1

Evaluation of Cathodic Disbondment Resistance of Pipeline Coatings – a Review 1

Min Xu a, *, Catherine Lam b, Dennis Wong b, and Edouard Asselina 2

aDepartment of Materials Engineering, The University of British Columbia, 6350 Stores Road, 3

Vancouver, BC V6T 1Z4 4

bShawCor Ltd. Research & Development, 25 Bethridge Road, Toronto, ON M9W 1M7 5

*Email: min.xu@ubc.ca 6

7

Abstract 8

For coated pipelines with cathodic protection, cathodic disbondment has been recognized as the 9

main cause of coating degradation. A thorough understanding of the cathodic disbondment 10

mechanism, a proper evaluation on the coating resistance to disbondment as well as a precise 11

monitoring of the cathodic protection level would minimize the occurrence of cathodic 12

disbondment and help to maintain a pipeline’s long-term integrity. This paper first reviews the 13

various mechanisms proposed in the literature for cathodic disbondment. To provide useful 14

guidance for selecting pipeline coating systems, cathodic disbondment test methods (categorized 15

as standard test methods or modified test methods) that have been developed over the past two 16

decades are then critically reviewed. In-situ techniques for assessing cathodic disbondment of 17

coatings, especially those having promise for field applications are also discussed. Finally, a 18

brief discussion on the mitigation of cathodic disbondment and the associated corrosion is 19

presented. 20

2

Keywords: pipeline coatings; cathodic disbondment; cathodic disbondment test; electrochemical 1

impedance spectroscopy; corrosion and corrosion control 2

1. Introduction 3

Pipelines are an integral component of our modern society as they provide the most practical and 4

the safest way of delivering oil and gas [1] to satisfy the energy demands of our daily life. Take 5

Canada for example: approximately 1.2 billion barrels (~ 190 billion liters) of liquid petroleum 6

products and 5.3 trillion cubic feet (~ 150 trillion liters) of natural gas are transported by 7

pipelines each year [2]. In 2015, the total economic impact from the operation of all of the 8

energy transmission pipelines in Canada on GDP was estimated to be about $11.5 billion (about 9

0.7% of the total GDP) [3]. The long-term integrity of pipelines is therefore of significant 10

economic importance. As steel pipes tend to corrode when exposed to moist soil or wet air, 11

applying coatings on the surface of pipes is the primary approach to defend them against 12

corrosion. Coatings provide a physical and electrochemical barrier between the steel surface of 13

the pipe and the surrounding environment. However, defects in pipeline coatings are unavoidable 14

at every stage of the coating process, the pipe installation and operation of the pipeline. A 15

discontinuity in pipeline coatings is referred to as a “holiday”. Once a holiday exposes bare metal 16

to the surrounding environment, corrosion can occur. This is where cathodic protection (CP) 17

becomes important. CP enforces a cathodic current at the coating defect sites to protect the steel 18

pipe from corrosion so that it cannot become an anode. Therefore, for immersed and buried 19

pipelines in the field, a combination of protective coatings and CP is almost always adopted to 20

defend against corrosion. Nevertheless, depending on the nature of the defect and the 21

surrounding environment’s chemistry, cathodic current provided by CP may result in reaction 22

products that can affect adhesion of the coating around the defect and cause so called cathodic 23

3

disbondment (CD), which is considered to be the most significant degradation mechanism for 1

organic coatings on submerged steel [4]. Since the performance of a coating system directly 2

affects the integrity of the pipelines it protects, it is essential to evaluate coating performance in 3

conditions encountered during the installation and operational life of pipelines. As a 4

consequence, ex-situ CD tests as well as in-situ CD assessment were developed through the 5

years to assess coating performance before and after its service in the field. The ex-situ CD test 6

provides a judgement on the resistance of coatings to CD and thus acts as a practical tool to 7

select pipeline coatings. In-situ CD assessment offers a way to monitor a coating’s performance 8

non-destructively and is useful in aiding the maintenance of pipelines in the field, such as via the 9

adjustment of applied CP potentials. 10

In this review, fundamental aspects of CD are first introduced, focusing on a discussion of the 11

proposed mechanisms. Following this, ex-situ CD test methods including both the standard and 12

modified methods are examined. The influence and the significance of test parameters on the 13

disbondment rate is analyzed to reveal the advancement of the standard test methods. The need 14

for developing modified CD test methods as well as their contribution to coating evaluation is 15

discussed. The potential use of these ex-situ CD test methods to compare the disbondment 16

resistance of coatings, e.g., fusion bonded epoxy (FBE) and high-performance powder coating 17

(HPPC) is also presented. The third part of this review introduces various in-situ techniques to 18

monitor the CD behavior of coatings, compares the pros and cons of each technique and 19

addresses the development of two techniques in particular, i.e., electrochemical impedance 20

spectroscopy (EIS) and electrochemical measurement using a wire beam electrode (WBE). As 21

the mitigation of CD as well as ensuring pipeline integrity are the ultimate goals of any research 22

in this field, the final section discusses strategies that have been used to achieve these goals in 23

4

terms of coating selection and cathodic protection adjustment. This discussion further highlights 1

the importance and necessity of using the proper CD test for a given situation. 2

2. The mechanisms of CD 3

CD occurs on coated metals that are cathodically protected. CD refers to the failure of adhesion 4

at the coating/substrate interface, which is directly related to the application of CP [5, 6]. The 5

mechanistic study of CD reveals how the disbondment commences, shedding light on the 6

methods that may mitigate it. 7

CD is often initiated by the formation of defects due to accidental coating damage during 8

pipeline handling and installation or imperfect application, leading to excessive permeability of a 9

coating [7] to reactive species including water and dissolved oxygen etc. With the application of 10

CP, two cathodic reactions, i.e. (1) water reduction/hydrogen evolution and (2) oxygen 11

reduction, can take place at the defect and coating/substrate interface. It is the applied potential 12

that determines which reaction is dominant. A potential of about −1.1 V vs. Ag/AgCl is 13

considered to be a critical value [6]. When the applied potential is more negative than this, 14

hydrogen evolution dominates. While if the applied potential is between the free corroding 15

potential and −1.1 V vs. Ag/AgCl, oxygen reduction is dominant. Under the disbonded area, the 16

potential becomes less negative due to IR drop in the narrow gap, where oxygen reduction is 17

considered to be the main cathodic reaction [8]. 18

2𝐻!𝑂 + 2𝑒" → 2𝑂𝐻" + 𝐻!(1) 19

𝑂! + 2𝐻!𝑂 + 4𝑒" → 4𝑂𝐻"(2) 20

21

Table 1 Mechanisms proposed for CD of organic coatings 22

5

Mechanisms

Coating examples Controversial aspects

Dissolution/reduction

of the metal oxide layer

Polybutadiene, commercial

epoxy powder coating, high-

density polyethylene/reactive

ethylene terpolymer blend

coatings [9-12]

The removal of metal oxide can

be a consequence of CD instead

of a cause, since it is not observed

throughout the disbonded area

Degradation of the

coating polymer

Alkyd based coatings, epoxy

based and polybutadiene

coatings [12-16]

Instead of hydroxyl ions that

cause chemical attack of the

coating, intermediate free radicals

and peroxides generated from

oxygen reduction may constitute

an alternate explanation

Aqueous displacement

of the coating

Oleoresinous and polybutadiene

coatings [17]

This mechanism was only

reported on untreated steel

substrates in ammonium

hydroxide solution (pH 11.7)

Mechanical lifting of

the coating by

hydrogen gas

Polyester [18] Vigorous hydrogen evolution at

the holiday area was proposed to

inhibit [19] rather than facilitate

CD

1

6

Table 1 presents the mechanisms proposed for CD of organic coatings and Fig. 1 shows the 1

schematics for each mechanism. Various mechanisms that have been proposed and that are 2

associated with the generation of hydroxyl ions have been categorized and discussed in [20, 21]. 3

It is generally agreed that hydroxyl ion (through both cathodic reactions (1) and (2)) results in a 4

highly alkaline environment, which is a major factor leading to coating disbondment. 5

