the analysis outcome will be unexpressed about the actual root causes. Also, the wrong root causes may accuse factors that are far from playing any role in the failure, thus, taking unsuitable recommendations which can accelerate the failure. The false root causes due to poor and inefficient failure analysis methodology can be likened to accusing an innocent person of murder; however, the real killer is free.

From my point of view, failure analysis in the oil and gas industry, specifically in the offshore environment, is considered a crucial case for the failure analysis society. Since these cases enrich the knowledge and information, thus, saving the marine environment against the consequences of failure and causing disasters such as pollution and marine life death in addition to economic loss. In April 2010, in one of the most significant pipelines containment failure accidents, that is, Deepwater Horizon, it was reported that approximately 3.19 million barrels of oil spilled into the ocean [7] and polluted at least 11,200 km2 of seawater [8], which was catastrophic to both the economy and environment.

The crude oil at offshore platforms is transferred to the processing plants onshore through pipelines, which are considered one of the safest and most effective ways to transport oil and gas, security and reliability of the transmission pipeline [9]. There are three types of pipelines: gathering lines, transmission lines, and distribution lines. Gas or crude oil gathering lines exist between a well and a treatment plant or collection point [2]. The offshore pipelines are much more critical due to their operational condition, inspection, repair difficulties, and environmental issue [10]. These pipelines are manufactured from carbon steel. The consequences of pipeline rupture could lead to loss of life, injury, fire, explosion, environmental pollution, economic loss, decreasing capacity, and increasing maintenance difficulty [11].

#### **2. Failure analysis methodology**

The standard methodology for the failure analysis will be discussed in detail for each step. Practical cases will support the steps. The following steps are common for most failure cases [2]:

1.Collection of background data

2.Preliminary examination of the failed part


It is worth pointing out that the above methodology steps can be applied to most components in the oil and gas industry, such as storage tanks, pressure vessels, and pipelines. In other words, the above steps can be considered a generic methodology for most facilities in the oil and gas industry. However, in the further discussion for each step, most examples of the practical cases of failure incidents will be confined primarily to the pipelines, whether liquid or gas services, since the book's topic is mainly concerned with the pipelines. Therefore, it will be helpful to give examples of pipelines specifically.

#### **2.1 Collection of background data**

The failure analysis methodology's first step is collecting the background data. The failure investigation should include gaining an acquaintance with all pertinent details relating to the failure. In this step, the failure analyst acts as the detective who investigates a crime case by collecting all available data since the useless information, according to the operators of the failed part, is considered very important to the failure analyst and can contribute to solving the case. The failure analyst only who can judge the importance of the data.

Preparing a checklist containing all the essential required data and questions you need to ask is recommended to do not to forget anything. The list can include: the drawing or the as-built, history of the anomalies for the failed portion, repair history, history of the service environment (i.e., oil/water/gas/multiphase), pigging schedule, type of pig, whether BIDI or foam pig, water analysis reports, gas analysis reports, and scale analysis reports, in addition to the interview with the persons who are responsible for the operating of the failed portion.

It is worth noting that the failure analyst must be decent and non-offensive during the interview with the persons responsible for the failed part in the field and not accuse or blame anyone; otherwise, most of these persons vanish the important data fearing the blame or the accusation.

#### **2.2 Preliminary examination**

The preliminary examination can be divided into two stages; the first is performing a site visit to the incident location before retrieving the failed portion. The site visit to the incident location aims to figure out the actual situation of the whole location (i.e., the platform or the plant), not only the failed portion. In some cases, the preliminary site visit to the failure location of the failed part was the clue. Sometimes, the site visit to the failure location revealed simultaneous works implemented at the moment of the incident; consequently, the root causes are highly believed to be external causes due to these operational activities, not the failed part's environment. This indicates that the analyst must enlarge the investigation area around the failed part. Gradually narrow it, especially when the damage is external, such as a dent, gouge, and scratch.

The second stage in the preliminary examination is a visual examination of the failed part that depicts the actual condition after the incident without any change. The visual inspection aims to detect abnormal features such as damage morphology, plastic deformation, cracks, thinning, erosion, dents, dimensions change, and scratches. During the visual examination, a magnifying glass can be used [12] to enlarge some significant features of the failed part, such as the origin of the crack, damage morphology, and crack arrest location. Using a ruler or measure tap is recommended to determine the aspects of the failure, mainly in the cracking or rupture incident before movement or cutting the failed part, as shown in **Figure 1**, which

**Figure 1.** *Measuring the dimensions of the rupture of 24-inch water injection.*

**Figure 2.** *Failed 2-inch bleeding valve due to drop of a mechanical tool.*

shows the crack's measurement process dimensions for rupture of 24-inch water injection pipeline.

**Figure 2** shows a failed 2-inch bleeding valve disconnected from the main line of high-pressure gas (i.e., 1200 psi), and the short nipple connection between the valve and the main line was found in a bent shape. On-site inspection indicated scaffolding had been installed around the failure due to concurrent work to replace an 8-inch water injection line directly above the failed valve. This means that the main reason for a defective valve to bend is that a mechanical tool fell and hit the valve, whereas the failed valve is a free end with no support. This failure case is an ideal example of the importance of on-site inspections at the fault location. Without on-site inspections, investigators will come up with unclear root causes and thus make incorrect recommendations.

The visual examination of the failed part must include determining the origin from which the crack or the corrosion started. The origin can be manifested as a bulge, such as the overloading failures, or stress concentration location, such as the sharp edges and weldment. For example, **Figure 3** shows the origin of the rupture, which occurred in an 18-inch gas pipeline, where the origin appeared in a bulge shape; the spout-like form (bulge) is indicative of an overloading incident. Eventually, during a visual inspection, failure analysts can use failure morphologies to brainstorm and imagine expected failure cause scenarios.

*Failure Analysis of Pipelines in the Oil and Gas Industry DOI: http://dx.doi.org/10.5772/intechopen.108140*

**Figure 3.** *Illustrates the origin (bulge) of fast-running ductile rupture of 18-inch oil pipeline.*

#### **2.3 Chemical analysis (sludge: water-liquid)**

It is a crucial step to obtain a sample from the failed portion directly after the incident. The collected sample can be sludge or liquid. The sludge sample will be analyzed to determine the predominant components, especially in the corrosion failure cases. The scale analysis can provide us with informative data about the damage mechanism. Additionally, samples from debris and water are obtained to perform sulfate-reducing bacteria (SRB) testing to determine the count of sessile and planktonic bacteria in the fluid.

If the predominant compound in the analysis is iron carbonate, the expected damage mechanism is CO2 corrosion. If the iron sulfide (FeS) is the dominant compound, the predicted damage mechanism is microbiologically induced corrosion (MIC) due to the microorganism activity of the sulfate-reducing bacteria (SRB) [13]. If the predominant component is the sand, the expected damage mechanism is erosion damage due to the abrasive particles of the sand. If the main compound is iron oxide, the anticipated damage mechanism will be Oxygen corrosion.

The complete water analysis, iron content, H2S dissolved, Oxygen dissolved, and CO2 dissolved are significant parameters in most failure cases, indicating how the environment can be harsh. Chlorides have a detrimental effect on the passive layer, destroying the protection formed by the corrosion inhibitor. Oxygen is one of the corrosive species in any corrosion reactions, which accelerates the corrosion rate; therefore, it is recommended to control the Oxygen dissolved to levels of 10 to 50 ppb (part per billion). With increasing the partial pressure of CO2, the dissolved concentration in the water increase, thus, lowering the pH and increasing the corrosion rate [14]. **Figure 4** illustrates the scale sample collection from a failed 24-inch oil pipeline.

#### **2.4 Nondestructive testing**

Nondestructive testing is a helpful and essential tool in most failure cases. Ultrasonic testing (UT) is often used to measure the remaining wall thickness to calculate the maximum pressure that can be applied. Also, UT is used to detect internal defects, whether base metal or weldment, like porosity, cracks, and laminations.

**Figure 4.** *Collection of scale sample from failed 24-inch oil pipeline.*

**Figure 5.**

*UT Technique (a) calibration of the UT device, and (b) Conducting of UT examination.*

Additionally, UT can give the location and size of the defects [12] and measure the remaining wall thickness, as shown in **Figure 5**.

The dye penetrant test (PT) is often used to detect the surface cracks and clearly show the damage's extent, as shown in **Figure 6** shows a fine crack in the fillet weld of the 2-inch line of the High-pressure gas. The magnetic particles test (MT) also performs the same function as the PT, detecting surface cracks.

#### **2.5 Mechanical testing**

It is a crucial step among the steps of the analysis to confirm the desired mechanical properties of the failed part according to the specification to facilitate the subsequent steps, especially the stress analysis step. Thus, the tensile test determines the yield strength, ultimate tensile strength, and elongation. Additionally, hardness testing is the simplest of mechanical tests; it can be used to assist in evaluating heat treatment [6]. Furthermore, the impact test is used to measure the toughness of the failed parts

*Failure Analysis of Pipelines in the Oil and Gas Industry DOI: http://dx.doi.org/10.5772/intechopen.108140*

**Figure 6.** *Crack detected in the fillet weld of 2-Inch HPG line by dye penetrant test (PT).*

#### **Figure 7.**

*Performing impact test for specimen cut from an 18-inch offshore pipeline.*

in the overloading cases and fracture when the notch or point of stress concentration is experienced in the failure case [15]. The impact test is Mainly used when the failed part operates at low temperature or a mechanical tool hit or falls on it, as shown in **Figure 7**. It can be concluded that the mechanical tests are considered a verification that the material of the failed part was convenient to the applied stresses or the environment during the in-service period or not, according to the specification and the design criteria.

#### **2.6 Chemical analysis of the material**

In failure analysis cases, routine chemical composition analysis is highly recommended [6]. The chemical analysis of the material is used to determine the failed part's chemical composition, the same as the mechanical testing. The alloying elements in the cast alloys are rarely distributed uniformly. Thus, the chemical analysis is considered a verification tool by ensuring that the chemical composition does not deviate from the nominal composition according to specification. The chemical analysis is conducted in the laboratory using the optical emission spectrometer. The deviation from the nominal composition at a specific location of the failed part is called segregation [2]. **Table 1** illustrates the deviation of the chemical composition of 2205 duplex stainless steel pipe using X-Ray Fluorescence (i.e., XRF) device as shown in **Figure 8**, which caused a premature failure and gas release [16].

In some cases, it is difficult to perform the chemical analysis in the laboratory; therefore, positive material identification (PMI) is a prompt tool that can be used in the field to identify the material with the chemical compositions of the alloying element. The PMI is a portable device used to determine the chemical composition of the material without needing to transfer the sample to the laboratory. For example, **Figure 9** shows embrittlement in the flare of low-pressure gas, where the burner's designed material is 310 stainless steel; however, the PMI analysis revealed that the material is 384 stainless steel. This case shows the importance of the PMI and how it can facilitate determining the clue in the field without needing the laboratory.

#### **2.7 Selection, sectioning, preservation and/or cleaning of specimens**

In my view, the steps of selection, cleaning, and preservation are very critical and crucial since any fault can destroy the fracture surface, thus, making the failure analysis process difficult, and the root causes may not be present in the actual failure. The significant and valuable portion of the failed part is the origin, whether it is cracking or corrosion damage. The origin (i.e., the start point of the failure) is the clue of most failure cases since it would contain a defect, and the damage starts. **Figure 10** shows the locations to be cut for examination (i.e., mechanical testing, chemical testing, macro examination, and micro examination). The proficiency of the failure analyst is shown in the selection of what specific portions are to be studied. Since the whole failed part is not used in the analysis, small selected specimens, such as the origin, are expressed and valuable.

Furthermore, the sectioning step of the selected portion must be performed far enough from the failure's origin to prevent destroying it, which could lead to false


#### **Table 1.** *Chemical composition of 2205 Duplex stainless steel pipe.*

*Failure Analysis of Pipelines in the Oil and Gas Industry DOI: http://dx.doi.org/10.5772/intechopen.108140*

#### **Figure 8.** *Performing chemical analysis for the weldment using XRF.*

#### **Figure 9.** *Embrittlement of flare due to unsuitable material.*

conclusions [2]. During the sectioning process, it must be considered the sample size to be suitable for macro and micro examination and optical and scanning electron microscopes, respectively. It is recommended to use cold cutting techniques for

**Figure 10.** *Selection of location to be cut and studied.*

**Figure 11.** *Cutting machine with cooling system.*

sectioning the samples, like the wire cut and the abrasive blade cutting, which do not introduce any heat to the failed portion, thus, preventing any alteration to the actual condition, as shown in **Figure 11**.

Moreover, handling the selected portion of the failed part is a significant step. Therefore, the selected samples are stored in special boxes made from plastic to avoid friction between the specimens and the container. It is recommended to coat the pieces with grease or apply a removable coating of oil or plastic compound to prevent further interaction between the cut samples and the surrounding environment [12]. The surrounding environment causes corrosion and oxidation to the specimens and confuses whether the source of this corrosion is due to an in-service environment or not.

The cleaning process of the specimens is a crucial step since improper cleaning can destroy the fracture surface. The cleaning process aims to remove corrosion products, debris, and grease from the fracture surface. The cleaning process is an inevitable step

*Failure Analysis of Pipelines in the Oil and Gas Industry DOI: http://dx.doi.org/10.5772/intechopen.108140*

**Figure 12.** *Ultrasonic cleaning for specimen before examination by SEM.*

in the microscopic examination, particularly when the scanning electron microscope is used. The cleaning program must be started with soft tools and gradually increase to aggressive cleaning tools based on the surface condition. Many cleaning methods can be used in the preparation of the samples. For instance, using the soft hair artist's brush as a preliminary step, and then using one or more methods from the following: using inorganic solvents, either by immersion or by jet, acetone or alcohol, cellulose acetate tape, replica, and use the ultrasonic cleaning bath. The ultrasonic bath is very useful in accelerating the cleaning process, as shown in **Figure 12**.

#### **2.8 Macroscopic examination and analysis**

The macro examination step is performed after the excellent cleaning of the sample in which the fracture surface becomes clear and free of corrosion products, dirt, grease, and residual species. The stereoscope is the most helpful tool in this step. The stereoscopic viewing has the same scope as visual inspection but is more detailed as typical stereoscopes allow 10X to 70X magnification [17]. The macroscopic examination does not require the fracture surface to be extremely smooth. The cleaning only is sufficient to perform a macroscopic analysis, unlike the microscopic examination (i.e., Optical microscope), which needs a polished surface to produce high contrast between the microstructural constituents.

