September 2014 

NASA/CR–2014-218521

AFC-Enabled Simplified High-Lift System 

Integration Study 

Peter M. Hartwich, Eric D. Dickey, Anthony J. Sclafani, Peter Camacho, 

Antonio B. Gonzales, Edward L. Lawson, Ron Y. Mairs, and Arvin Shmilovich 

The Boeing Company, Huntington Beach, California 
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September 2014 

NASA/CR–2014-218521

AFC-Enabled Simplified High-Lift System 

Integration Study 

Peter M. Hartwich, Eric D. Dickey, Anthony J. Sclafani, Peter Camacho, 

Antonio B. Gonzales, Edward L. Lawson, Ron Y. Mairs, and Arvin Shmilovich 

The Boeing Company, Huntington Beach, California 
Available from: 

NASA Center for AeroSpace Information 

7115 Standard Drive 

Hanover, MD 21076-1320 

443-757-5802 

Acknowledgments 

This study was carried out under NASA Contract: NNL10AA05B / NNL13AB65T. Dr. 

John Lin was the NASA LaRC Task Principal Investigator. In addition to the authors, the 

following Boeing personnel contributed to this work: J. Courtney, D. Lacy, Y. Yadlin, 

J.C. Vassberg, and E. Whalen. 

The use of trademarks or names of manufacturers in this report is for accurate reporting and does not constitute an 

official endorsement, either expressed or implied, of such products or manufacturers by the National Aeronautics 

and Space Administration.
Abstract

The primary objective of this trade study report is to explore the potential of using Active Flow 

Control (AFC) for achieving lighter and mechanically simpler high-lift systems for transonic 

commercial transport aircraft. This assessment was conducted in four steps. First, based on the 

Common Research Model (CRM) outer mold line (OML) definition, two high-lift concepts were 

developed. One concept, representative of current production-type commercial transonic transports, features leading edge slats and slotted trailing edge flaps with Fowler motion. The other

CRM-based design relies on drooped leading edges and simply hinged trailing edge flaps for 

high-lift generation.

The relative high-lift performance of these two high-lift CRM variants is established using Computational Fluid Dynamics (CFD) solutions to the Reynolds-Averaged Navier-Stokes (RANS) 

equations for steady flow. These CFD assessments identify the high-lift performance that needs 

to be recovered through AFC to have the CRM variant with the lighter and mechanically simpler 

high-lift system match the performance of the conventional high-lift system.

Conceptual design integration studies for the AFC-enhanced high-lift systems were conducted 

with a NASA Environmentally Responsible Aircraft (ERA) reference configuration, the so-called 

ERA-0003 concept. These design trades identify AFC performance targets that need to be met to 

produce economically feasible ERA-0003-like concepts with lighter and mechanically simpler 

high-lift designs that match the performance of conventional high-lift systems.

Finally, technical challenges are identified associated with the application of AFC-enabled highlift systems to modern transonic commercial transports for future technology maturation efforts. 

Nomenclature

D Angle of attack

AFC Active Flow Control 

AFRL Air Force Research Laboratory

AIAA American Institute of Aeronautics & Astronautics

APU Auxiliary Power Unit

ATT Advanced Tactical Theater

CFD Computational Fluid Dynamics

CRM Common Research Model 

DLR Deutsche Luft & Raumfahrt (German Air & Space)

ERA Environmentally Responsible Aircraft

LE Leading edge

MADCAP Modular Aerodynamic Computational Analysis Process 

NFAC National Full Scale Aerodynamic Complex 

lNPV Net Present Value

LLL
OEW Overall Empty Weight

QFD Quality Function Deployment 

RANS Reynolds-Averaged Navier–Stokes 

TE Trailing edge

TOD Takeoff Distance

TOFL Takeoff Field Length

TOGW Takeoff Gross Weight 

TRL Technical Readiness Level

WACC Weighted Average Cost of Capital 

LY
CONTENTS

1.0 Introduction............................................................................................................................ 1

2.0 Conventional and AFC-Enabled High Lift System Development ..................................... 1

2.1 Conventional High-Lift CRM Configuration .................................................................. 2

2.2 Simplified High-Lift CRM Configuration ....................................................................... 2

3.0 CFD-Based assessments of High-Lift AERODYNAMICS ................................................ 4

3.1 Setup of CFD Analyses .................................................................................................... 4