Nevertheless, the literature is not consistent about which mechanism(s) listed in Table 1 6

initiate(s) and govern(s) the process. Leidheiser et al. [9] proposed that the primary mechanism 7

for CD is the dissolution of oxides, which breaks the coating/substrate interfacial bond and the 8

high pH then leads to localized attack of the polymer through saponification or hydrolysis. On 9

the other hand, as suggested by Koehler [17], displacement of the coating with a high pH 10

aqueous film, which may be formed by the diffusion of water and electrolyte ions through the 11

coating to the coating/substrate interface initiates the disbondment process. Oxide dissolution 12

could be a consequence of exposure to the highly alkaline medium after disbondment rather than 13

a cause [20]. One interesting thing to note is that all of these alkaline environment-related 14

disbondment mechanisms are postulated only for coated metals that are cathodically polarized. It 15

was reported that without a cathodic potential, no disbondment was observed when immersing a 16

coated metal (with a defect) in a strong sodium hydroxide solution [22]. This observation 17

indicates that a negative open circuit potential at the steel surface (e.g. about − 0.6 V vs. 18

Ag/AgCl at the beginning of the CD test in 3 wt% sodium chloride solution) impedes transport 19

of hydroxyl ions from the solution to the coating/substrate interface. This would preclude attack 20

of the metal/coating interface by hydroxyl ions. However, through the cathodic reactions that 21

take place with the application of a cathodic potential, hydroxyl ions that are generated at the 22

interface may result in coating disbondment. This also implies that the loss of iron oxide at the 23

7

coating/substrate interface is primarily caused by an oxide reduction reaction, e.g., 𝐹𝑒!𝑂# +1

3𝐻!𝑂 + 2𝑒" → 2𝐹𝑒(𝑂𝐻)! + 2𝑂𝐻", which is driven by the application of a cathodic potential. 2

The work reviewed in the following subsections highlights the representative experimental 3

findings for supporting each mechanism proposed as well as contradictory views to each 4

mechanism shown in Table 1. 5

6

7

8

9

10

Fig. 1 Schematics of different CD mechanisms: (a) dissolution/reduction of the metal oxide layer 11

[9-12]; (b) degradation of the coating polymer [12-16]; (c) aqueous displacement of the coating 12

[17]; and (d) mechanical lifting of the coating by hydrogen gas [18]. 13

2.1 Dissolution/reduction of the metal oxide layer 14

This mode of failure states that the dissolution/reduction of the metal oxide layer in the alkaline 15

environment accounts for the loss of adhesion (Fig. 1 a). Through studying the CD of a 16

transparent polybutadiene coating on steel using an ellipsometric technique in [9] as per a 17

personal communication reported therein, a direct visual observation of the oxide dissolution as 18

indicated by the disappearance of the oxide’s straw color and appearance of metallic iron 19

suggests that the metal oxide layer is destroyed by the alkaline environment. Metal oxide 20

reduction at a powdered epoxy coating/steel interface was also examined by Watts and Castle 21

Coating

Oxide layer

Steel Holiday

Coating

Oxide layer

Steel Holiday

Coating

Oxide layer

Steel Holiday

H2 H2 Coating

Oxide layer

Steel Holiday

Na+, H2O

(a) (b)

(c) (d)

8

[10] using X-ray photoelectron spectroscopy (XPS) and was proposed as an important precursor 1

to CD. A silver-grey halo around the holiday was presented in all the test samples they studied 2

where, as suggested by compositional analysis, the metal oxides were reduced to metallic ions. It 3

is interesting to learn from their study that the dimensions of this halo are constant and are 4

independent of exposure time. As exposure time was prolonged, the disbonded area was 5

predominated by a darker area of thickened oxide. The adhesive failure was presumed to change 6

from one that occurs at the steel/oxide interface (through oxide dissolution) to one that occurs at 7

the oxide/polymer interface (through chemical attack of the coating). Fig. 2 illustrates the 8

possible CD process involving the changing of failure mechanisms. In a more recent study, the 9

iron oxide dissolution mechanism was also supported by Love et al. [11] when discussing the CD 10

behavior of high-density polyethylene/reactive ethylene terpolymer blend coatings using energy 11

dispersive X-ray spectroscopy. A high elemental iron concentration detected near the initial 12

defect area was also considered as evidence of iron oxide dissolution, accounting for the initial 13

stage of CD. 14

Clearly, metal oxide dissolution/reduction has been observed using different characterization 15

techniques and for different coating materials, however, since the high metallic iron 16

concentration area is only localized, and is not observed throughout the disbonded films, it is 17

arguable as to whether the removal of metal oxide initiates CD or occurs after CD due to an 18

attack of the highly alkaline environment. 19

2.2 Degradation of the coating polymer 20

Another proposed mechanism for CD is that the alkaline solution under the coating attacks the 21

coating polymer, resulting in a coating failure (Fig. 1 b). The evidence for this mechanism is 22

9

primarily from compositional analysis of failure surfaces between various coatings and steels 1

using XPS. For example, a study on the CD of three different epoxy-based coatings shows that 2

there is a predominance of polymer residues on the disbonded substrate surfaces relative to iron 3

oxide [13]. It was therefore suggested that the degradation of the polymer coating through 4

saponification or hydrolysis leads to cohesive failure of the coatings. Similar findings were 5

reported by Dickie et al. [14, 15] when studying the interfacial chemistry and adhesion between 6

steel and polybutadiene coatings and epoxy-based coatings. Even though surface analytical 7

methods like XPS are helpful in characterizing the composition of failure surfaces, limited 8

information has been obtained about the depth distribution of products in the disbondment area. 9

An alternative perspective to chemical attack of the coating by hydroxyl ions is that intermediate 10

oxygen reduction products, such as peroxide ions 𝐻𝑂!", superoxide radicals 𝑂!" and hydroxyl 11

radicals ∙𝑂𝐻, attack the coating and cause CD at the coating/steel interface. This mechanism was 12

concluded to occur through observation of the reduced CD rate after adding different free radical 13

scavengers to the epoxy coating [16]. On the other hand, as free radical scavengers also prevent 14

the formation of hydroxide, it was argued that they could prevent CD through, for example, 15

impeding metal oxide dissolution rather than coating degradation [4]. 16

2.3 Aqueous displacement of the coating 17

Other than the oxide dissolution and coating degradation mechanisms, interfacial separation or 18

the displacement of the coating by a high pH aqueous film at the coating/substrate interface was 19

also proposed as a root cause of CD by Koehler [17] when studying the CD of organic 20

oleoresinous and polybutadiene coatings on untreated steel substrates. As pointed out by 21

Koehler, the ammonium hydroxide solution used in the tests have an influence on the result. The 22

10

ammonia and water are both capable of diffusing directly through the organic coating and 1

resulting in uniform disbondment. Aqueous displacement of an organic layer from steel was also 2

suggested by Kendig [23] for initiating the disbondment process. It was found that when the 3

cathodically polarized steel has a negative zeta potential, it repels organic layers (a hydroxyl-4

terminated polybutadiene dissolved in xylene) with a zeta potential of the same polarity. The 5

author therefore recommended incorporation of a large molecular weight organic cation into the 6

coatings to stabilize them against aqueous displacement from a cathodically polarized steel. 7

2.4 Mechanical lifting of the coating by hydrogen gas 8

According to reaction (1), another product of the cathodic reaction is hydrogen gas. The role of 9

hydrogen formation on coating disbondment remains controversial. Some studies claim that 10

hydrogen gas develops at the coating/substrate interface and facilitates coating damage. For 11

example, Mahdavi et al. [18] investigated the effect of cathodic potentials (−0.76, −0.95, −1.1, 12

−1.2, and −1.4 V vs. Ag/AgCl) on the disbondment rate of a defective polyester coating that is 13

about 1000 µm in thickness. They found that at potentials less negative than −1.1 V, little or 14

slight disbondment was observed; when cathodic potentials reached −1.1 V where the hydrogen 15

evolution reaction became dominant, there was a significant increase in the disbondment rate. 16

Incidental development of elemental hydrogen was also reported as a cause of coating damage or 17

hydrogen embrittlement to the steel [24, 25]. 18

Other studies argued that hydrogen gas does not contribute to coating disbondment but rather 19

inhibits the disbonding process. To reveal the individual role of hydrogen and hydroxyl ion in 20

CD, Kamalanand et. al. [26] studied the CD behavior of coal tar-based coatings in four different 21

environments: acidic, neutral, alkaline and oxygenated alkaline. Their conclusion was that 22

11

hydrogen gas did not play any role in the disbondment process as no disbondment was observed 1

in acidic solution where a large quantity of hydrogen gas was generated throughout the test. 2

