August 2017 by Energy-Tech Magazine

VFD induced draft fan coupling failure 19 â&#x20AC;˘ Using mobile technology in the energy generation sector 23

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AUGUST 2017

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Evaluation of unintended high stresses for power piping systems operating in the creep regime

Steam and Gas Turbine Symposium Troubleshooting and Repair Options for Improved Reliability and Operational Flexibility

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Stephen R. Reid, P. E. is president and principal engineer of TG Advisers Inc. (TGA) Steveâ&#x20AC;&#x2122;s company has provided troubleshooting and condition assessments on more than 350 STG units including assessments on all nuclear units in the U.S., Spain and Belgium. TGA also provides expert witness services, failure analysis and maintenance optimization services on both steam and gas turbine units. He has numerous patent disclosures/awards, and has published more than 50 papers. He received the ASME George Westinghouse Silver Medal, is past chairman of the ASME Power Operations Committee and is a Registered P.E. He is a short-course instructor for all major conferences, EPRI and a number of OEMs.

Thomas R. Reid, P.E. is the manager of engineering for TG Advisers Inc. Tom and team provide engineering services to gas turbine, steam turbine and generator users worldwide. Tom is a graduate of General Electricâ&#x20AC;&#x2122;s Edison Engineering Development Program and has numerous patent disclosures related to turbine design and advanced repair technology. Tom holds a BSME from Virginia Tech and a MSME from Georgia Tech.

Course Description This course provides each delegate with a detailed and applied review of the most common turbine failure modes. Our instructors will focus on risk informed repair and life

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best practices and maintenance considerations more cycling, lower minimum loads and greater capacity ratings. All modules are supplemented with current case studies demonstrating the applied techniques.

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Evaluation of unintended high stresses for power piping systems operating in the creep regime

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Numerical study on flame mesoscopic characteristics of oxy-coal mild combustion Ruochen Liu, Enke An and Kun Wu Department of Mechanical and Energy Engineering, Tongji University, Shanghai, China

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EDITOR’S NOTE

Is it time for a mental health day? Did you see the recent story that has gone viral about a women who sent an email to her boss to tell him she was taking time off work to focus on her mental health? Her boss responded, thanking her for reminding him of the importance of using sick days for mental health. He went on to compliment her for helping to cut through the stigma so that employees can bring their whole selves to work. Think about that story. Whether you take a day and stay home or you take time to step away from your job to attend a conference, you are recharging your mind.You are able to return to work with a new energy, ready to tackle whatever lies ahead. I recently attended the ASME Power Conference and saw well over 2500 attendees racing around from session to session as they tried to pack as much in to their day as they could. They were, in fact, recharging their minds with new methods and processes learned from presenters and conversations as they networked with conference attendees during sessions and evening activities. I’m sure these attendees went back to work fully charged and ready to share their experience with their co-workers. It’s important to plan these recharging events into your schedule. I know everyone isn’t able to leave your job to attend a weeklong conference and for some, getting away is the only way they can separate themselves from their daily work. That’s why Energy-Tech offers both! Energy-Tech University, with its mix of online training and live conferences is a great source for you to use to clear your head and recharge with the energy needed to take your job to the next level. If you haven’t taken the time to recharge this year, check out the NEW Turbine Symposium hosted by Energy-Tech and TG Advisers coming up in October in San Antonio. This course will cover detailed and applied review of the most common turbine failure modes. Focus will be on risk informed repair and life extension strategies which have a significant bottom line impact. See page 2 for more information.

CALENDAR Aug. 22-24, 2017 Feedwater Heater Operation and Maintenance Seminar Sheraton Station Square Pittsburgh, PA. Contact Mary Jane Luddy www.powerfect.com Sept. 12-14, 2017 Turbomachinery & Pump Symposium George R. Brown Convention Center Houston, Texas http://tps.tamu.edu/ Oct.17-18, 2017 Turbine Rotor Design – Keys to Ensure Long Term Reliable and Flexible Operation Wyndham Garden, San Antonio Riverwalk San Antonio, Texas www.Energy-Tech.com/Turbine-Symp Dec. 5-7, 2017 Power-Gen International Las Vegas Convention Center Las Vegas, Nev. www.power-gen.com March 19-22, 2018 Electric Power Conference and Exhibition Gaylord Opryland Convention Center Nashville, Tenn. 2018.electricpowerexpo.com June 24-28, 2018 ASME 2018 Power & Energy Conference & Exhibition Disney’s Contemporary Resort Lake Buena Vista, Fla. www.asme.org/events/power-energy Submit your events by emailing editorial@WoodwardBizMedia.com

If attending a conference isn’t convenient for you, we have many online training courses ready for on-demand download from our website – go to the webinars and events tab at www.energy-tech.com to take advantage of these very informative training sessions. As always, if you have an idea for a new online or live training course or would like to be a presenter for a course, give me a call at 563-588-3857. I’d love to hear your ideas. Thanks for reading

Kathy Regan

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August 2017

FEATURES

Evaluation of unintended high stresses for power piping systems operating in the creep regime Since the catastrophic high energy line failures at the Mohave plant in 1985 and the Monroe plant in 1986, the electric utility industry has focused on the asset integrity management (AIM) of high energy piping (HEP) systems. An AIM/HEP program is important for the safety of plant personnel and reliability of the generating units. An as-designed piping system stress analysis is performed during the initial design of a piping system, considering the expected sustained and thermal expansion range loads. The as-designed piping system stress analysis is performed to determine the optimum locations and types of piping supports for the piping system during its operational life. This evaluation assumes that the piping system behaves as anticipated in the design configuration. However, during the actual unit operation, there may be off-design conditions, such as steam hammers, condensate issues, and unit trips. Furthermore, the supports may deteriorate over time such that they do not operate as originally designed and specified (e.g., spring hangers may migrate and become topped-out or bottomed-out). These service-related situations may create substantially different stresses that were not anticipated in the original as-designed analysis.

The ASME B31.1-2016 Code (Chapter VII, Paragraph 144) states that “The Operating Company shall develop and implement a program requiring documentation of piping support readings and recorded piping system displacements.” It also states that “Significant displacement variations from the expected design displacements shall be considered to assess the piping system’s integrity.” It is recognized that piping spring hangers seldom return to their exact original positions after each heat cycle. However, the ASME B31.1-2016 Code (Appendix V, Paragraph V-7.6) recommends that an additional evaluation should be

Figure 1

Creep is a material degradation mechanism that is stress-, time-, temperature-, and material–dependent plastic deformation under load. If a piping system is observed to displace significantly different than the expected design displacements, it is prudent to perform a piping system simulation of the observed thermal displacements in an as-found piping stress analysis. This analysis will provide a more accurate estimate of the actual weldment stresses and a more accurate ranking of the creep damage in the weldments. Figure 2 August 2017 ENERGY-TECH.com

FEATURES

Figure 3

performed by qualified personnel if there is a variation of the maximum of 1) 20% or 2) 0.5 inch from the expected spring hanger travel. If the expected travel is below 2.5 inches, the governing criterion is 0.5 inch.

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A piping displacement profile, illustrating the observed vertical pipe displacements at the pipe support locations, is compared to the expected (designed) pipe displacements in Figure 1. At Support S2, the predicted displacement is about 2.2 inches. The observed pipe displacement at Support S2 (about 0.9 inch) is about 1.3 inches less than the predicted displacement, so an additional evaluation is warranted.

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Figure 5 Figure 7

Figure 8

Figure 6

Evaluations of creep rupture curves for common power piping materials have revealed that the creep rupture times are strongly dependent on the applied stress. The range of stresses within a piping system typically vary by a factor greater than 1.7. Intertek has found that some high stress locations may have creep rupture lives less than 100,000 operating hours, the majority of locations may have creep lives beyond 250,000 operating hours, and some low stress locations may have creep lives beyond 1,000,000 operating hours.

