Several case studies are presented herein which highlight a few of the basic, yet important, concepts which provide the link between the laboratory and field conditions.
Case Study No. 1: Simulation of Conditions of Deaeration or Aeration
One of the most important and universally applicable situations that must be evaluated when conducting laboratory tests in simulated service environments is the need to produce a reasonable representation of the level of deaeration or aeration found in the actual environment. The main reason for the importance of this effect is that corrosivity, in many service applications, can change dramatically with changes in oxygen content. Aeration accelerates anodic corrosion processes with a concomitant increase in localized corrosion activity (i.e., pitting, crevice attack and stress corrosion cracking). Examples of oxygen effects can be seen in applications such as seawater injection, the use of heavy brine completion fluids in oilfield operations and desalination. As shown in Fig. 1, the corrosion rate of steel increases by an order of magnitude going from 10 ppb to just 100 ppb [8]. It only takes very low levels of oxygen contamination (about 1% of normal atmospheric saturation levels) to greatly accelerate corrosion.
Furthermore, due to the sensitivity of corrosion reactions even in low levels of aeration, oxygen contamination can produce excursions to higher corrosion rates that have prolonged effects [9]. The increase in localized anodic attack produced by aeration can be illustrated by its interaction with other species such as chlorides and sulfides. An example, in terms of susceptibility to stress corrosion cracking (SCC), is the interaction between dissolved oxygen and chlorides in elevated temperature applications involving alkaline phosphate treated boiler water [10]. As the availability of oxygen increases above the 0.1 ppm (100 ppb) level, the tolerance for chloride is reduced resulting in a dramatic increase in susceptibility to SCC.
Figure 1. Corrosion rate vs temperature for various oxygen levels
39
The situation in aqueous sulfide-containing environments is even more complicated since oxygen can result in the formation of elemental sulfur, acids, and in some cases, polysulfide species. These can synergistically interact with the oxygen effects mentioned previously to produce quite severe limitations on the corrosion and SCC performance of even very highly alloyed materials. Such conditions can be found in applications which involve pumping of sour oilfield brines, injection of wastewater and flue gas desulfurization. A comparison of the minimum required pitting resistance equivalent (PRE) for conditions involving a simulated sour oilfield service can be seen for identical situations (0.7 kPa H2S, 138kPa CO2, 2 meq/L HCO3, 30,000 ppm Cl-, 65oC) except one is aerated and one deaerated [11]. For the deaerated condition, the minimum pitting resistance equivalent [PRE = Cr + 3.3Mo + 11N + 1.5(W + Cb)] is only 12. This would indicate successful use of materials with > 12 Cr. However, this same environment under aerated conditions yields a minimum PRE value of 30. Under evaporative conditions, this can increase still further.
Understanding the conditions of aeration in the service application is necessary to reproduce similar conditions in the laboratory corrosion test. For example, most geochemical systems naturally contain less than 10 ppb oxygen. By comparison, mechanical deaeration techniques usually will not go below 100 ppm. Multiple vacuum, ultra-low oxygen inert gas purge cycles and prolonged gas purges are usually required to get below 50 ppb oxygen. In some cases, oxygen scavengers must be used to obtain complete deaeration. However, these must be used carefully because they may, in some cases, add other chemical species into the environment that can complicate electrochemical measurements.
Case Study No. 2: Simulation of Corrosion in Multiphase Environments
There are many factors that need to be considered when conducting corrosion assessments in multiphase environments. These include important factors related to the dynamic or flowing nature of the fluids which determine the mode of flow [12] and the kinetic shear forces that are imparted by the flowing fluids on the pipe wall. There have been several major studies involving very sophisticated simulations of three-phase flow. These studies are particularly capital intensive and costly since major investments must be made in the handling, pumping and disposal facilities required for such tests. However, there are no real alternatives for investigating questions involving the direct effects of flow regime such as measuring the shear forces developed by particular flow regimes and operating conditions and the movement of inhibitors [13].
On a more practical basis, more simple yet reasonable approximations of multiphase flow conditions can be obtained using pseudo-three phase systems such as the flow loop shown in Fig. 2 [14]. These systems provide for the establishment of three phase conditions (gas/oil/water) in a reservoir autoclave. Under these conditions, the primary corrodents are dissolved gases (e.g., CO2, H2S and sometimes O2) and an aqueous brine or condensed water phase. Facilities for replenishment of both the gas and liquid phases must be considered depending on the exact nature of the environment. Simulation of the affects of a flowing environment is usually based on modeling the shear stress produced in service on the metal surface by the flowing liquid containing the dissolved gases using the equations given in Table 1 [15]. The main assumption utilized in this approximation is that the major contribution to the wall shear stress is usually made by the liquid phase. In most cases, this
Kane
technique is valid since the contribution of liquid phase density and viscosity on the resultant shear stress predominates over that of the gas phase.
