*Article* **Effect of Sulfate Ions on Galvanized Post-Tensioned Steel Corrosion in Alkaline Solutions and the Interaction with Other Ions**

**Andrés Bonilla, Cristina Argiz, Amparo Moragues and Jaime C. Gálvez \***

Departamento de Ingeniería Civil, Construcción, E.T.S de Ingenieros de Caminos, Canales y Puertos, Universidad Politécnica de Madrid, c/Profesor Aranguren 3, 28040 Madrid, Spain; af.bonilla@alumnos.upm.es (A.B.); cg.argiz@upm.es (C.A.); amparo.moragues@upm.es (A.M.)

**\*** Correspondence: jaime.galvez@upm.es

**Abstract:** Zinc protection of galvanized steel is initially dissolved in alkaline solutions. However, a passive layer is formed over time which protects the steel from corrosion. The behavior of galvanized steel exposed to strong alkaline solutions (pH values of 12.7) with a fixed concentration of sulfate ions of 0.04 M is studied here. Electrochemical measurement techniques such as corrosion potential, linear polarization resistance and electrochemical impedance spectroscopy are used. Synergistic effects of sulfate ions are also studied together with other anions such as chloride Cl− or bicarbonate ion HCO3 <sup>−</sup> and with other cations such as calcium Ca2+, ammonium NH4 <sup>+</sup> and magnesium Mg2+. The presence of sulfate ions can also depassivate the steel, leading to a corrosion current density of 0.3 μA/cm2 at the end of the test. The presence of other ions in the solution increases this effect. The increase in corrosion current density caused by cations and anions corresponds to the following orders (greater to lesser influence): NH4 <sup>+</sup> > Ca2+ > Mg2+ and HCO3 − > Cl− > SO4 <sup>2</sup>−.

**Keywords:** corrosion current density; sulfate; galvanized steel; alkaline solutions; linear polarization resistance; electrochemical impedance spectroscopy

#### **1. Introduction**

The phenomenon of corrosion usually causes more severe damage to prestressed steel structures than to conventional reinforced concrete structures. Protection of prestressed galvanized steel wires is ensured by injecting alkaline grout into polyethylene ducts covering the strands. However, in areas not protected by these ducts, in deteriorated places, or in areas with insufficient grout, corrosion phenomena can occur. Accumulation in these areas of water contaminated with aggressive ions coming from the atmosphere, marine environments, with decomposition products of organic matter can cause corrosion of these wires and failures in the post-tensioned strands.

Galvanizing protects steel through two mechanisms. Firstly, it creates a physical barrier that isolates it and acts as a sacrificial anode. In addition, corrosion products create a second protective barrier. The behavior of zinc in alkaline media has already been considered in the literature [1–3]. Zinc in contact with the alkaline matrix of cement in its fresh state shows temporary chemical instability. High pH values of the aqueous phase inside concrete pores, usually above 12.5, cause zinc oxidation. The cathodic reaction is associated with water hydrolysis and generates hydrogen on the galvanized surface, according to Equation (1).

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

The possible transformation of molecular hydrogen into physically adsorbed atomic hydrogen, proposed by Riecke [4], increases the risk of hydrogen embrittlement in galvanized post-tensioned steel.

**Citation:** Bonilla, A.; Argiz, C.; Moragues, A.; Gálvez, J.C. Effect of Sulfate Ions on Galvanized Post-Tensioned Steel Corrosion in Alkaline Solutions and the Interaction with Other Ions. *Materials* **2022**, *15*, 3950. https://doi.org/10.3390/ ma15113950

Academic Editor: Vít Kˇrivý

Received: 22 April 2022 Accepted: 24 May 2022 Published: 1 June 2022

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**Copyright:** © 2022 by the authors. Licensee MDPI, Basel, Switzerland. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC BY) license (https:// creativecommons.org/licenses/by/ 4.0/).

A Pourbaix diagram of zinc indicates that, at a pH value around 12, it forms an insoluble oxide layer of ZnO/Zn(OH)2 more stable than the oxide layer formed at pH 13. With high alkalinities, Zn has an amphoteric behavior, forming soluble ions Zn(OH)3 − and Zn(OH)4 <sup>2</sup><sup>−</sup> [5]. Formation of Zn(OH)2 leads to hydrogen formation [6]:

$$\text{Zn} + 2\text{H}\_2\text{O} \rightarrow \text{Zn(OH)}\_2 + \text{H}\_{2(g)}\tag{2}$$

The risk of corrosion in an alkaline medium and in the presence of calcium ions can be limited due to the formation of a passive layer of calcium hydroxyzincate Ca(Zn(OH)3)2 · 2H2O, which is stable and protective. Some authors [7,8] have identified a value of pH 13.3 ± 0.1 as the limit for the passivation of galvanized steel. At a lower value of pH than 13.3, Ca(Zn(OH)3)2 · 2H2O crystals are small enough to form a thin, homogeneous and stable layer on the surface of the steel capable of keeping it passive. At a higher value of pH than 13.3 and when the calcium content is low, the size of the crystals increases, making it difficult to cover the entire surface of the galvanized steel. In this case, large, isolated Ca(Zn(OH)3)2 crystals that do not passivate galvanized steel are formed.

Other works [9] have studied the behavior of galvanized steel as a function of pH in the absence of Ca2+ ions. In the range of 12 < pH < 12.8, the galvanized layer dissolves at a slow speed. In the range of 12.8 < pH < 13.3, the galvanized layer is capable of being covered with a protective layer that insulates it. However, at pH > 12.8 ± 0.1, hydrogen release occurs. At a value of pH > 13.3, the galvanizing layer dissolves completely. It is worth noting that the role of sulfate ions in the corrosion of galvanized steel has been less studied. Acha [10] studied stress corrosion of prestressed steel immersed in saturated solutions of Ca(OH)2 with five concentrations of sulfate ions at various values of pH (0.01 M SO4 <sup>2</sup><sup>−</sup> at a pH of 12.1; 0.025 M SO4 <sup>2</sup><sup>−</sup> at a pH of 12.2; 0.05 M SO4 <sup>2</sup><sup>−</sup> at a pH of 12.4; 0.1 M and 0.2 M SO4 <sup>2</sup><sup>−</sup> at a pH of 12.85). Results showed a limiting sulfate concentration between 0.025 (pH = 12.2) and 0.05 (pH = 12.4). Above this limit, the steel surface presented severe localized corrosion, and below this concentration limit, the steel remained passive. Liu et al. [11] also showed that a sulfate concentration of between 0.02 and 0.03 mol/L, in a saturated solution of Ca(OH)2, produced steel corrosion.

Therefore, corrosion of prestressed steel in the presence of sulfates depends on the sulphate ion concentration in the solution and on the pH. Carsana and Bertolini [12] identified a pH dependence on the anodic behavior of steel in sulphate solutions for the corrosion of the steel. Acha's thesis also addressed the effect of bicarbonate ions (0.05 M concentration combined with pH = 11 and 0.1 M concentration with pH = 8.2) and of carbonate ions (CO3 <sup>2</sup>−) in saturated solutions of Ca(OH)2. In both cases, current density reaches values lower than 0.2 μA/cm<sup>2</sup> after 45 days. These ions do not cause corrosion problems in prestressed steel.

In alkaline media and in the presence of carbonates, the most common corrosion product is hydrozincite (Zn5(CO3)2(OH)6) [13,14], and in the presence of sulfates, zinc hydroxysulfate (Zn4(SO4)(OH)6 · 3H2O). After galvanized steel exposure to marine environments, the formation of a passive layer of hexagonal crystals of simonkolleite Zn5Cl2(OH)8 · H2O, and zincite ZnO has been identified as the main corrosion product. Simonkolleite is formed after hydrozincite, both being white crystalline compounds. Later, a more protective layer of gordaite NaZn4Cl(OH)6SO4 · 6H2O can be formed by incorporation of sulfate and sodium ions in the crystalline structure of simonkolleite. Soluble compounds such as ZnCl2 and ZnSO4 have also been identified in marine environments [13,15–19].

Xu et al. [20] studied the effect of cations from different sulfate salts (MgSO4, (NH4)2SO4, Na2SO4, CaSO4) added in a fixed concentration of 0.01 mol/l. The corrosion study was carried out in saturated calcium hydroxide solutions. Solutions of magnesium sulfate and ammonium sulfate showed higher rates of corrosion than solution with sodium sulfate. Solution pH was lowered with the addition of ammonium sulfate and magnesium sulfate. Neupane et al. [21] studied the effect of NH4 +, Na+, and Mg2+ cations on the corrosion of galvanized steel. Solutions of (NH4)2SO4, Na2SO4 and MgSO4 0.5 M were prepared in distilled water. The increase in corrosion current density caused by cations and anions

corresponds to the following order (greater to lesser influence): Na+ > NH4 <sup>+</sup> > Mg2+. Magnesium ions form finer, more compact and less porous corrosion products than the other salts.

There are numerous studies on the influence of potentially aggressive ions on the corrosion of galvanized steel. The results obtained show that the critical concentrations for each ion are often determined by the pH value of the solution. However, the effect that the joint presence of two or more potentially aggressive ions generates has not yet been determined.

In the present work, we studied the effect of sulfate ions on galvanized steel in alkaline solution and the synergistic effect of sulfate ions with various cations and anions found in seawater or marine environments (Ca2+, NH4 +, HCO3 <sup>−</sup>, Mg2+ and Cl−), or in the atmosphere, and their influence on the corrosion of the galvanized steel wires.

#### **2. Experimental Work**

*2.1. Solutions*

Concentrations of different ions were used, based on those obtained in real solutions produced by the action of rainwater together with degradation processes of living beings' waste. Table 1 indicates the ionic composition of six synthetic solutions prepared from the following salts: Na2SO4, Ca(OH)2, NH4COOH, NaHCO3, Mg(COOH)2 and NaCl. Solution pH was set at 12.7 by adjusting with NaOH 2M. The main component was the sulfate ion, followed by magnesium, chloride and ammonium. Ions were introduced incrementally in order to determine which ones were responsible for corrosion initiation.


**Table 1.** Composition of synthetic solutions prepared for corrosion tests.

#### *2.2. Corrosion Electrochemical Cells*

Electrochemical cells were made in polypropylene bottles using a three-electrode system (Figure 1). A Ag/AgCl electrode was used as a reference electrode. Stainless steel mesh was used as a counter electrode and galvanized wire (nominal diameter of 0.519 cm) as a working electrode. Galvanized wires were cleaned with alcohol. Adhesive tape was used for limiting an exposed attack area of 4891 cm2. To avoid carbonation, the solution was covered with a liquid paraffin layer.

**Figure 1.** Corrosion cell and Autolab PGSTAT 204 potentiostat/galvanostat assembly together with experimental connections.

#### *2.3. Techniques*

#### 2.3.1. Electrochemical Tests

Electrochemical tests were carried out with an Autolab PGSTAT 204 potentiostat/galvanostat from MetrohmAutolab BV®. NOVA 2.4.1 software with FRA 32 impedance module (Figure 1) was used. Linear polarization resistance (LPR) and electrochemical impedance spectroscopy (EIS) were also used. The LPR method was used to determine the instantaneous corrosion rate [22,23]. Measurement was carried out by applying a polarization scan from −20 mV to 20 mV around the open circuit potential (OCP) at a sweep rate of 0.1667 mV/s. Ohmic drop (RΩ) obtained by an electrochemical impedance technique is then subtracted from this resistance. Therefore, charge transfer resistance between zinc surface and solution (Rp) (Equation (3)) is calculated as:

$$\mathcal{R}\_{\mathsf{P}} = \mathsf{R}\_{\mathsf{P}\ (\mathsf{LPR})} - \mathsf{R}\_{\mathsf{O}\ (\mathsf{EIS})} \tag{3}$$

Corrosion current density (Icorr) was obtained from the polarization resistance (Rp) calculated as the slope of the polarization resistance curve around the corrosion potential according to the Stern and Geary relationship (Equation (4)) [24] with parameter B = 13 [8,25–27] and the procedure proposed in UNE 112072 standard [22].

$$\mathbf{I}\_{\text{corr}} = \mathbf{B} \cdot \frac{1}{\mathbf{R}\_{\text{p}} \cdot \mathbf{A}} \tag{4}$$

Impedance measurements (EIS) were carried out by potentiostatic control in a frequency range between 10 mHz and 100 Khz, taking 10 points per decade. Amplitude of the input AC voltage signal was ±10 Mv (rms). This technique consists of taking measurements by applying a small signal of alternating current and constant voltage to a working electrode, making frequency sweeps of the applied signal [25,26,28].

#### 2.3.2. Electron Microscopy SEM/EDS

At the end of the tests, the steel was extracted from the solution and dried in an oven at 40 ◦C for a week. Surface morphology was then observed using electron microscopy SEM/EDS. The attacked zone of the galvanized steel was observed by scanning electron microscope. A JEOL 6400 JSM microscope with EDS analysis was used with a resolution of 133 eV.

#### 2.3.3. Optical Microscopy

In addition, an OLYMPUS SZX7 optical microscope with an OLYMPUS SC50 camera was used to characterize and observe the surface of wires after exposure to the corresponding solutions.

#### **3. Results and Discussion**

#### *3.1. LPR Results*

Figure 2 shows polarization resistance curves resulting from LPR measurements at the end of the test. Curves show the cathodic and anodic branches and corrosion potential. All zinc wires showed their potentials with an intermediate corrosion probability (Ecorr > −332 mV) except for Solution 4, which produced somewhat higher values.

**Figure 2.** Polarization resistance curves of zinc wires in solution cells after 35 days of manufacture.

Table 2 shows values obtained for Ecorr, Rp and Icorr of all synthetic solutions made during the entire test period. Wires evolve from high corrosion risk to less electronegative potentials. Figure 3 shows evolution in time of Icorr and Ecorr of the wires.


**Table 2.** Electrochemical parameters obtained after 35 days of manufacture.

**Figure 3.** Evolution in time of (**a**) Icorr and (**b**) Ecorr of the wires.

Starting point results were very electronegative for all solutions (–1.4 V) and showed corrosion current densities around 100 μA/cm2. After 35 days of testing, the potential increased to less electronegative values, higher than −0.3 V (except for Solution 4). Corrosion current density decreased to 0.5 μA/cm2. These values of Ecorr were also recorded in other works. In particular, at a pH value of 13, Zn is actively dissolved, giving rise to a soluble phase of Zn(OH)4 <sup>2</sup><sup>−</sup> in the potential range between −1.35 to −1.45 VSCE [29]. The corrosion current density decreases until they reach values between 0.1 and 1 μA/cm2.