3.2 CFD Run Matrices ........................................................................................................... 5

3.3 CFD Results for Wing/Body/Nacelle/Pylon Configurations ........................................... 6

3.4 CFD Results for Wing/Body Configurations ................................................................... 7

3.5 Performance Targets for AFC-Enhancements to Simplified High-Lift CRM 

Concepts........................................................................................................................... 8

4.0 AFC Sizing .............................................................................................................................. 8

4.1 Inviscid Performance Limit Concept for AFC Sizing ..................................................... 9

4.2 CFD-Based AFC Sizing ................................................................................................. 10

5.0 Trade Study on AFC-Enabled Simplified High-Lift System ........................................... 11

5.1 Net Present Value (NPV) Assessment ........................................................................... 13

6.0 Analysis, Conclusions, and Recommendations ................................................................. 13

7.0 REFERENCES .................................................................................................................... 14

8.0 FIGURES.............................................................................................................................. 15

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1.0 INTRODUCTION

Wing sizing for current transonic transport aircraft represent trades between efficient cruise 

and effective high-lift characteristics. This includes trades of weight, complexity, supportability, 

realiability, and noise induced by complex high lift systems against wing planform sizing for 

cruise around maximum lift-to-drag ratio. Most of these trades lead to high-lift systems that 

combine leading edge devices (e.g., slats, Krueger flaps) with slotted single- or multi-element 

trailing edge flaps.

McLean et al. [1] explored a range of active flow control (AFC) applications for improving the 

efficiency of next-generation commercial transonic transport aircraft. They identified AFCenhanced simplified high-lift systems as a high payoff / high risk technology for significantly 

reducing weight and fuel consumption for the targeted aircraft class. The rationale was that AFC 

would remedy flow separations that are anticipated when current high-lift systems would be 

replaced with mechanically simpler and lighter variants with simply hinged single element 

trailing edge flaps.

Since that McLean report, much research has been directed at exploring the capabilities and 

limitations of AFC applications to simplifed high lift systems, starting with computational and 

experimental studies on two-dimensional airfoils [2] and some exploratory applications to nonconventional configurations such as the AFRL Advanced Theater Transport (ATT) [3] and the 

AFRL Speed Agile [4] concepts that were enabled only through AFC applications.

The encouraging findings in these studies on AFC-enhanced high-lift concepts motivated the 

current trade study which aims at exploring the feasibility of using AFC for achieving lighter and 

mechanically simpler high-lift systems for transonic commercial transport aircraft. This assessment was conducted in four steps. First, two baseline high-lift concepts were developed with 

one being representative of current high-lift wing designs and the other being a mechanically 

simplified high-lift configuration for AFC integration. In the next step, the high-lift aerodynamics of these two high-lift CRM variants were assessed using Computational Fluid Dynamics 

(CFD) solutions to the Reynolds-Averaged Navier-Stokes (RANS) equations for steady flow. 

These CFD assessments identified the high-lift performance that needs to be recovered through 

AFC to have the CRM variant with the lighter and mechanically simpler high-lift system match 

the performance of the conventional high-lift CRM.

The thus-established AFC aerodynamic performance targets guided conceptual design integration studies for the AFC-enhanced high-lift systems on a reference configuration. This project 

element aims at the identification of suitable AFC/airframe integration concepts that match takeoff and landing approach capabilities of current fleet transonic transport aircraft while yielding 

weight and fuel savings that translate into a positive net present value (NPV).

At the conclusion of the current study, which is rather exploratory in nature, technical challenges of AFC-enabled high-lift systems for modern transonic commercial transports were to be

identified for future technology maturation efforts.