The macroscopic examination can provide the failure analyst with beneficial information, for instance, the origin of the fatigue crack and secondary cracks, especially the fine cracks that could not be observed during the visual examination. Furthermore, the macroscopic shows a comprehensive view of the failure location compared to that of microscopic examination, thus, helping the failure analyst to imagine the scenario of what happened during the incident. **Figure 13** shows a cross-section macrograph of the failed pipe at the seam weld side. The joint configuration of the seam longitudinal weld is a double-V groove weld. A fusion welding process produces the seam weld. It is most likely a SAW process was used due to the considerable penetration depth in one pass (the filling pass). It is also evident that the crack propagated along the HAZ, HAZ/base metal boundary, or the base metal.

**Figure 13.** *Macrographs across the fracture surface at about 40 cm from the burst origin.*

#### **2.9 Microscopic examination and analysis**

The microscopic examination is the clue in most failure cases, especially for metallurgical investigation, and is typically performed by the scanning electron microscope (SEM), as shown in **Figure 14**. The scanning electron microscope is an effective and helpful tool to know what happened during the incident; it looks like a recorded camera of the events of the failure. The scan electron microscope can provide a large magnification ranging from 5000 to 10,000X [6].

Some fracture modes can be identified according to the microscopic characteristics, where the dimpled morphology indicates ductile fracture due to overloading, and cleavage facets refer to brittle fracture. The striations are the most characteristic microscopic evidence of fatigue fracture. **Figure 15** shows cleavage facets at the origin

**Figure 14.** *Scanning electron microscope (SEM).*

*Failure Analysis of Pipelines in the Oil and Gas Industry DOI: http://dx.doi.org/10.5772/intechopen.108140*

**Figure 15.** *Shows cleavage facets.*

**Figure 16.** *Shows quasi cleavage.*

of an 18-inch gas pipeline crack ruptured due to a welding defect. **Figure 16** illustrates the quasi cleavage of a fracture surface of an 18-inch oil pipeline, which ruptures in the form of rapid ductile fracture.

#### **2.10 Selection and preparation of metallographic sections**

The metallographic selection and preparation are crucial steps in metallurgical investigation cases. The selected region of the failed part must be chosen to present unique features of the failed part, which are selected for the characterization process. The selection of the samples which would be examined must be carried out carefully since these samples shall be near the edge of the fracture (i.e., around the origin). In other words, the significant sample which provides valuable information is the sample of the failure location. The sectioning of the metallographic specimens should be perpendicular to the fracture surface in edge view. Furthermore, selecting pieces far from the failure region (i.e., undamaged location) is helpful.

The selected specimens for the metallographic examination should be cut with cold manners. The cold cut prevents alteration of the specimens' microstructure, high precision cut, and deformation-free cutting for various workpiece sizes. Thus, the wire cut machine or the abrasive cut-off wheels are the standard practical methods for cutting the specimens examined in metallography. The samples are cut into small sizes to facilitate the handling and examination processes; these samples are ground using silicon carbide (SiC) foil and paper to produce a smooth surface before polishing.

#### **2.11 Examination and analysis of metallographic specimens**

Metallography is defined as the scientific discipline of examining and determining the constitution and the underlying structure of (or spatial relationships between) the constituents in metals, alloys, and materials (sometimes called materialography) [18]. The most common tool in the metallography examination is the optical or light microscope, with magnifications ranging from ~50 to 1000×, as shown in **Figure 17**. The optical microscope is used to identify the phases, constituents, and precipitations and determine the size and shape of the grains. The high contrast between microstructural constituents in light microscopy mainly depends on the quality of the specimen preparation process. The metallographic examination is also a verification of the heat treatment where the grains sizes are an indication of the heat treatment quality.

In **Figure 18**, it is evident that the HAZ at both sides of the joint contains coarse ferrite grains. In addition to the coarse ferrite grains, grain boundary austenite (GBA) has been observed along the grain boundaries of the coarse ferrite. It is believed that the excessive heat input of the welding process resulted in coarse-grained HAZ, which played a role in the degradation of HAZ zones.

#### **2.12 Analysis of fracture mechanics (stress analysis)**

It is sometimes quite apparent that excessive loading plays a detrimental role in the failure. Noticeable plastic deformation is observed at the failed pipeline due to

**Figure 17.** *The optical microscope for metallography.*

*Failure Analysis of Pipelines in the Oil and Gas Industry DOI: http://dx.doi.org/10.5772/intechopen.108140*

**Figure 18.** *Coarse grain HAZ of 2205 duplex stainless steel pipe.*

the overloading, which causes a change in the pipeline configuration from circle to oval shape. Additionally, the stress analysis is performed based on the specifications, standards, and codes, for instance, ASME B31.8, API 579, and ASME 31G.

For the components with very complex shapes and high thermal gradients, a finiteelement analysis (FEA) may be performed to estimate the most likely stress level in the failed part. These analyses can stand alone or can be used to help select critical locations for strain gauge attachment. Finite-element calculations can be time-consuming and expensive, but they are necessary for accurately assessing stress levels in areas

**Figure 19.** *Shows the topography of the ruptured wall and the wall bow-out.*

of the complex geometry of some components. This analysis is almost essential for determining stresses caused by thermal gradients such as those found in welding [6].

**Figure 19** shows an offshore pipeline ruptured due to overloading, where the outer surface of the ruptured wall is bowing out, indicating that the pipeline was highly inflated before it bursts. After failure, the surface is still bowing out, meaning that the deformation of the pipe wall was plastic deformation, and the level of stress encountered was very high.

#### **2.13 Determination of failure mechanism**

Determining the failure mechanism is almost the last step in the failure analysis methodology. The failure mechanism is determined based on the main finding of the visual examination, chemical analysis, mechanical testing, macro examination, and micro examination. It is believed that identifying the failure mechanism is easier and faster than determining the root causes of this failure damage. Since in most failure cases, the visual examination is preliminary and sufficient to figure out the failure damage.

**Figure 20.** *Erosion at the 4-inch control valve of discharge line.*

**Figure 21.** *Microbiological induced corrosion (MIC) at 30-inch discharge header.*

#### *Failure Analysis of Pipelines in the Oil and Gas Industry DOI: http://dx.doi.org/10.5772/intechopen.108140*

**Figure 20** shows a failed 4-inch control valve, and from the visual examination, it is clear that the expected damage is erosion-corrosion damage. **Figure 21** illustrates scattered pitting at the 30-inch oil discharge header. Based on the damage morphology and operating condition, it is believed that the microbiological induced corrosion (i.e., MIC) is the anticipated damage mechanism due to bacterial activity (i.e., SRB) and according to the cup-shaped damage.

**Figure 22** shows a rupture of the 24-inch water injection pipe, where the observed bumps indicate an overload event; visual inspection shows severe thinning due to corrosion damage; therefore, with a working pressure of 1600 psi, the anticipated failure mechanism is Stress corrosion cracking. The corrosion cracking mechanism combines the stress and the corrosive environment [5]. Also, **Figure 23** shows a fracture in the 6-inch well flowline, and site visits to the platform show that the tubing is undergoing significant movement; therefore, the damage mechanism is believed to be fatigue damage due to cyclic loading.

#### **Figure 22.**

*Stress corrosion cracking (SCC) at 24-inch water injection pipeline.*

**Figure 23.** *Crack due to fatigue at 6-inch flow line of oil well.*

#### **2.14 Determination testing under simulated service conditions (special tests)**

In some cases, simulation of operating conditions or special testing is strongly recommended to understand the effect of the environment on the same material of the failed component. It is important to note that most simulation tests are not feasible or practical because service conditions cannot be achieved, especially in the case of corrosion failure, which is not feasible in the laboratory. In addition, running simulations requires safety considerations, especially in the event of failures that occur under catastrophic conditions, such as overloading high-pressure gas lines. **Figure 24** shows pitting and crevice corrosion testing of an undamaged specimen cut from a defective 2205 duplex stainless steel weldment. This test is designed to determine the effect of the chloride (i.e., high salinity) on welded joints operating in harsh high salinity

**Figure 24.** *Pitting and crevice corrosion testing according to STM G-48.*

**Figure 25.** *The specimen after conducting the pitting and crevice corrosion testing.*

environments. The results show that preferential corrosion of the specimen occurs in the heat-affected zone and near the weld, as shown in **Figure 25**.

#### **2.15 Determination analysis of all the evidence, formulation of conclusions, and writing the report**

This step is the outcome of all stages in the failure analysis methodology. The main findings depict every step and are discussed and analyzed intellectually and scientifically to prove the damage mechanism and support the believed root causes. Also, all evidence is collected together during the investigation steps to complete the correct scenario about what has occurred.

The analyst's competence is also reflected in the preparation of the report, especially in the discussion and conclusion. Suppose the inspector has access to extensive laboratory facilities. In that case, best efforts should be made to analyze and discuss the results of mechanical testing, chemical analysis, fracture, and microscopy before formulating any preliminary conclusions [6]. Eventually, in investigations where the cause of failure is particularly elusive, searching for reports of similar cases may help identify possible root causes. Some references to other similar issues are suggested to support the discussion during the discussion.

#### **3. Conclusion**

The failure analysis methodology is considered the guide to perfect failure analysis in the oil and gas industry. Failure analysis in the oil and gas industry, specifically in the offshore environment, is a significant case for the failure analysis society. Since these cases enrich the knowledge and information, thus, saving the marine environment against the consequences of failure and causing disasters such as pollution and marine life death in addition to economic loss. It is helpful to publish the failure analysis cases and share them with others to enrich the knowledge and experience relevant to the society of failure analysis. The steps of the failure analysis methodology can be summarized as the following:


### **Author details**

Mohamed Mohamed Azzam Department of Metallurgical Engineering, Faculty of Engineering, Cairo University, Giza, Egypt

\*Address all correspondence to: eng\_azzam60@yahoo.com

© 2022 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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[7] United States of America v. BP Exploration & Production, Inc. et al. Findings of fact and conclusions of law: Phase two trial. In: RE: Oil Spill by the Oil Rig "Deepwater Horizon" in the Gulf of Mexico, on April 20, 2010. U.S. District Court for the Eastern District of Louisiana; 2015

[8] DWH-NRDA, Deepwater Horizon Natural Resource Damage Assessment (NRDA) – DRAFT. 2015 – section 4, injury to natural resources, DWH-NRDA. 2015. p. 685

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[10] Cheng YF. Stress Corrosion Cracking of Pipelines. Hoboken, New Jersey: John Wiley & Sons; 2013

[11] HSE. Guidelines for Pipeline Operators on Pipeline Anchor Hazards. Aberdeen: Health and Safety Executive; 2009

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[14] Palacios CA. Corrosion and Asset integrity Management for Upstream Installations in the Oil/Gas industry, CreateSpace Scotts Valley, California, 1st ed. 2016

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[16] Azzam M, Khalifa W. Investigation of duplex stainless steel flow line failure, engineering failure analysis (under review). 2022

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[18] ASM International. ASM Handbook. Metallography and Microstructures. Vol. 9. ASM International: Materials Park, OH; 2004

#### **Chapter 2**

## Prevention of Pipes Bursting by Using a Novel Deicing Technology

*Milad Rezvani Rad and Andre McDonald*

#### **Abstract**

Freezing of water inside above-ground steel pipes is an unwelcome phenomenon that leads to internal pressurization, bulging, and bursting of pipes and can cause noticeable financial losses, environmental pollution due to the resulted leakages, and damage to the property, equipment, and workers in the field. Thus, a practical deicing/antiicing system that is highly efficient must be developed to minimize the detrimental impacts caused by freezing of liquids inside pipes. First, numerous tests were carried out in a relatively large cold room in which the actual working conditions of bare pipes exposed to cold weather were simulated to comprehend the freezing mechanism of the pressurized water. In the second phase of the project, the performance of the novel heating system was assessed by conducting deicing tests in the cold room. It was concluded that the freezing of the enclosed water was heavily dependent on the pressurization extent of water that itself was a function of pipe size and material properties. It was also found that the novel heating system that was produced by using thermal spraying means was able to eliminate the ice that was formed inside the pipe even under harsh conditions that may not be experienced in the field.

**Keywords:** deicing, functional thermal-sprayed coatings, joule heating, materials characterization, pipe bursting

#### **1. Introduction**

#### **1.1 Steel pipe bursting**

Ice formation inside steel pipes is a widespread incident in both residential and industrial sectors at locations where the ambient air temperature is below the freezing temperature of the water. When pipes are exposed to cold air for a prolonged duration, the ice that was formed initially at the inner surface of the pipe gradually accumulates and grows inward. In extreme cases, this causes a partial blockage inside a pipe or a full blockage at the thinner sections near connections and fittings. Once a portion of water has converted to ice, this causes an expansion in the volume that it occupies and when there is no further room in the pipe that can accommodate this volume growth, the pressure of the entrapped water starts building up. It is wellestablished that the pipe bursting occurs at a location where pressurized unfrozen

water exists [1]. This incident is not only limited to the steel pipes, but it can also occur for plastic piping systems [2].

It is well-known that the negative consequences of bursting pipes are both costly and perilous. Furthermore, it can cause noticeable nonproductive downtime in industrial setups that results in loss of production. The financial losses because of this undesirable phenomenon are of great importance. According to the examination carried out by Insurance Information Institute [3], an average loss of \$1.2 billion annually resulted from freezing issues only in the United States during the past 20 years and freezing of water inside pipes has always been one of the major problems.

In addition to the financial losses, the tremendous amount of energy that is stored in pipes is released after bursting of pipes. These explosive ruptures accompanied by a possible projection of chunks of steel from the damaged portions of pipes and pressure equipment can pose a threat to the well-being of the workers. Considering all the negative impacts of this undesirable phenomenon, it stands to reason that developing a novel and functional deicing system that can minimize or eliminate accumulation of ice inside pipes and pressure equipment is of great importance.

The common practice to minimize formation of ice is the installation of insulation around pipes. However, this cost-effective approach can only be practical in reducing the heat loss rate when pipes are exposed to cold ambient air for a short period of time [4]. That said, an efficient deicing system should accompany the thermal insulation to bring about satisfactory results when pipes are located in frigid environments for a prolonged period.