The solution containing only sulfate ions SO4 <sup>2</sup><sup>−</sup> (Solution 1) begins with active Icorr (97.24 <sup>μ</sup>A/cm2) and a fairly electronegative Ecorr (−1.4 V). Initial values recorded for Ecorr and Icorr indicate that the galvanized layer upon contact with the alkaline medium dissolves anodically, with consequent evolution of hydrogen on the galvanized surface [6,7]. After ten days of testing, these electrochemical parameters changed. The wire became covered with a passive layer and the hydrogen evolution process slowed down. Icorr values decreased and reached 0.28 μA/cm2 by the end of the test. This value is within the limits representing a low corrosion state (0.1 > Icorr < 0.5 μA/cm2), as has also been shown in other works [10,12,25] that a passive layer is formed over time. This layer is capable of reducing the initial corrosion current density. This passive layer could be Zn4(SO4)(OH)6 · 3H2O [1] in the absence of Ca2+ ions.

According to Acha [10] and Liu [11], 0.04 M sulfate ions in saturated Ca(OH)2 with a pH value of 12.4 should behave as depassivating ions. An increase of 0.3 units of pH in the presence of sulfates is not enough to passivate the steel. Vigneshwaran et al. [30] considered fixed amounts of sulfate ions of 2000 and 20,000 ppm (0.02 and 0.2 M respectively) to study carbon steel corrosion at pH values of 12.6 and 13.3, respectively. While steel corrodes at a pH of 12.6, at a pH of 13.3 the passive layer is not destabilized. Variation of just one-tenth in the pH value (12.7 vs. 12.6) changes the sulfate ion aggressiveness. With a sulfate ion, the steel is passivated. The corrosion ability of the sulfate ion depends on the pH and the type of steel.

When Ca2+ ions (Solution 2) were added to sulfate solution, the initial Icorr value of 47.26 μA/cm2 was half of the Icorr value from Solution 1 and was the lowest value of all solutions. This initial corrosion decreased over time. However, it did so more slowly than in the rest of the solutions. The wire did not reach even a moderate corrosion until the 30th day of testing. The Ca2+ concentration was 100 times lower than that of the sulfate solution. The initial concentration of dissolved sulfate decreased by half when precipitating as calcium sulfate. The passive layer on the surface of the steel in the presence of calcium could be different from the previous solution and represent a slower development. Some authors [7,8] identified a passive layer of calcium hydroxyzincate Ca(Zn(OH)3)2. They indicated a stability limit for this layer at a pH value of 13.3. Above this pH, the larger size of the crystals does not allow them to cover the entire surface of the steel. The Icorr value at the end of the test (0.53 μA/cm2) was within the limits that represent moderate corrosion (0.5 > Icorr < 1 μA/cm2). At the studied pH, the presence of Ca2+ coating turned out to be less protective.

The protection of the wire in the presence of Ca2+ ions was modified when NH4 + ammonium ions were added to the solution (Solution 3). In the beginning, the Icorr (92.45 μA/cm2) was at least twice that of Solution 2 (47.26 μA/cm2) and was very similar to the sulfate solution. H Pan et al. [31] reported than the corrosion of NH4 <sup>+</sup> could be attributed to dissolution of the MgO inner layer and the Mg(OH)2 outer passive layer. A similar process could take place in the case of ZnO and of Zn(OH)2, increasing the initial corrosion rate. A higher ionic charge increased the solubility of calcium sulfate. The wire began to be covered by a passive layer after fifteen days of testing. The Icorr obtained after 35 days reached 0.62 μA/cm2, within the limits that represent a state of moderate corrosion (0.5 > Icorr < 1 μA/cm2).

When bicarbonate was added (Solution 4), an initial Icorr value of 72.48 μA/cm2 was recorded. This was lower than the previous solution but higher than Solution 2. The solubility of zinc carbonate (1.4 · <sup>10</sup><sup>−</sup>11) is lower than that of calcium sulphate (≈9.1 · <sup>10</sup><sup>−</sup>6). Therefore, Zn ions could react with bicarbonate ions according to Equation (5) to produce corresponding carbonates with CO2 released [13], thereby removing sulfate ions from the solution.

$$\text{Zn}^{2+} + 2\text{HCO}\_3^- \leftrightarrow \text{ZnCO}\_3 + \text{H}\_2\text{O} + \text{CO}\_2 \tag{5}$$

At the end of the test, the Icorr (0.73 μA/cm2) was the highest of the six solutions and it was within the limits that represent moderate corrosion with low tendency values (0.5 > Icorr < 1 μA/cm2). The corrosion potential shifted to more negative values (−537 mV), which were the most electronegative of the six solutions. A passive film could be formed by hexagonal crystals of ZnCO3 and monoclinic crystals of hydrozincite Zn5(CO3)2(OH)6. This passive layer would be the least protective.

Subsequently, initial corrosion current density adding Mg2+ (Solution 5) was 102 μA/cm2. This was higher than previous solutions. After 15 days, the passive layer began to form. Icorr values after 35 days (0.19 μA/cm2) were within the limits of low corrosion state with a negligible tendency ( Icorr < 0.1 μA/cm2). This corrosion current density was the lowest of all solutions, which means that any passive layer formed under these conditions would be the most protective. These results agree with those found by Neupane et al. [21]. They compared the dissolution effects of Na2SO4, NH4SO4 and MgSO4 on galvanized steel. They found that the corrosion rate in the presence of NH4 <sup>+</sup> ions is higher than with Mg2+ ions. However, if the pH of the solution is not buffered, the addition of magnesium and ammonium sulfate, accompanied by a decrease in the initial pH of the solution, causes an increase in the corrosion rate, as shown in the study by Xu [20].

Lastly, once the behavior of the wire in chloride-free media was known, its behavior in a medium contaminated with Cl− (Solution 6) was studied. The initial value of Icorr was 119 μA/cm2, the highest of the six solutions, despite the fact that the concentration of this ion was lower than the sulfate ion solution. Icorr values at 35 days (0.42 μA/cm2) were higher than the corrosion of the sulfate solution. This was within the limits of low corrosion state with a negligible tendency. A passive layer of simonkolleite Zn5Cl2(OH)8 · H2O would thus be less protective than zinc hydroxysulfate (Zn4(SO4)(OH)6 · 3H2O.

#### *3.2. EIS Results*

As noted above, impedance measurement is useful to complete the Rp calculation in the linear polarization resistance (LPR) method. In addition, it enables the determination of resistance of the different parts that make up a system.

One of the key aspects of this technique as a tool to research the electrical and electrochemical properties of systems is the direct relationship between the real behavior of a system and that of a circuit made up of a set discrete component of electrical components, called an equivalent circuit. The most accurate circuit will be the one with the fewest possible time constants, which would provide a clear physical meaning [32]. Nyquist and Bode diagrams are obtained with their respective adjustments through the equivalent circuit to determine the values of each parameter that make up the system.

Figure 4 shows the Nyquist and Bode diagram obtained with the adjustment through the equivalent circuit for Solution 1. The equivalent circuit used in this study consisted of two constants connected in series with the resistance of the electrolyte (Figure 5). Elements of the circuit had the following physical meanings: Rs was related to the resistance of the electrolyte (solution). Upon the addition of bicarbonate ions (Solution 4), the Rs increased due to the greater presence of solid carbonate species in the solution; the first time constants, Rc and Cc, were attributed to the resistance of the passive film on the steel surface [33] and their capacitance. The second time constants, Rct and Cct, were related to the charge transfer resistance or mass transfer resistance. The latter was comparable to that obtained by the linear polarization resistance method (Rp). A constant phase element (*Q*) was used instead of a pure capacitance in the adjustment because of the heterogeneity of the layer on the surface of the wire [32,34]. From the Nyquist diagram, values can be seen corresponding to the high-frequency zone (1·105 Hz), given by the diameter of the first semicircle. It corresponds to Rc and Cc. Rct and Cct correspond to the diameter of the second semicircle in the low-frequency zone (0.01 Hz).

**Figure 5.** Equivalent circuit.

Figures 6 and 7 show the Nyquist diagrams with equivalent circuit adjustments resulting from study cases after 35 days. Parameters obtained from adjustments may be observed in Table 3. From the Nyquist diagrams, it is possible to confirm the beginning of the formation of a passive layer on the surface of the wire. It is also possible to observe the diameter of the second semicircle increasing.

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**Figure 6.** Nyquist diagram in "Low-Frequency Zone" (0.01 Hz).

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**Figure 7.** Nyquist diagram in "High-Frequency Zone" (1 · <sup>10</sup><sup>5</sup> Hz).



The equivalent circuit used seems to be the correct one because it has the fewest possible time constants with clear physical meaning. The deviation (*χ*2) in all cases is less than 0.03. Charge-transfer resistance on the wire surface (Rct) is then compared (Table 4) with the obtained Rp through linear polarization resistance to validate the results obtained through EIS. Equivalence is maintained in all cases except for Solution 4. The percentage difference between both methods is less than 14%. Although the difference with Solution 4 is 24%, both magnitudes are equal, which represents a state of medium corrosion in both cases (0.5 > Icorr < 1 μA/cm2). Differences between outcomes are valid and are mainly attributed to the fact that LPR uses direct current while the EIS technique uses alternating current (sinusoidal disturbance of electric potential) of variable frequency to the studied material). The order of passive layer stability is confirmed by the two electrochemical techniques (Solution 5 < Solution 1 < Solution 6 < Solution 2 > Solution 3 > Solution 4).



#### *3.3. Morphology of Steel Surface*

Figure 8 shows images of different steel surfaces obtained by optical microscopy and SEM after electrochemical analysis in synthetic solutions. No traces of iron oxides are observed on the surface of the wires. The Solution 1 wire is homogeneously coated with prismatic crystals, possibly of Zn4(SO4)(OH)6 · 3H2O. The Solution 3 wire shows a distributed oxide layer, leaving large voids on the surface. The size and coating of the crystals in the rest of the wires varies. This can be attributed to insoluble crystalline products such as Zn5(CO3)2(OH)6, ZnCO3, Zn5Cl2(OH)8 · H2O, ZnO, or others with the ability to passivate galvanized steel to a greater or lesser extent [15]. Some areas of exposure are darker in color due to increased detachment of the zinc layer and oxidation of the steel.

**Figure 8.** *Cont*.

**.** 

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**Figure 8.** Optical microscopy and SEM results of the steel surfaces.

Only one isolated white crystal of calcium hydroxyzincate Ca(Zn(OH)3)2 was found on the surface of the wire exposed to Solution 2. Due to the high pH of the solution, the crystal becomes larger and does not have the capacity to homogeneously coat the surface of the wire (Figure 9) [14]. By observing the SEM appearance of this crystal, it was possible to identify a totally different morphology from those observed on the surface of the other wires. Small crystallized threads can be observed.

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**Figure 9.** Optical microscopy and SEM crystal results Ca(Zn(OH)3)2 Solution 2.

#### **4. Conclusions**

Galvanized steel in contact with a strongly alkaline solution (pH = 12.7) and in the presence of sulfates dissolves anodically with corrosion potentials around −1.4 V and corrosion densities around 100 μA/cm2. Over time, the surface of the steel is covered with a protective layer, and the corrosion potential increases until values of −0.25 V are attained, with a corresponding decrease in corrosion current density to 0.3 μA/cm2, but greater than 0.1 μA/cm2 in all cases. Consequently, the presence of sulfate ions enables the depassivating of galvanized steel at highly alkaline levels of pH.

The presence of other anions and cations together with the sulfate ions keeps the corrosion process active. The nature of the passive layer depends on the ions present. Cations and anions studied here contribute to the increase in the corrosion current density of the sulfate ions. The magnitude of this increase follows the following orders: NH4 <sup>+</sup> > Ca2+ > Mg2+ and HCO3 − > Cl− > SO4 <sup>2</sup>−.

At a pH of 12.7, the hydroxyzincate crystal formed on the surface of the steel immersed in the solution of sulfate and calcium ions is large and occurs in isolation without the ability to cover the entire surface of the steel.

**Author Contributions:** Conceptualization, A.B. and A.M.; methodology, A.B., C.A., A.M. and J.C.G.; software, A.B.; validation, A.B., C.A. and A.M.; formal analysis, A.B. and C.A.; investigation, A.B. and C.A.; resources, A.B. and C.A.; data curation, C.A.; writing—original draft preparation, A.B. and C.A.; writing—review and editing, A.B., C.A., A.M. and J.C.G.; visualization, A.M.; supervision, A.M. and J.C.G.; project administration, A.M. and J.C.G. All authors have read and agreed to the published version of the manuscript.

**Funding:** This research was funded by the Ministry of Science and Innovation of Spain by means of the Research Fund Projects RTI2018-100962-B-100 and PID2019-108978RB-C31R.

**Institutional Review Board Statement:** Not applicable.

**Informed Consent Statement:** Not applicable.

**Data Availability Statement:** Not applicable.

**Conflicts of Interest:** The authors declare no conflict of interest.

#### **References**


## *Article* **Corrosion Damage to Joints of Lattice Towers Designed from Weathering Steels**

**Vít Kˇrivý 1, Zden ˇek Vašek 2, Miroslav Vacek 1,\* and Lucie Mynarzová <sup>1</sup>**


**Abstract:** The article dealt with the load-bearing capacity and durability of power line lattice towers designed from weathering steel. Attention was paid in particular to the bolted lap joints. The article evaluates the static and corrosion performance of bolted lap joints in long-term operating towers, and also presents and evaluates design measures that can be applied in the design of new lattice towers, or in the reconstruction of already operating structures. Power line lattice towers are the most extensive realization of weathering steel in the Czech Republic. On the basis of the inspections carried out to evaluate the working life of the transmission towers in operation, it can be stated that a sufficiently protective layer of corrosion products generally developed on the bearing elements of the transmission towers. However, the development of crevice corrosion at the bolted joints of the leg members is a significant problem. In this paper, the corrosion damage of bolted joints was evaluated considering two basic aspects: (1) the influence of crevice corrosion on the bearing capacity of the bolted joint was evaluated, using experimental testing and based on analytical and numerical calculations; (2) appropriate design measures applicable to the rehabilitation of developed crevice corrosion of in-service structures, or the elimination of crevice corrosion in newly designed lattice towers, was evaluated. Calculation analyses and destructive tests of bolted joints show that the development of corrosion products in the crevice does not have a significant effect on the bearing capacity of the joint, provided that there is no significant corrosion weakening of the structural elements, and bolts of class 8.8 or 10.9 are used. The results of the long-term experimental programme, and the experience from the rehabilitations carried out, show that, thanks to appropriate structural measures, specified in detail in the paper, the long-term reliable behaviour of the lattice towers structures is ensured.