Figure 8a

August 2017 ENERGY-TECH.com

FEATURES As an example, the number of piping system locations versus applied stress is illustrated in Figure 3. In this case, there are only a few locations with an applied stress greater than 7.5 ksi (the minimum time to creep rupture, tr,m , is at the location with an applied stress of 8.0 ksi), many locations have piping stresses between 5.9 and 6.6 ksi (creep rupture lives between 3 x tr,m and 5 x tr,m ), and only a few locations have piping stresses below 4.8 ksi (creep rupture lives beyond 10 x tr,m ). Due to the fact that the weldment creep rupture lives are highly dependent on the applied stress, there are only a few high

priority locations to examine for possible in-service related creep damage. The impact of unintended high stresses due to malfunctioning supports has typically resulted in significant ranking differences between the as-designed and simulation as-found piping stress analyses. If there is a significant unintended high stress location, that location may become the highest priority as-found location. Field conditions have also revealed that a high priority as-designed location may have much lower actual applied stress and in such case there would be no cost-benefit of frequent reexaminations.

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In some cases, the high priority as-found location may be near the 50 percentile stress in the initial as-designed stress analysis. Considering the as-designed stress rankings, 50% of the welds would have to be examined to include the maximum stress weld determined in the as-found evaluation. Risk analyses using the as-designed piping stress analysis rankings could result in expensive reexaminations of locations with negligible creep damage. Intertek has also found that some piping system girth welds may actually have very low applied stresses. We have recently completed a project in which a dissimilar metal weld (DMW) had a minimum wall thickness about 30% below the ASME B31.1 Code allowable. During a 100,000 operating hours scheduled outage, this DMW was examined and had no creep damage. A fitness-for-service evaluation was performed to justify a couple of additional years of operation so that a thicker reducer fitting replacement would be available during the next scheduled outage. Using a Life Fraction Rule DMW evaluation that is more conservative than the API 579 requirements, it was determined that the remaining creep rupture life of the DMW was greater than 10 years and there was ample justification for continued operation until the next scheduled outage. Over the past 15 years, Intertek has been evaluating unintended high stresses in power piping systems operating in the creep regime. The field stresses are determined by modelling the pipe behavior observed during on-line and off-line piping system walkdowns. Both the elastic and inelastic stresses are considered to evaluate the axial and circumferential stress August 2017

FEATURES redistributions over the life of the piping system. Creep life consumption is calculated using the Life Fraction Rule as the stresses redistribute over time. This methodology, calibrated to historical girth weldment creep rupture failures, has been used to determine the top priority locations for the next set of girth weld examinations. To provide a greater reliability for an AIM/HEP program, it is recommended that a one-time examination of all weldments be performed to capture any significant fabrication flaws, operational problems, condensate management issues, and possible in-service material damage. Several examples of successful predictions of girth weld creep damage are illustrated in Figures 4 through 8. Figure 4 illustrates creep damage downstream of a main steam header connection at 158,000 operating hours. Figure 5 is an example of creep damage predicted and confirmed downstream of a main steam reducer fitting at 122,000 operating hours. Figures 6 and 7 are photomicrographs of microstructural creep damage predicted and confirmed at 68,000 operating hours. Figure 8 illustrates creep damage propagating subsurface (revealed at 163,000 operating hours). This AIM/HEP multidiscipline methodology has been used successfully to identify locations of high priority creep damage and approximately when the level of creep damage is sufficient to be detectable by nondestructive examination techniques.

The identification and mitigation of unintended high stresses can reduce risk, increase personnel safety, and increase system reliability. ■ Disclaimer Although the information presented in this work is believed to be reliable, this work is published with the understanding that Intertek and the author are supplying general information and are not attempting to render or provide engineering or professional services. Neither Intertek nor any of its employees make any warranty, guarantee, or representation, whether expressed or implied, with respect to the accuracy, completeness or usefulness of any information, product, process, method or apparatus discussed in this work, including warranties of merchantability and fitness for a particular or intended purpose. Neither Intertek nor any of its officers, directors, or employees shall be liable for any losses or damages with respect to or resulting from the use of, or the inability to use, any information, product, process, method or apparatus discussed in this work. Editor’s note: References are made to the code, ASME B31.1-2016 Code and API 579-1/ASME FFS-1 2016. To find more information, visit the ASME Digital Store at www.asme.org. Marvin J. Cohn is the Director of Piping Asset Integrity Management at Intertek AIM in Santa Clara, Calif. His primary assignments with Intertek AIM have been in field engineering services with a strong technical interest in asset integrity management of power piping systems. He is been a member of the ASME B31.1 Power Piping Section Committee for over 10 years and is a Registered Professional Engineer in several USA states and two Canadian provinces. He is an ASME Fellow and has written more than 50 technical papers. You may contact him by emailing editorial@WoodwardBizMedia.com

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ASME: Numerical study on flame mesoscopic characteristics of oxy-coal mild combustion Ruochen Liu, Enke An and Kun Wu Department of Mechanical and Energy Engineering, Tongji University, Shanghai, China

Abstract For achieving efficient oxy-coal combustion in a MILD (Moderate or Intense Low Oxygen Dilution) state, the optimum operating conditions with high-velocity jets in a labscale cylindrical furnace (δ200mmx2000mm) was determined. The mesoscopic characteristics of turbulent and flame behavior under different jet design and jet spacing were simulated and compared. The results show that L=30~60mm(O2 side) and L=60mm(O2 center) conditions are recommended as oxy-coal MILD combustion as well as IFRF furnace condition, the flame front locates in distributed regime, the global regime was depict as 1<l/lF <4, 60<ReT F<150 and 50<Ka<500; for flaming conditions, the flame front locates in small-scale turbulent regime or thin reaction zone, the global regime was depicted as 0.5<l/lF <4, 40<ReT<110 and 30<Ka<900; with high-velocity oxygen jet technology, the combustion process is in slow chemistry regime (Da<<1), governed by chemical-kinetic mechanism; large spacing (L-75mm) is not favored for co-flow burners due to poor radial mixing as well as the restriction of wall. Keywords: oxy-coal combustion, MILD combustion, flame regimes, numerical analysis Nomenclature E global activation energy/ kJ • mol-1 T inlet temperature/K in Cp specific heat/ J • kg-1 • K-1 W Molecular weight/ kg • mol-1 f Q reaction heat/ kJ • kg-1 Y mass fraction of inlet fuel/% f u’ turbulent intensity/m • s-1 l average integrated length/m lF laminar flame thickness/m SL laminar burning velocity/m • s-1 ST turbulent burning velocity/m • s-1 u velocity/m • s-1 T time/s k turbulent kinetic energy/m2 • s-2 E dissipation rate/m2 • s-3 v kinematic viscosity/m2 • s-1 Subscripts K Kolmogorov eddy C chemical reaction T turbulent flow δ inner layer August 2017 | ASME Power Division Special Section