#### **1.2 Usage of coating heating systems**

It is well known that electric heat tracing is a practical method for deicing or temperature control of pipes that are used for water distribution, especially when waste heat and process steam are not obtainable [5]. In addition, heat tracers are employed where decent flow of high-viscosity fluids is needed. Another benefit of using electric heat tracing compared to steam-based tracing is its cleanliness [6]. Furthermore, they offer accurate temperature control and environmental and economic advantages in comparison to steam tracing [7].

One of the great features of electric heat tracing is the ability to control the amount of power that is supplied to the tracers. This can simply be done by input voltage adjustment. Thus, the heating or deicing operation can be adjusted at any given moment according to requirements in the field or the environmental conditions such as wind gust and air temperature. The system can become adaptable to climatic conditions and be operated automatically by using a controller [8]. The advantage of using controllers is twofold. Apart from the reduction in energy consumption, it can also increase the efficiency of the electrically resistive heating systems. Another feature of electric heat tracing is the wide range of power that can be supplied to tracers. In this regard, operating voltages that start from 24 VAC up to 750 VDC and power from 4 Watt per foot (W/ft) up to 30 W/ft at 10°C were reported [9].

An alternative to electric heat tracers is thermally-sprayed coating heaters that have been developed recently. These heaters are composed of several layers that have favorable mechanical and electrical features for the given tasks. The usage of these coating-based heaters has been investigated and tested for a wide variety of applications with different compositions, geometries, and power requirements [10–15].

#### **1.3 Development and performance of coating-based resistive heating systems**

It has been about a century since the thermal spraying concept was presented initially. This idea was optimized and refined over time, and it has become one of the most popular and cost-efficient means of additive manufacturing. This process employs a high-temperature flame or plasma to melt and accelerate tiny powder particles and deposit a relatively dense coating onto given surfaces. These coatings have been used successfully in different fields such as automotive industry and electricity production so much so that it has become a reliable and essential element in today's industry. Fabrication of electric resistive heaters by using thermal spraying processes has received an increasing attention recently.

The application of thermal spraying methods in production of meso-electronics was studied by Sampath, et al. [16]. It was concluded that different electrical elements including resistors, insulators, and conductors can be manufactured by using thermal spraying processes to deposit proper ceramic and metal feedstock materials that have favorable electrical properties. Nickel-chromium alloy was deposited onto alumina by using high-velocity oxygen fuel and plasma processes to develop resistors. The fabricated resistors had a sheet resistance in the range of 17–54 KΩ/sq. The advantages of this production procedure including high production rate, low cost of production, flexibility on thermal spray method, and ability to deposit millimeter-thick coatings made this method an interesting means of production of electrical components [16]. Similarly, high-quality conductors and dielectrics were developed by using both highvelocity oxygen fuel and cold spray processes [17].

#### *1.3.1 Deposition of the electrically insulating layer*

Thanks to its unique dielectric properties, alumina has been used widely as electrical insulating layer. Capability of thermal spraying processes in easy and quick deposition of alumina on large surfaces made this method appealing to experts in electronics industry. The volume resistivity of alumina coating fabricated by plasma spraying was reported to be very high in the order of 10<sup>9</sup> –10<sup>10</sup> Ω cm [18]. It was found that the electrical resistivity of alumina depends on several factors, namely humidity, microstructural characteristics, phase composition of alumina, and applied pressure [19–23].

#### *1.3.2 Deposition of the heating element*

The heating element was made by depositing a conductive metallic alloy that possesses relatively high electrical resistivity onto electrical insulating layer (alumina). Several researchers reported successful deposition and operation of several materials, namely molybdenum, nickel, nickel-20 wt.% chromium, nickel-5 wt.% aluminum, iron-13 wt.% chromium, and iron-chromium-aluminum that were manufactured by using different thermal spraying processes such as wire arc, VPS, APS, HVOF, and wire flame spray for usage as the heating element [10–15, 24]. As an example, three thermal spraying methods, namely vacuum plasma spray, air plasma spray, and highvelocity oxygen fuel, were used to develop heating elements to produce an easily controlled and uniform heat flux. The maximum heat fluxes that were generated from the developed samples before they fail were relatively high in the range of 10.6–17.2 MW/m<sup>2</sup> [10].

**Figure 1.**

*Pipe assembly and the installed thermowells, pressure transmitter, and thermocouple used for freezing tests [25].*

#### **2. Bursting of steel pipes due to ice formation**

Given solidification of water inside pipes and pressure equipment is a widespread phenomenon in both industrial and residential sectors, and it brings about negative consequences every year, placing emphasis on further study of the freezing mechanism of pressurized enclosed water seems reasonable. In this regard, numerous experiments were conducted on two well-known pipe materials, namely ASTM A106- B and ASTM A333-6, to study the freezing of pressurized enclosed water and examine the damage exerted on the integrity of pipes that led to their ultimate failure. Then, the fracture surfaces of the pipes were cut and removed from the pipes for further analysis by using a scanning electron microscope.

#### **2.1 Experimental method**

The pipe assemblies that were used for the freezing tests were composed of 254-mm long and 51-mm diameter carbon steel pipes. The samples were made from both ASTM A333-6 and ASTM A106-B materials so that the impact of pipe material on the freezing behavior of the enclosed water can be investigated. Schedule 40 pipes that are used extensively in industry were selected for these experiments. The geometry of the pipe assembly completed with fittings and sensors is shown in **Figure 1**.

The pressure and temperature data were collected regularly at the rate of one reading per second from the pressure transmitter and thermocouples. Two separate data acquisition systems (cDAQ-9171, National Instruments, Austin, TX, USA and SCXI-1600, National Instruments, Austin, TX, USA) were used for collection and storage of entrapped water pressure and temperature data, respectively. The NI MAX software package was used for collection of the measurements.

#### **3. Development of multi-layered coating-based heater**

A multi-layered thermally-sprayed heating system was fabricated to eliminate formation of ice inside the pipes and control the temperature of the enclosed liquid. Given the steel pipe was a conductive material, alumina layer as the insulating layer was deposited onto the carbon steel pipe to prevent flow of electrons and short-circuiting between the pipe and the conductive heating element. In addition to flame spraying process, cold spraying method was also used to deposit dense copper coatings where proper electrical connections were required. The effect of various spraying parameters on the quality and heating performance of the coating heater was investigated.

After the samples were fabricated and their heating functionality was tested, they were sectioned for microstructural evaluation and elemental composition analysis. Important features of the coating system, namely continuity, homogeneity, and microstructural defects, were analyzed by using a scanning electron microscope.

#### **3.1 Experimental method**

#### *3.1.1 Feedstock powder*

Several powder feedstock materials, namely aluminum oxide (Al2O3, AMDRY 6060, Oerlikon Metco, Westbury, NY, USA), nickel-50 wt.% chromium (50 Nickel-50 Chromium, 1260F, Praxair, Concord, NH, USA), and copper (SST-C5003, CenterLine, Ltd., Windsor, ON, Canada), were utilized to produce the coatings by using flame spraying and cold spraying methods. The micrographs that were taken in secondary mode by using a scanning electron microscope (Zeiss Sigma 300 VP-FE, Carl Zeiss Canada Ltd., Toronto, ON, Canada) from these powders at 500x magnification are shown in **Figure 2a**–**c**. The alumina powder had angular/blocky morphology due to their manufacturing process, which was fusing and crushing, and their size distribution was between 5 and 45 μm [26]. The size distribution of Ni-50Cr powder particles was from 22 to 53 μm and their spheroidal shape was because of the gas atomization process that was used for fabrication of this powder [27]. The minimum purity of the copper powder that was used in this research was 99.7% and its size distribution was between 5 and 45 μm [28]. The dendritic structure of this powder was obtained through electrolysis process.

#### *3.1.2 Substrate preparation*

The coating layers were sprayed onto the pipe sections so that the performance of the heating system as a deicing element can be assessed. The pipe sample that was used as the substrate in this study is shown in **Figure 3**.

In order to deposit the alumina layer onto pipe sections, the surface of the pipe was cleaned and then grit blasted with #24 alumina grit (Manus Abrasive Systems Inc., Edmonton, AB, Canada) to create the roughness required to cause the adhesion between the substrate and alumina coating. An air pressure of 586 kPa (85 psig) was used to accomplish the grit blasting process. Only the central section of the pipe assemblies was grit blasted and the ends were covered by using a heat-resistant masking tape (170-10S Red, Green Belting Industries, Mississauga, ON, Canada) to ensure that coating layers would not be deposited at the welded end caps.

#### *3.1.3 Deposition of coating layers*

Once the pipe assemblies were grit blasted, they were coated with a layer of pure alumina (Al2O3 99.5+ wt.%.). The powder particles were consistently fed to the oxyacetylene torch (6P-II, Oerlikon Metco, Westbury, NY, USA). The spraying parameters for deposition of this layer were obtained through detailed trial and error procedure. Each time the microstructure of the obtained coatings was analyzed to ensure that the thickness of the coatings was uniform all over the substrate and that the coating layer was not damaged during the spraying process. Given the high melting point of alumina, relatively high flow rates for both acetylene and oxygen were selected to bring about a large high-temperature flame that is capable of melting the alumina powder particles. Furthermore, argon was selected as the carrier gas. The

**Figure 2.** *SEM images taken at 500 X magnification from powder materials, namely (a) Al2O3, (b) Ni-50Cr, and (c) Cu.* *Prevention of Pipes Bursting by Using a Novel Deicing Technology DOI: http://dx.doi.org/10.5772/intechopen.108972*

**Figure 3.**

*The components of the pipe assembly that were used as the substrate [29].*

flow rate of argon was set at 0.56 m<sup>3</sup> /h (20 standard cubic feet per hour) to carry the powder particles to the flame spray torch that was installed on a programmable robot (HP-20, Motoman, Yaskawa Electric Corp., Waukegan, IL, USA). The robot was used to ensure the uniformity and reproducibility of the fabricated coatings.

The samples were put into a rotating chuck to create a uniform coating. As the torch was passed in a linear manner along the pipe, the pipe was rotated at the angular speed of 600 rpm. To avoid making helical patterns for the alumina coating, relatively low linear speeds of 10 mm/s for alumina layer, 24 mm/s for nickel-chromium layer, and relatively high angular speed of 600 rpm for the rotating chuck were selected. To minimize the inconsistencies, the powder feed rate was reduced, and the number of passes was increased to achieve the most uniform microstructure for the deposited coatings. A relative term (flow meter reading parameter) was used to adjust the flow of powder particles to the torch.

After deposition of alumina, nickel-chromium alloy with a composition of Ni 53 wt.%, Cr 46 wt.%, and Fe 1.0 wt.% was deposited onto the alumina layer by using flame spraying method. The structure of the pipe assembly complete with bi-layered coating can be seen in **Figure 4**. The alumina coating was intentionally sprayed over a longer section to ensure that conductive nickel-chromium coating would not be deposited directly onto the substrate as it can result in short-circuiting and malfunction of the deicing system.

In order to connect the developed coating system to the power source, copper coatings in the form of rings were fabricated at both sides of the pipe assembly. Copper was selected for this purpose because of its proper electrical properties, which

**Figure 4.**

*The pipe assembly coated with flame-sprayed insulator and resistor layers and cold-sprayed connector [29].*

**Figure 5.**

*The coated pipe assembly used for deicing tests completed with all fittings and thermocouples [30].*

makes it an ideal option for electrical connections. The coating fabrication was accomplished by using the low-pressure cold spray system (SST series P, CenterLine, Ltd., Windsor, ON, Canada). The dendritic copper powder particles were accelerated inside the converging–diverging nozzle, and they formed a dense coating upon impact with the substrate. The working fluid that was used for this process to preheat and spray the copper particles was compressed air. The programmable robot was used for deposition of the copper rings as well. As the pipe was held in the chuck and was rotating, the cold spray nozzle passed the pipe assembly in the radial direction and generated the contact rings that are shown in **Figure 4**.

#### *3.1.4 Installation of fittings and sensors*

In order to read the temperature values of the enclosed water, two long thermocouples were inserted inside the thermowell at each end of the pipe assembly. The pipe assembly coated with the heating system is shown in **Figure 5**. In addition to the water/ice temperature, the temperature of the heating element was also measured by using a temperature sensor. The ambient air temperature inside the cold room was also measured by using a Type-T thermocouple.

#### *3.1.5 Deicing test*

The coated samples were put inside an 18.2 m3 cold room freezer (Foster Refrigerator USA, Kinderhook, NY, USA) to simulate the harsh environmental conditions. The temperature of the cold room was set at 25°C with a 2°C bandwidth. To control the airflow over the pipe, the samples were placed horizontally inside a closed galvanized sheet metal duct with dimensions of 2 m 0.66 m 0.48 m. The impact of circulation of air inside the cold room was minimized for the free convection scenario by placing the lead of the duct over the pipe assembly. Furthermore, the forced convection was simulated by turning on the fan at the end of the duct. In this regard, a 0.25 kW (0.33 hp) direct-drive tube-axial fan (DDA-12-10033B, Leader Fan Industries, Toronto, ON, Canada) was used to simulate the harsh environmental conditions when pipes are exposed to wind gusts during cold winter days. Further details about the structure and dimensions of the apparatus that was designed and fabricated for this investigation can be found in the previous study [31]. **Figure 6** shows the location at which the coated pipe assemblies were placed during the freezing and deicing tests.

#### **4. Results and discussion**

#### **4.1 Freezing behavior**

**Figure 7** shows the temperature and pressure traces that were obtained from the entrapped water inside Schedule 40 ASTM A333 Grade 6 pipe during the freezing test. The pipe assembly in this experiment was not fully filled with water and there was 5 vol.% of air inside the pipe. This graph demonstrates five distinct stages during the test, namely cooling of the water, the first solidification plateau, cooling of the mixture of water and ice, the second solidification plateau, and ice cooling. The water-cooling stage started when the fans inside the cold room started operating. This was followed by reduction in the temperature of ambient air inside the cold room. This stage was finished after a short super cooling period when plate-like solid crystals that are known as dendritic ice nucleated and grew [32]. Solidification of a portion of the enclosed water occurred in the next stage. The duration of this stage is directly impacted by the amount of air (gap) inside the pipe. It is well-known that the volume of water expands by 9% when it transforms into ice. Therefore, in case the vol.% of air inside the pipe is less than 9%, not all of the enclosed water can be converted to ice simply because there is not enough room for the volume expansion during the phase change. During this stage, the annular ice formed and grew inward [33]. Only a portion of water transformed into ice and once the pipe assembly was fully filled with the mix of water and ice, the pressure inside the pipe started building up. This noticeably reduced the rate of further water being transformed into ice because the pressure rise inside the pipe lowered the freezing temperature of water.