2.0 CONVENTIONAL AND AFC-ENABLED HIGH LIFT SYSTEM DEVELOPMENT

Prediction of high-lift characteristics has long been known to challenge the capabilities of 

CFD-based flow solvers. A good survey of the state of the art in CFD-based high-lift analysis 

can be found in the report on the latest AIAA high-lift prediction workshop [5]. The baseline 

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SMAAART Contract # NNL10AA05B

configuration for this series of AIAA workshops is the so-called DLR-F11 configuration (see 

Fig.1). After having given some serious consideration to adopting this reference geometry for 

the present study, it was decided to proceed instead with the development of a high-lift variant of 

the public-domain Common Research Model (CRM) [6], also shown in Figure 1. Developing a 

high-lift variant of the CRM configuration avoids drawbacks associated with the DLR-F11 model such as a non-functional integration of leading-edge and trailing-edge high lift devices, lack of 

tail and landing gear definition, and limited access to wind-tunnel test data.

2.1 Conventional High-Lift CRM Configuration

The CRM configuration is representative of a twin-aisle transonic civil transport aircraft 

which was designed only for a single transonic cruise condition. The static aeroelastic effects at 

a nominal cruise condition are accounted for by defining a wing shape showing the effects of 

aerodynamic loads rather than adopting its jig shape. To facilitate the aerodynamic design of a 

representative conventional high-lift wing design, the CRM wing was re-sheared to allow for a 

linear slat trailing edge while retaining total wing dihedral. (Figure 2).

In a planform view, Figure 3 shows the layout of the leading-edge slats and the trailing-edge 

single element flaps with Fowler motion. The arrangement and the geometric dimensions (e.g., 

chord ratios, span extents of the flaps) were chosen from a survey of current production civil 

transport aircraft. This is illustrated in Figure 4 which shows planform views for a cross section

of current fleet transonic transport aircraft along with an indication of the type of leading-edge 

(LE) and trailing-edge (TE) devices employed. From studying Figure 4, it appears that the trend 

in current fleet aircraft is toward high lift systems that combine slats with single-element slotted 

flaps with Fowler motion. Hence, this layout was selected for developing the high lift variant of 

the CRM configuration.

The planform views in Figure 4 also guided the selection of the span extents of the slats and 

trailing edge flaps. The slat and trailing edge flap chord ratios in Figure 3 were chosen to be in 

family with those for the current fleet aircraft listed in Figure 4. Here, ‘in family’ means that 

they were chosen to roughly fall in the middle of the bands of the leading-edge and trailing-edge 

device chord ratios shown in Figures 5 and 6.

Likewise, the range of the Fowler motion of the slotted trailing edge flaps of the conventional 

high-lift CRM was also chosen to be in family with the chord ratios defined by the distance between spoiler trailing edge and flap trailing edge for the fleet aircraft used to build the plot in 

Figure 7.

2.2 Simplified High-Lift CRM Configuration

The intent of defining a simplified high-lift CRM configuration is to reduce weight, part 

count, and excrescence drag due to any fairings needed to aerodynamically dress the mechanical 

systems for operating the high-lift devices, particularly, the slotted trailing edge flaps with 

Fowler motion. In that vein, a simply hinged trailing edge flap is the simplest possible trailing 

edge device for increasing wing camber. Allowing for a slotted flap configuration maintains the 

mechanical simplicity of the simply hinged trailing edge flap while making for aerodynamically 

more effective designs.

To protect against premature stall, these two types of trailing edge high-lift devices need to 

be paired with a leading edge device. One choice is a Krueger flap design for which some mechanically rather simple designs are available. A drooped leading edge is mechanically still sim2
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SMAAART Contract # NNL10AA05B

Table 1 All six simplified high-lift systems in this study aim for reductions in weight, 

part count, and excrescence drag.

pler than a slat while avoiding the careful rigging required for an effective slat installation. 

While both Krueger flaps and drooped leading edge designs are found on current fleet aircraft, 

morphing leading edges are sitting at low technical readiness levels (TRLs) but hold the promise 

of being simpler while being as effective aerodynamically if not more so than either of the other

two leading edge device concepts.

