Fast Startup Design For Cycling and
of Large HRSGs
Wesley Bauver, Ian Perrin, and Thomas Mastronarde ALSTOM, Heat Recovery & Plants, Windsor , CT, USA
Abstract
Large heat recovery steam generators (HRSGs) are increasingly required to endure significant cyclic operation. The ability of horizontal HRSGs to endure such operation has been questioned and a substantial body of literature exists discussing scenarios that might result in premature failures. This paper summarizes some of the results of a program of monitoring operational conditions coupled with detailed finite element modeling and fatigue evaluation of critical components. This paper outlines the monitoring and analysis methods and discusses the results for a few salient cases, highlighting the effect of design features on the long-term durability and reliability of large HRSGs.
POWER-GEN International December 9-11, 2003
Las Vegas, Nevada
Fast Startup Design For Cycling and
of Large HRSGs
Wesley Bauver, Ian Perrin, and Thomas Mastronarde ALSTOM, Heat Recovery & Plants, Windsor, CT, USA Introduction
Large heat recovery steam generators (HRSGs) are increasingly required to endure significant cyclic operation, particularly in power markets where the plant is dispatched daily according to system load demand requirements. Daily cycling operation can, potentially, result in fatigue damage to critical components if they are not designed correctly. The thermal and mechanical flexibility of an HRSG is strongly dependent on the fundamental layout of the pressure parts and the detailed design of components.
Potential fatigue damage depends on both design details of at-risk components and operation of the combined cycle power station. To avoid perceived risks associated with fast startup, some of these plants incorporate part load hold points or soak times for warm-up of HRSG components, degrading the fuel efficiency of the plant, or delaying the achievement of optimum emissions and dispatch power levels.
Over the past several years, ALSTOM has been engaged in a program of monitoring boiler operating parameters and associated temperature histories of critical
components. Data obtained from full-scale components operating in a real-world plant environment has significant value: (1) by providing actual rather than assumed thermal gradients and temperature histories for use in stress calculations, (2) by validating temperature histories predicted by sophisticated dynamic analysis models.
These temperature histories provide the essential boundary conditions for analysis and verification of fatigue life.
In this paper, the following issues will be explored:
• elimination of thermal stresses and damage from condensate flooding during hot restart
• reduction of the thermal stresses in superheater assemblies caused by row-to-row temperature differences during cold start
• limitations imposed by the HP drum for applications where the fast startup is desired Monitoring
Verification of design assumptions and evaluation of the response of HRSG
components to cycling operation requires detailed transient operational data. The type of information required is typically not available from standard plant instrumentation.
Programs were instituted at multiple sites to obtain local component temperatures in conjunction with and synchronized with plant operating data. Components instrumented included the tubes and headers of high-pressure superheaters and reheaters,
superheater and reheater manifolds, steam drums, desuperheaters, and support links.
Specific locations for thermocouples installed in the lower superheater area are shown in Figure 1.
Figure 1. Thermocouple Locations for Lower Superheater
Thermocouples were spot welded directly onto metal surfaces as shown in Figure 2 and covered with insulation as shown in Figure 3 to ensure accurate metal temperature readings.
Tube below fins Tube to header joint
Header top
Header bottom Manifold top
Manifold mid-height
Manifold bottom
Lower Superheater Manifold Lower Superheater Headers (Single-row harps)
Drain
Figure 2. Thermocouples on Tube and Header Figure 3. Insulated Thermocouples
Data from metal temperature thermocouples and the plant Distributed Control System (DCS) were integrated and stored on a stand-alone computer. Typical sampling rates were once per minute to facilitate transient analyses. Monitoring systems have been in operation at several HRSGs (operating behind GE 7FA gas turbines) for up to three years, generating a large database of information on both operating profiles and component responses.
The information obtained from the field monitoring programs has been used to evaluate and confirm mechanical and thermal design assumptions for various components. In particular, two areas related to the specific configuration of pressure parts were
explored in detail. In these two areas the design of ALSTOM HRSGs differs from most other HRSGs.
1. Superheater and reheater drain system where the drain manifold is separate from the lower harp headers. This configuration moves rapid water accumulation away from the tube entrances into the lower harp header and into a large manifold located below the harp headers, which functions as a generous sluice-way for the receiving condensate during hot restart. This arrangement precludes the possibility of water backing up into individual tubes, preventing damage from condensate flooding.