During the third stage, the pressure of the entrapped water sharply increased. This prevented the phase change of the rest of the entrapped water. A nonlinear relationship between melting temperature and pressure is obtained for hexagonal ice, which is given by Eq. (1) [34]. The purple curve in **Figure 7** shows the phase change pressure values for any given temperature according to this relation. As

**Figure 7.** *Temperature and pressure traces versus time for water/ice inside the pipe assembly.*

long as the pressure of the entrapped water was greater than the value obtained from this equation, water was not able to transform into ice even at subzero temperatures. This situation was discontinued when the pressure of the water was so high that it caused plastic deformation of the pipe. Afterward, given the volume of the pipe assembly was expanded during this inelastic deformation period and the pressure and the slope of pressure trace during this stage were much lower than that of the pressure trace during the elastic deformation stage, further water was able to undergo phase change process. Therefore, the second plateau that can be seen in fourth stage was indicative of the transformation of the remaining water inside the pipe into ice. Once the pipe was fully filled with solid ice at the end of the fourth stage, further exposure to the cold ambient air only reduced the temperature of the solid ice as shown in the fifth stage.

$$P\_{\rm f} = 6.11657 \times 10^{-4} - 414.5 \times \left[ \left( \frac{T\_{\rm f}}{273.16} \right)^{8.38} - 1 \right] \tag{1}$$

It is well-established that the plastic deformation of the pipe occurs when the hoop stress in the pipe reaches the yield strength of the pipe material. The hoop stress itself depends on the pipe diameter. Therefore, it can be concluded that the freezing point depression and the freezing behavior of the water that depends on the plastic deformation of the pipe, itself is a function of the pipe size, wall thickness, and material. The hoop stress values can be obtained by applying the relations developed by thickwalled pressure vessel theory. It is worth mentioning that the yield strength of the pipe is not a constant value, and it changes from one freezing–thawing cycle to another cycle due to the work hardening of the pipe material that is caused by movement of dislocations.

The comparison of the pressure values obtained from the freezing experiment with the results obtained from Eq. (1) is only valid when still water is entrapped in the pipe. That being said, the increasing deviation between the pressure experimental measurements and theoretical estimations is due to the fact that only ice exists in the pipe by the end of the fourth stage, and therefore, that comparison is not valid anymore.

#### **4.2 Plastic deformation and work hardening**

A PI tape was used to measure the inelastic deformation of the pipes during the freezing experiment. In order to calculate the transverse strain of the pipe, the outer diameter of the pipes was measured before and after freezing experiments. Obviously, the maximum deformation took place in the middle of the pipe assembly where it was farthest from the circumferential welds and the reinforcing impact of the welds was at its minimum level. It was concluded that the maximum strain values that A333-6 and A106-B steel pipes could accommodate before failure were 11.3% and 13.3%, respectively. Although the minimum transverse elongation before rupture has been reported to be 16.5% for ASTM A106 Grade B steel [35], the pipe did not meet the expected value under freezing experiments. However, the obtained value for A333-6 steel was very close to the reported minimum transverse elongation, which is 11.4% [36].

The maximum pressure values inside the pipe were greatly impacted by the number of times the pipe underwent plastic deformation during the freezing and thawing cycles. The increase in the peak pressure value from one test to another test, which was because of the work hardening of A333-6 and A106-B Schedule 40 pipes, can be clearly observed in **Figure 8a** and **b**.

#### **4.3 Pipe rupture**

**Figure 9** shows the failure pattern of the pipes after several freezing and bulging experiments, which is a common pattern for pipe bursting cases. The axial crack was formed in both pipes due to the developed excessive hoop stresses during the tests. It is well-established that the hoop stress is the maximum principal stress when pipes are pressurized. This pressurization was caused by gradual formation of ice accompanied by volume expansion and low compressibility of unfrozen water [1]. This pattern and location of the failure were expected as it is known that pipes fail where water freezes last. In this case, the remaining unfrozen water was entrapped in the middle of the pipe where final rupture occurred.

It can be seen in **Figure 9** that the extent of the deformation of the pipes was different for these two pipes. At the failure spot, the generated gap was wider and larger for the A106-B pipe compared to the A333-6 pipe. This was because of the amount of energy that was absorbed by the pipe material during the formation and propagation of the crack upon rupture. It is believed that the toughness of these pipe materials at low temperature has an important role to play. It is well known that the A333-6 pipe is the preferred option for installation at sites that are exposed to lowtemperature environments because of its acceptable low-temperature toughness. Therefore, the higher toughness of A333-6 pipe compared to A106-B pipe results in greater resistance against brittle fractured caused by internal pressurization and overload. In contrast, the degradation of absorption energy of A106-B pipe material at lower temperature was found by conducting Charpy impact test [37].

#### **Figure 8.**

*The impact of work hardening of the pipe materials on the peak pressure during the freezing test for (a) A106-B and (b) A333-6 pipes.*

In order to observe the fracture surfaces, the pipes were cut after the failure. To study the fracture patterns at different locations of the pipes, eight different rings were cut at 1 cm intervals from both pipes as shown in **Figure 9**.

#### **4.4 Failure analysis**

High-quality side images were taken from the rings shown in **Figure 9** to observe the macroscopic fracture features. These features, along with the geometry of the fractured surfaces are often used to determine the mechanism of the fracture and also *Prevention of Pipes Bursting by Using a Novel Deicing Technology DOI: http://dx.doi.org/10.5772/intechopen.108972*

**Figure 9.** *Failure site of (a) A106-B and (b) A333-6 pipes due to overload during freezing test [25].*

the pipe failure mode. The slant and double-slant fractures that can be observed in **Figure 10** are indicative of ductile failure due to plane stress loading conditions [38], which is a common form of failure for pressure vessels and pipes [39]. In these cases, the pipe failure begins with local thinning. Then localization of shear bands that occurs at 45°angle develops in the necked spot [40, 41]. The necked area that led to the rupture of the pipe in slant failure mode can be seen in **Figure 10**.

#### **4.5 Fractography**

The micrographs taken from the fracture surfaces confirm that the failure mode of both pipes was ductile tearing. This occurred because of the propagation of microcracks as a result of the coalescence of microvoids in the axial direction (along *z*-

**Figure 10.** *Side views of the rings cut from (a) A106-B and (b) A333-6 pipes [25].*

axis), as shown in **Figure 11**. Another feature that confirms the ductile failure mode of pipes is the fibrous and dull appearance of the fractographs that embodies numerous dimples.

#### **4.6 Coating characterization**

High-magnification images were taken in backscattered electron mode from the cross sections of the coated samples. These SEM images are shown in **Figure 12**. Although generation of microcracks was anticipated initially due to the brittle nature of the ceramic layer and thermal stress that can be generated at the interfaces due to the mismatch in material properties [42, 43], no major damage or delamination was observed during the microstructural evaluation of the coated samples even after conducting expedited deicing tests with high supplied power values.

The material for each layer of the coating system was selected based on the required mechanical and electrical properties to accomplish the given tasks. Alumina was selected due to its unique properties. While this material is a great electrical insulator, it conducts heat readily in comparison to other ceramic materials [26]. Therefore, this material was able to prevent short-circuiting between the pipe (substrate) and the heating element (nickel-chromium alloy), and at the same time, it was capable of transferring the generated heat by the heating element to the pipe to accomplish the deicing task [44, 45].

Although penetration of the nickel-chromium alloy inside the alumina layer was observed to some extent during the microstructural evaluation, this did not result in the malfunction of the coating-based heater. The penetration of the molten particles in the alumina layer was due to the presence of a network of open pores inside the ceramic coating, which is a characteristic of coatings fabricated by flame spraying process [45]. Given the effective thickness of alumina was reduced as a consequence

**Figure 11.** *Fractographs taken at 500X magnification from (a) A106-B pipe and (b) A333-6 pipe.*

of this incident, alumina coatings with at least 50 microns thickness had to be deposited onto the pipe substrates to prevent any possible short-circuiting. It is wellestablished that much denser coatings can be fabricated by using more sophisticated

#### **Figure 12.**

*SEM images taken from the coated pipe assembly: (a) bi-layered coating system, (b) NiCr-Al2O3 interface, and (c) Al2O3-steel interface [29].*

thermal spray means such as plasma spraying, high-velocity oxygen fuel, or suspension plasma spray to prevent this undesirable occurrence [46, 47].

The nickel-chromium material was selected for the heating element because of its relatively high electrical resistivity. Although the electrical resistivity is a material property, the electrical resistance of the heating element is a function of its cross-

**Figure 13.**

*Temperature traces from deicing tests for (a) free convection with supplied power of 80 W and (b) forced convection with supplied power of 500 W.*

sectional area, which itself can be affected by discontinuities and presence of imperfections such as cracks and pores [45]. Presence of pores in the coating increases the electrical resistance as a result of reduction in the cross-sectional area through which free electrons can move [48]. It has been observed that formation and propagation of microcracks can also bring about the same consequences [49]. Therefore, the performance of the heating element that depends on its electrical resistance can be varied by using different thermal spraying methods.

**Figure 14.**

*The presentation of (a) carbon steel pipes used as substrate and (b) pipes equipped with novel cold-sprayed tin and copper heating elements.*

#### **4.7 Heating performance**

The performance of the coating-based heater was assessed under different power inputs. For this purpose, when 20 V was applied to the coating, 80 W power was generated, which was sufficient to heat and melt the ice inside the pipe. The temperature traces for the coating surface and enclosed ice/water temperatures can be seen in **Figure 13a** for the case of free convection. The functionality of the developed deicing system was evaluated based on the time that was required to accomplish the deicing test. It was observed that even under harsh conditions where the bare pipe was exposed to the forced convection due to the circulation of the cold ambient air, the fabricated coating-base heater melted the ice inside the pipe successfully as shown in **Figure 13b**. It was found that the voltage and current of the heating element were proportional and therefore, the nickel-chromium layer was an ohmic material [45].

#### **4.8 Future work**

In addition to the technical aspects, the economic and environmental implications of fabrication and utilization of the coating-based heating systems have also been investigated [50–52]. The promising results obtained from these studies confirmed the possibility of using the coating-based deicing systems on mass scale, however further advancement and improvement in the geometry of patterned coatings, manufacturing process, and the selection of the materials can even bring about more encouraging outcomes. Therefore, in the next phase of the project, in order to reduce the production cost, minimize the emission of exhaust gases to the environment, and achieve a more consistent and dense structure, the heating elements are fabricated by using cold spray system in helical pattern as shown in **Figure 14**. Other advantages of fabrication of heating element via cold spraying are prevention of penetration of the heating element inside the electrically insulating layer, on-site repairability, and enhanced bonding to the insulating layer thanks to the applied compressive stresses during the spraying process.

### **5. Conclusions**

The freezing mechanism of the pressurized water inside steel pipes and the subsequent bulging and bursting were studied. In order to overcome this undesirable issue that brings about noticeable financial losses and environmental concerns, a novel deicing system was developed by using flame spraying and cold spraying processes. Once the performance of the developed heating system was assessed in simulated environmental conditions under different supplied powers, the coated pipes were sectioned so that their microstructure can be analyzed. The outstanding findings of this study are presented hereunder:


#### **Author details**

Milad Rezvani Rad1 \* and Andre McDonald<sup>2</sup>

1 University of Southern Indiana, Evansville, Indiana, United States of America

2 University of Alberta, Edmonton, Alberta, Canada

\*Address all correspondence to: milad.rad@usi.edu

© 2022 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.

*Prevention of Pipes Bursting by Using a Novel Deicing Technology DOI: http://dx.doi.org/10.5772/intechopen.108972*

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#### **Chapter 3**

## Evaluation of Electromagnetic Interferences Affecting Metallic Pipelines

*Denisa Șteț, Levente Czumbil and Dan Doru Micu*

#### **Abstract**

This chapter presents some analysis of the modeling techniques used to evaluate the effects of electromagnetic interference phenomena that could occur when metallic pipelines are placed close to high-voltage power lines. The electric and magnetic fields produced by overhead power lines could perturb the normal operation of the metallic pipelines through induced currents and voltages. These perturbations could be dangerous for both pipeline operating personnel (as electrical hazard) and pipeline structural integrity (due to accelerated electrochemical corrosion phenomena). The chapter depicts the electromagnetic coupling mechanisms behind the abovementioned interference phenomena and how the induced voltages could be evaluated. A parametric analysis is showcased to highlight the influence of various geometrical and electrical parameters.

**Keywords:** metallic pipelines, electromagnetic interferences, power lines influence, modeling and simulation, electromagnetic compatibility, induced voltages

#### **1. Introduction**

European regulations regarding environmental protection, as well as economic reasons aimed to reduce construction costs, have determined a significant decrease in the access of energy carrier (like oil products, electricity, methane gas, or/and water) infrastructure to new right of ways. Therefore, the transport and distribution network of water, natural gas, or crude oil pipelines must share on long distances (for several kilometers) the same distribution corridors with high-voltage power lines (HVPL) (see **Figure 1**).

Studies on high-voltage AC power lines operating in steady state or fault regime identified the presence of electromagnetic interferences in nearby metallic structures due to the electric and magnetic fields generated by the currents flowing through the HVPL conductors [1–4].

Therefore, in many situations, the metallic pipeline (MP) used for the transport and distribution of liquid or gaseous substances may be exposed to induced currents or voltages. These could be dangerous on one side for operating personnel, who coming into contact with the pipeline may be exposed to electrocution, and on the

**Figure 1.** *Cross section of a common HVPL-MP right-of-way.*

other hand, to the structural integrity of the pipeline due to accelerated corrosion phenomena [5–7].

In 1995, within the CIGRE Working Group 36.02, the document entitled Guide Concerning Influence of High Voltage AC Power Systems on Metallic Pipelines [8] has been developed, which addresses the influence of high-voltage energy devices on nearby metallic pipelines. It is a reference document in the field of electromagnetic interferences and summarizes the following aspects:


Regarding the protection of underground MP against electrochemical corrosion processes, a special attention has been paid to the effects caused by AC systems. Previously, AC corrosion was considered negligible compared to the interferences in DC state. But research in the field proved the opposite, resulting in pipeline protection guidelines [9, 10] and a European regulatory standard [11]. The main causes of AC corrosion are the leakage current densities from the pipeline into the surrounding soil. The source of leakage currents is the AC voltages induced in the pipelines due to various electromagnetic coupling mechanisms [6, 7].