**Keywords:** steel structures; crevice corrosion; lattice towers; bolted lap joints; weathering steel; experimental tests; numerical modelling

#### **1. Introduction**

Weathering steels have been used in the Czech Republic for steel structures since the mid-1970s. Bridges and electric power transmission systems are among the typical objects constructed from it [1]. Under appropriate atmospheric and design conditions, weathering steels form a layer of protective oxides on their surface that significantly slows the corrosion rate [2,3]. Weathering steels are used for structures with a design working life of up to 100 years, without additional corrosion protection measures [4]. These steels are mainly known by the trade name Corten [5], of which Atmofix steels are a similar variant [6]. For structures not affected by chloride deposition in coastal zones, two grades of weathering steels are practically applied—S355J2WP (Corten A) and S355J2W (Corten B).

With the correct application of weathering steels, it is possible to apply and utilise a number of significant advantages in the process of realization and long-term operation of structures made of these materials, in comparison with structures made of other grades of

**Citation:** Kˇrivý, V.; Vašek, Z.; Vacek, M.; Mynarzová, L. Corrosion Damage to Joints of Lattice Towers Designed from Weathering Steels. *Materials* **2022**, *15*, 3397. https:// doi.org/10.3390/ma15093397

Academic Editor: Mingchun Zhao

Received: 30 March 2022 Accepted: 6 May 2022 Published: 9 May 2022

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structural steels that need protection against corrosion in the long-term, using traditional corrosion protection systems. These potential advantages are summarised and expressed in the following basic terms:


Steel lattice towers are typically used as bearing structures in electric power transmission systems. These structures represent the most extensive realization of weathering steel in the Czech Republic. Between 1974 and 1992, approximately 4000 transmission towers, and 130 substations, of 110 kV, 220 kV, and 400 kV were built. Power line lattice towers are manufactured in various shapes, always in the form of four-sided lattice structures, and braced in the individual faces by bracing members made of simple angles. The leg members [14] of the angles are lapped along the height by bolted lap joints, and the bracing angles are connected to the leg members by a single bolt [15].

The corrosion behaviour of the weathering steels of transmission towers has been systematically evaluated since 1975 [16–18]. In general, the protective corrosion products develop favourably on the bearing elements of the transmission towers, even on surfaces not exposed to direct weathering.

However, for the use of weathering steels, the same details were originally applied as for zinc-galvanized or paint-protected transmission towers. No specific requirements were given for the operational control and maintenance of the transmission towers. Therefore, inadequately designed details lead to the unfavourable development of corrosion products [19,20]. These are mainly the following details of the lattice towers bearing structure:


Corrosion weakening of the bearing elements in the area of transition to the concrete foundation is a typical failure of all structures, not only of weathering steel transmission towers. There are known cases where the critical weakening of the structure in the area of transition to the concrete foundation was one of the main causes of the collapse of power line towers.

In the case of the bolted joint of the leg members, crevice corrosion [21,22] occurs in the crevice, between the splices and the angles to be connected [23], as illustrated in Figure 1. The time of surface wetting is prolonged in the crevice, impurities accumulate, and aeration and concentration differences are created, all of which favour the development of the corrosion process. The conditions necessary for the formation of a protective patina do not occur in the crevice. The accumulated corrosion products deform the splices in the connection.

**Figure 1.** Crevice corrosion in the bolted joint of the leg member.

Concern about the possible effects of adverse corrosion development on the bolted joints of leg members led transmission system operators in the Czech Republic to abandon the concept of using weathering steels in the construction of new power lines. In order to operate existing transmission powers responsibly, it is important to have verified technical data on the effect of crevice corrosion on the bearing capacity of the joints. In this paper, the results of loading tests of the leg members joints of two 110 kV power line transmission towers, taken from collapsed transmission towers, are presented. The main objective of the experimental programme, and subsequent computational analyses, was to verify the effect of crevice corrosion on the mechanical resistance of the transmission tower. The calculation analyses of bolted joints are performed using analytical equations, according to current standards [24,25]. For detailed numerical analyses of bolted joints using the finite element method, the recommendations given in technical publications [26–28] can be used.

High-voltage power lines with weathering steel transmission towers are still in operation. The transmission system operators carried out extensive rehabilitation work on most of the transmission towers, including the rehabilitation of the bolted joints in the lattice structure of the transmission towers. One of the objectives of the authors of the article was to provide concentrated information on the real technical condition of the operating transmission towers. The specific findings and recommendations are based primarily on the authors' long-term experience with the design, realization, and reliability assessment of these structures. The article deliberately provided mainly practically applicable information and recommendations for repair and maintenance.

In order to take advantage of the economic and environmental benefits of using weathering steels in the construction of lattice towers in the future, it is necessary to obtain demonstrable data confirming the functionality of the recommended design solutions for the bolted joints of leg members. For this purpose, a programme of experimental lattice towers was prepared, on which various bolted joints designs were tested in real conditions. After 9 years of the experimental programme, sufficiently reliable data are now available to confirm the suitability of the various design measures.

The issue of the corrosion behaviour of structures designed from weathering steels is very extensive, and this area received attention from researchers all over the world. Most of the scientific literature is devoted, in detail, to the study of the evolution of corrosion layers under different environmental conditions [29,30], or using accelerated tests [31,32]. The development of appropriate prediction models is important for the design of building structures [33]. This article, however, aimed at a somewhat different goal. The focus was mainly on the evaluation of the real influence of the development of corrosion products on the load-bearing capacity and durability of lattice truss towers with bolted joints, and on the design of appropriate structural measures. The presented results can be used by the scientific community, but also by experts from the construction industry, who are responsible for the design or operation of steel structures of power line transmission systems.

#### **2. Materials and Methods**

#### *2.1. Loading Tests of Bolted Joints*

The material for the test specimens was taken from the leg members of the collapsed power line transmission towers of 2 110 kV power lines *Neznášov—Týništˇe nad Orlicí* (overhead lines V1195 and V1196 in the Czech Republic). The transmission towers were made in the 1980s, using the weathering steel Atmofix made by Vítkovice Ironworks, Czech Republic. Eight test specimens were subsequently prepared from the collected leg members.

For the destructive tensile test of the joints, two pieces of the A-type joint, and six pieces of the K-type joint, were prepared. In all cases, L90/6 equal-leg angles were connected using P8 splices. The splices were connected to the profiles with 2 × 6 M20 bolts, of the class 5.6. The geometry of the test specimens is presented in Figures 2–4.

**Figure 2.** Detail of the A-type joint.

**Figure 3.** Detail of the K-type joint.

**Figure 4.** Details of the joint: an isometric view of the joints.

The chemical composition of the steel was determined using an optical emission spectrometer. Chemical analysis was carried out on two selected angles (test specimen A1 and K1), and one splice (test specimen K1). The chemical analysis confirms that the specimens taken from the collapsed lattice towers correspond to standard weathering steels, in accordance with the requirements of the standards [34,35]. The results are presented in Table 1. The mechanical properties of the steel were tested in the accredited laboratory of the Brno University of Technology [36]. The tensile testing, according to EN ISO 6891-1 [37], was carried out on eight flat test pieces. For the tensile testing, five pieces of leg members and three pieces of splices were made. The results are given in Table 2. The results of the chemical analysis and mechanical properties are compared with the original Czech national standard CSN 41 5217 [34], and the currently valid European standard EN 10025-5 [35]. The material meets the chemical and mechanical requirements for Atmofix steel (S355J2W).


**Table 1.** Loading tests of bolted joints: chemical composition of the test pieces.

**Table 2.** Loading tests of bolted joints: mechanical properties of test pieces.


During the tensile testing of the bolted joints, the relation between the tensile force *F* (kN) and the total deformation *u* (mm) of the test specimens was registered. At the same time, the gap opening (i.e., the change in the longitudinal distance between the ends of the connected angles) between the connected members was also measured. The uniform length of the test specimens between fixed ends is *L*<sup>0</sup> = 1010 mm, see Figure 5. The tensile testing was carried out until the specimen failed. Verification of the bearing capacity of the joint was performed using analytical relationships, in accordance with the valid European standards EN 1993-1-1 [24] and EN 1993-1-8 [25]. Numerical analysis was also carried out on the selected joint using ANSYS software. The joint was modelled using 3D solid elements and assuming the application of deformation load. The material properties of the steel and bolts were introduced using the Ramberg–Osgood stress–strain curve [38].

The microstructure of corrosion products in the crevice was analysed using a scanning electron microscope SEM and a EDAX analyser. The results are presented in Tables 3 and 4.

**Table 3.** Composition of the surface corrosion layer in the crevice (wt. %).


**Figure 5.** Test piece arrangement in the tensile testing machine.



#### *2.2. Assessment of the Technical Condition of the Transmission Towers in Operation*

The transmission towers that are part of very high-voltage power lines in the Czech Republic (overhead lines *V434 Slavˇetice—Ceb ˇ ín, V437/V438 Slavˇetice—Dürnrohr*) were selected for evaluation. The overhead lines were built in the 1980s, and have been in operation for approximately 40 years. In July 2021, an inspection of the steel bearing structure was carried out at 5 representative transmission towers, with a focus on the evaluation of the development of corrosion products in the area of the bolted joints of leg members.

The transmission towers are located in an agricultural area (corrosivity category C2). The structural design of the transmission towers can be seen from the photographs presented in Figure 6.

The transmission towers were visually inspected to evaluate the development of corrosion products, and the effectiveness of the implemented rehabilitation measures. The thickness of the corrosion products was measured on typical elements of the transmission towers structure, using the magnetic induction method with the Positector 6000 instrument. A total of 30 measurements were made on each of the evaluated surfaces. The actual thicknesses of the structural elements were continuously measured using a Positector UTG ME ultrasonic thickness gauge.

#### *2.3. Experimental Lattice Towers*

Experimental verification of the bolted joints in steel lattice structures designed from weathering steel was carried out in the premises of steelmaker Liberty Ostrava Inc. For this purpose, three experimental towers were built in 2012, on which the long-term monitoring of the development of corrosion products in the field of bolted joints with different structural design was carried out, as illustrated in Figure 7. This was a long-term experiment, the aim of which was to find a variant of the bolted joint of the leg members that minimized the formation of crevice corrosion in the long term, and at the same time be simple, inexpensive, and not complicate the assembly and maintenance of the structure.

**Figure 6.** The structural design of the evaluated transmission towers (Locality 1: Tasovice; Locality 2: Dyjákoviˇcky; Locality 3: Ivanˇcice; Locality 4: Rosice; Locality 5: Veverské Knínice).

**Figure 7.** View of the experimental towers' assembly at the time of installation.

During the experiment, the effectiveness of individual structural measures against the formation of crevice corrosion was continuously monitored. The design of the towers differed only in the joints of the leg members. In particular, the influence of the following factors was examined:


with a nominal thickness of *t* = 10 mm, and on tower 3 the splices were designed with a nominal thickness of *t* = 5 mm.


**Figure 8.** Application of silicone sealant ((**left**) application of sealant to the splices; (**right**) sealant smoothed around the perimeter of the splice).

**Figure 9.** Application of Teflon Vaseline to the splices.

All towers have a rectangular ground plan of 1200 × 1500 mm, and the height of the towers is 2000 mm. Each tower has two vertical faces, and two sloping faces inclined from the vertical plane at an angle of 10◦. The leg members are made of equal-leg angles L120 × 10, and the diagonals are made of equal-leg angles L50 × 5. Steel grade S355J2W was used. The angles used were produced on the rolling mills HCC and SJV of Liberty Ostrava Inc. The chemical composition of the leg members is listed in Table 5. The values of carbon equivalent, *CEV* = 0.42%, and atmospheric corrosion index, *I* = 6.3%, were calculated from the chemical composition [39,40].


**Table 5.** Chemical composition of the leg members.

The mechanical properties of the hot-rolled products are summarized in Table 6. The yield strength *R*eH exceeds the characteristic value of 355 MPa by 20%, and the tensile strength is in the required range of 470–630 MPa. The A5 elongation value is 35%, above the required limit of 22%. Although the melt was rolled in S355J2W grade, the impact strength KV with 2 mm V-notch at 0 ◦C, −20 ◦C, and −50 ◦C was also determined. According to the results obtained, very good notch toughness of the tested steel is evident, even at extreme negative temperatures.

**Table 6.** Leg members: the mechanical properties of hot-rolled products.


The diagonals of the towers were rolled on the rolling mill SJV of Liberty Ostrava Inc. The results of mechanical properties are listed in Table 7. The lower values of impact strength, compared to the L120 × 10 equal-leg angles, are due to the smaller non-standard tested thickness. However, the minimum value of 13.5 (J) is still met.

**Table 7.** Diagonals: the mechanical properties of hot-rolled products.


The splices were also rolled on the rolling mill SJV as P80 × 10 and P60 × 10 flat pieces, and L80 × 5 angle. The mechanical properties of the splices are provided in Table 8.


**Table 8.** Splices: the mechanical properties of hot-rolled products.

According to the analyses of chemical composition and mechanical properties, good compliance with the requirements of the EN 10025-5 standard for grade S355J2W can be stated. The experimental towers were built from material that meets the conditions of weathering steels of the required strength grade. All joints were designed with bolts. There are two bolted lap joints on each leg member. The towers are located in the premises of Liberty Ostrava Inc. (the area corresponds to the corrosivity category C3).

The surface of the angles and splices was left as rolled, including the mill scales. Only the coated splices were cleaned down to clean metal.

#### **3. Results**

*3.1. Loading Tests of Bolted Joints*

3.1.1. Calculation of the Resistance of the Bolted Connection

To compare with the results of the loading tests, the following section calculates the resistance of the bolted connection, according to the European standards EN 1993-1-1 and EN 1993-1-8.