When compared with that in alkaline solution without sparging of oxygen, a larger disbondment 3

is observed in the alkaline environment saturated with oxygen (8.1 mg/L). According to equation 4

(2), an oxygenated solution can bring the reaction from concentration polarization to activation 5

polarization, which is expected to produce a higher amount of hydroxyl ions in the vicinity of 6

holiday, leading to a larger disbondment. In a CD study of three-layer polyethylene coated steel 7

pipe at an elevated temperature, Kamimura et al. [19] also found that vigorous hydrogen 8

evolution did not accelerate disbonding, but instead inhibited disbonding due to its role in 9

decreasing dissolved oxygen concentration and restricting the current flow under the coating, 10

both of which contribute to the lower disbondment. 11

It is unlikely that a universal mechanism to explain all CD processes exists, as they are affected 12

by many factors, such as the type of coating, the coating thickness, type and pretreatment of the 13

substrate, and the chemistry used in the CD test itself. A change in the controlling mechanism 14

would also be expected over time as the disbondment area grows and the penetrating potential at 15

the edge of the disbonded area declines [27]. To explain this, as shown in Fig. 2, the initial stage 16

(stage I) of a CD process would be primarily due to the dissolution/reduction of an iron oxide. 17

Once oxide dissolution/reduction cannot proceed due to the decreased cathodic potential (less 18

negative) along the crevice, other mechanisms, such as coating degradation may control the 19

following disbondment rate (stage II). 20

21

22

12

1

2

3

4

5

6

Fig. 2 Schematic of the cathodic disbondment process on coated steels immersed in a sodium 7

chloride solution with cathodic protection. 8

As discussed above, CD is primarily caused by cathodic reactions under the coating, the 9

disbondment rate is therefore expected to be correlated to the rate of cathodic reactions. As 10

observed through experiments [9], the disbondment rate is proportional to the oxygen solubility 11

in the electrolyte and the applied cathodic potential, both of which affect the oxygen reduction 12

rate. For a given coating, and with a fixed oxygen solubility and cathodic potential, the rate of 13

CD is also found to be proportional to the diffusion coefficient of electrolyte cations 𝐷and 14

time𝑡. The disbonded length 𝑥 was expressed in [16] via equation (3) and found to follow 15

parabolic kinetics. The transport of cations to the disbonded area is considered to be the rate-16

determining step for CD [28, 29]. In order to achieve electro-neutrality with OH- that is 17

generated from water/oxygen reduction reactions at the substrate/coating interface, electrolyte 18

cations like Na+ are transported to the disbonded area and NaOH is formed. Due to the high 19

solubility limit of NaOH (about 25 M at room temperature), the activity of water under the 20

disbonded area is largely reduced, causing a very large osmotic pressure across the coating and 21

promoting further penetration of water through the coating [30]. 22

I: Dissolution/reduction of iron oxide

𝑒. 𝑔.𝐹𝑒!𝑂" + 3𝐻!𝑂+ 2𝑒# → 2𝐹𝑒(𝑂𝐻)! + 2𝑂𝐻#

2𝐻!𝑂 + 2𝑒# → 2𝑂𝐻# +𝐻!

II: Chemical attack on the coating by oxygen reduction products

𝑒. 𝑔. 𝑂! + 2𝐻!𝑂+ 4𝑒# → 4𝑂𝐻#

Steel substrate

Oxide layer

Coating

Half holiday

H2O, O2

Na+

Disbonded area Unaffected area

I II OH- OH-

13

𝑥 = 2(𝐷 ∙ 𝑡)$.& (3) 1

As pointed out by Knudsen and Forsgren [4], diffusion would lead to transport of both 2

electrolyte cations and anions to the disbonding front due to a concentration gradient, while 3

migration due to a negative electric field would be selective to cations. According to the fact that 4

no chloride anions were observed under a disbonded coating, the transport of cations should 5

occur by migration due to the cathodic potential applied - instead of diffusion. A general power 6

law of disbondment kinetics was therefore suggested as shown in equation (4), where 𝑘 is a 7

constant. 8

𝑥 = 𝑘 ∙ 𝑡' (4) 9

Knudsen and Forsgren [4] suggested that if the disbonding follows parabolic kinetics, then 𝑎 10

should be 0.5, corresponding to the square root of time, and the disbondment rate should be 11

controlled by the transport of cations by migration. If oxygen reduction is under pure activation 12

control, rather than limited by the transport of cations, the disbonding rate is determined by the 13

electrochemical potential, then 𝑎 would be 1. By testing the CD rate of various commercial and 14

model epoxy coatings on steel, 𝑎 was found to lie between 0.5 and 1. Thus, according to 15

Knudsen and Forsgren, the disbondment rate is affected by both applied cathodic potential and 16

migration of cations [4]. However, even though equation (4) does not explicitly include a 17

diffusion coefficient, the constant 𝑘 should reflect the mobility of electrolyte cations, which 18

depends on the cation type (e.g. Na+, K+, Cs+) and the applied electric field. Thus, for a given 19

electrolyte, the migration of cations would be determined by the applied cathodic potential. That 20

is to say, the disbonding rate based on equation (4) should only be determined by one factor, i.e., 21

the applied cathodic potential. This is consistent with what has been found in experiments: when 22

14

the cathodic potential becomes more negative, larger disbondment is observed. On one hand, the 1

more negative potential facilitates the oxygen reduction at the disbondment front; on the other 2

hand, it results in a larger electric field which promotes the migration of cations to the 3

disbondment area. 4

3. Ex-situ evaluation of CD resistance of coatings 5

Once the coatings have cathodically disbonded, corrosive gases, water, and reactive species may 6

enter the disbonded area and cause pipeline corrosion [31-34]. It is therefore essential to evaluate 7

the CD resistance of coatings when selecting desirable protective coatings for pipelines. CD tests 8

are designed with relevant test parameters to investigate their effects on the rate of coating 9

disbondment and to provide a reliable assessment of the coating quality. This section discusses 10

both standard CD test methods that are widely used to provide the basis for coating comparison 11

as well as modified CD test methods that are being developed for selecting suitable coatings to 12

accommodate particular field conditions. 13

3.1 Standard CD test methods 14

Standard CD test methods specify the parameters and procedures for conducting the CD testing. 15

Typical CD test parameters include applied potential/voltage, temperature, test duration, 16

electrolyte composition and concentration, coating thickness and anode configuration. Various 17

combinations of the parameters and procedures generate different standards and specifications. 18

Currently, there are at least 23 CD test standards released from ASTM, CSA, ISO, EN, NF, AS, 19

ARCO China, and NACE etc., for rating the CD resistance of coatings all over the world. In 20

2007, Holub et al. [25] published a critical analysis of 22 CD test methods such as ASTM G8 21

[35], ASTM G42 [36], ASTM G80 [37], ASTM G95 [38], CSA Z-245 [39], ISO 15711 [40], AS 22

15

3862 [41] and NF A 49-711 [42] and discussed the factors that affect the coating disbondment 1

under cathodic protection. According to their investigation, anode isolation and test duration are 2

two major parameters that can differentiate the results among the standards. Anode isolation 3

refers to a separation of anode and cathodic sites by, for example, placing the anode into a glass 4

tube (with a frit at the bottom). This can prevent the dissolved anolyte gases, such as chlorine, 5

from migrating to the cathodic sites and forming hypochlorite, which can attack coatings [43]. 6

For example, by using standards AS3862 [41] and related AS/NZS 4352 [44], in which anode 7

isolation was applied, smaller disbondment (about 15-20%) was observed when comparing to 8

those without anode separation. While larger disbondment using ASTM G95 [38] and ISO 15711 9

[40] was found due to their longer test duration (90 days and 182 days respectively, compared to 10

28-30 days in other standards). It is difficult to correlate varying results and compare coating 11

performance from different standards. To address this problem, the NACE International 12