To comply with the European norms and regulations, it is necessary to investigate the electromagnetic interference phenomena in detail and to precisely determine the

*Evaluation of Electromagnetic Interferences Affecting Metallic Pipelines DOI: http://dx.doi.org/10.5772/intechopen.108811*

magnitude of the voltages induced in MPs exposed to nearby power lines, both under normal and HVPL fault operating conditions. To identify and apply the proper protection techniques for the metallic pipelines, the geometrical and electrical parameters that influence the level of induced voltages must be determined. The difficulty in accessing buried MPs not only requires mathematical models for the numerical evaluation of induced voltages, instead of on-site measurements [12].

#### **2. Electromagnetic interference phenomena**

At low frequencies, when the wavelength of the electrical signal is long compared to the size of the perturbation source, the electromagnetic interference propagates through capacitive (electric), inductive (magnetic), or conductive (galvanic) couplings [8, 11, 13]. Conversely, at high frequencies when the wavelength is comparable to the dimensions of the perturbation source, electromagnetic radiation appears. The radiation transition limit is variable, but in most practical cases, it is approximately 10 m, which corresponds to a frequency of 30 MHz. Therefore, in case of interference problems between HVPL and neighboring metallic structures, only conductive, capacitive, and inductive couplings should be considered [8, 11].

#### **2.1 Galvanic coupling**

Galvanic (conductive) couplings occur when two electrical circuits have a common portion of the circuit. In this situation, electromagnetic interference propagates in the receiver (victim) through conduction, through one or more conductors (power line, cable shield, etc.), or even through passive elements (capacitors, transformers, etc.).

The energy transported through power lines can be transmitted conductively to adjacent metallic structures (underground metallic pipes, etc.) if these structures are connected to the AC circuit (the grounding grid of a power line tower or any other device that has a network extended earthing), either directly (metallic connection) or through the proximity of the pipeline and the grounding grid (see **Figure 2**).

**Figure 2.** *Galvanic coupling between a HVPL and a MP.*

The occurrence of a ground fault in the power system corresponds to a high-value short-circuit current flowing through earth. The electrical potential of the soil (close to the grounding grid) will increase considerably in relation with the pipeline potential (assumed to be at a reference potential) [14]. Any element connected between MP and the ground will be subject to this potential difference; in other words, the fault currents that flow through the earthing electrode of an HVPL tower, or substation, produce an increase in the potential of the electrode as well as the potential of the soil in its vicinity compared to the reference potential. Under these conditions, the metallic structures (pipelines) will be influenced if they are directly connected to the grounding grid of the electric energy transport/distribution system (for example, of a transformer station), or if they cross the influence zone near the grounding sockets. In these cases, the potential of the pipeline increases and can be transmitted over a relatively large distance (several km), depending on the degree of electrical insulation of the interfered metallic structure and the resistivity of the soil in the respective area.

#### **2.2 Capacitive coupling**

Capacitive (electric) coupling occurs between two circuits whose conductors are at different potentials. As a result of the potential difference between the conductors, an electric field is produced and modeled in an equivalent electrical scheme by a parasitic capacitance. **Figure 3** presents an example of electrical coupling of two circuits (1) and (2) by means of a quasi-stationary electric field, respectively, through parasitic capacities.

**Figure 3.** *(a) Field model, and (b) network model for capacitive coupling.*

**Figure 4.** *Capacitive coupling between HVPL and metallic pipeline.*

*Evaluation of Electromagnetic Interferences Affecting Metallic Pipelines DOI: http://dx.doi.org/10.5772/intechopen.108811*

In the case of capacitive coupling, determined by the electric field of the highvoltage power line, the hypothesis that the soil is a perfect conductor, of infinite conductivity, can be used, because its conductivity is much higher than that of air (see **Figure 4**).

Therefore, only aboveground metallic structures located in the vicinity of HVPL are subject to capacitive coupling perturbations; this effect appears both in normal operating conditions and power line fault conditions [15].

#### **2.3 Inductive coupling**

Inductive (magnetic) couplings occur between two or more circuits passed by electric currents producing time-varying magnetic fields that induce in the victim circuit a voltage that disturbs the desired signal (if exists). The action of the magnetic field of the disturbing circuit on the perturbed one can be represented in the equivalent electric circuit model by a mutual inductance or by an induced voltage source (see **Figure 5**).

A metallic pipeline near a high-voltage power line (as shown in **Figure 6**) is subject to induced voltages by magnetic coupling. In other words, a significant part of the energy transmitted through power lines surrounds the conductors and extends over large distances from their center. The nearby metallic pipelines, which have a common path with HVPL, can capture this energy in unfavorable parallelism and/or power line operation conditions (high

**Figure 5.** *(a) Field model, and (b) network model for inductive coupling.*

**Figure 6.** *Inductive coupling between HVPL and metallic pipeline.*

zero-sequence currents, phase to ground faults, unbalanced loads, non-symmetric current system, etc.)

As shown in some previous works, such as [16–18], the inductive influence from 50 Hz or 60 Hz electric power systems on pipeline networks is related to the timevarying magnetic field generated by the electric currents flowing in power lines conductors. The main parameters generally considered for this type of influence are the electric power system's current load level, the ground wire current, the exposure length, the separation distance between involved structures of both systems, and the soil model resistivity [19, 20].

Given that the pipeline is buried and assuming that the ground permittivity can be neglected, there is no capacitive influence between the overhead transmission line and the buried pipeline. In this regard, the next chapters emphasize aspects related to investigation of the induced AC voltage effects on a buried pipeline due to the currents in an overhead transmission line.

#### **3. Modeling techniques of the electromagnetic interference problems**

Numerous numerical simulations have been carried out over the past few decades to analyze interference problems involving power lines and earth-return conductors. There are two main approaches for assessing such phenomena: transmission line theory [1, 21–23] and a hybrid approach based on finite element method (FEM) and circuit analysis [2, 3, 24–26]. The first approach has the advantage that it can be used for any operation state, while the other is only applicable to steady-state conditions.

In either case, a fundamental requirement of such modeling technique is that they should consider soil resistivity variations [24, 27, 28], non-parallel pipeline-power line right of ways, and multiple metallic conductors, such as sky wires and/or mitigation wires. Various commercial computer software are available to assess the influence of electric power systems on pipeline systems, but their licensing and annual renewal fees could be very high.

The analytical forms of the solutions of the electromagnetic field problems, related to the cylindrical current-carrying conductors in the presence of the ground, are a rather laborious problem, even when the geometry of the conductor network is simple. There are several difficulties with buried pipelines:


In many cases [3, 15, 19, 20, 26, 28], the pipeline network is analyzed for lowfrequency energizing currents conditions, but problems could arise when the

#### *Evaluation of Electromagnetic Interferences Affecting Metallic Pipelines DOI: http://dx.doi.org/10.5772/intechopen.108811*

performance of the network at high frequency should be investigated and transient state studies are required [18, 21, 23]. For both situations, an equivalent electrical circuit model of the investigated interference problem (like the one in **Figure 7**) must be constructed and analyzed considering all the present metallic conductors and all the coupling mechanisms.

By solving the equivalent electric circuit model of the entire right-of-way for the proper HVPL operating conditions, the induced currents and voltages could be evaluated along the pipeline length, and it could be determined if the pipeline would operate normally or mitigation measures should be applied.

#### **3.1 Transmission line approach**

The transmission line approach is the most general and complex formulation of an HVPL-MP interference problem that could be applied for both steady state and transient state studies. It has been developed from the well-known lumped-element model for an infinitesimal (*Δx* length) single-wire conductor presented in **Figure 8**.

Imposing *Δx* ! 0 the following telegrapher's equations can be written:

$$-\frac{\partial u(\mathbf{x},t)}{\partial \mathbf{x}} = R\_0 \cdot i(\mathbf{x},t) + L\_0 \cdot \frac{\partial i(\mathbf{x},t)}{\partial t} \tag{1}$$

$$-\frac{\partial i(\mathbf{x},t)}{\partial \mathbf{x}} = G\_0 \cdot u(\mathbf{x},t) + C\_0 \cdot \frac{\partial u(\mathbf{x},t)}{\partial t} \tag{2}$$

where *R*<sup>0</sup> is per-unit length resistance, *L*<sup>0</sup> per-unit length inductance, *G*<sup>0</sup> per-unit length conductance, and *C*<sup>0</sup> per-unit length capacitance. If Laplace transformation (*s* ¼ *jω*) is applied to the above equations, we will obtain:

**Figure 7.** *Equivalent circuit model for a general case of HVPL–MP interference.*

**Figure 8.**

*Lumped-element equivalent circuit transmission line model.*

$$-\frac{\partial u(\mathbf{x},\mathbf{s})}{\partial \mathbf{x}} = (R\_0 + sL\_0) \cdot i(\mathbf{x}, \mathbf{s}) = Z\_0 \cdot i(\mathbf{x}, \mathbf{s}) \tag{3}$$

$$-\frac{\partial \dot{\boldsymbol{x}}(\mathbf{x}, \mathbf{s})}{\partial \mathbf{x}} = (\mathbf{G}\_0 + \mathbf{s}\mathbf{C}\_0) \cdot \boldsymbol{\mu}(\mathbf{x}, \mathbf{s}) = \mathbf{Z}\_0 = \mathbf{Y}\_0 \cdot \boldsymbol{\mu}(\mathbf{x}, \mathbf{s})\tag{4}$$

where *Z*<sup>0</sup> is the longitudinal impedance, and *Y*<sup>0</sup> is the shunt admittance The solution of Eqs. (3) and (4) can be written as:

$$u(\mathbf{x}, \mathbf{s}) = e^{-\mathbf{y}\cdot\mathbf{x}} \cdot u^+ + e^{\mathbf{y}\cdot\mathbf{x}} \cdot u^- \tag{5}$$

$$i(\mathbf{x}, \mathbf{s}) = \frac{\mathbf{1}}{Z\_{\mathcal{C}}} (e^{-\gamma \cdot \mathbf{x}} \cdot u^{+} - e^{\gamma \cdot \mathbf{x}} \cdot u^{-}) \tag{6}$$

where *u*<sup>þ</sup> and *u*� denote the voltage as incident and reflect waves, respectively, *γ* is the propagation constant, and *ZC* is the characteristic impedance, defined as:

$$
\gamma = \sqrt{Z\_0 \cdot Y\_0} \tag{7}
$$

$$Z\_{\mathbb{C}} = \sqrt{\frac{Z\_0}{Y\_{\mathbb{C}}}} \tag{8}$$

The substitution of boundary conditions *x* ¼ 0 and *x* ¼ *l* where *l* denotes the length, into (5) and (6) leads to the following line terminals equations:

$$
u\_k - Z\_C \cdot i\_k = e^{-r \cdot x} \cdot (u\_m + Z\_C \cdot i\_m) \tag{9}$$

$$
u\_k + Z\_C \cdot i\_k = e^{r \cdot x} \cdot (u\_m - Z\_C \cdot i\_m) \tag{10}$$

where *k* and *m* denote the terminals of the line corresponding to *x* ¼ 0 and *x* ¼ *l*, respectively. The currents *ik* and *im* are entering the line at both ends. The above equations are rewritten to:

$$\dot{u}\_k = Y\_C \cdot u\_k - H(Y\_C \cdot u\_m + i\_m) \tag{11}$$

$$i\_m = Y\_\mathcal{C} \cdot u\_m - H(Y\_\mathcal{C} \cdot u\_k + i\_k) \tag{12}$$

where *YC* <sup>¼</sup> <sup>1</sup>*=ZC* is the characteristic admittance, and *<sup>H</sup>* <sup>¼</sup> *<sup>e</sup><sup>γ</sup>*�*<sup>l</sup>* is the propagation function. Eqs. (11) and (12) can be rewritten using vector and matrices for a multi-conductor system:

*Evaluation of Electromagnetic Interferences Affecting Metallic Pipelines DOI: http://dx.doi.org/10.5772/intechopen.108811*

$$\left[I\_{k}\right] = \left[\mathbf{Y}\_{\mathcal{C}}\right] \cdot \left[\mathbf{U}\_{k}\right] - H(\left[\mathbf{Y}\_{\mathcal{C}}\right] \cdot \left[\mathbf{U}\_{k}\right] + \left[I\_{m}\right]) \tag{13}$$

$$[I\_m] = [Y\_C] \cdot [U\_m] - H([Y\_C] \cdot [U\_k] + [I\_k]) \tag{14}$$

To determine the frequency-dependent characteristic impedances and admittances required for transient state analysis, **Figure 9** illustrates the transition from a linear earth electrode to an equivalent transmission line with complex values, for the frequency-dependent parameters *ZC*ð Þ *ω* and *γ ω*ð Þ, which denote the characteristic impedance and propagation function, respectively:

$$\underline{\mathbf{Z}\_{C}}(\boldsymbol{\alpha}) = \sqrt{\frac{\mathbf{R}\_{0} + j\boldsymbol{\alpha}\mathbf{L}\_{0}}{\mathbf{G}\_{0} + j\boldsymbol{\alpha}\mathbf{C}\_{0}}} = \sqrt{\frac{\underline{\mathbf{Z}\_{0}}(\boldsymbol{\alpha})}{\underline{\mathbf{Y}}\_{0}(\boldsymbol{\alpha})}} \underline{\mathbf{y}}(\boldsymbol{\alpha}) = \sqrt{\underline{\mathbf{Z}\_{0}}(\boldsymbol{\alpha}) \cdot \underline{\mathbf{Y}}\_{0}(\boldsymbol{\alpha})}\tag{15}$$

$$
\underline{\Gamma\_0}(a) = j a \sqrt{\mu\_0 \cdot \varepsilon\_0} \tag{16}
$$

This form corresponds to any two-conductor transmission line. Sunde derived equivalent expressions for a single conductor in contact to the soil, with a current returning through the earth [29, 30]. Both the complex longitudinal impedance per unit length, *Z*0, and the complex transversal admittance per unit length *Y*0, of a horizontal conductor consist of an internal term and an earth return term in the following manner:

$$\underline{Z}\_0 \cong \underline{Z}\_{0i} + \frac{j o \mu\_0}{2\pi} \cdot \ln \frac{1.85}{\sqrt{\underline{\chi}^2 + \underline{\varGamma}^2} \cdot \sqrt{2ah}} \tag{17}$$

$$(\underline{Y}\_0)^{-1} \cong (\underline{Y}\_{0i})^{-1} + \frac{j o \mu\_0}{\pi \cdot \underline{\varGamma}^2} \cdot \ln \frac{1.12}{\underline{\varchi} \cdot \sqrt{2ah}} \tag{18}$$

Here *Z*0*<sup>i</sup>* denotes the internal complex impedance of a conductor of radius *a*, buried at depth *h*, and *Y*0*<sup>i</sup>* stands for the insulation admittance of an eventual coating gap, through which the conductor is in contact to the surrounding medium. The latter is characterized by the earth conductivity *σE*, relative electric permittivity *εrE*, and magnetic permeability *μ*0. All three lead to a propagation function [31–33]:

$$
\underline{\Gamma}(o) = \sqrt{j o \mu\_0 \cdot (\sigma\_E + j o \sigma\_0 \varepsilon\_{rE})} \tag{19}
$$

which would govern the transmission of impulses along the conductor if it were imbedded in a homogeneous soil with these parameters.