Figure 2. MAPS™ - Multiphase Autoclave Pipeline Simulator
Of significance in most flowing multiphase systems is the handling of slug flow which is the predominate flow regime for horizontal and near horizontal flow applications. The main attribute of slug flow is the very high shear stresses and accompanying high turbulence in the region of the flow just ahead of the moving slug [16]. This effect results in levels of shear stress much greater than those produced by the bulk fluid. It has been proposed that this is the location where excessive corrosion is generated as a result of the effect of locally high shear stress and turbulence on both corrosion and inhibitor films. Investigations have recently focused on techniques such as flow loops and jet impingement to reproduce accurate simulations of such highly turbulent conditions for assessment of corrosion resistance and inhibitor performance [17].
Another major effect that must be addressed in multiphase systems is the potential role of the oil phase as a possible mitigation factor in terms of reducing the corrosion rate [18].
The properties of the hydrocarbon/liquid phase significantly influence the severity of the environment with respect to weight loss corrosion (see Fig. 3). In a typical case, oil/water mixtures remain relatively non-corrosive under flowing conditions of up to about 30% water cut resulting mostly from the preferential wetting and persistence of the oil phase on the metal surface. However, depending on the nature of the liquid hydrocarbon phase, some cases become corrosive with very low water cuts (<5%) while other cases do not become corrosive until more than 50% water is reached. The exact variables that relate to these mitigating
41
effects have only been qualitatively investigated. Therefore, when simulating service applications, the exact nature of the oil phase should be considered since it may play a major role in the overall corrosivity of the system and in the efficacy of inhibitor treatments.
Furthermore, the presence of the nonconductive oil phase in multiphase tests can also produce confusing results when electrochemical techniques are used as the sole basis for evaluation.
Care must be utilized in the selection of corrosion monitoring techniques and the comparison of this data with the results of physical examination, mass loss and localized corrosion measurements.
Table 1. Flow/Shear Stress Relationships
Description Relationsh p i
Determine dimension-less parameters to describe fluid flow characteristics (e.g., Reynold’s number) to account for mass transfer effects
R VD
e =ρ
μ
Determine friction factor, f, to account for pipe wall roughness (from Moody diagrams)
f = z (Re, e/D)
Determine wall shear stress, t, as a function of the friction factor and other flow properties.
τ= f Vρ 2 2
Determine flow regime (annular, stratified, bubble, slug, etc.) to estimate correction factors (e.g., for slug flow, Jepson et al. use the Froude number as a basis to estimate turbulent intensity).
Laboratory simulation: Jet
impingement (Giralt and Trass) τ= ρ
⎡
Laboratory simulation: Rotating cylinder electrode (D. C.
Silverman)
τ=0 0791. Re−0 3. ρ ωr2 2
Summary: The corrosion rate in fully developed turbulent pipe flows computed from field parameters can be simulated in the laboratory produced. can be expressed in terms of wall shear stress. Wall shear stress through experimental methods, and hence similar corrosion rates.
Case Study No. 3: Simulation of Geometry of Exposure
One of the factors that can have a great impact on corrosion severity is the geometry of the service application. Obvious cases are those involving crevices, seams, laps and welds where the formation of an occluded cell can result in differences between local and bulk solutions. In some cases, the whole service condition may bring together somewhat unique combinations of solution and geometric variables which cannot be accurately simulated by
Kane
simple immersion or atmospheric tests. One such case that illustrates this situation is exhibited by corrosion under insulation (CUI).
Figure 3. Effect of hydrocarbon/liquid phase on weight loss corrosion
CUI can result from a build-up of water and contaminants in the annular space between the metal surface and the thermal insulation. It is compounded by situations such as hot wall effects and alternate wetting and drying. The problem typical of CUI is that corrosion rates are typically greater than predicted based on aqueous corrosion data produced from either open or closed system measurements [19]. In an open system, corrosion rates are generally low due to the decreasing solubility of oxygen with increasing temperature. The CUI situation more closely represents a closed system; however, prior studies attempting to simulate CUI by these methods have generally been unsuccessful.
Recently, experiments were conducted with a special test cell designed to model CUI (see Fig. 4) [20]. This novel approach included the use of an internally heated metal tube and isolated ring specimens surrounded by insulating material. The annular space was filled with a simulated atmospheric condensate. Corrosion was assessed using ring specimens that could be monitored using linear polarization resistance (LPR) techniques per ASTM G59, mass loss per ASTM G1 and localized corrosion rate per ASTM G46 [21-23]. Tests incorporated isothermal conditions, thermal cycling and alternate wet dry conditions.