2. The use of small diameter (thin-wall) headers in harps with a single tube row.
Small diameter headers in conjunction with step changes in pressure part thickness lead to minimum tube-to-header temperature differences at the weld joint. The use of a single row eliminates the bend in the tube near the header, and permits more rapid rates of temperature change than more conventional thick-walled headers used with multi-row harps.
Temperature histories obtained from the monitoring program have been used to validate the effectiveness of these two design features in eliminating fatigue damage.
Drain Effectiveness during Hot Restart and GT Purge
The problem of condensate flooding during hot re-start has been reported in the literature and in various symposia on a number of occasions. This phenomenon takes place when GT exhaust gas during purge is lower in temperature than the saturation temperature inside superheater tubes, typical when the boiler is still hot within a few hours of a GT trip. Under these conditions, a large amount of water will condense in a short time inside superheater tubes. With conventional multi-row harp construction, the superheater header acts as the drain manifold. Backup of condensate into individual tubes can occur when the lower superheater header floods. Such flooding can result in damaging temperature differences between flooded tube and tubes with steam flow, resulting in bowing of tubes or failure of tube-to-header welds.
When the lower superheater headers are segregated from the lower superheater manifold, as shown in Figure 1, backup of water into individual tubes can not occur.
Multiple links between lower superheater headers and the lower manifold assure that condensate can drain freely to the lowest point. The following example illustrates the verification of this approach. A full load gas turbine trip occurred at 7:37 AM. The gas turbine was restarted with a purge at 9:14 and reached base load at 10:50. Figure 4 shows temperatures on the front high pressure superheater tubes and header over this time period. See Figure 1 for thermocouple locations.
HPSH1 Harp bottom
500 600 700 800 900 1000
7:30:00 8:42:00 9:54:00 11:06:00 Time
Temperature, deg. F
Tube Below Fins Tube - Header Joint Header Top
Header Bottom -
GT Restart
Figure 4. High Pressure Superheater - Tube and Header Temperatures
As can be seen, metal temperatures at all locations remain within 28 degrees C (50 degrees F) of each other, indicating that the stepped metal thickness and thin headers minimize metal temperature variations and thermal induced stress.
Temperatures around the HP superheater manifold (also see Figure 1) over the same time period are shown in Figure 5. Temperatures decrease uniformly for about one hour. Then metal temperatures on the bottom of the manifold rapidly decrease from about 750oF to 550oF,while temperatures on the top and the sides at mid-height continue to decrease slowly. Bottom temperatures then rapidly rise back to the same range as the sides and tops. This indicates a rapid build up and draining of water in the manifold. The mid-height temperatures do not follow the temperatures at the bottom.
This indicates that the water level never reaches the middle of the manifold. The generous manifold size and the location and size of drain connections on the bottom promote free flow of water to the drain. Measured temperatures confirm the manifold never floods, proving effectiveness of the drain system. Temperatures around the circumference of the entire header decrease at the same time during GT restart. This demonstrates that the header is being uniformly cooled and then heated by steam, with no indication of local cooling by condensate.
HPSH Bottom Manifold
500 600 700 800 900 1000
7:30:00 8:42:00 9:54:00 11:06:00 Time
Te m pe ra tu re d eg . F
top
mid height bottom GT Restart
Figure 5. High Pressure Superheater - Manifold Temperatures
Modeling and Analysis
Analysis of the effects of fast cycling requires both sophisticated stress analysis models and realistic thermal boundary conditions. Data obtained from monitoring has been
used to provide boundary conditions for existing operating modes. Assessment of other proposed operating modes and evaluation of components for which monitoring data is not available requires validated dynamic performance models. Such a model has been developed based on the APROS-5 TM simulation platform. The model includes
superheater sections, reheater sections, HP evaporator sections, applicable control valves and associated control systems. Figure 6 shows a schematic of the modeled system.
The predicted behavior of several system parameters during the transient is compared with the field data in Figures 7 and 8. In the plots, the simulation results are shown using smooth curves and square or triangular markers are used for field data. In the simulation, the main steam flow rate increases rapidly and then begins to level off. At 50 minutes, the main steam flow again goes through a steep increase as the
superheater pressure levels off. In the field data, the flow initially shows many sharp increases and decreases. These may be related to the opening and closings of various drains, vents and valves not modeled. In the simulation, the modeled steam
temperature is just at the exit of HPSH1, and, in the absence of significant steam flow, equals the metal temperature. The field data recorded the temperature in the main steam piping. The thermal inertia connected to the metal mass of the steam pipe and the steam headers is not included in the simulation model, yielding a much steeper slope to the temperature rise curve.