**Figure 9.** *Characteristic impedance of a linear earth electrode.*

#### **Figure 10.** *Multiple conductors over a conductive earth.*

For a more complex geometry like the one from **Figure 10**, the generalized telegrapher's equations for the case of multi-wire system along the *x*-axis above an imperfectly conducting ground and in the presence of an external electromagnetic excitation are [33]:

$$\begin{cases} \frac{d}{d\mathbf{x}} \left[ \underline{U}\_i(\mathbf{x}) \right] + \left[ jo\underline{L}\_{\vec{\eta}} \right] \cdot \left[ \underline{I}\_i(\mathbf{x}) \right] + \left[ \underline{Z}\_{\mathbf{g}\vec{\eta}} \right] \cdot \left[ \underline{I}\_i(\mathbf{x}) \right] = \left[ \underline{S}\_{\mathbf{i}i}^\epsilon(\mathbf{x}) \right] \\\ \frac{d}{d\mathbf{x}} \left[ \underline{I}\_i(\mathbf{x}) \right] + \left[ \underline{G}\_{\vec{\eta}} \right] \cdot \left[ \underline{U}\_i(\mathbf{x}) \right] + \left[ jo\underline{C}\_{\vec{\eta}} \right] \cdot \left[ \underline{L}\_i^\epsilon(\mathbf{x}) \right] = \left[ \underline{S}\_{\vec{\omega}i}^\epsilon(\mathbf{x}) \right] \end{cases} \tag{20}$$

where:


In (20), there are neglected the terms corresponding to wire impedance and the ground admittance. Indeed, it has been shown in literature [30–33] that for a typical frequency range of interest (bellow 10 MHz) and for typical overhead lines, these parameters can be disregarded with reasonable approximation.

The expression for the mutual ground impedance between two conductors *i* and *j* has been derived by Sunde and is given by:

$$\underline{Z}\_{\text{g\ddot{y}}} = \frac{j\alpha\mu\_0}{\pi} \cdot \int\_0^\infty \frac{e^{-\left(h\_i + h\_j\right)\cdot x}}{\sqrt{\mathbf{x}^2 + \underline{Y}^2} + \mathbf{x}} \cdot \cos\left(r\_{\hat{\mathbf{y}}}\mathbf{x}\right) d\mathbf{x} \tag{21}$$

where:

$$\underline{\chi}\_{\rm g} = \sqrt{j o \mu\_0 \cdot \left(\sigma\_{\rm g} + j o \epsilon \epsilon\_0 \epsilon\_{\rm g}\right)}\tag{22}$$

in which *σ<sup>g</sup>* and *εrg* are the ground conductivity and relative permittivity. The diagonal terms of the ground impedance matrix are given by:

$$\underline{Z}\_{\rm gij} = \frac{j o \mu\_0}{\pi} \cdot \int\_0^\infty \frac{e^{-2h\_i \cdot \mathbf{x}}}{\sqrt{\mathbf{x}^2 + \underline{\chi^2}} + \mathbf{x}} \cdot d\mathbf{x} \tag{23}$$

Eqs. (22) and (23) are not suitable for a numerical evaluation since they involve Sommerfeld integrals. Several approximate expressions for *Zgij* have been presented in literature, but one of the simplest forms was proposed by Sunde himself and is given by the following logarithmic function [29]:

$$\underline{Z}\_{\rm gi} \cong \frac{j\alpha\mu\_0}{2\pi} \cdot \ln \frac{\mathbf{1} + \underline{\chi}\_{\mathbf{g}} \cdot h\_i}{\underline{\chi}\_{\mathbf{g}} \cdot h\_i} \tag{24}$$

It has been shown [33] that (24) is an excellent approximation of the general expression (23).

In the following, the logarithmic approximation is extended also to off-diagonal terms. Using Euler relation in (21) and after some simple mathematical manipulations, we can obtain [33]:

$$\underline{Z}\_{\text{g\ddot{y}}} = \frac{j\alpha\mu\_0}{\pi} \cdot \left(\frac{1}{2}\right)\_0^\infty \frac{e^{-2\underline{h}\_{\text{g}}\cdot\mathbf{x}}}{\sqrt{\mathbf{x}^2 + \underline{\chi}^2} + \infty} \cdot d\mathbf{x} + \frac{1}{2}\int\_0^\infty \frac{e^{-2\underline{h}\_{\text{g}}^\*\cdot\mathbf{x}}}{\sqrt{\mathbf{x}^2 + \underline{\chi}^2} + \infty} \cdot d\mathbf{x}\right) \tag{25}$$

in which *hij* is a complex quantity, and *h*<sup>∗</sup> *ij* is its complex conjugate [31]:

$$h\_{\underline{i}\underline{j}} = \frac{h\_i + h\_{\underline{j}}}{2} + j \cdot \frac{r\_{\vec{i}\underline{j}}}{2} \tag{26}$$

Using the earlier mentioned approximate identity between expressions (23) and (24) we can express:

$$\frac{j o \mu\_0}{2\pi} \cdot \int\_0^\infty \frac{e^{-2\underline{h}\_j \cdot x}}{\sqrt{\varkappa^2 + \underline{\chi}\_g^2} + \varkappa} \cdot dx \cong \frac{j o \mu\_0}{4\pi} \ln\left(\frac{1 + \underline{\chi}\_g \cdot \underline{h}\_{\dot{\imath}j}}{\underline{\chi}\_g \cdot \underline{h}\_{\dot{\imath}j}}\right) \tag{27}$$
 
$$\frac{j o \mu\_0}{2\pi} \cdot \int\_0^\infty \frac{e^{-2\underline{h}\_{\dot{\imath}}^\* \cdot x}}{\sqrt{\varkappa^2 + \underline{\chi}\_g^2} + \varkappa} \cdot dx \cong \frac{j o \mu\_0}{4\pi} \ln\left(\frac{1 + \underline{\chi}\_g \cdot \underline{h}\_{\dot{\imath}j}^\*}{\underline{\chi}\_g \cdot \underline{h}\_{\dot{\imath}j}^\*}\right)$$

Introducing (27) in (25), the following approximation can be derived for the general term *Zgij* of the ground impedance matrix:

$$\mathcal{Z}\_{\text{gij}} = \frac{j\alpha\mu\_0}{\pi} \cdot \left(\frac{1}{2}\right)\_0^\infty \frac{e^{-2\underline{h}\_j \cdot \mathbf{x}}}{\sqrt{\mathbf{x}^2 + \underline{\chi}^2} + \mathbf{x}} \cdot d\mathbf{x} + \frac{1}{2}\int\_0^\infty \frac{e^{-2\underline{h}\_j^\* \cdot \mathbf{x}}}{\sqrt{\mathbf{x}^2 + \underline{\chi}^2} + \mathbf{x}} \cdot d\mathbf{x}\right) \tag{28}$$

By introducing another type of approximation called the low-frequency approximation *σ<sup>g</sup>* > > *ωε*0*εrg* � �, valid for low-frequency analysis, the general expression will become [33]:

$$\underline{Z}\_{\rm gi} = \frac{j a \mu\_0}{\pi} \cdot \int\_0^\infty \frac{e^{-\left(h\_i + h\_j\right) \cdot x}}{\sqrt{\varkappa^2 + j a \mu\_0 \sigma\_\rm g} + \varkappa} \cdot \cos\left(r\_{\vec{\eta}} \cdot x\right) d\mathbf{x} \tag{29}$$

And in particular, the diagonal terms will be given by:

$$\underline{Z}\_{\rm gij} = \frac{j o \mu\_0}{\pi} \cdot \int\_0^\infty \frac{e^{-2h\_i \cdot \mathbf{x}}}{\sqrt{\mathbf{x}^2 + j o \mu\_0 \sigma\_\mathbf{g}} + \mathbf{x}} \cdot d\mathbf{x} \tag{30}$$

Therefore, the general expressions for the elements of the ground impedance matrix in the multi-conductor transmission line equations are in terms of infinite integrals. Accurate approximations for the diagonal terms of the ground impedance matrix have already been presented in the literature [30–33] and that proposed by Sunde is the most accurate and simple.

#### **3.2 Combined field and circuit method for evaluating inductive and capacitive matrices**

In case of steady state or phase to ground fault HVPL operating conditions when the fault is far away of the common HVPL-MP right-of-way and a single frequency analysis is enough (no transients study is required), a simpler distributed elements equivalent electrical circuit approach could be implemented. At the same time for more complex conductor shapes (not only cylindrical/cable type conductors), a hybrid/combined field and circuit method should be applied [2, 3, 5, 25].

The hybrid method [2] was developed for single frequency analysis but could be extended to multiple frequency studies also if necessary. It combines the finite element method (FEM) analysis of the electromagnetic field in the common distribution corridor, with Faraday's law and with electric circuits theory respectively to determine the self and mutual inductances/capacitances between any metallic structures present in a HVPL-MP interference problem [5].

This type of analysis consists of solving Maxwell's equations in nonuniform threedimensional space and consists of the following steps [5, 25]:

*Step 1.* A multi-conductor system that includes the pipeline, phase wires, and overhead ground wires together with electrical towers and grounding systems is modeled.

*Step 2.* A reference current/voltage is set on phase wire *A* ð Þ 0° and zero value currents/voltages on the rest of the metallic conductors (including the other phases).

*Step 3.* The magnetic vector potential *A* ! on the cross sections of all the metallic

structures (*A* ! *cond*) that make up the studied problem is determined by FEM analysis. *Step 4.* The electric charges *Q*, due to the imposed reference voltage, that appear on the surface of all the conductors, respectively on the surface of the ground, are determined by means of the FEM analysis.

*Step 5.* Self-inductance per unit length of the conductor on which the reference current *Iref* was imposed is determined, respectively, the mutual inductances per unit length given by this current in the rest of the metallic structures are computed.

$$L\_{self} \text{ or } L\_{mut} = \frac{\overrightarrow{A}\_{cond}}{\overrightarrow{I}\_{ref}} [\text{H/m}] \tag{31}$$

*Evaluation of Electromagnetic Interferences Affecting Metallic Pipelines DOI: http://dx.doi.org/10.5772/intechopen.108811*

*Step 6*. Based on the determined electrical charges *Q* or linear charge distributions *ρℓ*, the conductor-soil (self) capacity per unit length of the conductor on which the reference voltage *Vref* was imposed, respectively, the mutual capacities compared to the rest of the conductors per unit length, are evaluated

$$\mathbf{C}\_{\text{ref}}\text{ or }\mathbf{C}\_{\text{mut}} = \frac{\rho\_\ell}{U} = \frac{\rho\_\ell}{V\_{\text{ref}} - V\_0} = \frac{\rho\_\ell}{V\_{\text{ref}}} \,\mathrm{[C/m]}\tag{32}$$

*Step 7.* Steps 3 and 5 respectively 4 and 6 are repeated for the case where the imposed reference current/voltage is set one by one on all the metallic structures that constitutes the investigated interference problem.

Using this iterative algorithm [25], the matrixes of self and mutual inductances and capacities between all the present structures could be constructed. Thus, the obtained inductive and capacitive matrixes describe the inductive respectively the capacitive couplings mechanism between HVPL and MP. Therefore, they could be used to solve the equivalent electric circuit model [5, 25].

The FEM evaluates the total electromagnetic interference effect in a single step, avoiding the separation of the inductive, capacitive, and conductive components, which is necessary in the circuit model. However, the FEM limitation is that when the common corridor is very long and consists of many circuits, the modeling and computation can be extremely time-expansive.

Therefore, usually 2D cross-section FEM analysis is applied for HVPL-MP parallel exposures. The following system of equations describes the linear 2D electromagnetic diffusion problem for the *z*-direction (along the common right-ofway) components *Az* of the magnetic vector potential and *Jz*of the total current density vector [2]:

$$\begin{cases} \frac{1}{\mu\_0 \mu\_r} \cdot \left[ \frac{\partial^2 A\_x}{\partial \mathbf{x}^2} + \frac{\partial^2 A\_x}{\partial \mathbf{y}^2} \right] - j\alpha \sigma A\_x + J\_{sx} = \mathbf{0} \\\\ -j\alpha \sigma A\_x + J\_{sx} = J\_x \end{cases} \tag{33}$$
 
$$\iint\_{S\_i} J\_x ds = I\_i$$

where *Jsz* is the source current density in the *z* direction, and *Ii* is the imposed current on conductor *i* of *Si* cross section.

Eq. (33) could be solved by any dedicated finite element calculation software to compute the magnetic vector potentials on the surface of each metallic structure (phase wires, sky wires and pipeline).

#### **4. Analysis of main parameters that influence the induced voltages**

For the numerical evaluation of the induced voltages in metallic pipelines exposed to the electromagnetic fields produced by nearby high-voltage power lines, several software applications and software packages that implement the above presented approaches or similar methodologies could be used. There are to main types of such software applications:


To highlight the effect of the main parameters that could influence the level of induced voltages in metallic pipelines, a usual HVPL-MP electromagnetic interference problem is simulated and analyzed in the following using the EMTP-RV software, due to its user-friendly interface and high computational capabilities.

An underground metallic gas transportation pipeline is considered to share the same distribution corridor with a single circuit 220 kV/50 Hz overhead electrical power line.