The calculation is carried out for a K-type joint. Specimens for the loading tests are taken from transmission lattice towers made of Atmofix steel, which correspond to the current structural steel grade S355J2WP (yield strength *f* <sup>y</sup> = 355 MPa, ultimate strength *f* <sup>u</sup> = 510–680 MPa). The leg members of the L90 × 6 equal-leg angles are connected. In the connection, double-sided P8 × 90 splices are designed. The M20 non-preloaded bolts, of strength class 5.6, are used in the connection. The hole diameter for the bolts is *d*<sup>0</sup> = 22 mm. The bolt spacing and the end and edge distances are shown in Figure 3. The resistance of the following components of the connection is determined in sequence:


The design plastic resistance of the gross cross-section of the angle is determined, according to EN 1993-1-1, cl. 6.2.3(2a):

$$N\_{\rm Pl,Rd} = \frac{A \, f\_{\rm y}}{\gamma\_{\rm M0}} = \frac{1050 \cdot 355}{1.0} \cdot 10^{-3} = 372.8 \text{ kN} \tag{1}$$

The design ultimate resistance of the net angle cross-section at holes for fasteners is determined, according to EN 1993-1-1, cl. 6.2.3(2b):

$$N\_{\rm u,Rd} = \frac{0.9 \text{ Anet } f\_{\rm u}}{\gamma\_{\rm M2}} = \frac{0.9 \cdot 816 \cdot 510}{1.25} \cdot 10^{-3} = 299.6 \text{ kN} \tag{2}$$

The design plastic resistance of the gross cross-section of the splices is determined, according to EN 1993-1-1, cl. 6.2.3(2a):

$$N\_{\rm Pl,Rd} = \frac{A \, f\_{\rm y}}{\gamma\_{\rm M0}} = \frac{1440 \cdot 355}{1.0} \cdot 10^{-3} = 511.2 \text{ kN} \tag{3}$$

The design ultimate resistance of the net splice cross-section at holes for fasteners is determined, according to EN 1993-1-1, cl. 6.2.3(2b):

$$N\_{\rm u,Rd} = \frac{0.9 \, A\_{\rm ret} \, f\_{\rm u}}{\gamma\_{\rm M2}} = \frac{0.9 \cdot 1088 \cdot 510}{1.25} \cdot 10^{-3} = 399.5 \text{ kN} \tag{4}$$

The design block shear tearing resistance for a bolt group out of the angle is determined, according to EN 1993-1-8, cl. 3.10.2:

$$V\_{\text{eff,2,Rd}} = 2\left(\frac{0.5f\_{\text{u}}A\_{\text{nt}}}{\gamma\_{\text{M2}}} + \frac{\frac{f\_{\text{T}}}{\sqrt{3}}A\_{\text{RV}}}{\gamma\_{\text{M0}}}\right) = 2\left(\frac{0.5 \cdot 510 \cdot 204}{1.25} + \frac{\frac{355}{\sqrt{3}} \cdot 870}{1.0}\right) \cdot 10^{-3} = 439.9 \text{ kN} \tag{5}$$

The design shear resistance (shear plane passes through the unthreaded portion of the bolt) is determined, according to EN 1993-1-8, cl. 3.4.1, 3.6.1, and 3.7:

$$F\_{\rm v,Rd} = n \frac{\alpha\_{\rm v} f\_{\rm ub} A}{\gamma\_{\rm M2}} = 6 \cdot \frac{0.6 \cdot 500 \cdot 314}{1.25} \cdot 10^{-3} = 452.2 \text{ kN} \tag{6}$$

The design bearing resistance is determined, according to EN 1993-1-8, cl. 3.4.1, 3.6.1, and 3.7:

$$F\_{\rm b,Rd} = n \frac{k\_1 a f\_\rm u}{\gamma\_{\rm M2}} = 6 \cdot \frac{2.5 \cdot 0.606 \cdot 510 \cdot 20 \cdot 6}{1.25} \cdot 10^{-3} = 445.0 \text{ kN} \tag{7}$$

On the basis of a comparison of the resistances corresponding to the possible failure modes of the bolted joint of the leg member, it can be stated that the design ultimate tension resistance of the net angle cross-section at holes for fasteners determines the resistance of the joint *N*u,Rd = 299.6 kN.

3.1.2. Analytical Calculation of the Tensile Force in Bolts Due to Pressure of Corrosion Products in the Crevice

Bulky corrosion products develop in an imperfectly sealed crevice of a bolted joint. The corrosion products formed push the connected parts away from each other, so that the crevice opens up. The separation of the connected members is prevented by the bolts, where tensile forces are generated. The experiments carried out show that at the bolt location the gap opening is zero, and that with increasing distance from the bolts, the gap opening of the connection gradually increases. A permanent deformation of the connected members occurs.

A conservative estimation of the tensile force acting on the bolt *F*t,Ed can be determined on the basis of the assumption of plastic loading of the splice. For the calculation of the plastic bending resistance of the splice, the bolt spacing *L* = *p* = 80 mm, the splice width *b* = 90 mm, and the flange thickness *t* = 8 mm are considered. A simple beam model with uniform loading is assumed, and the tensile force in the bolt is equal to the reaction:

$$M\_{\rm pl,Rd} = \frac{W\_{\rm pl} f\_{\rm y}}{\gamma\_{\rm M0}} = M\_{\rm Ed} = \frac{1}{8} q\_z L^2 \tag{8}$$

$$F\_{\rm t,Ed} = q\_{\rm z}L = \frac{2 \,\text{b} \,\text{t}^2 f\_{\rm y}}{\gamma\_{\rm M0} L} = \frac{2 \cdot 90 \cdot 8^2 \cdot 355}{1.0 \cdot 80} \cdot 10^{-3} = 51.1 \text{ kN} \tag{9}$$

The bolts in the connection are assessed for combined shear and tension, according to EN 1993-1-8, cl. 3.6.1. The shear force into the bolt is determined from the resistance of the least bearing member of the connection, i.e., the design ultimate resistance of the net angle cross-section at holes for fasteners:

$$F\_{\rm t,Rd} = n \frac{k\_2 f\_{\rm ub} A\_s}{\gamma\_{\rm M2}} = \frac{0.9 \cdot 500 \cdot 245}{1.25} \cdot 10^{-3} = 88.2 \text{ kN} \tag{10}$$

$$\frac{F\_{\rm v,Ed}}{F\_{\rm t,Rd}} + \frac{F\_{\rm t,Ed}}{1.4 \, F\_{\rm t,Rd}} = \frac{49.9}{75.4} + \frac{51.1}{1.4 \cdot 88.2} = 0.662 + 0.414 = 1.076 > 1.0\tag{11}$$

The above assessment concludes that the combined action of extreme load effects, due to the axial stresses in the leg member, and the maximum tensile effects in the bolts, caused by the development of bulky corrosion products in the crevice, leads to bolts failure. This finding is valid for bolts of lower strength classes (4.6, 5.6) that were previously used in the design of transmission towers. Nowadays, bolts with strengths classes of 8.8 or 10.9 are commonly used, which minimizes the possibility of bolt failure in the connection under evaluation.

#### 3.1.3. Results of Loading Tests

The loading tests are carried out until the failure of the test specimens, which occurs either by rupture of the angle of leg member, or by shear of the bolts. Summary results for both 'A' and 'K' test specimens are shown in Table 9. The ultimate load at failure of the test specimen is in the interval 441 to 484 kN, the average value being 466.9 kN. The load at the beginning of the plastic deformation of the test specimen in all cases is close to 350 kN (visible reduction of the axial stiffness of the specimens, as shown Figure 10). Significant differences are observed in the deformation values of the individual test specimens. In the opinion of the authors of the paper, these differences are mainly due to the different values of slip in the individual joints, and probably also the different levels of friction in the joints. The failure of the test specimen occurs by rupture of net cross-sectional area at holes for fasteners, as illustrated Figure 11. In one case, there is a shear failure of the bolts.


**Table 9.** Summary results of tensile tests.

<sup>1</sup> *F*ma*<sup>x</sup>* ... the ultimate load at failure of the test specimen. <sup>2</sup> *u*max ... total deformation of the test specimen under load *F*max. <sup>3</sup> *F*el ... the elastic capacity of the test specimen. <sup>4</sup> *u*el ... total deformation of the test specimen under load *Fel*.

**Figure 10.** Tensile test results: the solid line shows the dependence of the tensile force *F* (kN) and the total deformation *u* (mm); the dashed line shows the dependence between the tensile force *F* (kN) and the gap opening between the connected angles.

**Figure 11.** Failure of the test specimens by rupture of the angles (test specimens K3 and K5).

#### 3.1.4. Results of Numerical Models

Numerical analysis is performed for the K-type joint in ANSYS software. Both cases of possible failure are considered, i.e., without the influence of the crevice corrosion, and then with the influence of the crevice corrosion on the joint taken into account. The models are loaded with deformation loads. In the lower part of the model, a support with zero displacements is placed on the L90 × 6 profile, and a linear displacement is applied to the upper cross-sectional area, in a direction parallel to the specimen under load. The magnitude of the displacement corresponds to the measured real magnitude (i.e., the average of the measured *u*max displacement values given in Table 9; i.e., 22.1 mm). The model mesh has a size of 2 mm (the major part is created using tetrahedron and hexahedron type of mesh; the minor part is created using the triangular prism and pyramid type of the mesh). Linear finite elements are used in the numerical model. The number of nodes in the model is 293,652. For contact between bodies, the *ANSYS contact tool* is used, with friction (a coefficient of friction 0.15).

For the K-type joint modelled without the effect of joint corrosion, the maximum von Mises equivalent stress of 497.7 MPa is obtained by numerical analysis. Figures 12 and 13 demonstrate the distribution of the von Mises equivalent stress, and the von Mises equivalent strain, and, therefore, the assumed location of the real failure is confirmed by destructive tests. The reaction force of the model is 474.0 kN, and the average value measured during destructive testing is 466.9 kN.

For a K-type joint modelled considering the effect of crevice corrosion, the von Mises equivalent stress of 499.9 MPa is obtained by numerical analysis. The consideration of the effect of crevice corrosion is based on a specific case of deformation of the splice of the K2 tested joint, where the measured maximum size of the crevice with corrosion products is equal to 10 mm. The model prescribes a 10 mm deformation of the vertical lines of the splices away from the L profile. The total elongation of the assembly is set, according to the K2 failure test, to 15 mm. The reaction force of the model is 486.7 kN, and the real measured force in the destructive test with bolt failure is 475 kN. Figures 14 and 15 below show, on the left, the results corresponding to the application of deformations from the crevice corrosion, and on the right, the results after the application of the subsequent load by the prescribed displacement.

**Figure 12.** The von Mises equivalent stress (MPa) in the K-type joint (model without the effect of the crevice corrosion).

**Figure 13.** The von Mises equivalent strain in the selected part of the K-type joint (model without the effect of the crevice corrosion).

**Figure 14.** The von Mises equivalent stress (MPa) in the K2 joint (model considering the effect of the crevice corrosion); left: results corresponding to the application of deformations from crevice corrosion; right: results after subsequent loading with the prescribed displacement.

**Figure 15.** The von Mises equivalent strain in the K2 joint bolts (model considering the effect of the crevice corrosion); left: results corresponding to the application of deformations from crevice corrosion; right: results after subsequent loading with the prescribed displacement.

These results show that the bolts are significantly stressed from the beginning of the development of corrosion products in the crevice. With subsequent static tensile loading, the stresses in the bolts increase further. The numerical analysis is in good agreement with the results of the analytical calculation presented in Section 3.1.2.

#### *3.2. Assessment of the Technical Condition of the Transmission Towers in Operation*

#### 3.2.1. Development of Corrosion Products on Leg Members and Diagonals

A protective adhesive compact layer of corrosion products, typical of directly wetted surfaces, develops on the surface of all the evaluated transmission towers. The average thickness of corrosion products is a suitable qualitative indicator of the favourable development of corrosion products on structures designed with weathering steels. The results of the thickness of corrosion products measurements (see Table 10) confirm the findings of the visual inspection. The thicknesses of the corrosion layers are similar on all the assessed transmission towers. The average corrosion thicknesses range from 136.3 μm to 197.7 μm, with a mean value of 161.0 μm. Thus, the criterion reported in the literature, [41,42] defining a maximum average thickness of corrosion products of 400 μm for sufficiently protective patinas, is reliably fulfilled. The value of the coefficient of variation ranges from 0.17 to 0.29, with a mean value of 0.22.

**Table 10.** Measured thicknesses of corrosion products.



**Table 10.** *Cont*.

#### 3.2.2. The Bolted Joints of the Leg Members

For all lattice towers situated at sites locality 1 to locality 5, as presented in Figure 6, the development of corrosion products in the crevice between the angle and the splices is identified. The bolted joints rehabilitation had not yet been performed on the transmission tower placed at locality 5. In the area of the bolted joint of the leg member, a typical development of corrosion products in the crevice between the splices and the leg member is observed, associated with the development of permanent plastic deformation of the splices, as illustrated in Figure 16. The thickness of corrosion products in the crevice reaches up to 10 mm.

**Figure 16.** Crevice corrosion at the bolted joint of leg members (Locality 5).

For the other transmission towers located at sites locality 1 to locality 4, the rehabilitation of the bolted joints was already carried out. The rehabilitation of the transmission towers was undertaken approximately 10 years ago. The repair of the transmission towers is carried out by workers wearing full body harnesses, and who are authorised to work at heights. In these difficult working conditions, it is very complicated to comply flawlessly with all the requirements recommended for the rehabilitation of the bolted joints of the leg members. For the transmission towers evaluated, failures of the additionally applied coating system are more frequently identified, especially on surfaces that the primer was applied to without prior proper surface cleaning (refer to Figure 17). However, from the point of view of the long-term reliable functioning of the structure, these failures of the coating system do not represent a significant problem, as the original protective layer of corrosion products remain under the peeling coating.

**Figure 17.** Failure of a paint system applied to a poorly prepared surface (the original protective layer of corrosion products remains under the peeling paint).

The protection of the crevice by the sealant is still functional, and no significant failures are identified on the transmission towers evaluated (see Figure 18). Thus, the bolted joint of the leg members is protected from further development of crevice corrosion, even after about 10 years following the application of the sealant.

**Figure 18.** Functional crevice protection with sealant (selected examples).

#### *3.3. Experimental Lattice Towers*

#### 3.3.1. Development of Corrosion Products on Leg Members and Diagonals

The development of corrosion products on leg members and diagonals is monitored at regular intervals. The increase in the average corrosion thicknesses is noted in Figure 19. It is expected that the average thickness of the corrosion products will continue to increase gradually in the coming years (for the long-standing transmission towers in operation, an average thickness of corrosion product of 161 μm is found, see Chapter 3.2.1). During the manufacture of the towers, the surface of the steel elements is deliberately left as rolled, including the mill scales. The thin surface mill scale has mostly fallen off after 9 years of exposure. The rolled thicker scales still remain on some surfaces, but their surface is gradually diminishing.