Introduction Recently, the improvement of energy efficiency, reduction of NOx emission and carbon capture and storage (CCS) are highlighted in energy research field. Oxy-fuel combustion is one of the most promising technologies to realize CCS, and MILD combustion is known to produce low NOxemission and high combustion efficiency. In oxy-fuel combustion, the flame characteristics will change significantly. However, due to the complexity of gas solid interaction, the researches focus on gaseous fuel more than pulverized coal so far. X.J. Zhong et al.[1] studied the laminar flame speed of C1~C7 n-alkanes by detailed chemical kinetic models for the atmosphere of O2/CO2, O2/N2 and O2/Ar respectively, revealing that CO2 had a negative effect on flame propagation, while higher O2 concentration was beneficial for n-alkanes combustion process. J.H. Wang et al. [2] detected the convex and concave structures of the flame front in methane combustion with CO2 dilution. The local radius of curvature, fractal inner cutoff scale and local flame angle were calculated from the OH-PLIF images. The results showed that flame front at high temperature and high pressure was a wrinkled flame front, in the case of CO2 dilution, the fractal inner cutoff scale decreased, the flame brush became thicker and the mean flame volume enlarged.Yuen et al. [3] studied methane and propane turbulent flame stabilized on a buesen type burner using planar Rayleigh scatting and particle image velocimetry. The results showed that the non-dimensional turbulence rms velocity covered from 3 to 24, corresponding to conditions of corrugated flamelets and thin reaction zones regime; the flame front thickness increased as non-dimensional turbulence rms velocity increased.Y.Minamoto et al. [4] studied the flame structure and flame interaction through three-dimensional direct numerical simulation (DNS) of methane turbulent combustion with CO2 dilution in MILD state. The results suggested that the commonly used canonical flamelets was not fully representative for MILD combustion; the reaction rate and the scalar gradient can be used as a marker for the flame interaction. MILD combustion was developed from excess-enthalpy combustion of F.J Weinberg. Oxy-fuel MILD combustion is a novel technology, owning three major advantages: the stability of oxy-coal combustion can be guaranteed; the efficiency of heat transfer develops; NOx emission can be further reduced. ENERGY-TECH.com

ASME FEATURE The key to maintain MILD state is that the energy provided by the oxidizer jets, with a high preheating temperature and low oxygen dilution, is greater than the energy loss by radiation to supply enough energy for ignition of fuel [5]. Therefore, preheating oxidizer jet to high temperature and developing oxidizer jet velocity are generally applied to establish oxy-fuel MILD combustion. By far, the researchers developed some typical methods to judge MILD combustion state. A.Cavaliere [6] proposed that a combustion process is named MILD combustion when the inlet temperature of the reactant mixture is higher than the mixture self-ignition temperature whereas the maximum allowable temperature increase with respect to inlet temperature during combustion is lower than mixture self-ignition temperature. M.Oberlack et al. and N.Peters [7, 8] thought that MILD state is established in swirl chamber when (Eq1)

S.Kumar et al. [9] thought that MILD state is established when temperature fluctuation is within 15%. G.Szego et al. [10] thought that MILD combustion is characterized by no substantial flame front and low CO, NO emissions (<100ppm). Wunning et al.[11] thought that MILD state would be achieved when the furnace temperature and flue gas recycle ratio are high enough. Here, oxy-coal MILD combustion was achieved with high-velocity oxygen entrainment technology in a cylindrical furnace (δ200mmx2000mm). Combustion characteristics under different jet designs and jet spacing were simulated using the commercial software package FLUENT. The modeling methodology has been validated with the experimental data from the 0.58MW IFRF furnace (6.25mx2mx2m). The mesoscopic characteristic parameters were calculated, and the specific flame regime was worked out.

ASME Power Division: Fuels & Combustion Technology

A Message from the Chair Electric power is vital to our homes and businesses, our local communities and nation, our way of life. The ASME 2017 Power and Energy Conference focused on every aspect of power, including generation, transmission, distribution and consumption. Improvement of standard techniques and devices as well as the emergence of new ways to generate and harness energy were on display at the conference through vendor exhibits, keynotes, panels, and presentations on the most recent R&D developments in the power industry. The Computational Fluid Dynamics and Thermal Hydraulics track for 2017 focuses on analysis of every part of power systems from nanofluids to whole plants, from fossil fuels to renewables, from modern techniques like finitedifference/finite-elements to all-new methods developed for specific systems, and from development of theory to analysis of operating systems. These R&D advances are being used today in modeling and design of the power systems of today and tomorrow. Dr. George L Mesina, Division Fellow Idaho National Laboratory George.Mesina@inl.gov Phone: 1-208-526-8612

1 Turbulent flame regimes 1.1 Danker’s theory In 1940s, German scholar Danker [12] supposed that turbulent flame is distinguished into small-scale turbulent flame corresponding to l/lF<1 and large-scale turbulent flame corresponding to l/lF>1 through comparing the laminar flame thickness lF and turbulent eddies’ average integrated length. Large eddies can play an important role due to their large circulation and long residence time. For small-scale turbulent flame, the mass transportation within flame front is promoted, flame front just thickens rather than being deformed greatly. For large-scale turbulent flame, it is further distinguished by u' ‹/ S=1. When u' ‹/SL<1, it is weak large scale turbulent flame, the turbulent intensity is not large enough to compete against the laminar flame propagation, so flame front is wrinkled rather than broken up; otherwise it is strong large-scale turbulent flame. Fig.1 shows the turbulent flame regime based on Danker’s theory. 12 ENERGY-TECH.com

Figure1: Turbulent flame regime based on Danker’s theory

ASME Power Division Special Section | August 2017

ASME FEATURE 1.2 Schelkin’s theory However, for the strong large-scale turbulent flame, Danker’s theory was not confirmed by experiment. The reason is that the strong mixing process between products and fresh mixture is neglected in Danker’s theory. In 1950s, Schelkin [12] proposed the theory of volume combustion, which emphasizes on the interaction between eddies. It considers that in case of strong turbulent intensity, the micro-diffusion is rapid, some eddies are easily broken up into more new eddies within their lifetime. For each eddy, the turbulent intensity, migrating distance, ignition delay is all different, so there is no obvious front distinguishing the unburned and the burned. It indicates that chemical reaction happens in a more distributed area rather than limited to the flame front. In terms of the strong large scale turbulence, the comparison between Schelkin’s theory and Danker’s theory is displayed in Fig.2.

lent intensity and laminar burning velocity. When Schmidt number is assumed as unity in reaction zone, namely Sc=v/D-1, so

Figure 2: The comparison of two theories

1.3 Kalrovitz number To identify what the structure of flame front exactly is, Karlovitz introduced a non-dimensional number Ka, in addition, turbulent Reynolds number ReT and Damkohler number Da are also significant. According to literatures [13, 14], the formulas are presented in Table 1, and the flame regime based on non-dimensional numbers is displayed in Fig.3. Table 1 The definitions of non-dimensional numbers

In Fig.3, the abscissa is the ratio of integral length scale and laminar flame thickness, and the ordinate is the ratio of turbuAugust 2017 | ASME Power Division Special Section

Figure 3: The turbulent flame regimes

As mentioned above, Danker’s theory dominates for weak large-scale turbulent flame corresponding to l/lF>1 and u’/ SL <1 ; Schelkin’s theory dominates for strong large-scale turbulent flame corresponding to l / lF >1 and u'›/ SL>1. In the strong large-scale turbulent flame regime, the line Ka=1 means the laminar flame thickness is equal to the Kolmogorov length, called the Klimov-Williams criterion. When Ka<1, the flame front is embedded within Kolmogorov eddies, therefore the flame structure is not disrupted by turbulent fluctuations and remains quasi-laminar. When Ka>1 and Kaδ <1, namely lδ < lK< lF, it is the thin reaction regime. Usually the flame structure is consisted of preheat zone and reaction zone. The first layer of reaction zone is the inner layer where fuel is consumed and the radicals are depleted by chain-breaking reactions. The thickness of inner layer is l δ , so Kaδnumber is (EQ12)

The thickness of inner layer is estimated as one tenth of laminar flame thickness, namely l δ= lF /10 [14]. So the line Kaδ =1 is equal to Ka=100. In the thin reaction regime, the Kolmogorov eddies enter into the preheat zone but not the reaction zone, so the preheat zone is broadened while the reaction zone is thin. When the Kolmogorov length is even ENERGY-TECH.com

ASME FEATURE smaller than laminar flame thickness, namely Kaδ > 1, it is the distributed regime, where Kolmogorov eddies can penetrate into inner layer and disturb it so that the flame front is broken up. This regime is easily recognized that flame extinguishes and combustion ceases. However, it is like a volume combustion mode, the heat release is more homogeneous rather than confined to the flame front.