To model the electromagnetic coupling between HVPL and the underground MP, the new *Line/Cable Data* component is used from EMTP-RV, which allows to combine overhead conductors (HVPL phase wires and sky wires) with above or underground cables (the pipeline is considered as an unenergized insulated hollow cable with the same geometrical dimensions).

Using this *Line/Cable Data* component, the equivalent *wideband* transmission line model of the common distribution corridor is created (see **Figure 11**). The *wideband* transmission line model is based on the Universal Line Model (ULM) introduced in [34] applying the transmission line approach described in section 3.1. The *wideband* model

#### *Evaluation of Electromagnetic Interferences Affecting Metallic Pipelines DOI: http://dx.doi.org/10.5772/intechopen.108811*

form EMTP-RV uses complex poles and zeros for the rational approximation of the frequency-dependent characteristic admittance and wave propagation matrices, while it also takes fully into account the frequency dependence of the modal transformation matrix [23]. Therefore, it can be considered the most accurate time-domain model, its applicability being extended to any type of overhead or underground transmission line.

The HVPL consists of three phase wires in a horizontal layout with a 7.4 m spacing between each other placed at 24 m above ground and two sky wires placed at 26.5 m above ground with a 4.4 m spacing from HVPL axis. The metallic pipeline is buried at 1.2 m depth. The geometrical and electrical parameters of the conductors involved in the analyzed HVPL-MP common distribution corridor (phase wires, sky wires, and the metallic pipeline) are presented in **Table 1**.

**Figure 11** presents the electrical circuit model implemented in the EMTP-RV software of the investigated HVPL-MP electromagnetic interference problem considering a single equivalent transmission line element for the entire common distribution corridor. Such an implementation allows us to evaluate only the maximum values of the induced voltage, which is reached at the ends of the analyzed pipeline.

To be able to evaluate the variation of the induce voltages along the pipeline length and not just the maximum values, the common distribution corridor must be divided into several consecutive transmission line sections (segments). Usually, the length of such a transmission line segment is considered equal the span between two consecutive power line towers. This approach allows also to investigate the influence of power line towers and their grounding in case of transient state faults like the ones generated by lightning strikes to HVPL. **Figure 12** presents a detailed representation of a 1 km long HVPL-MP common right-of-way using 250 m transmission line segments as the distance between two consecutive power line towers.

To be as close as possible to real-life situations, at both ends of common distribution corridor, the pipeline is continued with 5 km long transmission line segments representing the pipeline outside of the HVPL zone of influence. Being an underground pipeline, the inductive coupling will prevail in the electromagnetic interference problem that occurs.

#### **4.1 Steady-State operating condition**

As a first step, the induced AC voltages are evaluated in the underground MP during power line steady-state conditions with a 300 A symmetrical current load (around 115 MVA power load with a 0.9 power factor). The HVPL-MP distribution


#### **Table 1.**

*Parameters that describe the system under investigation.*

**Figure 12.** *Detailed representation of a HVPL-MP common distribution corridor implemented in EMTP-RV.*

corridor has a length of 2 km. The underground MP follows a parallel route within the right-of-way at 30 m separation distance from the power line axis on the right side. The HVPL phase wires are placed in the phase A (0°), phase B (120°), and phase C (240°) order on the power line towers. Therefore, phase C (240°) is the nearest phase wire to the underground MP. The soil is considered uniform with a resistivity of 100 Ωm for the entire right-of-way.

The time-domain variation of the induced voltages at different locations along pipeline length in the investigated common distribution corridor evaluated through EMTP-RV simulation and modeling is showcased in **Figure 13**.

The graphical representation from **Figure 13** highlights values around 6.5 V regarding the induced AC voltages, which are in dangerous range from corrosion

point of view. The critical values of the induced voltages, regarding the start of electrochemical corrosion reaction, are around 1.2 V as in [9, 10, 22] are presented.

Due to the induced voltages in the metallic pipeline, leakage currents will appear that depend directly on the induced voltage levels, the pipeline conductivity, and its insulation. These leakage currents could highly increase if pipeline insulation defects are present, accelerating the corrosion process.

**Figure 14** shows the variation of the induced AV voltages along the pipeline length as RMS values. It can be observed that the "V"-shaped curve as reported in several literature studies [3, 17, 23–26], with the highest induced RMS voltage values recorded at right-of-way ends. It must be mentioned that at each time moment, the pipeline has opposite electrical potential at its end as in **Figure 13** can be noticed.

#### **4.2 Parametric analysis**

To highlight the effect that different geometric and electrical parameters can have on the values of the induced voltages in the case of a HVPL-MP interference problem, a parametric analysis is made, by taking into account: the length of the common distribution corridor, the HVPL-MP separation distance, the HVPL load current, and the soil resistivity variation.

**Figure 15** highlights the induced voltages along the MP in the case of a right-ofway with a length that varies from 0.5 km to 10 km. It is found that the induced voltage increases quasi-proportionally with the length of the common distribution corridor, up to a "critical" (or characteristic) length that depends on the geometrical and electrical parameters of the pipeline, the insulation, and the surrounding environment (soil), being defined based on the propagation constant of the electromagnetic wave along the pipeline. Among these parameters, various studies have shown that soil resistivity has the least influence.

Another important aspect in an HVPL-MP interference problem is the distance between the electrical line and the metallic structure. **Figure 16**.a highlights the induced voltage values in the MP if it is positioned at up to 1000 meters from the power line (on either side of it). While **Figure 16**.b is a more detailed representation of the induced voltages in MP recorded near the metallic towers of the power line, up to 100 m on either side HVPL axis.

**Figure 14.** *Induced voltages along the MP, in case of steady-state condition.*

**Figure 15.**

*(a) Induced voltages along pipeline length for different right-of-way length, and (b) Maximum induced voltage variation with right-of-way length.*

**Figure 16.**

*(a) Maximum induced voltage variation with HVPL-MP separation distance, and (b) Highlight for close HVPL-MP exposures (under 100 m).*

**Figure 17.**

*(a) Induced voltages along pipeline length for different HVPL load currents, and (b) maximum induced voltage variation with HVPL load current.*

To investigate the induced voltage levels that can occur for different power flows, the symmetrical current load on HVPL was considered in the range of 150–450 A. According to **Figure 17**, the levels of the induced voltages in MP increase with the

value of current load (an increase from 300 A to 450 A of the load current leads to 50% increase of the maximum value of the induced AC voltage).

Evaluation of induced voltages for different soil resistivity values has shown than even when the soil resistivity varies in a large range like from 10 Ωm to 10 kΩm, the maximum induced voltage values vary only with 2–3% with regard to the steady-state results. The soil resistivity has a higher influence on the pipeline leakage currents in case of insulation defects or in case of lighting generated HVPL faults when soil ionization phenomena could occur around power line tower groundings.

#### **4.3 Oblique exposures and intersections HVPL-MP**

The above-presented EMTP-RV implementation and the parametric analysis consider perfect parallel exposure between HVPL and MP. However, in practice [20, 35, 36], there are frequent situations in which above or underground MPs present oblique exposures to the HVPL axis (see **Figures 18** and **19**) [8].

Therefore, to analyze real-life HVPL-MP interference problems using electromagnetic transients programs like EMTP-RV, each oblique HVPL-MP exposure section has to be replaced with an equivalent parallel HVPL-MP exposure section for which the HVPL-MP separation distance *deq* is given by the Eq. (34) as long as the ratio between *d*<sup>2</sup> and *d*<sup>1</sup> (see **Figure 18**) is less than or equal to 3 [8].

**Figure 18.** *Oblique HVPL-MP exposure.*

**Figure 19.** *Induced AC voltage in case of a real HVPL-MP right-of way [37].*

$$d\_{eq} = \sqrt{d\_1 \cdot d\_2} \tag{34}$$

If the ratio between the distances at which the two ends of the oblique exposure *d*2*=d*<sup>1</sup> is greater than 3, the oblique exposure must be divided into two subsections, both sections verifying the imposed condition:

$$\frac{d\_3}{d\_1} \le 3 \quad \text{and} \quad \frac{d\_2}{d\_3} \le 3 \tag{35}$$

On the other hand, if MP sub-crosses the HVPL, this section is replaced by an equivalent section, in which the MP is located at *d* ¼ 6 *m* distance from the HVPL, for a length equal to the length of the MP projection segment considered in around the crossing point up to a separation distance less than or equal to 10 *m* on either sides of the HVPL.

**Figure 19** presents the AC induce voltage variation along a complex real-life HVPL-MP common distribution corridor applying the abovementioned right-of-way segmentation procedure [37].

#### **5. Conclusions**

An evaluation of the electromagnetic interferences phenomena effects that affect metallic pipelines placed in the vicinity of high-voltage power lines is presented in this chapter highlighting different analyzing and modeling techniques.

As it is shown, the presence of AC power supply systems may cause voltages to build up in nearby metallic pipeline systems, due to one or more of the following mechanisms: inductive, conductive and capacitive coupling. Such voltages may put the pipeline operating personal in danger, damage the pipeline, disturb the electrical/ electronic equipment connected to the pipeline.

The induced voltages and currents on the buried pipelines may be dangerous to pipeline security due to the AC corrosion and deteriorated the cathodic protection devices. Consequently, a mitigation system is required to be designed to reduce the effects of the corrosion of which main causes are the leakage current densities from the pipeline into the surrounding soil due to AC-induced voltages.

Since even very high-quality insulation of underground metal pipelines could present insulation defects, to reduce the probability of occurrence of corrosion phenomena, European standards [14] suggest limiting the induced voltages depending on the electrochemical characteristics of the soil, so that they do not exceed:


Different techniques can be applied to mitigate induced AC voltages and gas, such as cancelation wires, gradient control wires, and insulating joints [9, 10, 26].

#### **Thanks**

Energy Transition Research Center (EnTReC) of Technical University of Cluj-Napoca is an EMTP-RV official educational partner and would like to thank POWERSYS SOLUTIONS for offering the academic license of the software.

*Evaluation of Electromagnetic Interferences Affecting Metallic Pipelines DOI: http://dx.doi.org/10.5772/intechopen.108811*

### **Author details**

Denisa Șteț\*, Levente Czumbil and Dan Doru Micu Energy Transition Research Center (EnTReC), Technical University of Cluj-Napoca (TUCN), Cluj-Napoca, Romania

\*Address all correspondence to: denisa.stet@ethm.utcluj.ro

© 2022 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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#### **Chapter 4**

## Assessment of the Structural Integrity of the Pipes with Anomalies Such as Local Elastic-plastic Deformations

*Alin Dinita*

#### **Abstract**

Pipelines are one of the most practical and economically efficient ways to transport dangerous and/or flammable substances, for which road or rail transport is often impossible. The evaluation of the processes that can negatively influence the performance of the pipelines is particularly important for assessing the risk associated with the operation of these technical systems and the potential for technical accidents. The anomalies that can be found on the pipes can be classified into two main categories. Imperfections that do not inadmissibly affect their load-bearing capacity and defects with significant negative influences on the correct operation and load-bearing capacity of the piping, which require supervision and maintenance measures. The influence of these anomalies and the processes that lead to the decrease of the pipeline-bearing capacity constitutes the main objectives of the analysis performed. The local elastic-plastic deformation anomalies are considered, for which the parameters of the geometric model of the defect profile, the conditions for generating these anomalies, and the evaluation methods, respectively, were analyzed, both analytically and experimentally.

**Keywords:** pipeline, defect assessment, elasto-plastic deformations, burst pressure, dent and gouge

#### **1. Introduction**

The probability of the appearance of anomalies on the pipelines is closely related to the mechanical resistance and tenacity characteristics of the material (steel) from which the pipelines are made. Anomalies that can be found on pipelines can be classified into material loss anomalies, which consist of the thinning of the pipeline wall through the loss of metal in the presence or absence of a corrosive process, and anomalies such as cracks and anomalies such as local elastic-plastic deformations.

The steels from which the pipelines are mainly made have the ferrito-pearlitic structure typical of carbon or low-alloy hypoeutectoid steels, and the increase in their mechanical resistance characteristics is achieved mainly by increasing the carbon

concentration, which has the effect of increasing the percentage content of perlite in the structure. Modern steels for pipes have a structure with acicular ferrite (low carbon bainite), their high mechanical resistance characteristics being achieved mainly by obtaining a very fine grain and ensuring hardening effects by the precipitation of some intermetallic compounds. The weldability of these steels is satisfactory if the carbon concentration is not increased excessively and if the welding procedure and regime are chosen appropriately.

The carbon concentration of these steels does not exceed 0.30...0.31%, to obtain the higher degrees of resistance resorting to the use of manufacturing recipes with manganese concentrations higher than those typical of carbon steels. Flat strip semifinished products intended for the manufacture of longitudinally or helically welded pipes are made from such steels by controlled rolling or thermomechanical rolling (rolling with high degrees of deformation in which working temperatures, heating, cooling, and deformation speeds are strictly controlled), which emphasize the presence in these steels of a wide range of microalloying elements (Nb, V, Ti, Mo, etc.).

Defining and classifying the factors that lead to the degradation and failure of pipelines is a problem for which specialized literature provides a multitude of solutions. To carry out the analysis of the factors and processes that lead to the progressive degradation and failure of pipelines, it was considered that the factors to be considered correspond to the potential hazards specified by the ASME B31.8S standard for risk assessment and development of pipeline integrity management plans intended for natural gas transport.

The classification of these factors (potential hazards) is presented synthetically in **Table 1**; as can be seen, the ASME B31.8S standard recommends the division into three classes of factors that can determine the failure of pipelines, and for each class, there are three categories of factors, each category having one or more factors.

Imperfections are anomalies in configuration, dimensions, microstructure, etc. present in the pipelines, which do not inadmissibly affect their load-bearing capacity, and the defects are the imperfections with significant negative influences on the correct operation and the load-bearing capacity of the pipelines, which require maintenance measures. The classification of pipe imperfections and defects can be based on the configuration criterion that defines the following four categories of pipe imperfections and defects:

#### **1.1 Imperfections and geometric defects produced by the local deformation of the pipes**

*Local deformations or dents* are deviations from the circular shape of the crosssection of the piping, obtained by local deformation of the piping, inwards, without removing the material and, as a result, without reducing the wall thickness. Indentations influence the flow of gases in the pipelines and can cause major difficulties in performing cleaning or washing operations and in the internal inspection of the pipelines, by blocking the movement of working tools or PIG devices. Indentations of an elastic nature, produced, for example, by the interaction of pipes with pieces of rock can be eliminated by simply removing the cause (pieces of rock that produced the deformation); *Gouges* are areas of the pipes where the wall thickness has been locally reduced due to the removal of material through a mechanical action, and gouges are defects or imperfections of great severity because their presence can lead to the initiation of brittle cracking processes, representative types of such defects are shown in **Figure 1**.