Technical Exchange Group TEG 349X was formed to investigate the difference in test 13

parameters selected in 8 international standards as well as their influences on the disbondment 14

results [6]. 13 test parameters were covered in this investigation and their impact on the 15

disbondment rate are presented in Table 2. In general, dissolved oxygen concentration, applied 16

potential, coating film thickness, hypochlorite formation, and test temperature and duration are 17

those factors of significant influence. Detailed studies and discussion of the effects of different 18

test parameters on disbondment can be found in [5, 25, 27, 45-47]. 19

20

21

22

16

1

Table 2 The dependence of CD rate on various test parameters [48] 2

Test parameters

Disbonding rate

1 Oxygen concentration in the electrolyte Increases with dissolved oxygen content

2 Electrolyte type and concentration Proportional to the molar conductivity of the

cations

3 Applied potential Increases with decreased (more negative)

potential

4 Dry film thickness Decreases with thicker film

5 Test duration Proportional to test time

6 Pre-treatment of the substrate and surface

profile

Decreases with increasing surface roughness

and with phosphoric acid wash pre-treatment

7 Effect of hypochlorite Hypochlorite formation attacks the coating

8 Effect of temperature Increases with temperature

9 Specimen configuration – Flat and curved No significant difference

10 Holiday size – 3 mm and 6 mm No significant difference

11 Holiday shape – cone or straight hole No significant difference

12 Specimen orientation No significant difference

13 Selection of reference electrode and its

calibration

N/A (High temperature and environmental

friendly Ag/AgCl reference electrode was

recommended)

3

17

Following the findings of TEG 349X, a universal CD test method, i.e., NACE TM0115 [43], was 1

developed by NACE Technical Committee TG470. Table 3 compares the main test parameters of 2

NACE TM0115 with that of ASTM G8 [35] and ASTM G95 [38], the two standards that are 3

most widely adopted in CD testing. As pointed out in [49], special enhancements of NACE 4

TM0115 standard include the use of Ag/AgCl reference electrode, with a check of its accuracy; 5

the description of the test temperature measurement procedure; and the development of an 6

inexpensive anode isolation method to eliminate the hypochlorite formation. This new standard 7

covers all the test parameters in the CD test and includes both the attached cell method (a 8

plastic/glass tube is centered over the drilled holiday and adhered to the coating sample surface 9

to hold test solutions) and salt bath method (a vessel is used to contain the test solution and to 10

accommodate a coated sample, a reference electrode and an anode) in the test setup. For 11

example, in the attached cell method, a specimen with 6 mm holiday is used to simulate a 12

coating defect in the field. To mimic the field cathodic protection and to accelerate the 13

disbondment process, a constant negative potential of −1.38 V vs. Ag/AgCl is applied to the 14

coated specimen by a power supply. The anode (platinum or other appropriate materials) are 15

inserted into a glass tube to isolate it from the cathode (holiday), avoiding the contact of anode 16

reaction products, i.e., chlorine gas (2𝐶𝑙" → 𝐶𝑙! + 2𝑒"), with the cathode reaction product, i.e., 17

hydroxyl ions (2𝐻!𝑂 + 2𝑒" → 2𝑂𝐻" + 𝐻!), and the formation of hypochlorite (𝐶𝑙! + 2𝑂𝐻" →18

𝐻!𝑂 + 𝐶𝑙" + 𝐶𝑙𝑂"). This anode isolation setup represents field conditions where the 19

cathodically protected pipelines are separated from the anode. After the CD test, the 20

disbondment is determined by first making four radial direction cuts through the drilled holiday 21

and then using a rigid pointed knife to lift the coating at the holiday [43]. Fig. 3 shows a 22

representative disbonded specimen with FBE coating. The disbonded area includes the bright 23

18

circle area (sliver grey halo), where the oxide is believed to be removed [25], and the dark brown 1

circle area. The CD length is calculated by equation (5). 2

Disbondment radius = (average of four disbondment diameters – holiday diameter) / 2 (5) 3

Table 3 Comparison among representative CD test methods [6] 4

Test

Parameter

NACE TM0115 ASTM G8 ASTM G95

Applied

potential

-1.38 ± 0.02 V vs.

Ag/AgCl; adjust daily

-1.45 to -1.55 V vs.

Cu/CuSO4

-3 V vs. Cu/CuSO4;

adjust min. twice weekly

Temperature Room temperature (RT)

or elevated (heating

method specified)

RT (21–25°C) RT (21–25°C)

Duration 28 days or as specified 30 days or as specified 90 days

Electrolyte

type and

concentration

Deionized/distilled

water with 3 wt% NaCl

Potable tap water with

1 wt% each NaCl,

Na2SO4, and Na2CO3;

maintain daily

Deionized/distilled water

with 3 wt% NaCl;

maintain daily

Anode type Platinum, mixed metal

oxide (MMO), or other

appropriate anode

which does not corrode

during test period

Magnesium anode,

impressed current

anode

Platinum wire 0.02 in

diameter

19

Disbondment diameter

Holiday diameter

Anode

isolation

Glass tube with glass

wool plug

None Immersion tube with

fritted disk

Reference

electrode

Ag/AgCl (saturated

KCl), check accuracy

Cu/CuSO4 or calomel Cu/CuSO4 or calomel

1

2

3

4

5

6

7

8

Fig. 3 Disbonded FBE coated specimen. 9

As an example of typical CD test results, NACE TM0115 was used to evaluate the performance 10

of two coating systems, FBE and HPPC, in our lab. The samples were provided by Shawcor Ltd. 11

FBE coating is currently the primary coating of choice for new transmission pipelines in North 12

America [50] at operating temperature lower than 60°C. For service above 60°C, HPPC has been 13

extensively used in Canada. It outperforms FBE for corrosion and mechanical protection. In the 14

CD tests at room temperature, the initial pH of the 3wt% NaCl electrolyte was about 5, while it 15

quickly increased to around 12 after 1 day of testing and remained thereafter throughout the 28-16

day test period (Fig. 4). This is a typical high alkalinity environment obtained in the various 17

standard CD tests. The HPPC samples, which are typically 2-2.5 times thicker than FBE, 18

exhibited smaller disbondment radius and showed better disbondment resistance behavior as 19

expected (Fig. 5 and Fig. 6). 20

Holiday diameter

Disbondment diameter

20

The standard CD tests are developed to replicate field conditions. However, it is not possible for 1

the standards to replicate all of the environments that coated pipelines encounter in the field. 2

When it comes to selecting a proper coating system in a particular environment, using a standard 3

CD test method may not be sufficient or reliable. This stimulates a need to generate modified CD 4

test methods, which may better predict coating performance in the field. 5

6

Fig. 4 Variation of electrolyte pH during the 28 days CD test at room temperature. 7

2

4

6

8

10

12

14

0 5 10 15 20 25 30

Ele

ctro

lyte

pH

Number of days

FBE HPPC

21

1

Fig. 5 Comparison of CD behavior of FBE and HPPC using NACE TM0115 standard at room 2

temperature. 3

4

Fig. 6 Representatives of disbonded (a) FBE; and (b) HPPC samples using NACE TM0115 5

standard at room temperature. 6

3.2 Modified CD test methods 7

A modified CD test is primarily based on a standard CD test method. However, it alters one or 8

more parameters from the standard test and implements specific and alternate procedures to 9

better simulate the field condition. 10

One example is the work of Al-Borno et. al. from Charter Coating Service (2000) Ltd. that was 11

published in 2010 [51]. At that time, few CD standards included tests applied at temperatures 12

(a) (b)

22

close to or above 100 °C, while there was an increasing use of pipelines and other vessels at high 1

temperatures. Motivated by such an urgent need, they modified the salt bath method used in 2

ASTM G42 [36] (Fig. 7), and the attached cell method specified in ASTM G95 [38] (Fig. 8), in 3

an effort to operate the tests at higher temperatures of 80 °C, 120 °C and 150 °C. In the modified 4

ASTM G42 method, with a thermostatically controlled heater, iron filings (with good 5

conductivity and low air content) were put inside the test pipe and facilitated the pipe to achieve 6

the required temperature. While in the modified ASTM G95 method, the sample was placed on a 7

sand bed and heated by a heat pad. In order to keep the electrolyte from evaporating and boiling 8

at temperatures up to 180 °C, a cooling jacket was further added into the modified ASTM G95 9

test setup, as displayed in Fig. 9. Using this setup, the authors studied the effect of oxygen levels 10

on the disbondment rate of a single layer epoxy pipeline coating with the sample temperature 11

maintained at 180 °C. During the tests, the oxygen levels were controlled by bubbling air or 12

oxygen through the electrolyte. The results they obtained were contrary to what is normally 13

observed, i.e., ~ 1 mm smaller disbondment was observed in their tests with higher oxygen levels 14

achieved in the electrolyte. The authors suggested that this could be because the lower oxygen 15

levels in the tests were already sufficient to accelerate the disbondment and the additional 16

oxygen did not significantly add to the acceleration, but instead affected the solution chemistry 17

that resulted in less disbondment (no further analysis was done to elucidate the reason). As 18

indicated from their disbondment figures, only one sample was tested at each condition (no error 19

bar presented), which is not regarded as sufficient [43] to ascertain the disbondment behavior of 20

the coating and could lead to the discrepancy regarding the effect of oxygen level on coating 21

disbondment. It is worth mentioning that this modified ASTM G95 method also had the merit of 22

controlling the electrolyte temperature and the sample temperature separately. For example, a 23