Statically significant corrosion weakening is not yet observed; the differences in the thickness of the flanges at the time of installation of the lattice towers and after 9 years of exposure are minimal. As the experimental lattice tower project is designed to be long-term (assumed to be at least 25 years), detailed analyses of corrosion products sampled from the crevice of the de-installed bolted joints are not yet available.

**Figure 19.** Experimental towers: development of thickness of corrosion products over time.

3.3.2. The Bolted Joints of the Leg Members

No phenomena related to the formation of crevice corrosion are observed in tower 1, which is designed with 10 mm thick splices, and with minimum bolt spacing and end/edge distances. No deterioration of the sealant is observed in the connections protected by sealant, as shown in Figure 20. For the other connection types, no mechanical damage to the splices is observed so far, although a lighter strip of corrosion products is visible on the top edge of the connection. No differences are observed between joints with preloaded and non-preloaded bolts.

**Figure 20.** Tower 1: Selected examples of joints (functional sealant along the edge of the splice).

For tower 2, which is designed with 10 mm thick splices, and with normal bolt spacings and end/edge distances, phenomena related to the formation of corrosion products in the crevice between the bolt and the leg member are already observable. For bolted joints without treatment, the initial development of corrosion products in the crevice are observed (the thickness of the corrosion products at the top edge of the splice is still very small: up to about 0.5 mm, as demonstrated in Figure 21). For the other connections, the development of crevice corrosion is not yet identified, although a lighter strip of corrosion products is visible at the top edge, especially for the joints with coatings. All joints protected with sealant are functional. No significant differences between preloaded and non-preloaded bolts are observed. There is a visual influence on the development of corrosion products around the joints protected with Vaseline (this is only a visual failure, with a slowing of the development of corrosion products).

**Figure 21.** Tower 2: Selected examples of joints ((**left**): initial development of corrosion products in the crevice; (**right**): visual influence of the development of corrosion products in the area around the joints protected with Vaseline).

In tower 3, which is designed with 5 mm thick splices, and with normal bolt spacings and end/edge distances, development of crevice corrosion occurs at some bolted joints, as illustrated in Figure 22. For untreated bolted joints, corrosion products with a thickness of up to 3 mm are found at the top edge of the splice (the thickness of corrosion products is slightly less for preloaded bolted joints, approx. 2 mm).

**Figure 22.** Tower 3: Selected examples of joints (the development of corrosion products in the crevice of joints without treatment, and in joints with coating on the inner surface of the splices).

Corrosion products with a thickness of 1 mm are also found on joints (both nonpreloaded and preloaded bolted joints) protected by coating system. The joints protected with sealant are functional, as indicated in Figure 23.

**Figure 23.** Tower 3: Selected examples of joints (functional sealant along the edge of the splice).

#### **4. Discussion**

The environmental and economic benefits of using weathering steels for power line constructions can only be realised if the necessary technical data on the long-term behaviour of the structural elements, and their connections, are available. Therefore, in this paper, attention was paid to the issue of long-term reliable functioning of bolted joints of leg members, in relation to the possible development of corrosion products in the crevice between angles and splices. This paper examined the bolted joints of leg members from two basic aspects:


#### *4.1. Verification of the Static Resistance of Bolted Joints*

The tension resistance of the bolted joint of the leg members was verified experimentally, on undamaged specimens taken from the collapsed transmission towers (the cause of the collapse of the towers was corrosion damage in the area of anchoring the leg members to the foundation). Analytical and numerical calculations of the resistance of the selected connection were performed for comparison. The performed calculations are in good agreement with the results of the loading tests, where in most cases the failure of the test specimen occurred due to tensile damage of the net angle cross-section at holes for fasteners. In only one case was the shear of the bolts responsible for the failure of the specimen (in this case, the increased tensile stress on the bolts from the pressure of bulky corrosion products in the crevice may have contributed to the failure mode). For all tested specimens, the value of the elastic capacity of the joint *F*el ∼= 350 kN is higher than the resistance calculated, according to the applicable standards *F*T,Rd = *N*u,Rd = 299.6 kN.

On the dismantled joints used for testing, it is observed that the thickness of the corrosion layer is greatest around the perimeter of the splice, decreasing towards the bolts. There is no significant corrosion weakening of the bolts. Thus, permanent deformation of the splices does not have a significant effect on the shear resistance of the bolts. This observation is consistent with the results of earlier studies carried out on bolted lap joints made of weathering steel [17,43].

The calculations presented in Section 3.1.2, as well as the numerical analyses presented in Section 3.1.4, indicate a possible negative effect related to the increase in tensile stresses on the bolts, caused by the development of corrosion products in the crevice. This process represents the direct effect of corrosion development in the crevice on the static load capacity of the structural member. When the extreme load effects from the axial stresses in the leg member are combined with the maximum tensile effects in the bolts, caused by the development of bulky corrosion products in the crevice, failure of the bolts may occur, provided that the bolts are designed using lower grades steels. This conclusion follows from the quantification of Equations (8)–(11).

The calculation analyses given in Sections 3.1.1, 3.1.2 and 3.1.4 do not include possible corrosion weakening of the cross-sections of the individual components of the joint. It is assumed that possible corrosion weakening may have the following consequences for the resistance of the connection:


The results of the experimental and computational analyses of the bolted joints of the leg members correspond to the experience with the operation of the transmission towers. The cause of accidents of lattice towers designed from weathering steels in the Czech Republic is usually the loss of stability of the compressed elements of the lattice structure subjected to increased bending stresses (e.g., from the effects of wind loading), or significant corrosion damage in the area of the anchorage of the transmission tower to the concrete foundation. Adverse microclimatic conditions can occur in the anchorage area: the structure is affected by surrounding vegetation; structural elements are often permanently covered with soil; and frequent degradation of the foundation structures also contributes to increased wetting. In the case of transmission towers designed from weathering steel, a continuous cycle of wetting and drying of the surface is not ensured in locally affected areas and, therefore, a protective layer of corrosion products does not develop on the structural elements.

This results in corrosion weakening the bearing elements in the anchorage area. Selected typical examples of corrosion damage in the anchorage area of transmission towers are provided in Figure 24.

#### *4.2. Design Solutions for New Transmission Towers*

The experimental towers located on the premises of Liberty Ostrava Inc. are planned as a long-term corrosion test, which focuses primarily on the verification of design measures to eliminate the development of corrosion products in the crevice at the bolted joints of the leg members. Based on the results obtained after 9 years of exposure, the functionality of the individual design measures can already be reasonably evaluated.

The basic design measure is the design of sufficiently rigid splices, together with the use of minimum recommended bolt spacing and end/edge distances. Both of the above principles are applied to tower 1, where corrosion products at the crevice in any of the bolted joints evaluated have not yet been observed to develop. The effectiveness of the proposed measure is based on a basic static assumption. By increasing the thickness of the splices, and reducing the spacing and end/edge distances, a component with higher bending stiffness is installed in the bolted joint, which is able to resist transverse pressures from corrosion products arising in the crevice.

**Figure 24.** Selected examples of corrosion damage in the anchorage area of transmission towers.

The application of silicone sealant seems to be an appropriate measure. The principle of the method is to seal the edge between the splice and the angle against the direct effects of the atmosphere. The application of silicone sealant on the splice takes a maximum of 30 s, and the sealant is cheap. After placing the splice into the structure, and the subsequent tightening of the bolts, the sealant is pushed beyond the edges of the splice for both nonpreloaded and preloaded joints, and the sealant is easily smoothed around the perimeter of the splice. After 9 years of exposure, all the sealant-coated joints are functional, and the sealant intact, with no signs of cracking or other degradation. The suitability of this design measure was also verified for operational transmission towers, where crevice sealing was one of the partial steps in the rehabilitation of bolted joints affected by the development of crevice corrosion (see Section 3.2.2). The use of sealants for the renovation of bolted joints is also recommended in the guidelines for the use of steels with increased resistance to atmospheric corrosion [42].

Based on the results obtained after 9 years of exposure of the towers, it is not possible to clearly evaluate the beneficial effect of bolt preloading. However, the results obtained from tower 3 suggest that bolt preloading in the connection reduces the risk of the development of corrosion products in the crevice. The beneficial effect of preloading is documented by mathematical numerical models [26,27]. The real effect of bolt preloading in weathering steel lattice towers is also planned to be monitored in detail in the remaining years of the experimental measurement.

On the other hand, measures based on coating the contact surfaces of the splices appear to be less effective. Although the development of bulky corrosion products in the crevice has not been observed in bolted joints with Teflon Vaseline-protected splices after 9 years of exposure, the adverse visual effect on the development of corrosion products around the connection significantly limits the applicability of this measure.

#### *4.3. Recommended Method of Rehabilitation of Bolt Joints of the Leg Members*

While a sufficiently protective layer of corrosion products usually develops on the common elements of the steel bearing structure of weathering steel lattice towers, locally unfavourable behaviour related to the development of crevice corrosion occurs in the bolted joints areas. Loading tests carried out on specimens taken from the collapsed transmission towers show that the static load capacity of the bolted joints is not significantly affected by adverse corrosion development in the crevice between the leg member and the splices (see Section 3.1). The results of load tests are very important for steel structure designers and transmission line administrators when evaluating the reliability of in-service steel lattice towers [44]. This finding is confirmed by experience with collapsed transmission towers, where the cause of failure is never the bolted joints of leg members, even though they show crevice corrosion. The collapse is usually caused by critical weakening of the cross-sections of the leg members at the anchorage to the foundations, or by overloading and buckling in strong winds. However, the significant plastic deformation of the splices, caused by the effects of crevice corrosion, is visually unfavourable, and raises concerns among the owners or administrators of the structure about the possible limitation of the bearing capacity or working life of the structure. Rehabilitation of these joints is required for this reason.

It is possible to rehabilitate the connection without dismantling and replacing the splices and bolts with new ones. In order to ensure the longest possible working life of the connection rehabilitated by repair without removing the splices and bolts, it is recommended to use to the following repair procedure [42]:


Compatibility between the sealant and the coating system is necessary to achieve the expected life of the rehabilitation system. It is recommended to check the technical condition of the rehabilitation system at regular intervals (optimally, a period of 5 years).

#### **5. Conclusions**

The paper dealt with the bearing capacity and durability of steel lattice towers made of weathering steels. The main focus was the corrosion damage of the bolted joints of the leg members, and the effect on the load-bearing capacity and durability of the steel structure. Using experimental testing, and on the basis of analytical and numerical calculations, the effect of crevice corrosion on the bolted joint resistance was evaluated. Appropriate design measures, applicable for the rehabilitation of developed crevice corrosion of in-service structures, or the elimination of crevice corrosion in newly designed lattice towers, were evaluated. The main conclusions of the paper are as follows:

• The capacity of the bolted joint of the leg member is not significantly affected by the development of corrosion products in the crevice, provided that higher strength bolts (8.8 or 10.9) are used, and no significant corrosion weakening of the corner bolt occurs. This finding was verified both by tensile tests on specimens taken from actually operated transmission lattice towers, and by analytical and numerical calculations and static assessments.


**Author Contributions:** Conceptualization, V.K.; methodology, V.K. and Z.V.; software, M.V.; validation, V.K.; formal analysis, V.K. and L.M.; investigation, V.K. and Z.V.; resources, V.K. and Z.V.; data curation, V.K.; writing—original draft preparation, V.K.; writing—review and editing, L.M.; visualization, V.K.; supervision, V.K.; project administration, V.K.; funding acquisition, V.K. All authors have read and agreed to the published version of the manuscript.

**Funding:** The evaluation of the influence of corrosion damage on the durability of the steel structure of the lattice towers was carried out within the project of the Grant Agency of the Czech Republic (project GA21-14886S Influence of material properties of high strength steels on durability of engineering structures and bridges). Financial support from VSB-Technical University of Ostrava, by means of the Czech Ministry of Education, Youth, and Sports, through the Institutional support for conceptual development of science, research, and innovations for the year 2021 (project No. SPP IP2071111) is also gratefully acknowledged.

**Institutional Review Board Statement:** Not applicable.

**Informed Consent Statement:** Not applicable.

**Data Availability Statement:** Data are contained within the article.

**Conflicts of Interest:** The authors declare no conflict of interest.

#### **References**


## *Article* **Localized Corrosion Occurrence in Low-Carbon Steel Pipe Caused by Microstructural Inhomogeneity**

**Yun-Ho Lee 1,†, Geon-Il Kim 1,†, Kyung-Min Kim 1, Sang-Jin Ko 1, Woo-Cheol Kim <sup>2</sup> and Jung-Gu Kim 1,\***


**Abstract:** In this study, the cause of failure of a low-carbon steel pipe meeting standard KS D 3562 (ASTM A135), in a district heating system was investigated. After 6 years of operation, the pipe failed prematurely due to pitting corrosion, which occurred both inside and outside of the pipe. Pitting corrosion occurred more prominently outside the pipe than inside, where water quality is controlled. The analysis indicated that the pipe failure occurred due to aluminum inclusions and the presence of a pearlite inhomogeneous phase fraction. Crevice corrosion occurred in the vicinity around the aluminum inclusions, causing localized corrosion. In the large pearlite fraction region, cementite in the pearlite acted as a cathode to promote dissolution of surrounding ferrite. Therefore, in the groundwater environment outside of the pipe, localized corrosion occurred due to crevice corrosion by aluminum inclusions, and localized corrosion was accelerated by the large fraction of pearlite around the aluminum inclusions, leading to pipe failure.

**Keywords:** failure analysis; low-carbon steel pipe; pitting corrosion; aluminum inclusions; pearlite inhomogeneity

#### **1. Introduction**

District heating (DH) systems produce heat that is used to provide steam and hot water to the residents of large cities [1,2]. These heating systems provide higher thermal efficiency and lower heating costs than small private boiler units [3]. The steam and hot water that are transported in DH systems expose pipes to corrosion risks. Failure of system pipes due to corrosion reduces thermal efficiency and negatively impacts the system's cost and reliability [1]. As such, failure analyses and pipe corrosion prevention are important for the maintenance of DH systems.

Low-carbon steel has been widely used as a pipe material in various industrial plants such as those that manage oil, gas, and water, as well as in DH systems [1,4–8]. Several studies have been conducted on the failure of low-carbon steel pipes. Kim et al. reported that crack propagation occurs due to stress concentrations and high hydrogen susceptibility in the weld zone (WZ) and heat-affected zone caused by poor welding, resulting in the failure of low-carbon steel pipe [1]. Lee et al. reported that the failure of low-carbon steel pipes are caused by stress corrosion cracking due to chloride presence and residual stress in the WZ [5]. Heyes et al. observed the fatigue cracking as a result of oxygen-induced pitting in the pipe [9]. Various other failure types and mechanisms can be found in available literatures [10].