*Assessment of the Structural Integrity of the Pipes with Anomalies Such as Local... DOI: http://dx.doi.org/10.5772/intechopen.108437*


#### **Table 1.**

*The categories of factors that can affect the integrity of natural gas pipelines.*

**Figure 1.** *Imperfections and defects of local elastic-plastic deformation type.*

#### **1.2 Imperfections and defects such as material loss**

These imperfections or defects consist in the general or local thinning of the wall of the pipeline through the loss of metal in the presence or absence of a corrosive process. The most common imperfections or defects in this category are as follows: *areas of local thinning* (areas on the surface of a pipe element, having the axial extension or length of the same order of magnitude as the circumferential extension or width, in which the material has been removed by corrosion and/or erosion), *pinching or pitting* (traces of local corrosion on the surface of a pipe element, in the form of cavities or holes, having the surface diameter of the ordinal size of the wall thickness of the respective pipe element), and representative types of such defects are shown in **Figure 2**.

#### **1.3 Cracks**

Cracks are the anomalies with the greatest harm, which produce strong mechanical stress concentration effects and significantly reduce the carrying capacity of the pipes.

**Figure 2.** *Imperfections and defects such as material loss.*

#### *Assessment of the Structural Integrity of the Pipes with Anomalies Such as Local... DOI: http://dx.doi.org/10.5772/intechopen.108437*

Their dimensions can change over time through stable growth, until they reach critical dimensions, at which unstable propagation and rupture of the tubing can occur. Depending on the toughness characteristics of the piping material, cracks can generate brittle fracture phenomena (which occur at high speeds and propagate over large distances, giving rise to damage with important consequences) or ductile fracture phenomena (which occur at small speeds and are preceded by plastic deformation processes, which consume an important part of the available energy and thus contribute to stopping the phenomenon and limiting its consequences), and representative types of such anomalies are presented in **Figure 3**.

**Figure 3.** *Cracks.*

#### **1.4 Other types of defects or imperfections**

This category includes imperfections or defects that cannot be attached to one of the previously specified categories, for example: defects in the sealing systems of valves or fittings mounted on pipes, defects in threads, defects in the manufacture (lamination, welding) of pipes.

#### **2. Characterization of the geometry of anomalies produced by local elastic-plastic deformation of pipes**

The way of generating and modeling the geometry of the anomalies (imperfections and defects) produced by the local deformation of the pipes is a very delicate problem, considering both the complexity of the phenomena that take place in the deformed area and the need to obtain some methods for evaluating the bearing capacity of the pipes that present such anomalies.

In the specialized literature, you can find various geometric modeling (semielliptical, hemispherical, or hexagonal cavities) that take into account both the characteristics of the pipeline and the cause of these anomalies (third-party interventions). **Figure 4** shows two of the most important geometric models that characterize the imperfections and defects produced by local plastic deformation of pipes [1–3].

Local plastic deformation represents a change in the circular section of a pipe or a distortion of the pipe in a circular section. In **Figure 5**, a geometric modeling model of the shape of these elastic-plastic deformation type anomalies is proposed. The depth of this type of anomaly is defined as the maximum reduction of the diameter in the defect area compared to the initial nominal diameter of the pipe. This definition of deformation depth includes both local deformation and any distortion in the circular section of the pipe (ovality).

To determine the stress state in the area of an imperfection and defects produced by the local deformation of a pipe, it is necessary to carry out an analytical description of the geometry of the anomaly. For the indentations with axial orientation on the pipelines, the analytical description of the shape was proposed by using a cylindrical coordinate system, of the type defined in the sketch in **Figure 5** and a function with the following analytical form [4, 5]:

$$r(\varphi, z) = R - d\_p(z) \left[ \mathbf{1} + \cos \left( 2\pi \frac{\varrho}{\rho\_0(z)} \right) \right] |\varphi| \le \rho\_0 |\varphi| > \rho\_0 \tag{1}$$

#### **Figure 4.**

*Two models for defining the geometric profile for an anomaly of the type of local elastic-plastic deformations.*

*Assessment of the Structural Integrity of the Pipes with Anomalies Such as Local... DOI: http://dx.doi.org/10.5772/intechopen.108437*

**Figure 5.** *Scheme of the description of the configuration of a pipeline in the area of a local elastic-plastic deformation type anomaly.*

In Eq. (1) the circumferential anomalies profile are given as a deformation from the pipe radius *R*, the maximum of the deformation is given by the term *dp(z)*, and the circumferential deformation of the local elastic-plastic deformation defect is limited to the angle 2*φ*0ð Þ*z* , the measure of which is correlated with the defect width *cp(z),*2*φ*0ð Þ¼ *<sup>z</sup> cp*ð Þ*<sup>z</sup> <sup>R</sup>* [6–8].

A better modeling of the profile geometry of local plastic deformation defects is obtained if an expression corresponding to a Gauss curve (in polar coordinates) is used in Eq. (2):

$$r(\rho, z) = R\_{\epsilon} - d\_p(z) \exp\left[ -\frac{1}{2} \left( \frac{z}{z\_0(z)} \right)^2 \right] \left[ -\frac{1}{2} \left( \frac{\rho}{\rho\_0(z)} \right)^2 \right] \tag{2}$$

where the magnitude of the deviation is given by the term *dp(z)*, the circumferential extension of the local deformation type defect is limited to the angle 2*φ*0ð Þ*z* , the measure of which is correlated with the width of the defect *cp(z)*, 2*φ*0ð Þ¼ *<sup>z</sup> cp*ð Þ*<sup>z</sup> <sup>R</sup>* , and the axial extension variation is given by the term *z* limited to � *Sp/2*, *Sp* represents the axial length of the anomaly.

To describe with this analytical expression, the shape of the cross sections of a pipeline in the area of an indentation, a computer application, PROFIL\_DEF, was created using the Mathcad computing environment. Results obtained by using this computer product and which correctly and objectively describe, from an analytical point of view, the geometric profile of an imperfection or local deformation type defect, are presented in **Figure 6**.

**Figure 7** shows a semi-model of the geometric profile corresponding to an imperfection and defect of the local deformation type, made by considering several dimensions of the connection radius that the geometry of the anomaly presents in relation to the geometry of the pipelines that thus present anomalies. The geometric profile was

#### **Figure 6.**

*Results obtained when describing the geometric profile of an imperfection or local deformation type defect.*

divided into three contact zones corresponding to the three regions of the profile geometry (AB, BC, and CD).

Two examples of modeling the geometric profile of imperfections and local deformation type defects are also presented in **Figure 7**, considering the parameters and the *Assessment of the Structural Integrity of the Pipes with Anomalies Such as Local... DOI: http://dx.doi.org/10.5772/intechopen.108437*

#### **Figure 7.**

*The general semi-profile of the geometry of an imperfection and local deformation type defect; a–AB zone, b–BC zone, c–CD zone. (a) a - semi-model of the geometric profile for elastic-plastic anomaly, r1 - main deformation radius in the contact area, r2- the secondary deformation radius in the contact area, α - main angle corresponding to radius r1, β - the secondary angle corresponding to radius r2, rm-the average radius of the pipeline. (b) case study.*

method described above, with the observation that the method presented involves a rigorous and difficult determination in practice of the parameters that define defect geometry, modeling that was done using CAD software (AutoCAD).

The verification of the models designed to define the geometric profile of imperfections and local deformation type defects was carried out through experimental tests using ring samples taken from pipes for pipelines. Pipe rings with a diameter of

**Figure 8.**

*An example of the application of the proposed model for the analytical description of the configuration of local elastic-plastic deformation anomalies.*

*De* = 114.3 mm, wall thickness *t* = 4.6 mm; 5.4 mm and 8.0 mm and with a length of 50 mm were used for the experimental tests.

The tested pipe section was rested on a prism, and the application of force on the indenter was done gradually with the help of a testing machine, measuring the depth of the anomaly with the help of a dial gauge.

The graphic representations in **Figure 8** summarize the comparisons of the results obtained, depending on the type of indenter used and the wall thickness of the pipe ring used to simulate the appearance of an imperfection and/or local deformation type defect.

From the analysis of the experimental program, the proposed analytical modeling ensures good fidelity for anomalies with relatively small depth, of the type that must be evaluated to be accepted on pipelines.

#### **3. Experimental determination of burst pressure of pipes with anomalies such as local elastic-plastic deformations**

The experimental verification of the behavior of pipelines with different types of anomalies under mechanical stress [9, 10] is one of the methods by which their residual bearing capacity is established and leads directly to conclusions regarding the level of confidence that must be associated with the results of the assessment of the severity of the anomalies through analytical methods available. The research carried out concerned the pressure test until bursting of some pipes on which anomalies were

#### *Assessment of the Structural Integrity of the Pipes with Anomalies Such as Local... DOI: http://dx.doi.org/10.5772/intechopen.108437*

made (imperfections and/or anomalies of the type of local elastic-plastic deformations; dents with gouges, considered to be the most dangerous of the anomalies). Two samples were made for the internal pressure test, and each sample consisting of a pipe and two ellipsoidal bottoms welded to its ends, on which three connections were mounted: one for the manometer, one for filling the sample with water and pressurizing to it, and one for venting the sample before pressurization.

The anomalies realized in the two samples subjected to the internal pressure test until breaking/smashing are suggestively presented in **Figure 9**, they are of the indentation type for the first sample, respectively of the indentation type with a gouge for the second sample, and their geometric characteristics are shown in **Table 2**.

The samples that were subjected to the internal pressure test and were made from pipe sections with a diameter of *De* = 163 mm, the wall thickness being *t* = 8 mm. The stand on which the samples were tested is reproduced schematically in **Figure 10**, in

#### **Figure 9.**

*Realization of the anomalies on the two samples subjected to the test at internal pressure until bursting.*


#### **Table 2.**

*The geometric characteristics of the anomalies made on the samples.*

**Figure 10.** *Scheme and main components of the stand used for internal pressure testing of pipe samples with local surface anomalies.*

which the constructive elements of the high-pressure pump and the work platform from the composition of this stand are presented.

When carrying out the experimental research on the samples subjected to the tests, four strain-resistive transducers were applied around the anomaly, two in the circumferential direction (TER 1 and TER 3) and two in the axial direction (TER 2 and TER4) and two strain-resistive transducers, one in the circumferential direction (TER 5) and one in the axial direction (TER 6). During test, the computer controls the acquisition of data with the help of the SPIDER 8 device through the dedicated software CATMAN.

The method of carrying out the tests and the results obtained (quantitative and qualitative), regarding the behavior of the samples during the tests, are presented below:


The processing of the experimental results (see **Figure 13**) was carried out by determining the mechanical stresses in the circumferential direction and in the axial direction, using the known formulas:

*Assessment of the Structural Integrity of the Pipes with Anomalies Such as Local... DOI: http://dx.doi.org/10.5772/intechopen.108437*

#### **Figure 11.** *Images regarding the behavior of sample 1 in the internal pressure test.*

**Figure 12.** *Images regarding the area where the burst occurred for sample 2.*

**Figure 13.** *The results of the experimental analysis by the resistive tensometry method. a. results for sample, b. results for sample 2.*

$$
\sigma\_{\theta\ddagger\dot{\jmath}} = \frac{E}{\mathbf{1} - \mu^2} \left[ \varepsilon\_{\theta\ddagger} + \mu \varepsilon\_{\dot{x}\dot{\jmath}} \right]; \sigma\_{x\ddagger\dot{\jmath}} = \frac{E}{\mathbf{1} - \mu^2} \left[ \varepsilon\_{\dot{x}\dot{\jmath}} + \mu \varepsilon\_{\theta\ddagger} \right], \tag{3}
$$

where *E* represents the longitudinal modulus of elasticity, and μ—Poisson's coefficient for the sample steel; the specific deformations in the circumferential direction εθ<sup>i</sup> or in the axial direction εzj (i and j being the identification numbers of the transducers).

For the determinations, tensor-resistive transducers were used, with a base of 10 mm and resistance 120 Ω. With each of these transducers, at different pressures, during the test of the sample, the specific deformations in the circumferential direction and in the axial direction were determined.

#### **4. Conclusions**

The main causes that determine the degradation of pipelines, usually installed underground, are corrosion, third-party interventions and manufacturing, and construction defects; in the last period, there was a significant increase in concessions produced by third-party interventions/interferences.

The principles of codification of imperfections and defects of different types and categories ensure the synthetic and, at the same time, comprehensive specification in the technical documentation regarding the operation, exploitation, maintenance of pipelines of any anomalies found, and their causes; also, the coding of imperfections and defects is particularly useful for the development of maintenance procedures for pipelines.

#### *Assessment of the Structural Integrity of the Pipes with Anomalies Such as Local... DOI: http://dx.doi.org/10.5772/intechopen.108437*

Anomalies of the elastic-plastic deformation type may have the character of dents (deviations from the circular shape of the cross section of the pipe tubing, obtained by local deformation of the pipe, inward, without removing the material and, as a result, without reducing the wall thickness), gouges (areas of pipe where the wall thickness has been locally reduced due to material removal by mechanical action) or pitted indentations (indentations that have gouges at the bottom of the deformed zone).

Modeling the geometry of anomalies produced by local elastic-plastic deformation is particularly important because it serves to develop evaluation procedures and characterize the severity of these anomalies.

Experimental research into the operational behavior of pipelines with anomalies is usually carried out on stands that allow pressure testing of pipe or pipe sections, with or without anomalies. Usually, these stands can ensure the determination of the stress states of the tested samples and the pressure at which the respective pipe samples fail.

The stand designed and built allows the investigation of the behavior of pipes with or without anomalies, being able to provide both results obtained with the help of electro-tensometry transducers applied to the sample and the bursting pressure of the sample.

#### **Author details**

Alin Dinita Petroleum-Gas University of Ploiesti, Ploiesti, Romania

\*Address all correspondence to: adinita@upg-ploiesti.ro

© 2022 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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Section 2