23

test was run using a sample temperature of 108 °C and an electrolyte temperature of 6 °C to 1

simulate a subsea pipeline. 2

The examination of modified high temperature CD test method mentioned above and other 3

similar studies and reviews such as [52-54] played an important role in the development of new 4

CD standard to meet the emerging need of performing high temperature CD tests. As pointed out 5

previously, the NACE TM0115 standard released in 2015 incorporated CD tests at accelerated 6

temperatures ≥ 95 °C for both the salt bath method and the attached cell method. For the latter 7

method, it was advised that a hot plate with a sand bath or a heating mat with a temperature 8

controller should be used for heating the sample. An immersed cooling coil or other adequate 9

device for onshore coatings should be applied to maintain the electrolyte at 95 °C to avoid 10

boiling [43]. 11

12

13

Fig. 7 Modified ASTM G42 apparatus using a large Teflon bath (adapted from [51]). 14

Teflon bath

Heater unit Mg anode

Cathode (sample pipe)

Test water (3% NaCl)

24

1

Fig. 8 Modified ASTM G95 apparatus with a heating system (adapted from [51]). 2

3

Fig. 9 Modified ASTM G95 apparatus with both heating and cooling systems (adapted from 4

[51]). 5

Another example of a modified CD test was conducted in our lab. This relates to the Trans 6

Mountain Expansion Project, and to the selection of adequate pipeline coatings for application in 7

areas of the Rockies in British Columbia, Canada, where the pipelines will encounter cold rough 8

terrain (Fig. 10 a) with corrosive and abrasive rocks (Fig. 10 b). In order to evaluate and compare 9

the performance of FBE and HPPC coatings in those rocky and acid-rock generating areas, a 10

modified CD test protocol based on the NACE TM0115 standard was developed. In this 11

modified CD test (Fig. 11), a propeller (overhead agitator) was introduced in the CD test setup, 12

Test water

Pt anode

Heat pad Cathode (sample)

Pyrex vessel

Sand

Coolant in Coolant out

Pt anode Cooling jacket Cathode connection

Sample Heated sand bath

25

which was used to suspend pyrite into the 3 wt% NaCl (aq) electrolyte and above the coating 1

surface. These powders impinge on the coated surface during testing and result in abrasion, 2

which might simulate the effect of soil movement during pipeline operational life. Furthermore, 3

the reaction of the pyrite with water and oxygen added the acidification effect of the leaching 4

minerals, in-situ. This new protocol improves the standard CD test method by combining them 5

with quantitative measurements of resistance to abrasion/impact and aggressive chemistry. The 6

feasibility of applying this new test protocol to evaluate pipeline coating performance was first 7

proved on the FBE coating system. For FBE coatings tested at room temperature (RT), decreased 8

coating impedance was observed after increasing stirring rate and reducing the pyrite 9

concentration. In addition, pyrite powder abrasion contributed to decreased coating impedance 10

and dry film thickness (coating degradation). A suitable overhead stirring condition was 11

identified as: adding 1 wt% pyrite into the 3 wt% NaCl solution, using a stirring rate of 1400 12

rpm. 13

14

Fig. 10 (a) Rocky terrain where pipelines are constructed; and (b) rocky areas with pyrite ores – 15

acid rock drainage. 16

(a) (b)

26

1

Fig. 11 A schematic of the modified NACE TM0115 CD test using a propeller to suspend pyrite 2

powder into the 3 % NaCl electrolyte. 3

Following the establishment of the modified CD test protocol, the performance of FBE and 4

HPPC was compared. The results provide insights into the selection of pipeline coatings in field 5

conditions. A thorough investigation of FBE and HPPC coatings with a defect of 3 mm diameter 6

at both RT and an elevated temperature of 65 °C was performed, which included the coating dry 7

film thickness, the morphology change of the coating surface before and after the tests, as well as 8

the disbondment measurement. In the modified CD tests, due to the addition of pyrite, the 9

electrolyte remains acidic at RT, while it becomes alkaline when the test temperature is 65 °C 10

(Fig. 12 a). As shown in Fig. 12 b, HPPC shows much less disbondment (3.93 mm) compared 11

with FBE (9.7 mm) in acidic solutions tested at RT; when the temperature increases to 65 °C 12

with the electrolyte being alkaline, HPPC shows a slightly better performance. Compared to the 13

standard test method, the modified CD test demonstrates the remarkably superior performance of 14

HPPC over FBE in an acidic environment. It is apparent that HPPC coatings show higher 15

resistance to the transport of water and ionic species to the coating/substrate interface, 16

27

contributing to less observed disbondment. As reported in [55], the non-polar polyethylene top 1

coat contributes to the improved resistance of HPPC to water and chemical penetration. In an 2

acidic environment, hydrogen evolution becomes severe and its mechanical damage to the 3

coating appear more remarkable on thinner FBE coatings. 4

5

6

7

Fig. 12 (a) Comparison of electrolyte pH; and (b) disbondment resistance between FBE and 8

HPPC coatings using the modified NACE TM0115 CD test. 9

0

2

4

6

8

10

12

0 5 10 15 20 25 30

pH

Number of days

FBE_65C FBE_RT HPPC_65C HPPC_RT

0

2

4

6

8

10

12

RT 65

Dis

bond

men

t rad

ius (

mm

)

Temperature (°C)

FBE HPPC

(b)

(a)

Electrolyte was acidic throughout

28 days

Electrolyte became alkaline

after 7 days

28

4. In-situ assessment of CD of coatings 1

CD test methods provide an efficient way for predicting the coating performance and therefore 2

offer a useful guidance for coating selection. On the other hand, it is important to note that the 3

present CD test methods are all destructive ex-situ laboratory based tests, thus have limitations in 4

accurately predicting the disbondment behavior of coatings in service. For field monitoring of 5

coating disbondment, the use of in-situ techniques (non-destructive methods) is preferred, as 6

these are able to detect and map the location, and make quantitative measurement of coating 7

heterogeneities, facilitating the identification of the source of failure. 8

Techniques that have been investigated for in-situ assessment of CD behavior of pipeline 9

coatings include electrochemical impedance spectroscopy (EIS), localized electrochemical 10

impedance spectroscopy (LEIS), scanning Kelvin probe (SKP), scanning vibrating electrode 11

(SVE), scanning acoustic microscopy (SAM) and electrochemical measurement using wire beam 12

electrode (WBE). A critical review on the working principles and limitations of most of these 13

techniques can be found in [56, 57]. Table 4 lists the pros and cons of each technique. Among 14

these methods, LEIS, SKP, SVE and SAM are considered as scanning techniques, i.e., a 15

probe/tip is scanned over the coating surface to collect information, usually 16

impedance/current/potential images or acoustic amplitude images as in the case of SAM. These 17

spatially resolved images reflect the coating condition and can be used for CD evaluation. SKP is 18

frequently adopted together with SAM or SVE for verifying the visual results [58-60]. It should 19

be noted that LEIS, SKP, SVE and SAM are all limited to thin coatings. For example, coating 20

thickness should not exceed 1000 µm in the LEIS technique; for applying SKP, SAM and SVE 21

methods, coating thickness is generally less than 100 µm. While the two other in-situ techniques, 22

i.e., EIS and WBE, are applicable to thick coatings over 1000 µm, making them desirable for 23

29

field applications [56]. EIS has been applied in monitoring CD for a long time while WBE is a 1

relatively new technique that has recently been investigated. The following section focuses on 2

discussing representative studies of these two techniques. 3

Table 4 In-situ techniques (non-destructive methods) for assessing CD 4

In-situ

technique

Pros Cons Lab or field

tested

EIS[61-66]