The permutation of any of these five high-lift device concepts gives the matrix of six possible 

high-lift design concepts shown in Table 1. To accommodate resource and schedule constraints,

a Quality Function Deployment (QFD) approach was adopted to identify the preferred simplified 

high lift concept among the options shown in Table 1. The QFD approach employed in this 

study is illustrated in Figure 8. On the left-hand side of the House-of-Quality in Figure 8, the 

programmatic requirements are listed. As they were provided by NASA, they are being referred 

to as “NASA Requirements.” Along the ceiling underneath the pitched roof, a set of derived requirements is spelled out. They are referred to as technical requirements that a design must meet 

to achieve the system or programmatic requirements. For each technical requirement, a QFD 

score is assigned. Adopting the nonlinear scale commonly used in QFD assessments, the ratings 

are in powers of three (i.e., 1 for low importance, 3 for intermediate importance, and 9 for high 

importance). A weight is applied to each QFD score, reflecting NASA’s prioritization of their 

system performance requirements. Summation of the weighted technical requirement scores 

gives a technical priority score.

In the basement of the House of Quality, the six possible combinations of leading-edge and 

trailing-edge high lift devices are listed as ‘concepts.’ A score is assigned to each ‘concept’ in 

each technical requirement. These score reflect grading on a curve with ‘1’ being the lowest and 

‘5’ being the highest score. The total score for each high-lift concept is computed from summing 

up each concept score, scaled with its technical priority score. This QFD approach identifies a

preferred simplified high-lift CRM concept that features drooped leading edge devices and simply hinged trailing edge flaps.

The pitched roof of the House of Quality permits taking into account synergistic or canceling

interactions of any two or more particular technical requirements. This option was not exercised

in the present study. This additional evaluation dimension is considered to add a layer of refinement that seemed out of line with the rough order of magnitude estimation approach adopted 

here to identify a preferred simplified high lift CRM system.

For the same wing planform shown in Figure3, Figure 9 shows the layout of the simplified 

high-lift CRM wing. The chord ratio and span extents of the drooped leading edge devices are

close to that of the slats of its conventional high-lift CRM sibling, as are chord ratios and span 

extents for the conventional and the simplified trailing edge flaps. This similarity in planform 

layout for both simplified and conventional high-lift wing concepts aims to minimize the impact 

of any geometric differences on the high-lift aerodynamics.

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Table 2 Listing of all computational grids utilized in CFD-based assessment of CRM

high-lift aerodynamics.

Case Description Nacelle/

Pylon

# of 

Zones # of Points

Conventional, landing, low/high RN On 61 62.9 million

Conventional, landing/takeoff, high RN Off 38 52.2 million

Conventional, landing/takeoff, low RN Off 38 48.8 million

Simplified, landing, high RN Off 20 36.7 million

Simplified, landing, low RN Off 20 33.7 million

Simplified, takeoff, high RN Off 18 36.4 million

Simplified, takeoff, low RN Off 18 33.5 million

Note that the simply hinged flap will require a spoiler design different from that adopted for 

the conventional high-lift design. In absence of a Fowler motion capability, the trailing edge of 

the spoiler for the simplified high-lift system has to lie upstream of the leading edge of the simply hinged trailing edge flap.

3.0 CFD-BASED ASSESSMENTS OF HIGH-LIFT AERODYNAMICS

3.1 Setup of CFD Analyses

A series of overset grids was built to model the various high lift CRM configurations. Surface grid generation was accomplished using a Boeing-developed program called MADCAP. 

Building of the volume grids and establishing connectivity for zone-to-zone communication relied on NASA tolls such as Chimera Grid Tools and Pegasus5. Table 2 summarizes the grids 

employed in the present high-lift performance assessments. There were grids for the CRM 

wing/body and wing/body/nacelle/pylon configurations. The former configuration was gridded 

for takeoff and landing approach configurations, at wind-tunnel (i.e., ‘low’ Reynolds number) 

and fight conditions (i.e., ‘high’ Reynolds number). The number of grids points used varied with 

complexity of the configuration and with the freestream condition (i.e., wind-tunnel and flight 

conditions). As for the latter, care was taken to resolve the viscous sublayer by using a grid 

spacing of y+

=1 along and normal to solid walls.

All computations were carried out with the OVERFLOW flow solver. This is a Reynolds 

Averaged Navier-Stokes (RANS) code widely used to compute subsonic, transonic, and supersonic flow over complex configurations. This node-centered flow solver is designed for structured overset grids systems like those built for this high-lift CRM study. The OVERFLOW 

solution system can be run in many modes. Those employed for the present analyses are summarized as follows

x Code version 2.2g

x Spalart Allmaras turbulence model with curvature correction (SA-RC)

x Flux-difference splitting for right hand side (RHS) of discretized RANS equations

x Central differencing for left hand side (LHS) of discretized RANS equations

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AFC-Enabled High-Lift System Study

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Table 3 Run Matrix for the conventional high-lift CRM configuration with angle-of-attack 

schedules.