After the steam flow starts, the field data rapidly approaches the simulation prediction.
Prediction of drum pressure rise up to 17 bar (250 psig) matches field data extremely well. At this point, the bypass systems begins to control drum pressure and the simulation uses the ramp rate dictated by the bypass valve set point schedule.
Figure 6. Diagram of HRSG Dynamic Model
Model Validation- Cold Start Case - Main steam flow and Temperature -
0 20000 40000 60000 80000 100000 120000 140000 160000 180000 200000
0 10 20 30 40 50 60 70 80 90 100
Time (min)
Steam flow (lb/h)
0 100 200 300 400 500 600 700 800 900 1000
Steam Temperature (F)
MAIN_STM_FLOW (lb/h) F_HP_STM (lb/h)
STM_TEMP_HPSH1_OUT (F) T_HP_EAST_OUTL_A (F) ,,,
Figure 7. Main Steam Flow and Temperature Simulation
Model Validation- Cold Start Case - HP Drum Pressure -
0 100 200 300 400 500 600 700 800 900
0 10 20 30 40 50 60 70 80 90 100
Time (min)
Pressure (psia)
HPSH1_PRESSURE (psia) P_HP_DRUM (psia)
Figure 8. HP Drum Pressure Simulation
With the high quality field data and accurate transient models, detailed stress analysis and life assessment can be performed. This is discussed in the next section.
Analysis and Life Assessment of HRSGs
Analysis and life assessment of HRSGs is becoming increasingly important as the cyclic duty increases and as gas turbines impose greater demands on the HRSG. In the technical press, and in various symposia and workshops, the lack of a technically mature approach to life assessment of HRSGs has been criticized. Most HRSG
manufacturers are only gradually beginning to address design and operational concerns for cycling service. Many purchasers of new equipment have now become sensitized to this issue and are now routinely requiring a cyclic life assessment as part of the project documentation.
Life assessment of HRSGs is a complex exercise because the equipment consists of large assemblies of interconnected parts with different characteristic dimensions and flexibility, all of which are subjected to complex thermo-mechanical transients. To adequately represent the functional aspects of HRSG design, while providing sufficient fidelity to compute component lifetimes, requires experience in HRSG design and operation. It also requires expertise in advanced design and life assessment methods more commonly employed for gas turbine components than boilers.
The first step in a life assessment study of an HRSG is to determine which of the many components and assemblies actually warrant analysis. This is most easily achieved by a screening methodology to ascertain which components are subjected to the greatest thermal and mechanical loads. This screening must consider the connectivity between parts and assemblies that could result in load transfer or temperature differentials not apparent on first inspection. Furthermore, dynamic simulation or field data, such as that described earlier in this paper, is needed to define the temperature transients
experienced by components. The screening will identify components and assemblies, such as thick-walled drums or superheater outlet headers, or superheater assemblies with row-to-row temperature differences. These are then subjected to rigorous analysis, starting with the most severely loaded.
Although ALSTOM has performed numerous analyses of HRSG components, the discussion that follows focuses on the effect of cold startups on harp assemblies, highlighting some of the differences between single-row and multi-row constructions.
This also illustrates the type of analysis that is now being performed to understand the structural integrity of HRSG components and assemblies. A detailed discussion of modeling geometry, derivation of boundary conditions, or numerical details of the analysis is outside the scope of this paper. The discussion below will emphasize the technology employed and focus on the principal conclusions and recommendations from analysis.
Flexibility and Stepped Component Thickness
High-pressure superheater assemblies, Figures 9 and 10, are subjected to significant thermal transients due to the rapid rate of rise of gas turbine exhaust gas temperature on startup, and due to the rapid rate of rise of steam temperature that occurs as steam flow is first established. The superheater assembly, therefore, provides a good example of a complex assembly subjected to thermo-mechanical loads, which can be used to highlight the effect of design and layout of the assembly. For the purpose of analysis, the thermo-mechanical history can be broken into a number of phases to determine the stresses, which can then be appropriately compounded to arrive at a stress range for a fatigue life estimate. The key phases of the cold startup are:
• Tube row-to-row temperature difference on initial gas turbine firing.