Informative on

electrochemical reactions

occurred at the coating/metal

interface

Extended CD area might not

be detected as the collected

EIS data is dominated by

defect area with low

impedance

Field tested

LEIS [67-

69]

Provides the localized

quantitative electrochemical

information over the coated

metal

Limitations in measuring the

CD of thick coatings under

CP (< 1000 µm)

Lab

SKP [70,

71]

Allows correlation of

potential mapping with

disbonded/intact coating

Lab

30

SVE [72,

73]

Permits correlation of anodic

and cathodic current

processes with visual features

on the substrate surface

Only applicable to thin

coatings (< 100 µm) and

limited to laboratory tests

Lab

SAM [74,

75]

Enables the visualization of

the degraded area of coatings

Only applicable to thin

coatings (< 100 µm)

Lab

WBE [76,

77]

Each wire or mini-electrode

in a WBE acts as an

individual sensor to measure

the electrochemical

parameters from local areas

of the electrode surface

Complexity in preparing the

WBE and may require new

WBE for each single test

Lab

1

For an EIS measurement, an equivalent circuit is established to interpret the impedance data, 2

revealing the changes of interfacial properties including the resistance and capacitance of coated 3

electrodes, which can be used for studying the degradation of organic coatings. For example, Fig. 4

13 shows a widely adopted equivalent circuit for analyzing the impedance behavior of coated 5

electrodes [78]. In this model, RΩ is the electrolyte resistance, Cc represents the coating 6

capacitance, Rpo is the pore resistance of the coating, Rp is the polarization resistance or the 7

charge transfer resistance and Cdl is the double-layer capacitance of the metal substrate beneath 8

31

the coating. Table 5 lists the correlation between changes in EIS parameters and the degradation 1

of organic coatings. 2

Table 5 Indication of coating degradation from the change of EIS parameters 3

EIS parameters Change of the parameter Coating degradation

Rpo: pore resistance

Decrease

Increase of disbondment area Ad

[30] Rp: polarization resistance or

the charge transfer resistance

Decrease

Cdl: double-layer capacitance

Increase Increase of disbondment area Ad

[30, 61]

fb: break-point frequency

Increase Linear increase of the

disbondment ratio D [30]

Low frequency resistance

(e.g. at 10 mHz)

Decrease to below a

threshold (e.g. 107 Ω-cm2

for epoxy polyamide-

coated steel)

Indication of coating failure [79]

Phase angle A sharp transition from

−90 ° (capacitive

Indication of porosity in the

coating [80]

32

behavior) to 0 ° (resistive

behavior)

fb = fb0D, where fb0 = 1/2πɛɛ0ρ and is determined by the dielectric constant of the coating ɛ and 1

the specific ionic resistance of the coating ρ; D = Ad/A, A is the total sample area [30]. 2

3

Fig. 13 An equivalent circuit proposed for degraded coated metals having pores. 4

In attempting to apply EIS to monitoring the performance of field coatings, in-situ 5

electrochemical sensors were developed to detect the early stages of coating degradation and 6

substrate corrosion [81-84]. More recently, Been et al. [85] developed a modified EIS, namely, 7

EISPlus probe that can be easily mounted on field pipes. Fig. 14 (a) depicts the working 8

principle of this technique. This EISPlus probe contains two electrodes: a ‘high’ electrode (HE) 9

consisting of a stainless steel cylinder which is screwed to a polished pipeline steel sample with 10

the same curvature as the coated specimen; and a ‘low’ electrode (LE) that is the pipeline steel. 11

The HE and LE are separated by the thickness of the coating and the impedance properties of the 12

coating is measured by a frequency response analyzer (FRA) together with a dielectric interface. 13

The use of the EISPlus probe to evaluate coating performance was validated by the consistent 14

impedance results obtained from it with that obtained from a conventional EIS technique. This 15

probe avoids the use of an electrochemical cell containing electrolyte solution and it can be put at 16

all positions around a coated pipeline. Fig. 14 (b) shows an example of mounting the probe to a 17

33

pipe in the field with a belt. An outer Plexiglass cylinder was used to provide a means to attach 1

the probe/cell to the pipe in the field. The authors suggested to use a magnet to replace the belt, 2

as it can easily adhere the probe to the pipe. The magnetic field does not have an effect on the 3

impedance response of the coating. Two high impedance coatings, FBE and a high performance 4

composite coating (HPCC), were investigated by this EISPlus probe. It was shown that this new 5

technique can monitor the change of coating capacitance and resistance, as well as the low 6

frequency impedance (at 0.1 Hz), possessing the capability to detect holidays and coating non-7

uniformities. In addition, by analyzing the low frequency impedance of the coated FBE and 8

HPCC samples at 0 years (lab samples) and 11 years (field samples), the authors showed the 9

ability of using this EISPlus probe to predict the service life time of these two coating systems. 10

As seen in Fig. 15, based on an excellent coating barrier criterion of 109 Ω∙cm2 at 0.1 Hz [86], the 11

FBE and HPCC field coatings would cross below the threshold at approximately 30 years and 50 12

years of service, respectively. This prediction was based on two assumptions: first, the lab and 13

field data can represent the initial impedance and that after 11 years, respectively; second, the 14

coating degradation is linear. It should be noted that the wide spread data, especially those 15

regarding the original HPPC coating, would introduce inaccuracy in the prediction. Furthermore, 16

a linear behavior may not be valid for predicting the coating degradation over time. As derived 17

by King et al. [86] in analyzing the degradation behavior of field FBE-coated pipelines (5-21 18

years in service) using EIS, the coating degradation rate was expressed as below. 19

− ()*+()*-/012314567891('451':$.;<=)(:

= 0.64𝑡"$.?@; (6) 20

According to equation (6), a high coating degradation rate was expected within the 1st year of 21

service (the ratio is above 1 per year), followed by a slower long-term degradation process (with 22

an average ratio of about 0.15 per year after 5 years’ service). 23

34

1

Fig. 14 (a) A schematic of electrode arrangement for EISPlus measurements; and (b) EISPlus 2

probe is mounted on a pipe with a belt. An electric connection to the pipe is made with a small 3

magnet in a spot where the coating has been removed (adapted from [85]). 4

5

Fig. 15 Impedance data for the (a) FBE; and (b) HPCC coated pipe at 0 years and 11 years 6

extrapolated into the future assuming linear behavior (adapted from [85]). 7

As a non-destructive field and laboratory technique, EIS may quantitatively predict pipe coating 8

life. However, when using EIS to measure the CD of coated pipe specifically, it does have 9

limitations in its sensitivity toward detection of the extended disbondment area. In a study 10

correlating EIS parameters with the disbonded area of coated steel, Mahdavi et al. [66] found 11

(a) (b)

(a) (b)

35

that the charge transfer resistance decreases as the disbonded area increases, but this correlation 1

was only observed for the first 15 days of the test. After that, the charge transfer resistance tends 2

to reach a plateau even though the disbonded area keeps increasing (Fig. 16). It was argued that 3

with the presence of a defect, the EIS data collected was mostly due to the bare metal at the 4

defect area instead of the disbonded metal surface under the coating, which makes EIS lose its 5

sensitivity in measuring the changes of electrochemical parameters under the disbonded coating. 6

This limitation is expected to be more significant with the increase of coating thickness. 7

8

Fig. 16 The changes of charge transfer resistance and disbonded area of coated steel under −1.4 9

V vs. Ag/AgCl (replotted using data from [66]). 10

An alternative technique that allows the in situ local measurement of electrochemical parameters 11

under the disbonded area of thick coatings is WBE. WBE was firstly developed By Tan [87] to 12

characterize the inhomogeneity of a coating and for avoiding the effect of such inhomogeneity 13

on electrochemical measurements. WBE usually consists of 100 or more electrochemically 14

integrated metallic wires that are embedded in an insulating polymer matrix. Each wire is 15

isolated from those adjacent and serves as a mini-electrode/sensor. After polishing and cleaning, 16

0

1

2

3

4

0

5

10

15

20

25

30

0 5 10 15 20 25 30 35

Dis

bond

ed a

rea

(cm

2 )

Cha

rge

trans

fer r

esis

tanc

e (o

hm)

Time (day)