Table 5 Setting of mathematical models

2. Numerical simulations 2.1 The burners The furnace employed is used for measuring burnout and slagging characteristics of coal in Shanghai boiler works limited Co., which is a cylinder with an inner diameter of 200mm and length of 2000mm. The primary inlet is low velocity CO2 jet carrying pulverized coal, and the secondary inlet is a highvelocity pure oxygen jet. Three kinds of burner are applied in the simulation as Table 2 shows, and their dimensions are presented in Table 3. The inlet velocity is 20 m•s-1 for primary jet and 100 m•s-1 for secondary jet, the momentum ratio is approximately 4. The primary jet is preheated to 573K. The coal mass flow rate is 0.0005 kg•s-1 , the excess of oxygen is 20%, and the solid-gas ratio of the primary wind is 0.35kg/kg. Proximate and ultimate analysis of Datong bituminous coal employed is presented in Table 4. 2.2 The grid Due to the symmetry of the furnace, a quarter of the geometry was created using the GAMBIT 2.4.6 preprocessor. To ensure the accuracy of modeling results and least calculation time, the minimum number of hexahedral cell employed is approximately 1760000 for co-flow condition, including 65000 cells in wall faces, 101000 cells in symmetry faces, 13 cells in secondary inlet face and 32 cells in primary inlet face. Besides, the minimum cell number is 2130000 for coaxial condition, including 81000 cells in wall faces, 119000 cells in symmetry faces, 11 cells in secondary inlet face and 36 cells in primary inlet face. 2.3 Mathematical method The mathematical models employed are set as shown in Table 5. Coal particle trajectory was simulated by the discrete phase model (DPM). A total of 3200 coal particle trajectories were tracked at each step of the DPM coupling. The particle size was assumed to obey the Rosin-Rammler distribution with a spread parameter of 1.36 and 10 groups. The temperature of the furnace wall was set as 1273K and the internal emissivity as 1. In addition, convection terms were simulated with the firstorder upwind difference scheme, and SIMPLE algorithm was applied for coupling between velocity and pressure. The models were verified by the experimental data (including velocity, temperature, species concentration) of 0.58MW IFRF furnace under coal MILD combustion [16].

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Figure 4: Temperature fields on the X-Z section

3. Results and discussions 3.1 Temperature fields During the ignition delay, there is an internal mixing between coal flow, oxidizer flow and recirculated flue gas. As shown in Fig.4, the temperature peak decreases slightly as jet spacing increases to 60mm. Because the radial mixing between coal flow and oxidizer flow is so fast that the oxidizer flow is not well diluted internally under narrow jet spacing, like coaxial jet condition, the ignition delay is short, local reaction zone is obvious near the inlet, the peak temperature is highest as 2200K approximately. It should be noted that it is not the larger spacing the better, due to poor radial mixing between coal flow and oxidizer flow, as well as the restriction of wall on the inner recirculation, L=75mm is not favored in neither case. 3.2 Flame mesoscopic characteristics The non-dimensional numbers can be obtained through equations (5), (10), (11). As is known, flame front is a narrow transition zone of two gas masses with different physical ASME Power Division Special Section | August 2017

ASME FEATURE

Table 2.3.4

properties, where the horizontal gradient of temperature is largest, marked in red line in Fig.5. The mesoscopic characteristic parameters in the flame front are so significant that should be considered in first. Taking IFRF furnace condition as example, its maximum gradient of temperature is 500 K m-1.The mesoscopic characteristic parameters on the red line were calculated, as listed in Table 6.

Figure5: Maximum temperature gradient on the X-Z section under IFRF furnace condition

From Table 6, in terms of flame front where temperature gradient is largest, ReT is around 40~80, Ka is higher than 100, indicating that Kolmogorov eddies can penetrate into inner layer, which facilitates diffusion and heat transfer rate to the reaction zone, so that combustion temperature is uniform and the flame front is broken up. Affected by the strong inner recirculation of flue gas, TC increases while TT decrease, Da is even smaller than 0.1, so the combustion process is governed by chemical-kinetic mechanism. The calculated values of each condition were all drawn in Fig.6. Taking line l / lFand Kaδ=1 as boundaries. The scatters of L=30~60mm (O2 side) and L=60mm (O2center) conditions marked in red, all locate where l / lF>1 and Ka >1, namely distributed regime. For IFRF furnace condition, the scatters also locate in distributed regime, which has been proved experimentally to be pulverized coal MILD combustion. It confirms that the theory of turbulent flame regime is appropriate for pulverized coal combustion. Flame front usually forms during homogeneous stage, so the homogeneous stage of MILD combustion should locate in the distributed regime corresponding to l/lF1 , ReT>1 , Kaδ >1 and Da<1. ReT>1 means combustion occurs under high-intensity turbulence; l/lF >1 means it presents large-scale turbulent August 2017 | ASME Power Division Special Section

Figure 6: The mesoscopic characteristic parameters of the flame

flame; Kaδ >1 means that Kolmogorov eddies can penetrate into inner layer so that the flame front is broken up. Da<1 means chemical reactions rates is slower than turbulent mixing rate, so the combustion process is governed by chemical-kinetic mechanism. However, for other conditions, the scatters locate in small-scale turbulent regime or thin reaction zone, namely flaming combustion. L=75mm is not favored for co-flow burners, due to poor radial mixing between coal flow and oxidizer flow, as well as the restriction of wall on the inner recirculation. Furthermore, the distribution of Ka number and the ratio of mesoscopic length scale on X-Z section were shown in Figs.7 and 8. l/lF =1 and Ka=100 were marked as boundaries in bold lines. From Fig.7, as the z-axial distance increases, the turbulent kinetic energy weakens, so Ka number decreases, even smaller than 100 in the downstream from 1/3~2/5 furnace height where the char heterogeneous combustion may dominate. While the homogeneous combustion may dominate in the area of Ka>100 , which is relatively broader under MILD conditions than that under flaming conditions. It indicates that for MILD condition, the Kolmogorov eddies of lK<lδ diffuse to the ENERGY-TECH.com

ASME FEATURE wider space, so the flame front is more easily broken up by Kolmogorov eddies. For MILD conditions, the ratio of mesoscopic length scale l/lF is all beyond unity, indicating it locates in large-scale regime. However, for flaming conditions, there is a narrow area corresponding to l/lF <1, as shown in Fig.8. When the flame front occurs in this narrow area, it should present smallscale turbulent flame. The mass transportation within flame front is promoted, while flame front just thickens rather than being deformed greatly.

Table 6 The mesoscopic characteristic parameters of the flame front under IFRF furnace condition

In Figs.7 and 8, except IFRF furnace MILD condition, the distribution of mesoscopic characteristic parameters under other conditions is corresponding to z-axial distance not beyond 1m. The global specific regime was worked out as shown in Fig.9.

Figure 8: The distribution of l / lF on the X-Z section

For each condition, due to strong turbulent mixing and entrainment, the combustion process occurs in high-intensity turbulence ( ReT>1) and in slow chemistry regime( Da<<1 ), governed by chemical-kinetic mechanism. For MILD combustion conditions of L=30~60mm (O2 side) and L=60mm (O2 center), the regime was furtherly depicted almost as 1<l/ lF <4, T 60<ReT <150 and 50<Ka<500 , marked in red; for flaming conditions, the regime was depicted as 0.5 l/lF <4 , 40<ReT <110 and 30<Ka<900 , marked in blue.