• Tube-to-header and header top-to-bottom temperature difference due to differential heating of tubes and header.
• Component-to-component temperature differences due to rapid internal heating as steam flow is first established.
Each of these phases will be examined in the sections that follow.
Figure 9. (a) View of an entire “single-row” superheater assembly, (b) zoomed view of the lower portion of the assembly showing “single-row” harps with stepped components using links from harp headers to manifolds. Note that, for clarity, the extended surface (finning) of the tubes has been omitted.
Figure 10. (a) View of an entire “multi-row” superheater assembly, and (b)
zoomed view showing multiple rows of tubes entering a single header. Note that, for clarity, the extended surface (finning) of the tubes has been omitted.
Two superheater assembly configurations are examined here: (1) “single-row” harp assembly (Figure 9) used by ALSTOM in which each row of tubes enters a single
header which is linked to a common manifold, and (2) “multi-row” harp assembly (Figure 10) in which multiple rows of tubes enter a common header.
For the calculations reported here, the key dimensions of the components of the lower portion of the assembly are reported in Table 1. This confirms that the superheater tubes are the same and that the manifold in the single-row harp assembly has similar dimensions to the header in the multi-row harp assembly.
Table 1. Comparison of component sizes for harp assembly analyses.
Component Tube Lower Header
Single-row Harp Lower Manifold
Single-row Harp Lower Header Multi-row Harp Outer diameter 38 mm (1.5”) 114 mm (4.5”) 273 mm (10.75”) 273 mm (10.75”) Wall thickness 5.2 mm (.205”) 17 mm (.67”) 35 mm (1.3”) 35 mm (1.3”) Ratio of header wall
thickness to branch attachment
3.3 2.1 6.7
Tube Row-to-Row Temperature Differences during Cold Start
The effects of a tube row-to-row temperature difference are evaluated first. The row-to- row temperature difference is most severe on a cold start when the superheater tubes are initially cool (near ambient temperature) and are then subject to warm exhaust gas which rapidly reaches 370C (700F). A row-to-row temperature difference is established as the first tube row extracts heat from the exhaust gas such that a little less heat
reaches the second row, etc. The row-to-row temperature difference depends on the exhaust gas mass flow, temperature, and tube-fin configuration, and the loading characteristic of the gas turbine. Typical values are:
Conventional Startup: typical range 20C to 30C (36F to 54F) Fast Startup: typical range 30C to 50C (54F to 90F)
In the discussion below, a temperature difference of 25C (45F) has been assumed. For an accurate analysis of the deformation and stress in the single-row and multi-row superheater harp assemblies, finite element models have been carefully designed to capture the global geometry, thereby giving a good representation of the overall
flexibility of the assembly, while also modeling local stress concentrating features. This avoids the use of, often ill-defined, stress concentration factors from tables to
compensate for the local features that are not represented by beam or shell elements.
The tubes, header and manifolds are all constructed from grade 91 steel (9Cr-1Mo-V- Nb) and material properties have been assigned accordingly.
The results of the analyses highlight a number of design features. Firstly, in the present analysis the peak stresses in the single-row and multi-row harp assemblies are quite similar and do not exceed the yield stress of the grade 91 material, therefore the moderate 25C (45F) row-to-row temperature difference alone will not cause fatigue failure of either assembly. However, it is evident from Figure 11 that the highest stress in the multi-row harp assembly is concentrated in the tube (a thin walled component) at the toe of the weld. There are many of these locations on the harp assembly and, therefore, for a larger temperature difference many potential failure sites that are
generally difficult to access for inspection and repair. For the single-row harp assembly, Figures 12 and 13 illustrate that the stepped component thickness distributes the stress between the harp header and link to manifold connection. The highest stresses are therefore in relatively heavy wall components (links and harp headers). The link-to- manifold and link-to-header connections can be inspected with relative ease.
Figure 11. Color contour plot of Tresca stress for a portion of the lower assembly of the multi-row harp assembly subjected to a 25C (45F) row-to-row temperature difference.
Figure 12. Color contour plot of Tresca stress for a portion of the lower assembly of the single-row harp assembly subjected to a 25C (45F) row-to-row temperature difference.
Figure 13. Zoomed, cutaway, view of the Tresca stress contours at the header-to- link region. The highest stresses occur at the saddle of the inside surface of the tube holes in the header. Similar stresses are also evident on the outer surface of the header at the toe of the tube to header weld on the header side, and in the fillet of the nozzle to header weld.