Plateauing

36

the WBE surface is coated with the coating of interest. WBE applied to CD on pipe steel was 1

introduced by Le Thu et al. in 2005 [76]. To make WBE applicable to investigate the coating 2

disbondment under cathodic protection, the authors modified the WBE (Fig. 17) by firstly 3

coating each steel wire with the coating that was under study (a modified acrylic epoxy resin) in 4

order to form an insulating thin layer, and then inserting each wire in a hole drilled through a 5

steel sheet of the same steel grade. The WBE was generated by embedding the wire inserted steel 6

sheet into the same epoxy resin. Finally, the WBE surface was painted with the modified acrylic 7

epoxy coating after polishing and cleaning treatment. A defect was made at the center of the 8

coated WBE and the electrode was immersed in artificial seawater for up to 336 hrs. As shown in 9

Fig. 18, by considering a current density around ± 1 µA/cm2 as a criterion for a delaminated area, 10

the evolution of coating delamination with immersion time was reflected from the distribution 11

maps of the measured galvanic current. Following this, EIS measurement in addition to current 12

mapping was used in the WBE technique for monitoring and visualizing the coating 13

degradation/delamination and carbon steel corrosion beneath a defective coating [88, 89]. A 14

recent study [77] compared the use of EIS mapping and current mapping in WBE to detect 15

coating disbonded area at two CP potentials, i.e., −1.40 V and −0.95 V vs. Ag/AgCl. The 16

measurement of local impedance provided more accurate information in monitoring coating 17

disbondment. As pointed out by the authors, local impedance mapping can clearly indicate the 18

propagation of coating disbondment at both CP potentials, while current mapping lost its 19

sensitivity at −0.95 V vs. Ag/AgCl. The WBE technique offers a way to conduct in situ local 20

current and impedance measurements on coatings of various thickness under CP, making it 21

desirable in field applications. 22

37

1

Fig. 17 Schematic representation of the WBE [76]. 2

3

Fig. 18 Evolution of distribution maps of galvanic current with immersion time [76]. 4

38

As seen in Fig. 19, a differential aeration sensor (DAS) based on the WBE technique has been 1

designed for studying localized corrosion of pipe steel beneath disbonded coatings [90]. The 2

sensor consists of 100 wires that are closely packed but electrically isolated. A polymethyl 3

methacrylate (PMMA) cover with an incorporated rubber seal is used to simulate the disbonded 4

coating. A multiplexer is programmed to connect the selected electrode to the working electrode 5

2 (WE2) and the remaining 99 electrodes to the working electrode 1 (WE1) terminal. A Zero 6

Resistance Ammeter (ZRA) is used to measure local current flowing between the selected 7

electrode and the remaining 99 electrodes. A potentiostat is used to exert and control the CP. It 8

was suggested that this DAS has the potential to monitor localized corrosion under disbonded 9

coatings by correlating current density distribution with the corrosion damage observed at the 10

sensor electrode surface. The DAS sensor was also proposed for field adoption to monitor 11

pipeline coating performance [91]. As the reader will note from this discussion, most 12

investigations of in-situ techniques for monitoring coating CD are still conducted in the 13

laboratory. 14

15

(a) (b)

39

Fig. 19 (a) Schematic diagram of the differential aeration sensor (DAS); and (b) the experimental 1

configuration using a DAS to monitor the electrochemical processes under CP (adapted from 2

[90]). 3

5. Mitigation of CD and corrosion under the disbonded coating 4

For cathodically protected pipelines with coatings, an appropriate amount of CP is able to reduce 5

corrosion to less than 0.01 mm per year [92] and a good quality coating can decrease the current 6

required for protection by a factor of 1000 or greater [93]. However, excessive CP can lead to 7

coating CD, and insufficient CP cannot provide effective corrosion protection [94]. Therefore, 8

the coating resistance to disbondment, the applied CP level, as well as the compatibility of 9

coating and CP after coating failure are all critical factors in mitigating CD and the associated 10

corrosion problem. 11

5.1 Coating 12

The type of coating largely determines its resistance to CD. The development of coating 13

materials from the 1940s to the present day is illustrated in Fig. 20. From the first use of coal tar 14

coatings to the widely applied multilayered polyolefin-based coatings in the oil and gas industry 15

today, it can be seen that the modern pipeline coatings are not only required to provide external 16

protection against corrosion, but also to protect against mechanical/structural damage such as 17

that encountered during transportation and field installation [95]. With the assistance of CD test 18

methods, thick coatings like multilayered polyolefin-based coatings are also found to exhibit 19

superior resistance to CD, making them a desirable choice in the field. It is worth mentioning 20

that there have been many studies in the last 10 years to improve the anti-corrosion properties of 21

coatings and their CD resistance. One type of investigation focuses on pre-treatment of the metal 22

40

substrate with conversion coatings such as phosphate and rare earth metal salt films [96, 97]. The 1

deposition of a thin conversion film over the metal surface was claimed to not only enhance 2

adhesion between the applied coating (e.g. epoxy based coatings) and the substrate by increasing 3

the interface roughness, but also seemed to impede the transport of ions, water and oxygen to the 4

coating/substrate interface. Another type of study concentrates on incorporation of anti-corrosion 5

pigments into the organic coatings. From a work by William and McMurray [98], cation 6

exchanged bentonites were found to greatly improves the resistance of polyvinyl butyral-co-7

vinyl alcohol-co-vinyl acetate (PVB) to corrosion-driven CD. Cations like Ca2+ in the pigment 8

exchange with Na+ that migrated from the electrolyte to the coating/substrate interface, and then 9

precipitate as a Ca(OH)2 solid that is electrically non-conductive, leading to decreased electrolyte 10

conductivity under the disbonded film and reduced CD rate. Phosphate-based compounds have 11

also been proposed as promising anti-corrosion pigments to enhance CD resistance of organic 12

coatings [99-101]. These compounds have a different working principle from that of the cation 13

exchanged pigments, i.e., they restrict active electrochemical reaction zones at the disbonding 14

front by forming a protective layer at the coating/substrate interface. 15

41

1

Fig. 20 The development of coating materials used for pipelines (adapted from [95]). 2

5.2 CP level 3

Regarding the applied CP levels, there are codes and standards to set a maximum CP potential, 4

with the purpose of minimizing the rate of disbondment and most importantly of avoiding 5

damages to metal substrates including hydrogen embrittlement and stress corrosion cracking. An 6

accepted criterion in the industry is that polarized (instant off) potentials should not be more 7

negative than −1.0 to −1.05 V vs. Ag/AgCl. Depending on the type of pipe materials, the upper 8

limit of CP potential would also be adjusted to more positive values. For example, a titanium (Ti) 9

substrate would have a maximum polarization potential of −0.65 V vs. Ag/AgCl in order to 10

42

prevent the formation of hydrides [92]. Table 6 lists the limit of CP polarized potentials for 1

different substrates. 2

Table 6 Recommended minimum and maximum potentials for cathodic protection [92]. 3

Substrates Limit of polarized

potentials (vs. Ag/AgCl)

Problems to be prevented

Steels Min. −0.8 V

Max. −1.15 V

Stress corrosion cracking

Amphoteric materials (e.g,

aluminum)

Max. −1.15 V Accelerated corrosion

Prestressed concrete cylinder Max. −0.95 V Cracking of the steel reinforcing

Ti Max. −0.65 V Formation of hydrides

4

5.3 The effectiveness of CP under a disbonded coating 5

When disbondment occurs on coated pipelines, it is essential to evaluate the effectiveness of CP 6

penetration and its mitigation of corrosion. As shown in Fig. 21, at least two scenarios of disbonded 7

coatings should be considered. One is cathodic disbondment that occurs around a holiday, which is 8

the focus of this paper. Another is coating disbondment without the presence of a holiday, but 9

through, for example, water absorption caused blistering. 10

43

1

Fig. 21 Disbonded coatings with CP protection: (a) cathodic disbondment; and (b) coating blistering. 2

For disbonded coatings around a holiday, the local pH and oxygen concentration of the solution in 3

contact with the steel surface and the local steel potential were studied to reflect the effectiveness of 4