Figure 7: The distribution of Ka number on the X-Z section

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ASME Power Division Special Section | August 2017

ASME FEATURE

Figure 9: The specific turbulent flame regime

4. Conclusions The mesoscopic characteristics of turbulent and flame behavior under different jet design and jet spacing were simulated and compared. It concludes that L=30~60mm(O2 side) and L=60mm(O2 center) conditions are recommended as oxycoal MILD combustion as well as IFRF furnace condition, the flame front locates in distributed regime, the global regime was depict as 1<l/lF <4, 60<ReT <150 and 50<Ka<500; for flaming conditions, the flame front locates in small-scale turbulent regime or thin reaction zone, the global regime was depicted as 0.5<l/lF <4, 40<ReT <110 and 30<Ka<900: large spacing(L=75mm) is not favored for co-flow burners due to poor radial mixing as well as the restriction of wall. ■

5. X. Lv> Investigation of flameless combustion technology for hydrogen-rich fuels in micro gas turbine. Chinese Academy of Science, 2010. 6. Antonio Cavaliere,Mara de Joannon. MILD Combustion [J]. Progress in Energy and Combustion Science, 2004, 30(4): 329366. 7. Martin Oberlack, R. Arlitt,Norbert Peters. On stocastic Damkohler number variations in a homogeneous flow reactor [J]. Combust Theory Model, 2000, 4: 495-509. 8. Norbert Peters. Principles and potential of HiCOT combustion. Proceedings of the Forum on High-Temperature Air Combustion Technology. 2001. Hosui Kaikan. 9. Sudarshan Kumar, P. J. Paul,H. S. Mukunda. Studies on a new high-intensity low-emission burner [J]. Proceedings of the Combustion Institute, 2002, 29(1): 1131-1137. 10. [G. Szego, B. Dally,G. Nathan. Operational characteristics of a parallel jet MILD combustion burner system [J]. Combustion and Flame, 2009, 156(2): 429-438. 11. J. A.Wunning,J. G.Wunning. Flameless oxidation to reduce thermal NO-formation [J]. Prog. Energy Combust. Sci, 1997, 23: 81-94. 12. K.F.Cen, Q.Yao, Z.Y.Luo. Advance combustion theory [M]. Zhenjiang University Press.2002 13. Nedunchezhian Sanminathan,K. N. C. Bray.Turbulent Premixed Flames [M]. UK: Cambridge University Press. 2011. 14. Norbert Perters.Turbulent Combustion [M]. UK: Cambridge University Press. 2004. 15. E.K. An, X. Feng, L.J. Zhang, et al.The numerical simulation of pulverized coal MILD combustion. Journal of Tongji University, 2014, 42(7): 11055-1110

Acknowledgements The work has been carried out with the opening project of Key Lab of Power Engineering Multi-flow and Heat Transfer of Shanghai, which is financed by Shanghai Science and Technology Committee (NO.13DZ2260900). The authors acknowledge thanks with the financing. References 1. X.J.Zhong, H. Liu, R. Zhao, et al. Analysis of influence factors on flame speed under O2/CO2 atmosphere. Proceedings of the CSEE, 2011, 31(23):54-60 .Jinhua Wang, Futoshi Matsuno,Masaki Okuyama, et al. Flame front characteristics of turbulent premixed flames diluted with CO2 and H2O at high pressure and high temperature[J]. 2. Proceedings of the Combustion Institute, 2013, 34(1): 14291436. 3. Frank T. C.Yuen,Ömer L. Gülder. Premixed turbulent flame front structure investigation by Rayleigh scattering in the thin reaction zone regime [J]. Proceedings of the Combustion Institute, 2009, 32(2): 1747-1754. 4. Y. Minamoto,T. D. Dunstan,N. Swaminathan, et al. DNS of EGR-type turbulent flame in MILD condition[J]. Proceedings of the Combustion Institute, 2013, 34(2): 3231-3238.

August 2017 | ASME Power Division Special Section

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VFD induced draft fan coupling failure By Patrick Smith

Torsional vibration problems in rotating machinery can be difficult to recognize. Unlike radial vibration which can be easily measured with readily available sensors, torsional vibration is more difficult to measure because it involves the twisting of shafts while the machine is rotating. It can typically only be measured with special devices such as strain gauges or torsional lasers. As a result, torsional problems typically go unnoticed until something fails. Variable frequency drives (VFDs) are commonly used to vary the speed of various types of rotating machinery to efficiently control the capacity. Typical machinery applications include pumps, fans, blowers and compressors. However,VFDs can induce dynamic torques which can excite torsional natural frequencies, leading to undetected failures. The purpose of this article is to present a case study of a VFD driven induced draft (ID) fan that suffered several torsional failures believed to have been caused by unusual coupling wear that resulted from dynamic torques generated by the VFD. The failure, investigation and corrective action will be discussed.

Introduction This case study pertains to single stage induced draft (ID) fan driven by a 1185 RPM, 450 HP induction motor. The motor is connected to the fan by a disc pack type spacer coupling. The fan arrangement is a dual width, dual inlet (DWDI) configuration. The impeller is mounted on a shaft between two pillow block type bearing housings. The pillow blocks are bolted to steel pedestals which are mounted on a concrete foundation. The non-drive end (NDE) bearing is a split, Babbitt lined

sleeve type journal bearing. The drive end (DE) bearing consists of a split Babbitt lined sleeve type journal bearing and split, Babbitt lined fixed pad type thrust bearings. The fan rotor includes a single thrust collar that is integral with the fan shaft. The fan arrangement is shown in Figure 1. The fan bearing housings are fitted with a single velocity type transmitter mounted in the horizontalÂ direction which is used to continuously monitor vibration. The fan journal bearings are fitted with thermocouples which are used to continuously monitor bearing temperatures. In addition, the motor bearings are fitted with temperature probes and the stator windings are fitted with temperature probes which are used to continuously monitor these temperatures as well. The equipment condition monitoring system signals an alert if the bearing temperatures and/or vibration levels exceed predicted values, which vary based on actual operating conditions. The machinery protections include high bearing vibration alarms, high bearing temperature alarms, and a high motor amp alarm and trip. The machinery protection set points used fixed values. The ID fan bearings, seals and coupling are typically inspected during major plant outages that occur every few years. The ID fan is used in a steam methane reformer plant that produces hydrogen. The ID fan capacity varies with plant rates. This is done by varying the ID fan speed in order maintain a constant vacuum pressure in the reformer. The other type of ID fan capacity control that is commonly used is inlet dampers. However, varying the speed using a VFD is more efficient and results in a power savings. There are some installations that

Figure 1: Compressor configuration August 2017 ENERGY-TECH.com

MACHINE DOCTOR include a VFD and dampers. This allows for more operating flexibility, particularly at start-up and at high levels of turndown. However, this is costlier and more complex. In this application, only a VFD is used for capacity control.

History After the ID fan was commissioned it was put into continuous service. The fan initially ran at around 1000 RPM, but after several years, plant rates changed and the fan often ran in the 550 to 600 RPM range. There were no significant fan reliability issues until a coupling bolt failure was discovered during a routine coupling inspection after 9 years of operation. No other damage was uncovered and there were no changes in fan vibrations leading up to the plant outage when the broken bolt was discovered. A picture of the failed bolt is shown in Figure 2. A year later there was a catastrophic coupling spool failure. See Figure 3. The DE fan vibration had been trending higher for about 6 weeks leading up to the failure.

Figure 2

Based on the unexplained failures, it was decided to review the coupling design and revisit the torsional analysis.

Investigation Information on the coupling is shown in Table 1. Based on the motor rating, the coupling has a service factor of 2.3. For a fan application without a VFD, the coupling manufacturer’s catalogue recommends a minimum service factor of 1.5. For VFD applications, this would be increased by 0.5, bringing the recommended service factor to 2.0. Since the actual service factor is greater than this, the coupling size was deemed acceptable.