Horiz Altitude Temp Alph

Series Run Priority Tail IB Slat OB Slat IB Flap OB Flap (ft) (deg R) Re/in Remac Mach Flag

1 1 2 off on retracted retracted retracted retracted flight 10,000 483.01 89268.3 24,620,197 0.20 A5

1 2 2 off off retracted retracted retracted retracted flight 10,000 483.01 89268.3 24,620,197 0.20 A5

2 1 2 off on retracted retracted retracted retracted flight 10,000 483.01 89268.3 24,620,197 0.20 A1

2 2 2 off on retracted retracted retracted retracted flight 10,000 483.01 89268.3 24,620,197 0.20 A2

2 3 2 off on retracted retracted retracted retracted WT sea level 518.67 11834.5 3,263,955 0.20 A1

2 4 2 off on retracted retracted retracted retracted WT sea level 518.67 11834.5 3,263,955 0.20 A2

2 5 2 off off retracted retracted retracted retracted Flight 10,000 483.01 89268.3 24,620,197 0.20 A5

3 1 1 off on 12.5 45 30 30 WT sea level 518.67 11834.5 3,263,955 0.20 A3

3 2 1 off on 12.5 35 30 30 WT sea level 518.67 11834.5 3,263,955 0.20 A4

3 3 1 off on 12.5 25 30 30 WT sea level 518.67 11834.5 3,263,955 0.20 A4

3 4 1 off on 12.5 45 30 30 Flight 10,000 483.01 89268.3 24,620,197 0.20 A3

3 5 1 off off 12.5 25 30 30 WT sea level 518.67 11834.5 3,263,955 0.20 A3

3 6 1 off off 18 25 30 30 WT sea level 518.67 11834.5 3,263,955 0.20 A3

3 7 1 off off 18 25 30 30 Flight 10,000 483.01 89268.3 24,620,197 0.20 A3

4 1 3 off off 18 25 15 15 WT sea level 518.67 11834.5 3,263,955 0.20 A3

4 2 3 off off 18 25 15 15 Flight 10,000 483.01 89268.3 24,620,197 0.20 A3

Alpha Flags

A1: D = 0 to 20 deg in 4 deg increments

A2: D = 11, 13 deg

A3: D = 0, 4, 8, 12, 13, 14, 15, 16, 18, 19, 20, 21 deg

A4: D = 12, 15, 18, 19, 20, 21 deg

A5: D = 4, 8, 12, 13, 14 deg

TE Deflection (deg)

Landing

Clean Wing: Low Speed Loft (re-sheared for straight slat TE)

Takeoff

High-Lift CRM OVERFLOW Runs - CONVENTIONAL

LE Deflection (deg)

Clean Wing: High Speed Loft (original geometry from DPW-IV)

Nacelle / 

Pylon

Reynolds No.

x Low-Mach number preconditioning for mitigating stiffness of compressible discretized 

RANS equations at low-speed freestream conditions

All cases were run with the assumption of fully developed boundary layers (i.e., without laminar/turbulent transition modeling). For computations for flight conditions, the Reynolds number 

was 24.6 million, corresponding to a freestream Mach number of 0.2 at an altitude of 10,000 ft. 

For computations at wind-tunnel conditions, the Reynolds number was set at 3.3 million based 

on an assumed model scale of 10%, sea-level conditions, and a freestream Mach number of 0.2.

The geometric parameters used for computing all force and moment coefficients are summarized as follows:

x Cref = 275.8 in (MAC)

x Sref/2 = 297,360.0 in2

x b/2 = 1156.75 in

x Xref = 1325.9 in, Yref = 0.0 in, Zref = 177.95 in

3.2 CFD Run Matrices

Tables 3 and 4 provide the run matrices for all CFD analyses including angle-of-attack 

schedules. As summarized in Table 3, for the conventional high-lift CRM configuration, 157 

flow field solutions were computed for 10 configurations including nacelle/pylon on or off, clean 

wing set up(i.e., all leading-edge and trailing-edge devices are assumed in a fully retracted position), and landing and takeoff configurations with different slat and trailing edge flap settings. 