CP [33, 102-104]. A potential difference is always found to exist between the holiday area and inside 5

the crevice/disbonded area, which reduces the CP’s effectiveness. As revealed by Chen et al. [104], 6

due to the limited diffusion through the disbonded crevice, oxygen concentration inside the crevice 7

decreases with increased distance from the holiday. This leads to corrosion at the crevice area as it 8

serves as an iron oxidation zone where the holiday is the oxygen reduction zone. A more negative 9

potential is required to decrease the local potential of steel in the crevice, which is expected to 10

enhance the local solution alkalinity through water reduction and passivation of the steel surface. Yan 11

et al. [105] investigated the flow of CP current to the disbondment crevice on three types of coating 12

systems, i.e., a two-part epoxy coating, HPPC and FBE. CP potentials from −0.776 V to −1.126 V 13

vs. Ag/AgCl were applied. The calcareous deposits formed on top of the coating holiday were 14

expected to block the CP current. Very limited, and in some instances unsteady, CP current (several 15

to tens of μA) was detected, and no conclusions could be drawn through comparison of the CP 16

penetration for different coatings studied. The high pH environment generated by CP (e.g., pH 17

reached 12 at −1.126 V vs. Ag/AgCl) was also considered as the primary reason for corrosion 18

protection of the steel inside the disbondment. On the other hand, it should be noted that high 19

alkalinity resulted from high CP may promote cathodic disbondment through the mechanisms 20

(a) (b)

44

discussed in section 2. Furthermore, CP should not be too negative so as to avoid problems including 1

hydrogen embrittlement as mentioned above. Hydrogen gas formed at the more negative potentials 2

were also found to block the flow of current into the disbonded crevice, which may impair the 3

effectiveness of CP [106]. Through theoretical calculation, Schwenk [107] pointed out that over-4

protection has very small effect on the increase of the protection length of a pipe and therefore does 5

not bring any economic benefit. In order to achieve an effective adjustment of the CP level (neither 6

under-protection nor over-protection), obtaining in situ information such as the potential/current 7

distribution from the holiday area to the disbonded area is key. In this regard, in-situ techniques such 8

as WBE with EIS [108], that can reveal the evolution of local electrochemical properties of coated 9

pipe, offer a good approach to monitoring the potential/current distribution under the disbonded area; 10

thus assisting in the adjustment of the CP potential and the maintenance of pipelines in the field. 11

For disbonded coatings with no holiday, the CP effectiveness refers to whether the CP current can 12

penetrate through the coating and effectively polarize the steel under the disbonded coating. −0.8 V 13

vs. Ag/AgCl is the widely accepted CP criterion for a well coated pipes where reference electrode 14

position and soil IR drop do not play critical roles. When there is a need for considering soil voltage 15

drop and crevice voltage drop (e.g. disbonded coatings), a minimum CP of −1.45 V vs. Ag/AgCl was 16

recommended in soils with low resistivity (1 to 10 Ω∙m) and a CP of −4.95 V vs. Ag/AgCl was 17

suggested in higher soil resistivity (100 to 1000 Ω∙m) conditions [109]. Many studies have been done 18

on simulating disbonded coatings and their capabilities in conducting CP current, providing 19

implications for CP potential adjustment and coating evaluation. For example, Kuang et al. [110] 20

monitored the potential of steel under high density polyethylene (HDPE) and FBE coatings under 21

two CP potentials, i.e., −0.8 V and −0.95 V vs. Ag/AgCl, respectively. As shown in Fig. 22, a gap 22

between the intact (defect-free) coating and the steel substrate was created to simulate a disbondment 23

crevice. Both test chambers separated by the coating were filled with salt solutions. It was revealed 24

45

that the HDPE coating blocked cathodic protection current to the steel. The FBE coating allowed CP 1

to penetrate, protecting the steel from corrosion. By investigating the molecular structure of the 2

coatings before and after testing, the authors suggested that the non-polar structure of HDPE makes 3

the coating impermeable to water molecules, preventing CP current penetration. As for the FBE 4

coating, it was found that a more negative CP potential of −0.95 V vs. Ag/AgCl improved the CP 5

permeation. Latino et al. [111] questioned the accuracy of the instrumentation and methodology used 6

in the study of Kuang et al. by pointing out the inconsistency in results regarding the increased pH 7

measured and the net anodic current densities (which is supposed to cause local acidification). They 8

designed a side-by-side two-chamber cell set up (Fig. 23), where each chamber was separated by the 9

studied pipeline coating. A constant CP potential of −0.73 V vs. Ag/AgCl was applied between 10

anode and cathode, and the electron flow between the two chambers were correlated to hydroxyl ions 11

generated at the cathode, which was measured by a pH probe. One important improvement of their 12

experimental set up is that a high-sensitivity potentiostat as well as a Faraday cage were used to 13

capture the very small current (pA/cm2) involved. According to their study, both FBE and HPPC 14

coatings blocked the CP-driven ionic current. The CP blocking result for FBE contradicts the 15

common understanding in the industry that FBE is a non-CP shielding coating. The authors further 16

studied the ageing effect on CP shielding of FBE using the same set up and found that water uptake 17

associated with ageing (up to seven months) enabled the conduction of CP current through the FBE 18

films [112]. As revealed by several previous studies [113, 114], water uptake in FBE leads to 19

blistering of the coating and the further development of cracking, which allows CP current to 20

penetrate. Latino et al. [112] also conducted measurements for the unaged film with a pinhole defect 21

and revealed that CP current through it was at least four orders of magnitude higher than that 22

measured for aged samples. The failure mode as well as the defect formed on a given coating system 23

appears to be an essential factor in evaluating the effectiveness of CP, which requires detailed study 24

46

and review. However, it seems clear from the literature that disbonded FBE coatings are conductive 1

to CP (i.e. non-shielding) only when they have microscopic cracks or other defects that allow ionic 2

conduction. 3

4

Fig. 22 Experimental setup for measuring permeability of coatings to CP (adapted from [110]). 5

6

Fig. 23 Side-by-side two-chamber cell set up for measuring CP-driven ionic currents through 7

pipeline coatings (adapted from [111]). 8

Electrochemical station

Working electrode

Platinum sheet

Reference electrode Carbon rod

Micro pH meter

DC power supply Disbondment

Coating

Potentiostat

Reference electrode pH electrode

Inert electrode

Coating film

Anode-chamber Cathode-chamber

pH

47

6. Summary 1

Protective coatings and CP are used simultaneously on pipelines to prevent them from corrosion. 2

However, CD, i.e., the loss of adhesion of coatings on the metal substrate, appears as an adverse 3

side effect of the combination of coatings and CP. The application of ex-situ CD tests as well as 4

in-situ techniques to assess the performance of pipeline coatings plays a vital role in mitigating 5

the CD and helps to guarantee and maintain the pipeline integrity. Fig. 24 illustrates the effects 6

of coating and CP on CD and corrosion, as well as the importance of using ex-situ CD test 7

methods and in-situ techniques. The ex-situ CD test methods evaluates the resistance of coatings 8

to disbondment, providing a good reference for selecting desirable coatings for field application, 9

and the in-situ CD assessment offers a way to monitor the disbondment behavior of coatings in 10

the field and to assist the timely adjustment of CP potentials. 11

An overview of the advancement of standard CD test methods as well as examples of using 12

modified CD test methods to qualify the coating resistance to CD was given in this paper. The 13

modified CD test methods better replicate particular environment that pipeline coatings 14

encounter in the field and thus can provide more reliable evaluation on coating performance. In-15

situ techniques including EIS, various scanning techniques and WBE for assessing the field 16

performance of pipeline coatings were also reviewed. Efforts on developing in-situ 17

sensors/probes using EIS and WBE were addressed. An incorporation of EIS into WBE 18

technique seems as a promising approach to develop practical tools in monitoring the CD of 19

pipeline coatings in the field. 20

48

1

Fig. 24 Applying ex-situ CD tests and in-situ CD assessment to mitigate CD and related 2

corrosion of coated pipelines. 3

4

Acknowledgement 5

The authors acknowledge funding support from Natural Sciences and Engineering Research 6

Council (NSERC) of Canada [NSERC CRDPJ 503725-16]. The funding and in-kind support 7

from Shawcor Ltd. and Specialty Polymer Coatings Inc. [Industry portion - NSERC CRDPJ 8

503725-16] is also greatly appreciated. 9

10

CD

Protective coating

Cathodic protection

Corrosion of pipelines

Cause

Prevent

Provides a physical barrier to aggressive environment

Provides a negative potential to suppress metal oxidation

Ex-situ CD tests

In-situ CD assessment

Qualify and select proper coating system

Mitigate

Adjust the CP potential

Mitigate

Promote

Cathodic reaction products attack the metal oxides, coatings, and or coating/metal interface

49

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