In most studies, the failure of low-carbon steel pipe occurs in the form of stress corrosion cracking (SCC) or corrosion fatigue cracking (CFC) when stress is applied. However,

**Citation:** Lee, Y.-H.; Kim, G.-I.; Kim, K.-M.; Ko, S.-J.; Kim, W.-C.; Kim, J.-G. Localized Corrosion Occurrence in Low-Carbon Steel Pipe Caused by Microstructural Inhomogeneity. *Materials* **2022**, *15*, 1870. https:// doi.org/10.3390/ma15051870

Academic Editor: Vít Kˇrivý

Received: 23 January 2022 Accepted: 28 February 2022 Published: 2 March 2022

**Publisher's Note:** MDPI stays neutral with regard to jurisdictional claims in published maps and institutional affiliations.

**Copyright:** © 2022 by the authors. Licensee MDPI, Basel, Switzerland. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC BY) license (https:// creativecommons.org/licenses/by/ 4.0/).

the failure shown in this study was due to pitting corrosion. Low-carbon steel is not passivated as stainless steel is, and corrosion generally occurs uniformly over the pipe. Pitting corrosion is rarely observed on the low-carbon steel pipe [11]. Therefore, it is necessary to investigate the cause of pitting corrosion of low-carbon steel pipe to prevent specific corrosion and sudden leakage.

This study analyzes the failure of a low-carbon steel pipe due to corrosion, which has not been reported in the actual use of low-carbon steel pipe. Compliance with material specification was evaluated using inductively coupled plasma atomic emission spectroscopy (ICP-AES) component analysis. The cause of the failure was then analyzed through visual inspection, optical microscopy (OM), metallographic examination, scanning electron microscopy (SEM) with energy dispersive spectroscopy (EDS), electron probe microanalyzer (EPMA) and atomic force microscopy (AFM). Furthermore, the electrochemical properties of the failed low-carbon steel were evaluated using the potentiodynamic polarization and galvanostatic polarization tests.

#### **2. Materials and Methods**

#### *2.1. Description of the Pipeline*

Before conducting the failure analysis of the low-carbon steel pipe, the specifications and environment were investigated. According to the user's description, the failed lowcarbon steel pipe transported hot water through underground pipeline in DH systems. The low-carbon steel pipe satisfied standard KS D 3562 (ASTM A135, Electric-Resistance-Welded Steel Pipe Grade A) [1,12–14]. The outer diameter of the pipe was 457.2 mm with a wall thickness of 6.4 mm. As shown in Figure 1, the buried pipe is surrounded by polyurethane foam as a heat insulator and a high-density polyethylene (HDPE) pipe as an outer casing. Internal corrosion can occur inside of the pipe caused by the transported water in the DH system. External corrosion can occur on the exterior surface of the pipe due to degradation of the HDPE pipe and subsequent penetration of groundwater [15,16]. Tables 1 and 2 present the chemical composition of the DH water inside of the pipe and the synthetic groundwater outside of the pipe, respectively [15,17]. In DH systems, the design service life of a low-carbon steel pipe is 40 years; however, the pipe observed in this study failed after only 6 years of operation [1].

**Figure 1.** A schematic illustration of the heat transport pipe used in a DH system.

**Table 1.** Chemical composition of district heating water (ppm) used in a district heating system.



**Table 2.** Chemical composition of synthetic ground water (ppm).

#### *2.2. Metallurgical Analyses*

In this study, it was investigated whether the material used for the pipe satisfies the KS D 3562 standard through chemical composition analysis. The shape of the failed pipe was visually investigated via OM. The metallographic examination was performed to confirm the microstructure uniformity of the pipe metal. To investigate the microstructure of the pipe metal, the specimen was polished using a 1-μm diamond suspension, and thereafter etched by a 2% Nital etching solution for 20 s [1,4]. The volume fractions of pearlite and ferrite phases in the microstructure according to the location of corrosion in the failed pipe were measured by OM using Image J software (version 1.8.0) [18]. For the OM image, the fraction of each phase was calculated by counting the number of pixels in ferrite and pearlite and dividing by the total number of pixels. For microanalysis of the microstructure, topography and surface potential were measured using AFM and kelvin probe force microscopy (KPFM), a mode of AFM. AFM measurements were performed using a commercial AFM system (NX10, Park Systems, Suwon, Korea). KPFM measurements were performed using a conductive Pt/Cr coated tip (Multi75E-G, BudgetSensors, Sofia, Bulgaria) in lift mode with a tip-to-sample distance of 20 nm, and an AC modulation voltage of 2Vrms at 17 kHz. Measurements were performed at 10 μm × 10 μm and 2 μm × 2 μm, respectively. The fracture properties were analyzed using SEM/EDS (SEM-7800F Prime, JEOL Ltd., Tokyo, Japan) and EPMA (JXA-8530F, JEOL Ltd., Tokyo, Japan). Prior to SEM/EDS and EPMA analyses, specimens were pickled, polished with 1000-grit size silicon carbide paper, and then rinsed with deionized water and cleaned with ethanol.

#### *2.3. Electrochemical Tests*

Potentiodynamic polarization and galvanostatic polarization tests were performed using a VSP 300 (Bio-Logic SAS, Seyssinet-Pariset, France). To conduct these electrochemical tests, a three-electrode system comprising low-carbon steel pipe specimen taken from the failed pipe as the working electrode (WE), two pure graphite rods as the counter electrodes (CE), and a saturated calomel electrode (SCE) with a Luggin capillary as the reference electrode (RE) was used. To confirm the cause of the pitting corrosion, specimens were prepared for the pitting corrosion part (specimen A) and the uniform corrosion part (specimen B) of the failed pipe. For electrochemical tests, all specimens were polished with a 1000-grit silicon carbide paper, rinsed with ethanol, and dried with nitrogen gas. The area of all specimens was controlled to a size of 1 cm<sup>2</sup> using a sealant. A groundwater solution was used for the electrochemical tests (Table 2) because it was more corrosive to the pipe. The temperature was maintained at 60 ◦C to reflect the temperature of the distinct heating system [2,14,16]. Before conducting the electrochemical tests, the WE was immersed in the test solution for 6 h to obtain a stable open-circuit potential (OCP) as the corrosion potential (Ecorr). The potentiodynamic polarization tests were conducted using a potential sweep of 0.01 mV/s from −0.25 V vs. Ecorr to 1.6 V vs. Ecorr. The galvanostatic polarization tests were performed at 5 mA/cm<sup>2</sup> to accelerate corrosion. The total amount of coulombic charge was 143.86 mAh, which is equivalent to 6 months of uniform corrosion in specimen B. To evaluate the same amount of corrosion, the same coulombic charge was applied to both specimens.

#### **3. Results and Discussion**

#### *3.1. Chemical Composition Analysis of the Low-Carbon Steel Pipe*

Table 3 shows the chemical composition analysis of the low-carbon steel pipe material, as well as the KS D 3562 standard. It was confirmed that the composition of the failed pipe material complied with the KS D 3562 standard. A key finding of note was the detection of

a small amount of Al component in the failed pipe, where KS D 3562 does not dictate any Al component.

**Table 3.** Chemical compositions of the failed low-carbon steel pipe and KS D 3562 standard (wt. %).


#### *3.2. Visual and Macroscopic Inspections*

Based on visual inspection of the failed pipe, leakage occurred in the form of pitting, and the severe pitting corrosion occurred near the part of the pipe where the leaking occurred (Figure 2). Figure 3 shows surface images of the pitting corrosion part where leakage occurred, as well as pitting corrosion near the leakage. Figure 4 shows crosssectional images of the pitting corrosion part where the leakage occurred and the pitting corrosion near the leakage. As stated previously, pitting corrosion occurred on both the exterior and interior of the failed pipe wall. However, since low-carbon steel is a material that does not have passivation, pitting corrosion cannot occur due to the failure of passivation. In low-carbon steel, localized corrosion can be caused by crevice corrosion by specific inclusions, galvanic corrosion between dissimilar metals, and under deposit corrosion (UDC) by solid particles (sand, debris, and iron oxides) [14,19]. In other words, it is necessary to investigate the above possibilities openly.

**Figure 2.** Photographs of the failed pipe: (**a**) water leakage resulting from pitting in the failed pipe, (**b**) external surface of the failed pipe around the leakage area.

**Figure 3.** Surface image of the pitting part: (**a**) pitting corrosion area of the leakage (the outside of the pipe), (**b**) pitting corrosion area near the leakage (the outside of the pipe), and (**c**) pitting corrosion area near the leakage (the inside of the pipe).

**Figure 4.** Cross-sectional images of the pitting part: (**a**) pitting corrosion where the leakage occurred, (**b**) pitting corrosion near the leakage, and (**c**) pitting corrosion near the leakage.

Moreover, the depth of the pitting corrosion on the exterior of the failed pipe was significantly deeper than the interior. It is considered to have been due to the difference between the environment of the inside and outside of the pipe (Tables 1 and 2). In addition, to reduce the risk of corrosion, the DH water that flowed through the pipe was maintained at a dissolved oxygen level below 200 ppm through periodic water quality management [15]. In low-carbon steel, localized corrosion hardly occurs when the concentration of dissolved oxygen is low [20]. In addition, in the case of the exterior of the pipe, localized corrosion may be further accelerated due to the non-uniform distribution of groundwater penetrating into the heat insulator. Therefore, it is considered that the corrosion perforation of the pipeline is induced by the external corrosion.

#### *3.3. Metallographic Examinations and Atomic Force Microscopy*

Figure 5 shows the microstructure of the pitting corrosion region (specimen A) and the uniform corrosion region (specimen B) of the failed pipe. The darker-colored section is pearlite, which is a layered structure composed of ferrite and cementite, and the brightercolored section is ferrite [21]. Figure 5 and Table 4 show that there is a larger fraction of pearlite in specimen A than in specimen B. Pearlite is susceptible to corrosion as its constituent phases, ferrite and cementite, have dissimilar electrochemical potentials that cause microgalvanic corrosion when exposed to corrosive electrolytes [21–23].

**Figure 5.** The optical microstructures of specimens: (**a**) specimen A (pitting corrosion region), and (**b**) specimen B (uniform corrosion region).


**Table 4.** Volume fraction of the pearlite and ferrite phases according to the specimen A (pitting corrosion region) and specimen B (uniform corrosion region) in the failed pipe.

Figure 6 shows the AFM and KPFM analyses of the pearlite section of the failed pipe. Figure 6a,b show the correlation between pearlite and the surrounding pro-eutectoid ferrite by measuring topology and surface potential within a size of 10 μm × 10 μm, respectively. In Figure 6a, pearlite has a lamella structure, and the surface potential of pearlite (A site) is approximately 44.4 mV higher than that of the surrounding pro-eutectoid ferrite (B site) in Figure 6b and Table 5. This indicates that the corrosion of pro-eutectoid ferrite is locally accelerated by microgalvanic corrosion between pearlite and the surrounding pro-eutectoid ferrite [19,24]. Figure 6c,d show the correlation between cementite and ferrite in the pearlite by measuring topology and surface potential within a size of 2 μm × 2 μm, respectively. Figure 6c shows the lamella structure of pearlite, indicating that the bright protrusions are cementite while the dark region is the ferrite [25]. This is because the corrosion of ferrite was more corroded than that of cementite during the etching process using 2% Nital etching solution. Figure 6d and Table 5 show that the surface potential of cementite in the pearlite (C site) is approximately 34.78 mV higher than that of ferrite in the pearlite (D site). Thus, the corrosion of ferrite is locally accelerated by microgalvanic corrosion between ferrite and cementite in the pearlite [19,24].

**Figure 6.** Topography and surface potential of pearlite (specimen A): (**a**) topography of pearlite, (**b**) surface potential of pearlite, (**c**) topography inside of pearlite, and (**d**) surface potential inside of pearlite.


**Table 5.** Surface potential of different phase area and their differences.

Therefore, the inhomogeneity of the microstructure forces a difference in corrosion rate. A pipe section exhibiting pitting corrosion with a lot of pearlite experiences severe corrosion, whereas a uniform corrosion section of the failed pipe with a low pearlite phase incurs only minor thickness reduction.

#### *3.4. Microscopic Analyses*

To identify the cause of severe pitting corrosion, microscopic analyses were performed on the pitting corrosion near the leakage area using SEM/EDS and EPMA (Figure 7). Figure 7a,b showed several inclusions around the pitting corrosion site. The EDS elemental mapping of the inclusions in Figure 7c reveals that the particle is primarily composed of Al. The size range of these Al inclusions is 10–20 μm. In the chemical composition analysis, 0.04% Al was contained in the failed pipe, as shown in Table 3.

Al inclusions were considered to have been the result of Al use as a deoxidizer in the steelmaking process [26]. Al is known to have uniform distribution and fine particle size (less than 2 μm) upon proper heat treatment; accordingly, inclusions typically have little influence on corrosion behavior [27]. However, the failed pipe appeared to have largersized Al inclusions as well as an uneven distribution compared with the results in the literature [27,28]. This may have been the result of improper or incomplete homogenization during the steelmaking process. Furthermore, during the pipe manufacturing process, microcrevices formed at matrix–inclusion interfaces due to dissimilarities in the strain values and thermal expansion coefficients, which may have caused crevice corrosion, thereby accelerating localized corrosion [19,29].

Figure 8 shows the EPMA analysis of a cross section where pitting corrosion occurred in the failed pipe. Carbon agglomerations were partially observed in the pitting corrosion part. It appeared that the pitting corrosion region had a larger pearlite fraction than the uniform corrosion region in the failed pipe. Accordingly, the carbon agglomerations occurred due to the high fractional presence of cementite in the pearlite. As shown in the AFM and KPFM analyses results (Figure 6 and Table 5), when there are many pearlite phases, the material is vulnerable to corrosion due to microgalvanic corrosion between pearlite and surrounding pro-eutectoid ferrite, ferrite, and cementite [19,24,30]. The ferrite is corroded locally by microgalvanic corrosion around cementite.

**Figure 8.** The EPMA analysis of a cross section where pitting corrosion occurred in the failed pipe.

#### *3.5. Open-Circuit Potential Measurement*

Figure 9 shows the open-circuit potential (OCP) of specimen A and specimen B in the groundwater solution. Specimen A had a higher Ecorr than specimen B, and had a relatively large potential fluctuation of approximately 50 mV. Specimen A has a higher Ecorr due to a larger fraction of pearlite, which has a relatively noble potential compared to ferrite. Pearlite has a higher Ecorr than ferrite due to an increase in cathodic sites that cause oxygen reduction (O2 + 2H2O + 4e = 4OH−) [31,32]. In addition, due to the higher corrosion activity caused by larger fraction of pearlite and presence of Al inclusions, OCP fluctuation is shown on specimen A [33,34].