Table 1

The torsional analysis was redone and performed without VFD excitations. Despite repeated requests, the VFD manufacturer was unable to provide the VFD generated torque excitation data. The undamped torsional natural frequency (TNF) for the first mode was determined to be 1455 cpm, which is 21% above the rated motor speed. This first mode is characterized by twisting that occurs primarily in the coupling. However, this is well above the 10% separation margin required by API-673. Excitations from 1X running speed, 2X running speed, 1X line frequency, 2X line frequency and fan vane passing frequency were used in a damped response analysis. This analysis did not uncover any flaws in the mechanical design or identify a cause for the observed coupling failures. Thus, it was concluded that the observed coupling failures were most likely not due to any “conventional” torsional excitations. Instead, it was believed that the failures were due to excitations generated by the VFD. The VFD works by first taking AC and rectifying it to DC. The DC is then filtered and connected to a network of power transistors which inverts the DC into AC. The control method is known as Pulse Width Modulation (PWM). This works by switching the DC on and off very quickly using transistor switches to create a sine wave of motor current. For this application, the drive was configured to operate at a constant volts per hertz ratio. This is common for fan drives. However,VFDs can create torque pulsations, which can excite the TNF. This was the suspected cause of the observed coupling failures. Since the VFD manufacturer was unable to provide the frequency and amplitude of torque pulsations needed for the torsional rotordynamics analysis, it was decided to perform a field test. Kelm Engineering was contracted to take static and dynamic torque measurements on the ID fan drive train over the entire operating speed range. Kelm Engineering is a consulting firm specializes in analytical and field testing of turbomachinery.

Figure 3

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August 2017

MACHINE DOCTOR

Figure 4

Kelm attached a pair of temporary strain gauges with wireless transmitters to the motor shaft end to measure the torque. In addition, Kelm used a laser tachometer for speed indication. The gages were mounted on the motor shaft rather than on the coupling spacer due to the limited axial space available for instrumentation between the spacer flanges. The gages were powered by wireless radio transmitters mounted on the shaft with a wireless receiver near the machine. Both static and dynamic torques were measured from start-up to the minimum speed of 390 RPM. The fan was then accelerated to the maximum speed of about 1192 RPM using the VFD to increase the speed over a period of approximately 30 minutes. The speed was then decreased to the minimum operating speed for repeatability of the measurements. The motor was then tripped from the minimum operating speed in an attempt to excite the TNF.

The first mode TNF, based on the ring down plot of the torque verses time after the motor was tripped was determined to be approximately 1411 CPM. This correlated well with the calculated TNF. The dynamic torque varied from 210 to 4,310 in-lbs. during the testing. The highest sustained dynamic torque occurred at about 686 RPM, though the overall amplitude was low. Trends of the measured static and peak dynamic torque verses time as well as speed are shown in Figure 4. As the VFD incremented the speed, the measured static torque showed a step change that was typically accompanied by a short duration increase in dynamic torque. The maximum dynamic torque occurred at a speed of 999 RPM and the spectrum showed a peak at a frequency corresponding to the first mode TNF identified in the ring-down time waveform plot. After the completion of the test, coupling damage was discovered. The damage included a couple of failed bolts. The coupling assembly was then sent to the coupling manufacturer for evaluation. In addition to the failed bolts, the coupling manufacturer found multiple fractured washers, and several fractured discs. It was also observed that there was significant fretting around some of the bolt holes in the adapter, center member, and discs. See Figure 5. This type fracture was observed on approximately half of the discs in the pack. This type of fracture, which progresses radially through the center of a bolt hole, is indicative of a loose fastener condition. The fretting around the holes in the

Figure 5 August 2017 ENERGY-TECH.com

MACHINE DOCTOR adapter, spacer, and discs also supports loose fastener conditions were a contributing factor in the failure. In addition, examination of the fracture surface of the bolt indicated a low cycle fatigue condition which suggests that the bolt failure occurred in a short duration and most likely after the loose bolt condition.

Discussion This failure was not a result of excessive torsional vibration, such as the one reported in the paper by Feese and Maxfield (reference 1). In this failure, the dynamic torque exceeded the transmitted torques in certain speed ranges. In the testing that was done on the fan that is the subject of this case study, the dynamic torque levels are relatively low and the coupling has sufficient service factor. The coupling is a very mature coupling design and this style coupling has been used in many VFD applications. A close inspection of the coupling did not reveal any manufacturing defects or deficiencies. However, it appeared that the clamping force used to drive the disc packs was insufficient and led to disc fretting damage which then led to loosening of the coupling bolts and then to the coupling bolt and spacer failures. There were no coupling failures in the early years of operation when the fan tended to operate at higher speeds. Later, when the fan speed was run more often in the 550 to 600 RPM speed range, the failures seemed more prevalent. Although the dynamic torque levels are relatively low, the dynamic torque levels as a percentage of the static torque

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are higher than at higher speeds. In the worst condition, the dynamic torque is as much as 38% of the static torque. This is higher than typically recommended and it is possible that in this condition, there were minor torsional displacements of the discs, which led to the fretting damage that seems to be first step in the failure mechanism. The coupling bolt torque was reviewed with the coupling manufacturer and can’t be increased. It may be possible to change to higher strength bolts to allow for higher torques to increase the clamping force. Other possibilities include changing to a larger disc pack type coupling or changing a torsionally resilient coupling, such as a rubber block or viscous shear type. As described in the paper by Feese and Maxfield, changing to a larger coupling could simply cause something else to fail, especially when the cause is not well understood. The one option that has the most potential to solve this problem is to change to a torsionally resilient type coupling. These couplings have a lot of torsional damping and based on the field tests, the overall dynamic torque levels are low and well within the capability of these type of couplings.

Conclusion It is not likely further testing will help with the understanding of the failure mechanism. While it may be possible to make VFD configuration changes, the VFD manufacturer has not been helpful in the failure investigation. So, this might end up being a trial and error process. It is also possible to change to a different VFD manufacturer or model. However, changing to a torsionally resilient type coupling is a simple change and has a high chance of solving the problem. Based on this, the decision was made to change to a torsionally resilient type coupling. A couple of different coupling types are being evaluated. It will be necessary to do a torsional rotordynamic analysis as part of the retrofit. VFDs can also cause torsional vibration problems. And, these can be difficult to analyze. In some cases, field testing is needed. Even though the testing led to inconclusive results, it did show that high dynamic torque was not the problem. This provides the confidence that installing a more tolerant coupling will solve the problem. ■

References 1. Feese, Troy and Maxfield, Ryan, “Torsional Vibration Problem With Motor/ID Fan System Due to PMW Variable Frequency Drive, Proceedings of the Thirty-Seventh, 1985, pp 45 - 56. 2. API-673, “Centrifugal Fans for Petroleum, Chemical and Gas Industry Services,” 3rd Edition, December 2014, API, Washington DC. 3. Pavelek, Dustin, PE, “ID Fan Torsional Test,” Report Number: 11727- Report, October 4, 2016 Patrick J Smith is lead machinery engineer at Air Products & Chemicals in Allentown, Pa., where he provides technical machinery support to the company’s operating air separation, hydrogen processing and cogeneration plants. You may contact him by emailing editorial@ WoodwardBizMedia.com August 2017

MAINTENANCE MATTERS

Using mobile technology in the energy generation sector By Matthew Buck, Electric Power Research Institute

For decades, work management in the energy generation sector has been completed using paper-based data collection and reporting. Since the emergence of handheld mobile technology, such as smartphones and tablet computers, utility adoption of these devices has been slow. However, given the power, versatility, and productivity gains that mobile devices and software have demonstrated in other industries, utilities need to consider the near-term adoption of mobile solutions for fossil fuel generation plant operations and maintenance. A number of factors are playing a role in the introduction of this new technology to the industry. The utility industry workforce is transitioning to a cadre of younger, tech-savvy professionals. Wireless voice, data, and video connectivity is becoming prolific, and the skills to exploit these resources are becoming more abundant. The new world of this “digital worker” is transforming what is possible. Mobile devices are increasing in power and decreasing in size. The capabilities and performance of mobile applications supporting production field crews and management are improving. These factors raise a number of questions: What are the challenges to introducing mobile technology to generation facilities? Which mobile options offer the most advantages?