Figure 5 shows the comparison of isothermal and cyclic tests. The mass-loss corrosion rates show values comparable to those associated with CUI in field and plant operations. Of particular interest is the variation in corrosion rate with time for the cyclic tests. The trend indicates that periods of maximum corrosivity involve the periods during re-wetting of the metal surface following the dry cycle. The peaks in corrosion rate are 2 to 3 times the steady state corrosion rates. Furthermore, for cyclic wet-dry conditions, the steady state corrosion rate also increases with time. The benefit of protective surface treatments which results in much lower rates of corrosion versus time can also be seen.
Case Study No. 4: Need for Environment Replenishment
43
Two major concerns when simulating service conditions in the laboratory are the changes in severity that may be caused by depletion of reactive constituents in the corrosive environments and build-up of corrosion products or by-products. Both can modify the corrosivity of the environment to produce a variation in the severity of corrosion from that in the service application. Therefore, periodic monitoring of the laboratory environment may be required to determine the rate of consumption or build-up of various species. Additionally, replenishment may be necessary to eliminate the undesirable effects that they can produce.
Figure 4. Corrosion under insulation (CUI) cell designed by CLI International, Inc.
Figure 5. Instantaneous and mass-loss corrosion rates for a corrosion under insulation (CUI) system
An example illustrating these situations is industrial applications having high partial pressures of carbon dioxide (100-600 psia) in combination with low to moderate hydrogen
Kane
sulfide partial pressures (0.01 to 1.0 psia). These conditions illustrate the importance of both liquid and gas phase replenishment in these high CO2/low H2S systems. In such systems, the steel corrosion product (Fe+2) is soluble in the aqueous phase at low to intermediate temperature (< 60 oC). Additionally, this reaction is accompanied by an increase in the HCO
-3 concentration, which has a buffering effect resulting in increased pH. These factors can result in the premature formation of a protective FeCO3 scale. Furthermore, corrosion will also tend to consume the initial low-level supply of hydrogen sulfide. These three effects, if not controlled, will generally result in an artificially low corrosion rate for steel when compared to the service application. An autoclave procedure is needed for replenishment of the gaseous and/or liquid phases so that the test duration can be prolonged and accurate corrosion assessment can be achieved [24]. Additionally, in most cases, replenishment procedures should be combined with the careful use of a large solution volume to specimen surface area ratio to achieve optimum results.
In cases where a low hydrogen sulfide partial pressure is being utilized, special care must be taken to maintain the intended amount of this reactive constituent. The difficulty in this process increases directly with the corrosion rate of the materials being tested, the total specimen surface area and decreasing hydrogen sulfide partial pressure. The case shown in Fig. 6 is for replenishment involving tests of corrosion resistant alloys [25]. It can be seen that following the first gassing, the desired hydrogen sulfide partial pressure was achieved, but it decreased to a very low level after a short exposure period. At least two more replenishments were required to achieve acceptably constant levels of hydrogen sulfide in the test environment.
Figure 6. H2S partial pressure vs time Case Study No. 5: Acceleration of Corrosion Processes
One of the greatest needs in laboratory corrosion testing is the ability to attain accurate simulation of the test conditions while simultaneously achieving acceleration. The test conditions must produce a reasonable mechanistic simulation yet achieve a degree of acceleration which allows the laboratory test to predict future in-service events in a reasonably short period of exposure time. This is perhaps the most difficult combination of requirements. Oftentimes, tests that are accelerated produce artifacts in the data related to the influence of corrosion mechanisms which are not present in the actual service. Such tests
45
must be approached and conducted with caution. Examples of beneficial acceleration techniques are electrochemical techniques such as controlled polarization, mechanical techniques such as slow strain rate and fracture mechanics testing, and elevated temperature short duration tests on polymeric materials using Arrhenius. modeling techniques.
An example of electrochemical acceleration in a simulated environment was used to produce corrosion films on wear test specimens in a short period of exposure (< 10 days) that were comparable to those produced on actual components over a long period of service (1.5 years). The first step was to achieve accurate simulation of the service environment which, in this case, was high purity cooling water for a boiling water reactor (BWR). Since the intended fluid flow rate was low (2.2-2.8 ft/sec), a rotating cage setup located inside of the autoclave reservoir was utilized [26] (see Fig. 7). The cage employed a special contactor system to allow for application of a controlled anodic current while monitoring electrochemical current. To minimize corrosion of the internal fixtures and application of the anodic current to only the specimens, the fixture was constructed from pre-oxidized Zr-alloy parts.