**Figure 9.** Open-circuit potential of specimen A (pitting corrosion region) and specimen B (uniform corrosion region) with immersion time in the groundwater at 60 ◦C.

#### *3.6. Potentiodynamic Polarization Test*

Figure 10 and Table 6 show the results of the potentiodynamic polarization test of specimen A and specimen B in the groundwater solution. The corrosion current density was controlled by the oxygen reduction reaction [31]. The corrosion current density was analyzed using the Tafel extrapolation method. Once the corrosion current density was determined, the corrosion rate can be calculated using the following equation [11]:

$$\text{Corrosion rate (mm/year)} = 0.00327 \frac{a \cdot i\_{corr}}{n \cdot D} \tag{1}$$

where *a* is the atomic weight, *icorr* is the corrosion current density, *n* is the number of equivalents exchanged, and *D* is the density of the low-carbon steel. Specimen A had a corrosion current density twice that of specimen B. The high corrosion rate in specimen A is due to the accelerated corrosion caused by Al inclusions and the larger pearlite phase fraction. In the vicinity of Al inclusions, localized corrosion occurs due to crevice corrosion, increasing the corrosion rate [19]. Larger pearlite phase fraction accelerates the corrosion rate via galvanic corrosion between pearlite and pro-eutectoid ferrite and between cementite and ferrite in pearlite [31]. In addition, when a larger fraction of pearlite exists around the Al inclusion, corrosion is accelerated by the larger fraction of pearlite, and aggressive ions, such as the Cl− ion, are concentrated around the Al inclusion. This further accelerates crevice corrosion in Al inclusions. If there is no crevice around the Al inclusions, crevice corrosion does not occur. In addition, when pitting formed as crevice corrosion and galvanic corrosion progressed around the Al inclusions and the pearlite, the surface area became wider than the initial area due to morphological changes [29,35]. Equation (2) shows the anode current density according to the anodic overpotential in the activation polarization [11].

$$i\_a = i\_0 \exp\left(\frac{a \cdot n \cdot F \cdot \eta\_a}{2.3 \cdot R \cdot T}\right) \tag{2}$$

**Table 6.** The electrochemical parameters resulting from the polarization measurements of specimen A and specimen B in the groundwater at 60 ◦C.


**Figure 10.** Potentiodynamic polarization curves of specimen A (pitting corrosion region) and specimen B (uniform corrosion region) in the groundwater at 60 ◦C.

In Equation (2), *ia* is the current density by the anodic overpotential, *i*<sup>0</sup> is the exchange current density, *α* is the fraction of *η<sup>a</sup>* taken by the ionization reaction, *n* is the number of equivalent exchanged, *F* is the Faraday's constant, *η<sup>a</sup>* is the anodic overpotential, *R* is the gas constant, and *T* is the temperature. Equation (3) shows the correlation between area and current.

$$i = \frac{I}{A} \tag{3}$$

In Equation (3), *i* is the current density, *A* is the reaction area, and *I* is the current. When the same overpotential was applied, the generated current (*I*) increased in proportion to the increased area. However, to obtain the current density, the reaction area (*A*) is equally divided by 1 cm2 for the generated current (*I*). Therefore, the change in roughness due to localized corrosion of specimen A causes higher current density on the potentiodynamic polarization curve.

#### *3.7. Galvanostatic Polarization Test*

The galvanostatic polarization test was performed to accelerate corrosion. The acceleration time for the galvanostatic polarization test was calculated using the Faraday's law as shown below [15]:

$$i\_{real} \cdot t\_{real} = \frac{m \cdot \vec{r} \cdot \mathbf{n}}{a} = i\_{accelerated} \cdot t\_{accelerated} \tag{4}$$

where *m* is the reacted mass (g), *i* is the current density (A/cm2), *t* is the time (s), *a* is the atomic weight (g/mol), *F* is the Faraday's constant (96,500 C/mol), and *n* is the number of electrons exchanged. The same coulombic charge was applied to observe the corrosion behavior for the same amount of corrosion. The total coulombic charge was 143.86 mAh, and the applied current was 5 mA/cm2. Figure 11 shows surface and crosssectional images of specimen A and specimen B after galvanostatic polarization test. Pitting corrosion occurred in specimen A, and uniform corrosion occurred in specimen B. This indicates that pitting corrosion is related to the presence of Al inclusions and inhomogeneity of the pearlite.

**Figure 11.** Surface and cross-sectional images of the specimens after galvanostatic polarization test: (**a**) surface image of specimen A (pitting corrosion region), (**b**) surface image of specimen B (uniform corrosion region), (**c**) cross-sectional image of the specimen A, and (**d**) cross-sectional image of the specimen B.

#### *3.8. Mechanism*

Figure 12 shows the failure mechanism of the failed low-carbon steel pipe due to the Al inclusions and a large amount of pearlite formed locally during the steelmaking process. Due to the coefficient of thermal expansion differences between Al and Fe, microcrevices form around Al inclusions during the pipe manufacturing process. As crevice corrosion is initiated, pH drops and Cl− ions are concentrated in the microcrevice to maintain charge neutrality [19]. The concentration of Cl− ions further accelerates crevice corrosion and corrosion products accumulate on this part of the pipe. Additionally, oxygen-concentration cells form, which accelerate localized corrosion, and eventually the Al inclusions fall off [19,29].

The large fraction of pearlite has a higher corrosion rate due to microgalvanic corrosion between the surrounding pro-eutectoid ferrite and pearlite, and between cementite and ferrite in pearlite [19,24]. Due to the relatively accelerated corrosion rate of the large fraction of pearlite near the Al inclusions, the concentration of Cl− ions into Al vicinity and the accumulation of corrosion products are accelerated. This promotes the formation of an oxygen-concentration cell in the vicinity of the Al inclusion.

However, the presence of a large fraction of pearlite alone cannot cause localized corrosion such as pitting. The cementite inside the pearlite locally accelerates the surrounding ferrite corrosion by microgalvanic corrosion, but over time, the rust is covered by the dissolution of ferrite, and the corrosion proceeds in the form of a uniform corrosion [19]. In other words, a large fraction of pearlite accelerates corrosion locally, but cannot lead to pitting corrosion. That is, the large fraction of pearlite accelerates crevice corrosion via Al inclusions and promotes the formation of oxygen-concentration cell, thereby accelerating localized corrosion in the vicinity of the Al inclusions.

**Figure 12.** Failure mechanism of the failed low-carbon steel pipe based on aluminum inclusion and the larger phase fraction of the pearlite: (**a**) initial stage, and (**b**) later stage.

#### **4. Conclusions**

In this study, the failure analysis of a low-carbon steel pipe used in DH system was investigated using visual inspection, ICP-AES, OM, AFM, SEM/EDS, EPMA, and electrochemical tests. According to the results of the failure analysis, the following conclusions were drawn.


#### **5. Recommendations**


**Author Contributions:** Conceptualization, Y.-H.L. and G.-I.K.; methodology, Y.-H.L. and G.-I.K.; validation, Y.-H.L., G.-I.K., K.-M.K., S.-J.K., W.-C.K. and J.-G.K.; formal analysis, Y.-H.L. and G.-I.K.; investigation, Y.-H.L. and G.-I.K.; resources, Y.-H.L. and G.-I.K.; data curation, Y.-H.L., G.-I.K., K.-M.K. and S.-J.K.; writing—original draft preparation, Y.-H.L. and G.-I.K.; writing—review and editing, Y.-H.L., G.-I.K., K.-M.K., S.-J.K., W.-C.K. and J.-G.K.; visualization, Y.-H.L. and G.-I.K.; supervision, J.-G.K.; project administration, J.-G.K.; funding acquisition, W.-C.K. and J.-G.K. All authors have read and agreed to the published version of the manuscript.

**Funding:** This research was supported by the Korea District Heating Corporation (No. 0000000014524).

**Institutional Review Board Statement:** Not applicable.

**Informed Consent Statement:** Not applicable.

**Data Availability Statement:** Not applicable.

**Conflicts of Interest:** The authors declare no conflict of interest.

#### **References**


## *Article* **Method for Mitigating Stray Current Corrosion in Buried Pipelines Using Calcareous Deposits**

**Sin-Jae Kang 1, Min-Sung Hong <sup>2</sup> and Jung-Gu Kim 1,\***


**Abstract:** Stray current corrosion in buried pipelines can cause serious material damage in a short period of time. However, the available methods for mitigating stray current corrosion are still insufficient. In this study, as a countermeasure against stray current corrosion, calcareous depositions were applied to reduce the total amount of current flowing into pipelines and to prevent corrosion. This study examined the reduction of stray current corrosion via the formation of calcareous deposit layers, composed of Ca, Mg, and mixed Ca and Mg, at the current inflow area. To verify the deposited layers, scanning electron microscopy (SEM), energy dispersive X-ray spectroscopy (EDS), and X-ray diffraction (XRD) were performed. The electrochemical tests revealed that all three types of calcareous deposits were able to effectively act as current barriers, and that they decreased the inflow current at the cathodic site. Among the deposits, the CaCO3 layer mitigated the stray current most effectively, as it was not affected by Mg(OH)2, which interferes with the growth of CaCO3. The calcium-based layer was very thick and dense, and it effectively blocked the inflowing stray current, compared with the other layers.

**Keywords:** stray current corrosion; pipeline; calcareous deposit; corrosion mitigation; cathodic protection

#### **1. Introduction**

Stray current corrosion, which is a drastic corrosion phenomenon due to external current sources, can cause serious damage to buried pipelines in a short period of time. With the increasing number of buildings, facilities, and subways that use high voltages in modern society, the amount of stray current is also increasing, inducing more stray current corrosion in buried pipelines [1–6]. In particular, stray current from subways, power towers, and high-voltage facilities flows into buried pipelines that have lower resistances than soil. The area into which the stray current flows is negatively charged, resulting in anticorrosion, while the area out of which it flows is positively charged, resulting in corrosion [1,7,8]. The inflow and outflow of the stray current occur at random areas throughout pipelines, making it difficult to detect, and difficult to prevent the associated corrosion [9–11]. In addition, most pipelines are usually installed in urban areas, where stray current can be introduced to the pipelines, causing both economic and human-related losses [12,13]. In the UK, GBP 500 million is spent annually on infrastructure restoration and repair due to stray current corrosion [14]. To solve this issue, drainage systems and electrical shields have been applied to buried pipelines to prevent corrosion [1,15–17]. However, they are expensive and cannot be applied to all pipelines. Since it is hard to predict where stray current corrosion will occur, it is difficult to ensure that all pipelines are protected.

To overcome these issues, we applied calcareous deposits for protection against stray current corrosion. Generally, calcareous deposits are a type of combined deposit based on calcium (Ca) and magnesium (Mg), and they are generated under cathodic protection in

**Citation:** Kang, S.-J.; Hong, M.-S.; Kim, J.-G. Method for Mitigating Stray Current Corrosion in Buried Pipelines Using Calcareous Deposits. *Materials* **2021**, *14*, 7905. https:// doi.org/10.3390/ma14247905

Academic Editor: Vít Kˇrivý

Received: 10 November 2021 Accepted: 18 December 2021 Published: 20 December 2021

**Publisher's Note:** MDPI stays neutral with regard to jurisdictional claims in published maps and institutional affiliations.

**Copyright:** © 2021 by the authors. Licensee MDPI, Basel, Switzerland. This article is an open access article distributed under the terms and conditions of the Creative Commons Attribution (CC BY) license (https:// creativecommons.org/licenses/by/ 4.0/).

seawater. Generally, calcareous deposits act as electrical barriers, and they therefore have the potential to be an excellent solution with regard to mitigating stray current corrosion. The cathodic sites of pipelines, where the inflow current is introduced, have the same conditions when cathodically protected. Therefore, the supply of Ca and Mg can generate a calcareous deposit in a soil environment [18–22].

In this study, potentiostatic polarization tests were performed to form calcareous deposits with different compositions [23]. After forming the calcareous deposits, the surface morphology was analyzed using scanning electron microscopy (SEM), energy dispersive X-ray spectroscopy (EDS), and X-ray diffraction (XRD). Electrochemical impedance spectroscopy (EIS) experiments were also performed after each potentiostatic polarization test. Finally, potentiostatic acceleration tests were undertaken to verify the effects of the calcareous deposits on the stray current corrosion, and to determine the most effective composition for protection against it.

#### **2. Materials and Methods**

#### *2.1. Specimen and Solution Preparation*

As shown in Table 1, the specimens used in this study were made of SPW-400 (lowcarbon steel), which is the most common material for pipelines.

**Table 1.** Chemical composition of SPW 400 (wt.%).


The SPW-400 was cut into square sections (10 mm × 10 mm × 2 mm), which were used as the working electrodes (WE). The specimens were ground with SiC paper (600-grit), after which they were cleaned with deionized water, and then dried with N2 gas. Table 2 lists the chemical composition of the synthetic soil solution used in the experiments. For the formation of the calcareous deposits using potentiostatic polarization tests, Ca and Mg, which are the main components of these deposits, were added to the synthetic soil solution, separately and together. The elements were based on the following chemicals: Mg(OH)2 (Mg: 1000 ppm), CaCO3 (Ca: 1000 ppm), and Ca and Mg (500 ppm each).

**Table 2.** Chemical composition of synthetic soil solution (ppm).


#### *2.2. Formation of Calcareous Deposits*

A potentiostatic polarization test using a three-electrode system was performed to form the calcareous deposits. The specimens were connected to a WE, a carbon rod was used as the counter electrode (CE), and a saturated calomel electrode (SCE) was used as the reference electrode (RE). The area of each test specimen exposed to electrolytes was 1 cm2. The open-circuit potential (OCP) was established within 30 min, after which the electrochemical tests were performed. Potentiostatic polarization tests were undertaken to form the calcareous depositions. The tests were performed at −1.0 VSCE to put the specimens in the cathodic state. The current was inflowed for over 30 h at room temperature (25 ◦C), while the solution was rotated at 350 rpm.