What applications in generation facilities are best suited to mobile technology? How can the technology be implemented in generation facilities in the most practical and efficient way? In recent years, the Electric Power Research Institute (EPRI) has conducted research in mobile technology to explore how these exciting new digital trends can be applied in generation operations and maintenance (O&M) use cases. The goal is to filter through the many options, determine what is available and practical, and then guide utilities in their evaluation and adoption of these powerful tools. (See Figure 1.) Among the activities of this research has been a survey of utility members of EPRI’s Generation Sector to assess their current use of mobile technology. As a practical tool for electric utilities, EPRI has also developed a Mobile Technology Guidebook: 2016 Update (EPRI report no. 3002008762), which compiles descriptions of handheld devices and software applications, identifies challenges to use of mobile technology in the utility industry, and provides strategies for implementing practical pilot programs. In addition, EPRI launched and held the first meeting of a Digital Worker Interest Group (DWIG), which offers an opportunity for members to discuss challenges and share experiences regarding the integration of mobile technology into their facilities. This article provides highlights of the research. Mobility survey In 2016, the EPRI research team conducted a survey of EPRI Generation Sector members to evaluate the role of mobile technology in their organizations. The survey consisted of ten questions covering the implementation of mobile technology, issues encountered, challenges for work crews, concerns regarding integration, and software applications in use. According to answers provided by 37 respondents, power generation organizations are progressively adopting mobile solutions. Among the results:

Figure 1: Mobile technology options for today’s digital workers. August 2017 ENERGY-TECH.com

MAINTENANCE MATTERS

Figure 2: Developer making 3D scan of boiler

• Usage. 60% of respondents indicate that both company executives and end users used some type of mobile device for limited job functions, with 50% indicating that linelevel supervisors have company-issued devices. Only 11% indicate that end users have integrated mobile devices fully into their daily activities. Meanwhile, 66% say that field personnel still use paper and pen to record data for later transfer to workstation PCs. • Software. Respondents mentioned 66 different software packages in use across their organizations. The greatest software challenges that respondents encountered during mobile project development and implementation related to creating a clear scope and coding challenges, along with gaining technical access to the correct data and integrating the programs with other company systems. • Concerns. Concerns related to mobile technology were highest regarding the lack of integration with other systems and across multiple job functions, as well as unreliability of the devices, applications, and network connectivity. • Work crews. When asked whether respondents saw mobile/digital technology as a possibility to meet the needs of work crews, there was nearly unanimous agreement that mobile solutions could be extremely valuable. Most promising were the capabilities for real-time data access (e.g., work package guidance and procedures, lockout/ tagout status, and equipment operating data), data capture (e.g., inspections, readings, and equipment condition reports), and data communication (e.g., consistency between work teams and activities), assuming both the applications and connectivity were reliable. • Device value. Ninety percent of respondents felt that pads or tablets should be an integral part of an O&M worker’s tool kit, primarily because of the efficiency these devices afford. Approximately 80% said that mobile technology could help operators perform inspection rounds, and that field workers could use personal devices to take pictures, as needed. 24 ENERGY-TECH.com

The mobility team One key to the success of mobile technology integration is convening an organized mobility team under the strong leadership of a mobility program director. Another key is to address all stakeholders’ interests from the beginning in the design, implementation, and monitoring of mobility programs by assembling representatives from each stakeholder group—for example, the corporate office, IT, plant and department management, engineering, instrumentation and control, field supervisors, members of field crews, and participating vendors and contractors. From a kick-off meeting, through status meetings and field trials, the mobility program must consider their respective viewpoints and feedback. Together, the mobility team can create the common vision and values guiding the program, then identify and prioritize the key program objectives. Mobility programs are often deployed first in a pilot phase to one location and with minimum viable functionality. Once applications are stabilized, a coordinated rollout to each additional location can be completed. Value propositions Mobile technology’s value to power generation operations derives from the ability to communicate voice, data, and video information between enterprise workers, their teams, and the enterprise assets in a fashion that enhances employees’ safety and increases efficiency by providing the right information to the right person at the right time. Among the potential areas of value: • Condition documentation. Real-time entry of inspection readings and documentation of equipment condition save substantial time over the traditional method of clipboard documentation and later data entry. • Access to information. Field access to procedures and online resources facilitates more accurate equipment repair. Real-time dispatch of work orders and field reporting through the use of digital tablets is replacing paper workorder systems. • Access to documentation. Using smartphones and tablets, it is feasible for construction and maintenance specialists to pull up relevant documents related to any facility, piece of equipment, and business process. • Process management. location services on mobile devices can provide asset information via geographic information system (GIS) data. Interactive maps integrated with real-time process data can quickly draw attention to August 2017

MAINTENANCE MATTERS anomalies and provide at-a-glance “cockpit” displays of realtime operational system status. (See Figure 2.) • Asset availability, maintenance and repair. Asset management applications accessed through handheld devices help workers manage the asset life cycle from planning and budgeting to operational deployment and ultimately to retirement and disposal. • Access to expertise. Power generation expertise is leaving plants as older workers retire. Perhaps one of the highestvalue functionalities of handheld wireless devices for power generation enterprises is providing voice, camera image, and video remote access to subject-matter experts across plant geographies. • Workforce management. Mobile devices aid in monitoring and analysis of workflow data to decrease work process timelines and improve productivity. Handhelds help to improve workforce communication, enabling team leaders and managers to stay in closer touch with work teams.

track inventory and can be used to monitor resources that are aging and increasingly erratic. High-value use cases The highest-value use cases might logically be expected to appear where the most time is being spent in the plant—either in completing work tasks or in transit time to get and report data from/to various sources. According to EPRI research, three high-value use cases are as follows: • Work management. mobile work management solutions enable the rapid identification and documentation of equipment problems, and digital guidance on repair procedures with links to online assistance. If replacement parts are required, online inventory may be queried and parts ordered from the field, greatly reducing repair time.

• Supply chain management. Central warehouse processes are aided through wireless handheld access to supply chain data and reporting.

• Lockout-tagout (LOTO). All work crews tasked with performing maintenance duties must carefully follow lockout-tagout (LOTO) procedures when working on plant equipment and systems. Utilizing mobile technology to validate the proper work location and affected equipment helps reduce the risk of errors that could jeopardize safety or affect production.

• Asset management. Asset management data from the Internet of Things (IoT) and radio frequency identification device (RFID) tags on equipment in generation plants helps

• Equipment inspections. Many O&M crews still make their rounds with clipboards and paper checklists, recording data and capturing notes to be entered in a relevant program

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MAINTENANCE MATTERS on an office-based PC. They manually fill in spreadsheets and export the data to operations, maintenance, and engineering databases. This process is not efficient and is usually only completed by a select group of subject-matter experts. Capturing inspection data electronically on location is a more desired approach. Modern mobile inspection applications can provide work instructions, allow checklist updates, and make the process more efficient. These applications can also improve accuracy through validation functionality and improve efficiency through paperwork reduction and less time spent gathering and searching for data.

Sidebar: Digital worker interest group In 2017, EPRI launched a Digital Worker Interest Group (DWIG), which is a collaborative effort among members of EPRI’s Generation Sector who have an interest in use of mobile digital technologies and applications. The initial meeting of the DWIG was held in May and drew plant-level management for new combined-cycle plants and corporatelevel managers with responsibilities for IT, management of change, equipment reliability, work management, and mobile implementation. Utilities represented included Duke Energy, Associated Electric Cooperative Inc., Tennessee Valley Authority, and ENMAX Corporation.