Header
Manifold Link
Tube
Analyses of other attached piping arrangements demonstrate that the magnitudes of the stresses do depend on the stiffness of the attached piping, which reacts to the global motion of the harp assembly. A bounding case is that of rigid piping (no rotation of the lower header or manifold) which shows that the stresses in the multi-row assembly can be significantly higher (by 25%) than those in the single-row assembly. Even such a modest stress increase can halve the fatigue life of a component, which further serves to highlight the value of this type of analysis and the need to consider the relevant factors during pressure part design and layout.
Differential Heating by Exhaust Gas
During the initial period of gas turbine firing and loading, as the tube row-to-row temperature difference is established, additional temperature differentials are set up between the tube and the header and between the top and bottom of the header. The thin-walled finned tubes warm more rapidly than the header, which causes a tube-to- header temperature difference. This difference is more pronounced for the multi-row assembly because the tubes enter a thicker header which takes longer to warm, thereby compounding the stress due to the tube row-to-row temperature difference. In the case of the single-row assembly, the temperature difference between the tube and the
header is less significant because of the smaller difference in wall thickness between the tube and the header.
A temperature difference also develops between the top and bottom of the headers, both for the single-row and multi-row assemblies, because the top of the header is warmed by the combined effect of exhaust gas and conduction from the tubes. The resulting temperature difference between the top and bottom of the header is much more significant for the multi-row harp assembly because the header has a much larger diameter and is thicker, thereby prolonging the time for circumferential heat conduction.
The single-row construction places the larger diameter, thick-walled, manifold out of the exhaust gas flow and therefore only the smaller diameter, thinner-walled, harp headers are subjected to this differential heating. Field data confirm that conduction within the small diameter, thin walled, header is sufficient to ensure that the top-to-bottom
temperature difference on the header is no more than a few degrees.
Temperature Differences of Branch Connections
The final phase of the significant thermal transients on a cold startup is the rapid rate of rise of steam temperature that occurs as steam flow is first established. Generally, this does not occur until sufficient heat has been absorbed in the evaporator by which time the temperature differences described earlier have died away. Therefore, the rapidly rising temperature of the steam heats components of differing wall thickness at differing rates, which, in turn, creates local thermal stresses at connections between
components. The rates of temperature rise can be quite large 70C/min (125F/min) and heat transfer coefficients are also large because the cool components (especially headers and manifolds out of the exhaust gas flow) are warmed by condensation heating.
It is relatively straightforward to compute the temperature difference that can be established between components by first performing a transient heat transfer analysis and then evaluating the mean through-wall temperature for each component. This mean through-wall temperature can then be compared for connected components to provide an estimate of the typical temperature difference between the components.
These computations have been performed for the single-row and multi-row harp assemblies discussed earlier, the dimensions of which are given in Table 1. Of particular note is the ratio of header wall thickness to branch connection thickness, which is two to three times greater for the multi-row harp.
The results for the components of the multi-row assembly are shown in Figure 14, which presents the mean through-wall temperature of tube and header as a function of time and also plots the temperature difference between these components. The
temperature difference between the tube and header reaches a peak value of 103C (185F).
The results for the components of the single-row assembly are shown in Figure 15, which presents the mean though-wall temperature of the tube, header/link and manifold as a function of time and also plots the tube-to-header and link-to-manifold temperature difference. The tube-to-header temperature difference reaches a modest 41C (74F) and the link-to-manifold temperature difference reaches a value of 68C (122F).
Although the tubes and manifolds are of similar thickness in the two cases, the single- row assembly adds the links and harp headers (which are of equal thickness and, therefore, do not have any temperature difference). The stepped component
thickness of the single-row construction therefore reduces the temperature difference between parts which ensures that components subject to high stresses during the row- to-row temperature difference are not also subject to high stresses as steam flow is first established. This is critical for highly cycled units.