Figure 7. Rotating cage setup
An extensive literature/experience survey was conducted to determine the rate of steel corrosion with time in the BWR environment and the chemical structure of oxide that would be expected [27]. Based on this information, it was estimated that the corrosion film would be composed of a-Fe2O3 near the surface and Fe3O4 near the metal/oxide interface and, after 18 months of exposure, it would be about 24,000 Å thick. Using the simulated environment, a test was developed that utilized a slight anodic current to accelerate corrosion. Following a series of qualification tests, corrosion films were produced which were a the mixed iron oxide composition very close to the requirements. After an exposure period of ten days, the films were found to be between 15,3000 and 30,600 Å in thickness. These test specimens were subsequently utilized to conduct frictional wear tests so that the actuating force required to manipulate control valves could be estimated.
Kane
ACKNOWLEDGMENTS
I wish to thank the staff of CLI International, Inc. for their hard work and dedication and also the CLI clients which have provided financial support for many technical investigations.
Both have been essential, and have contributed greatly to the development and use of the techniques highlighted in this paper. I also give my appreciation to Ms Delia Cuellar, who has worked and collaborated with me for many years. Thanks are also given to those key people who have helped to provide a guiding light into the practical applications of laboratory testing: Dr. J Brison Greer, Mr. Walter K. Boyd, Professor Joe Payer, Dr. Peter Rhodes and Mr. Bill Ashbaugh.
REFERENCES
1. K. Lewis, Corrosion and Failure Prevention Using Appropriate Materials Selection during Design: Better, Cheaper, Faster, Safer, Presentation at the Golden Gate Materials Technology Conference, San Francisco, California, February 1-3, 1995.
2. J.H. Payer, Increased Reliability and Useful Life through Better Understanding of Corrosion Processes, Plenary Lecture, Seventh Middle Eastern Corrosion Conference, Bahrain Society of Engineering/NACE International, Manama, Bahrain, February 26-28, 1996, pp. 50-53.
3. B.C. Syrett, Cost Effective Corrosion Control in Electric Power Plants, Plenary Lecture, Seventh Middle Eastern Corrosion Conference, Bahrain Society of Engineering/NACE International, Manama, Bahrain, February 26-28, 1996, pp. 1-18.
4. J.W. Spence, et.al., Planning and Design of Tests, Corrosion Tests and Standards, R.
Baboian, ed., MNL 20, ASTM, West Conshohocken, Pennsylvania, 1995, pp. 33-39.
5. ASTM G31, Standard practice for laboratory immersion corrosion testing of metals, ASTM Annual Book of Standards, Section 3.02, West Conshohocken, Pennsylvania.
6. ASTM A262, Standard practices for detecting susceptibility to intergranular attack in austenitic stainless steels, ASTM Annual Book of Standards, Section 3.02, West Conshohocken, Pennsylvania.
7. ASTM G48, Standard test methods for pitting and crevice corrosion resistance of stainless steels and related alloys by the use of ferric chloride solution, ASTM Annual Book of Standards, Section 3.02, West Conshohocken, Pennsylvania.
8. J.W. Oldfield and B. Todd, Corrosion considerations in selecting metals for flash chambers, Desalination 31, 1979, pp. 365-383.
9. A. Ikeda et.al., Corrosion Behavior of Low and High Alloy Tubular Products in Completion Fluids for High Temperature Deep Wells, Paper No. 46, NACE Corrosion/92, March 1992, NACE, Houston, Texas.
10. M. Fontana, Corrosion Engineering, 3rd edition, McGraw-Hill, Inc., New York, 1986.
11. R.D. Kane and S. Srinivasan, Socratestm: Selection of Corrosion Resistant Alloys through Environmental Specification, CLI International, Inc., Houston, Texas, 1996.
12. C.A. Palacios and J.R. Shadley, CO2 Corrosion of API N-80 Steel at 71oC (160oF), Paper No. 476, NACE Corrosion/91, March 1991, NACE, Houston, Texas.
13. X. Zhou and W.P. Jepson, Corrosion in Three Phase Oil/Water/Gas Slug Flow in Horizontal Pipes, Paper No. 26, NACE Corrosion/94, March 1994, NACE, Houston, Texas.
47
14. Course Materials, MAPStm - Multiphase Autoclave Pipeline Simulator, Short Course on Corrosion Test Methodologies for Inhibitor Evaluation, CLI International, Inc., Houston, Texas, 1995.
15. S. Srinivasan, Internal report on flow modeling for laboratory simulation of field effects, CLI International, Inc., Houston, Texas, 1995.
15. S. Srinivasan, Internal report on flow modeling for laboratory simulation of field effects, CLI International, Inc., Houston, Texas, 1995.