PERPETUAL MOTION

The consensus was that mobile technology is rapidly developing and that many utilities are on the cusp of adoption. Utility presentations described roll-outs of mobile technology, pilot projects, experience with specific applications and devices, and recommendations for integration with current systems. Vendor presentations described new technologies and software applications. Attendees expressed the need for defined business cases for mobile technology in order to demonstrate safety enhancement, efficiency gains, or return on investment for senior management. They are also looking for a starting point to begin mobile implementation, and desire established standards so individual utilities do not have to determine how to build-out applications for enterprise software on their own.

What every plant manager wants… To achieve high productivity and financial success, today’s complex process operations require unyielding power that won’t quit. That’s why thousands of smart process operators have chosen custom-built Skinner Steam Turbine Generator packages to produce power & savings in such diverse applications as refineries, petrochemical plants, food processing and other facilities.

A future meeting is planned for November. ■ Matthew Buck (mbuck@epri.com) is Senior Technical Leader in EPRI Generation Operations & Maintenanc

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TURBINE TECH

Dynamic turbine reinspection intervals – A tool to optimize outage schedules By Rachel Sweigart, TG Advisers

The current North American generation marketplace presents many challenges including limited O&M budgets and pressures to extend major overhaul intervals. Now, more so than ever, there is significant value in performing major inspections only when necessary based on actual operating profiles and sound engineering analysis. Today’s operating profiles likely include frequent on-off cycling and low load operation. Further complicating matters is the fact that future run profiles are largely unknown. In an effort to better define and provide a flexible tool for future outage planning, “Dynamic Reinspection Intervals” were developed. The aim of these tools was to input actual operating profiles and operating data from plant historians and output updated reinspection intervals based on engineering principles and analysis. This article will discuss two dynamic reinspection analysis methodologies—rotor blade attachments in the Low Pressure (LP) section of the turbine and rotor blade attachments in the high pressure (HP) section. The LP dynamic reinspection anal-

ysis will consider cracking from stress corrosion cracking (SCC) and low cycle fatigue (LCF) and the HP dynamic reinspection will focus on creep. LP rotor blade attachment dynamic reinspection This dynamic reinspection analysis for low pressure turbine blade attachments considers two methods of crack propagation – stress corrosion cracking and low cycle fatigue. SCC is a well-documented industry phenomenon that occurs in wet steam environments with increased probability as the unit’s service hours exceed 150,000. The blade attachment area acts as a natural trap where chlorides can collect and concentrate. Crack growth rates are driven by a number of key factors including material yield strength, stage operating temperature, and steam chemistry. SCC is dependent on operating time and not cycles. On the other hand, LCF is caused by repeated stresses due to start and stop cycles. Because of this, different operating profiles lead to varied crack growth rates based on the phenomenon that is driving the crack.

Figure1: Crack propagation rates August 2017 ENERGY-TECH.com

TURBINE TECH The method for evaluating duty cycle dynamically utilizes fracture mechanics principles and assumes there is an initial flawâ&#x20AC;&#x201D;either the largest flaw found in a prior inspection or the minimum detectable value from that inspection technique if the prior inspection yielded no reportable indications.

summed and compared as a total crack growth amount to the reinspection crack size. The reinspection crack size is calculated using fracture mechanics principles and conservative material properties and includes a safety factor to allow for safe operation between inspections.

Crack propagation is considered as a combined rate from low cycle fatigue due to regular on/off cycles, low cycle fatigue from overspeed cycles, and hour dependent stress corrosion cracking. The resulting crack growth from each phenomenon is

The next two scenarios are provided to illustrate the concept and importance of considering both operating hours and cycles when determining low pressure turbine reinspection intervals. One of the most significant benefits in this type of automated analysis program is that only the operational hours and stop/starts need to be input in order to output a visual representation of crack progression. Figure 1 shows an example of the crack propagation rates. The reinspection crack size is indicated with the solid red line. The dashed red line shows the calculated crack propagation based on the user inputted duty cycle. When the two lines meet is when a reinspection is required. For this specific example (not a generic curve), the two major contributing crack growth rates (SCC and LCF) are shown in green and blue lines respectively. The circles reflect the amount of crack propagation due to SCC/operational hours (green) and LCF/cycles (blue) experienced to date. The scenario shown in Figure 1 shows a base loaded operation where most of the crack propagation has come from SCC (green) and there is only a small amount from LCF (blue). In this instance, the solid black line shows the calculated crack propagation is currently 73.8% of the way to the reinspection crack size. To compare, Figure 2 shows the same unit but with an operational profile that focuses on cycling. It has three times as many cycles as the previous scenario and significantly fewer operational hours. However, crack propagation is still in a similar range. With a dynamic reinspection interval methodology, outages can be scheduled while taking into account the changes in operating profile. HP rotor blade attachment dynamic reinspection Because HP blades are much smaller and typically do not see wet steam during operation, LCF and SCC are not traditionally the limiting mechanisms for rotor blade attachments. However, because HP blades operate in an elevated-temperature environment, they are susceptible to creep. Creep is a form of slow, continuous deformation that is

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TURBINE TECH

Figure 2: Same unit as Figure 1, but with an operational profile that focuses on cycling.

inevitable with high temperature operation and based on operational time, stresses, and temperature. In a steam turbine, the HP and IP inlet stages are most at risk for creepâ&#x20AC;&#x201D;in fact any stage that sees temperature in the range of 900Â°F or higher is likely to accumulate creep damage over its operational life. Inspection intervals for high temperature creep are often based off the assumption of full load operation at design steam temperatures. Thus, operation at low loads and below design steam temperatures can greatly reduce the rate of creep damage. Similarly, exceeding design steam conditions can significantly increase creep rates and reduce creep life. With a dynamic reinspection evaluation the actual operational hours and steam temperatures are considered to calculate expended creep life. Using the operational stress and the actual temperature data, the Larson-Miller method can be used to estimate how much creep life has been consumed. The Larson-Miller equation relates temperature, time, and creep rupture stress to predict time to failure based on a stress level. This relationship is shown in Figure 3. By varying inlet temperature new creep consumption rates or creep lives can be calculated for a given stress level. Completing this type of analysis allows for evaluation of actual creep life expenditure based on a unitâ&#x20AC;&#x2122;s operation rather than a generic assumption. It can both prevent premature replacement of rotors and indicate if a rotor needs to be inspected more frequently to ensure reliable operation.

August 2017 ENERGY-TECH.com

TURBINE TECH

Figure3: The Larson-Miller equation relates temperature, time, and creep rupture stress to predict time to failure based on a stress level.

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Conclusion Both the LP and the HP dynamic inspection methods are designed to evaluate areas of life expenditure. Based on a unit’s actual operational profile and not assumed duty cycle, more representative inspection intervals can be determined frequently and updated inspection intervals set based on a specific life limiting parameter for a given design. If the unit experiences lower temperature operation than design or has a more base-loaded operating profile than the reinspection interval was set on, then the current reinspection intervals may be too frequent. However, if the unit is operating at higher than design conditions or is cycling much more frequently, then the reinspection intervals may need to be more frequent. A proactive approach to operational based damage mechanism evaluations allows for more realistic reinspection intervals for a given design type. ■

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Rachel Sweigart joined TG Advisers in 2014 as a consulting engineer. She has provided life assessment, torsional and fracture mechanic analytical modeling and troubleshooting services for main turbine generators located throughout the country. Sweigart is a mechanical engineering graduate from Lafayette College. You may contact Rachel by emailing editorial@WoodwardBizMedia.com

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August 2017 ENERGY-TECH.com

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