Combination of Row-to-Row and Thermal Differences at Branch Connections Combining the stresses from the three phases of the startup transient shows that a continual compounding of stress occurs at the tube-to-header connection of the multi- row assembly. The single-row assembly, however, minimizes temperature differentials and hence thermal stresses, such that the temperature difference between parts is significantly lower (e.g. reduced from 103C (185F) for the multi-row assembly to 41C (74F) for the single-row assembly, at the tube-to-header connection). Therefore,
although detailed analysis of the single-row and multi-row assemblies subjected to row- to-row temperature differences showed that the peak stresses were of similar
magnitude, it is apparent that the complete thermal cycle should be considered in a fatigue life evaluation. The flexibility and stepped thickness of the single-row assembly avoids compounding of temperature difference and stress from the various phases of the startup transient in the same geometric location.
0 50 100 150 200 250 300 350 400
0 1 2 3 4 5 6 7 8 9 10
Time (mins)
Temperature (C)
Steam temperature Tube temperature Manifold temperature
Temperature difference (Tube-Manifold)
Figure 14. Temperature histories for thick-walled header in a multi-row assembly.
0 50 100 150 200 250 300 350 400
0 1 2 3 4 5 6 7 8 9 10
Time (mins)
Temperature (C)
Steam temperature Tube temperature
Link & Header temperature Manifold temperature
Temperature difference (Manifold-Link) Temperature difference (Tube-header)
Figure 15. Temperature histories for single-row assembly with stepped thickness tube/header/manifold construction.
Benefits of Stepped Component Construction with Single-Row Harps
The benefits of the single-row assembly, from a cycling perspective, are its inherent flexibility with load being shared between tubes, harp headers, links and the manifold (compared to the multi-row assembly which only has tube-to-header connections) and the stepped component thickness which minimizes temperature differences between parts.
This combination offers fatigue life three to five times greater than conventional multi- row harp construction, and is particularly suited to fast startup applications. Note also that single-row harp assemblies facilitate high quality tube-to-header welds due to complete 360 degree access provided to welders and inspectors.
HP Drum under Cycling Conditions
The high-pressure (HP) drum is often cited as a life limiting component because it is relatively thick-walled , 100 to 150 mm (4 to 6 inches) thick, and must endure relatively rapid rates of internal fluid temperature change, particularly during cold startups. This causes a through-wall temperature gradient, which combined with the stresses due to internal pressure and stress concentrations at nozzle details, can indeed result in fatigue damage. Often, overly simplistic methods of analysis can lead to erroneous
results – either giving a false sense of security or undue pessimism. The reasons are that methods such as EN12952 are often used which are generally highly conservative for the basic effect of the through-wall temperature difference due to the internal
temperature change but neglect the assembly loads such as reactions and moments at downcomers and risers. The limitations of these analysis methods can be overcome with appropriate dynamic models to simulate the actual rate of temperature rise in the HP drum while also giving information on the temperature differentials between the drum, downcomers and evaporator tubes which give rise to assembly stresses.
Although a detailed analysis of this topic is beyond the scope of this paper,
dynamic modeling at ALSTOM has shown that the initial rates of rise of temperature in the drum-evaporator system are primarily controlled by the heat input from the gas turbine and are not affected by steam bypass operation. As a result, the HRSG drum must meet the transient heating imposed by the gas turbine, or the gas turbine exhaust gas transient must be modified to meet the needs of the HRSG.
It can demonstrated that an HP Drum with a thickness of less than 140 mm (5.5”) will typically be suitable for cycling service with today’s gas turbine startup practices.
With fast loading of the gas turbine, however, the HP drum becomes a limiting component for fatigue life in cycling service. It is expected that fast loading of gas turbines will be become more prevalent in the future as plant owners strive to reduce the time to achieve emissions compliance and to reduce startup fuel consumption.
For this type of plant, elimination of the HP steam drum by using a once-through evaporator system will provide the thermal flexibility to withstand daily cycling for the thirty-year life of the plant.
Conclusion
Horizontal HRSGs have become the technology of choice and the dominant configuration world-wide for large HRSGs in the past ten years. To truly address
customer’s demands for operational flexibility through dispatch and cycling of combined cycle plants, key features in the design of HRSGs can greatly enhance the fatigue tolerance of certain components.
Adopting pressure part features such as single-row harp, stepped component thickness, and enhanced drain systems will improve the thermal response of large HRSGs and eliminate the need for hold points specifically to accommodate the HRSG during the conservative gas turbine loading ramps that are in use today in many plants. In the future, as fast loading of large gas turbines becomes economically advantageous, single row harps, coupled with stepped component construction, will accommodate temperature transients that would otherwise damage and limit the life of an HRSG built with multi-row harps.