#### *2.3. Surface Analyses*

The surface analysis of the experimental specimens was performed after the potentiostatic polarization tests. The morphology and the cross-sectional images of the calcareous deposits were observed using SEM and SEM/EDS (JSM-7900F, JEOL Ltd., Tokyo, Japan) to verify the type of calcareous deposit on the specimen. XRD (Dmax-2500V/PC, Rigaku, Tokyo, Japan) measurements were also performed on the calcareous deposits to verify their types. The XRD analysis was conducted using Cu Kα radiation (λ = 1.54056 Å), in a 2θ range of 0–60◦, at a scan rate of 0.02.

#### *2.4. Electrochemical Test*

EIS measurements were performed in a frequency range of 100 kHz–10 mHz, with a 10-mV amplitude. The impedance plots were interpreted on the basis of an equivalent circuit, using a fitting procedure performed by ZsimpWin software (ZsimpWin 3.20, Echem Software, Warminster, PA, USA). Stray current corrosion tests were performed in a stray current simulation cell, as shown in Figure 1. The 304 stainless steel rods used as the CE were enclosed with insulating tape to reduce the current dispersion, and the SCE was used as the RE. The specimens used for the inflow part of the current and those used for the outflow part of the current were electrically connected to each other. The tests were conducted in a synthetic soil solution, and 3.5 VSCE was applied for 100 h at room temperature (25 ◦C). The specimens for the outflow part of the current were weighed and recorded before the potentiostatic acceleration tests. After the tests, the specimens were cleaned, rinsed, and reweighed. All electrochemical tests were performed using a VSP-300 model potentiostat (Biologic SAS, Seyssinet-Pariset, France).

**Figure 1.** Schematic of stray current simulation cell.

#### **3. Results**

#### *3.1. Formation of Calcareous Deposits*

The potentiostatic tests were performed at −1.0 VSCE for 30 h to form three kinds of calcareous deposits on the carbon steel. Figure 2 shows the potentiostatic test results over the 30-h period. The current density decreased with time for all solution types. This can be explained by the electrochemical reactions on the surfaces of the cathodic site. When the specimens were negatively charged in the solution containing both Ca and Mg, the dissolved oxygens were converted into OH− ions, leading to an increase in the pH on the surface. Because of the increasing number of OH− ions, Mg ions reacted with them, forming an Mg(OH)2 deposition on the metal surface. In addition, the increase in OH− ions affected the carbonate equilibrium at the metal surface. Thus, a CaCO3 layer was deposited on the metal surface. These processes can be described by the following reactions [21–25]:

$$\text{CO}\_2 + 2\text{H}\_2\text{O} + 4\text{e}^- \rightarrow 4\text{OH}^- \tag{1}$$

$$\text{Mg}^{2+} + 2\text{OH}^- \rightarrow \text{Mg(OH)}\_2\text{ (s)}\tag{2}$$

$$\rm{OH}^{-} + \rm{HCO}\_{3}^{-} \rightarrow \rm{CO}\_{3}^{2-} + \rm{H}\_{2}\rm{O} \tag{3}$$

$$\text{Ca}^{2+} + \text{CO}\_3^{2-} \rightarrow \text{CaCO}\_3 \text{ (s)}\tag{4}$$

**Figure 2.** Current density vs. time curves during potentiostatic test (applied voltage: −1.0 VSCE; solutions: synthetic soil solution with Mg(OH)2 CaCO3 and Mg(OH)2, CaCO3; testing time: 30 h).

These calcareous deposits decreased the O2 diffusion to the metal surface as a physical and electrical coating layer, and hindered the oxygen reduction reaction [23,26]. Therefore, the current density was decreased because of the formation of calcareous deposits on the metal surface, as shown in Figure 2. However, in the case of the solution with only Mg, the Mg(OH)2 layer was porous and gel-like rather than solid. It offered a relatively lower protective property at the surface of the metal compared with the other layers, and it did not significantly decrease the current inflow to the metal [27]. Therefore, the specimen that deposited only Mg(OH)2 had the highest current density. In contrast, the CaCO3 layer has the property of forming a solid and dense layer. Because the CaCO3 layer with these properties grew without any interference, the current density decreased rapidly to the lowest current density value measured in this study [28]. The specimen that deposited both CaCO3 and Mg(OH)2 had a current density higher than that of the specimen with only CaCO3, and a current density lower than that of the specimen with only Mg(OH)2. This is because the Mg(OH)2 hindered the growth of the CaCO3, meaning that the formed CaCO3 layer was thin and unstable [28,29].

#### *3.2. Surface Analysis*

The cross-sectional SEM images and EDS mapping results of the calcareous deposits after the 30-h potentiostatic polarization tests are shown in Figures 3–5. Figure 3 shows this information for the calcareous deposit based only on Mg, revealing a thin Mg(OH)2 layer deposited on the carbon steel. Figure 4 shows this information for the deposit based only on Ca, revealing a relatively thicker CaCO3 layer deposited on the carbon steel than the other deposit layers. Figure 5 shows the SEM image and EDS mapping results of the mixed CaCO3 and Mg(OH)2 deposit. This calcareous deposit layer included both Ca and Mg, and can therefore be regarded as a combined CaCO3 and Mg(OH)2 layer. In addition, it was confirmed that the Mg(OH)2, which hindered the growth of the CaCO3, resulted in a thinner CaCO3 layer compared with the specimen containing only CaCO3 [28,29].

**Figure 3.** The cross-sectional SEM images and EDS mapping results of Mg(OH)2: (**a**) cross-sectional image; (**b**) Fe (EDS mapping); and (**c**) Mg (EDS mapping).

**Figure 4.** The cross-sectional SEM images and EDS mapping results of CaCO3: (**a**) cross-sectional image; (**b**) Fe (EDS mapping); and (**c**) Ca (EDS mapping).

**Figure 5.** The cross-sectional SEM images and EDS mapping results of CaCO3 + Mg(OH)2: (**a**) cross-sectional image; (**b**) Fe (EDS mapping); (**c**) Ca (EDS mapping); and (**d**) Mg (EDS mapping).

Figure 6 shows the XRD patterns used to verify the three types of calcareous deposits on the surfaces of the specimens. It was confirmed that the calcareous deposit layer from the Mg(OH)2-added soil solution was Mg(OH)2. In addition, the calcareous deposit layer from the CaCO3-added soil solution was CaCO3. Finally, the calcareous deposit layer from the CaCO3 and Mg(OH)2-added soil solution consisted of CaCO3 and Mg(OH)2. When the calcareous deposition layer is formed in a solution to which CaCO3 and Mg(OH)2 are added, not only CaCO3 and Mg(OH)2 are formed, but (Ca,Mg)CO3 (JCPDS 43-0697), which is similar to the peaks of CaCO3 (JCPDS 05-0586), is also formed as the product. The width of the peaks of the calcareous deposit layer from the CaCO3 and Mg(OH)2-added soil solution are greater than that of the peaks in the other two patterns because of the overlapping XRD peaks of the (Ca,Mg)CO3.

**Figure 6.** XRD results for the 3 types of calcareous deposits after 30 h of potentiostatic polarization.

#### *3.3. Electrochemical Impedance Spectroscopy*

After the 30-h potentiostatic polarization tests, EIS measurements were performed. The Nyquist plots of the data from the electrodes giving different types of calcareous deposits are shown in Figure 7a. The Nyquist plots consist of a capacitive semicircle at a high frequency. Figure 7b presents the equivalent electrical circuit for bare steel, where Rs is the solution resistance, Rct is the charge transfer resistance, and CPE1 is the doublelayer capacitance formed by the electrical double layer that exists at the interface between the electrolyte and electrode [26,30,31]. Figure 7c shows the equivalent electrical circuit that describes the formation of porous calcareous deposits on the surface of the steel. In Figure 7c, Rs is the solution resistance; CPE2 is the dielectric nature of the calcareous deposits, which is associated with the thickness of the calcareous deposit layer; Rfilm is the pore resistance; CPE1 is the capacitance generated by the metal dissolution reaction and by the electric double layer at the solution/metal interface; and Rct is the charge transfer resistance caused by the metal dissolution reaction [25]. Here, a CPE is used instead of a capacitor to compensate for the nonhomogeneity of the system frequency. The impedance of a CPE is described by the following equation:

$$Z\_{\rm CPE} = A^{-1} \left( \mathbf{j} \omega \right)^{-\mathbf{n}} \tag{5}$$

where *A*−<sup>1</sup> is the proportionality coefficient (with units, Ω−<sup>1</sup> sn cm−2); *ω* is the angular frequency (rad s−1); j2 <sup>=</sup> −1 is an imaginary number; and n is an empirical exponent (0 ≤ n ≤ 1) that measures the deviation from the ideal capacitive behavior [32–34].

**Figure 7.** (**a**) Result of EIS test according to the types of calcareous deposits. (**b**) Equivalent circuit diagram of bare carbon steel. (**c**) Equivalent circuit diagram of carbon steel covered by calcareous deposits.

The results of the EIS fitting using the ZSimpWin software are shown in Table 3. The Rfilm values were the largest in the CaCO3 layer, followed by the CaCO3 and Mg(OH)2 mixed layer, and then the Mg(OH)2 layer. Similar to the results for the Rfilm value, CaCO3 had the largest Rct value, followed by the CaCO3 and Mg(OH)2 mixed layer, and then the Mg(OH)2 layer. These results indicate that the CaCO3 layer worked better as a protective layer than the others. At the same time, the CPE1 values tend to be the opposite of Rct. This is because the active area of metal dissolution decreases as the calcareous deposition becomes wider and thicker on the metal surface. The CPE2 values of the CaCO3 layer were higher than those of the other calcareous deposit layers, meaning that the CaCO3 layer was the thickest. This is in agreement with the SEM image results. Figure 8 shows the total resistance values of the bare steel and the three types of calcareous deposits. All specimens with calcareous deposits had a higher total resistance than the bare specimen, demonstrating that the calcareous deposits provided protection to the bare specimens.


**Table 3.** The results of the EIS fitting using the circuit.

**Deposit (Ω**·**cm2) CPE Y0 (0 < n < 1) (Ω**·**cm2) CPE Y0 (0 < n < 1) (Ω**·**cm2)** Bare 393.7 1.912 <sup>×</sup> <sup>10</sup>−<sup>4</sup> 0.7527 - - - 4785 Mg(OH)2 511.3 2.955 <sup>×</sup> <sup>10</sup>−<sup>4</sup> 0.8693 523.5 3.893 <sup>×</sup> <sup>10</sup>−<sup>4</sup> 0.7953 6100 CaCO3 568.2 5.955 <sup>×</sup> <sup>10</sup>−<sup>4</sup> 0.7615 865.9 1.702 <sup>×</sup> <sup>10</sup>−<sup>4</sup> 0.7231 15,210 CaCO3 + Mg(OH)2 375.9 2.923 <sup>×</sup> <sup>10</sup>−<sup>4</sup> 0.7533 643.4 1.795 <sup>×</sup> <sup>10</sup>−<sup>4</sup> 0.7959 8872

**Figure 8.** Comparison of total resistance on the bare and 3 types of calcareous deposited specimens.

#### *3.4. Corrosion Acceleration Test*

A potentiostatic test was performed at 3.5 VSCE for 100 h to verify the stray current corrosion mitigation of the calcareous deposits [35]. Figure 9 shows the total electric charge at the current inflow area when 3.5 VSCE was applied for 100 h, along with the mass loss of the specimen at the current outflow area. The total electrical charge value was obtained using the following equation [14,36]:

$$\mathbf{Q} = \int\_{ti}^{tf} I \, dt \tag{6}$$

**Figure 9.** Mass losses of corrosion specimen and quantities of inflow current after potentiostatic acceleration test.

Since the calcareous deposits decreased the inflow current, the total electric charge values of all specimens with a calcareous deposit were lower than that of the bare specimen. During the corrosion acceleration tests, a crack occurred in the unstable Mg(OH)2 layer. Therefore, the total electric charge of the specimen with the Mg(OH)2 layer is higher than those of the other specimens with calcareous deposits. In addition, the total electric charge

of the specimen with the CaCO3 layer is lower than those of the other specimens, meaning that the CaCO3 layer was the most protective against inflow current. Figure 10 shows that the specimen on which CaCO3 is deposited receives the lowest inflow current and acts as a stable electrical barrier layer. This is because the CaCO3 layer grows in a solid and stable form, compared to the Mg(OH)2 deposition layer, and without the hindering of Mg(OH)2, it forms a thicker deposition layer than other layers [28,29]. Therefore, it acts as an electrical and physical barrier that blocks the external inflow current more efficiently than other layers. The total electric charge of the specimen with the CaCO3 and Mg(OH)2 mixed layer is higher than that of the specimen with the CaCO3 layer, and lower than that of the specimen with the Mg(OH)2 layer. Generally, more current flowing in means more current flowing out. Therefore, as current inflow increases, the areas where the current outflows become more susceptible to corrosion. As a result, among the specimens representing the current outflow area, the mass loss was the lowest in the specimen with the CaCO3 layer deposited. The mass reduction then increased in the following order, with respect to the deposit layer composition: CaCO3; the CaCO3 and Mg(OH)2 mixed layer; the Mg(OH)2 layer; and the bare specimen.

**Figure 10.** Current vs. time curves during potentiostatic test (applied voltage: 3.5 VSCE; solutions: synthetic soil solution; testing time: 100 h).

#### **4. Conclusions**

This study evaluated the stray current corrosion mitigation of calcareous deposits on carbon steel in a synthetic soil solution using the electrochemical tests, SEM, EDS, and XRD. On the basis of the experiments, the following conclusions can be drawn:


Consequently, stray current corrosion can be effectively mitigated if CaCO3 powders are buried together with the pipelines and deposited when the soil solution and stray current are introduced to the pipelines.

**Author Contributions:** Conceptualization, S.-J.K.; methodology, S.-J.K.; validation, M.-S.H. and J.-G.K.; formal analysis, S.-J.K.; investigation, S.-J.K.; resources, M.-S.H.; data curation, S.-J.K. and M.-S.H.; writing—original draft preparation, S.-J.K.; writing—review and editing, S.-J.K. and M.-S.H.; visualization, S.-J.K.; supervision, J.-G.K.; project administration, J.-G.K.; funding acquisition, M.-S.H. and J.-G.K. All authors have read and agreed to the published version of the manuscript.

**Funding:** This work was supported by the National Research Foundation of Korea (NRF) grant, funded by the Korean Government (MEST) (No. NRF-2019R1A2B5B01070453).

**Institutional Review Board Statement:** Not applicable.

**Informed Consent Statement:** Not applicable.

**Data Availability Statement:** Not applicable.

**Conflicts of Interest:** The authors declare no conflict of interest.

#### **References**