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Table 4 Run matrix for the simplified high-lift CRM configuration with angle-of-attack schedules.

Horiz Altitude Temp Alpha

Series Run Priority Tail Nac / Pyl IB Slat OB Slat IB Flap OB Flap (ft) (deg R) Re/in Remac Mach Flag

1 1 1 off off 20 25 30 34 WT sea level 518.7 11835 3,263,955 0.20 A0

1 2 1 off off 20 25 30 34 Flight 10,000 483.0 89268 24,620,197 0.20 A0

1 3 1 off off 0 0 30 34 WT sea level 518.7 11835 3,263,955 0.20 A0

1 4 1 off off 10 15 30 34 WT sea level 518.7 11835 3,263,955 0.20 A0

2 1 2 off off 20 25 15 17 WT sea level 518.7 11835 3,263,955 0.20 A0

2 2 2 off off 20 25 15 17 Flight 10,000 483.0 89268 24,620,197 0.20 A0

Alpha Flag

A0: D = 0, 4, 8, 12, 13, 14, 15, 16, 18, 19, 20, 21 deg

Takeoff

High-Lift CRM OVERFLOW Runs - SIMPLIFIED

LE Droop (deg) TE Droop (deg)

Landing

Reynolds No.

There are 55 solutions for the clean wing setting, 80 solutions for landing and 22 solutions for 

takeoff configurations.

As summarized in Table 4, for the so-called simplified high-lift CRM configuration, 52 flow 

field solutions were computed for four wing/body configurations in landing and takeoff configurations with different dropped leading edge and trailing edge flap settings. There are 30 solutions for landing and 22 solutions for takeoff configurations.

3.3 CFD Results for Wing/Body/Nacelle/Pylon Configurations

Figure 10 shows the variation of lift coefficient with angle-of-attack for a cleanwing/body/nacelle/pylon configuration at freestream flight conditions. At 12 degrees angle of 

attack, a glitch in the lift curve is observed. This phenomenon is explained with the liftversus/pitching moment plot along with wing surface flow visualizations in Figure 11. The initial kink in the lift/pitching-moment curve is associated with an angle of attack of 12 degrees. 

Comparing wing surface flow plots at angles of attack of 12 and 13 degrees indicate that separated flow from the nacelle is swept over the wing, causing the wing flow to separate as well downstream of the installed nacelle. The second kink in the lift-versus/pitching moment plot is 

associated with the onset of outboard wing stall. This is inconsistent with established high-lift 

design practice that aims to have the inboard wing sections stall first to maintain roll control.

To pinpoint the root causes for these undesirable flow separation characteristics, the conventional wing/body/nacelle/pylon configuration was re-run in a landing approach configuration at 

both wind-tunnel and flight conditions. The results for these runs are summarized in Figure 12. 

Figure 12 shows the variation of lift with angle-of-attack for a clean and a landing configuration, 

and surface flow visualizations in top views for maximum and post-stall conditions. As expected, lift increases with higher freestream Reynolds number as the Navier-Stokes solutions 

comes closer to the optimum inviscid flow solution as the Reynolds number is raised. The benefits of increased camber (i.e., shift in lift curve slope) and improved stall protection due to the 

deployed slats (i.e., increase install angle of attack) are obvious. For the landing configuration, 

one would expect higher values for maximum lift and the associated angle of attack based on the 

performance of current fleet aircraft which the conventional high-lift CRM configuration was 

intended to be representative for. This can be understood by inspection of the surface flow images. At maximum lift, a rapid growth of the region of separated flow across the mid-wing section id observed due to a poor nacelle/wing integration at maximum and immediate post-stall 

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conditions. Note that both surface flow images indicate stalled flow over the outboard wing section. It appears that while this is undesirable from an aircraft control perspective, it is of secondary importance to maximum lift performance.

In view of the results in Figure 12, a decision had to be made whether to improve the nacelle/pylon/wing integration or to bypass this issue by reverting to wing/body-only CFD analysis. Considering resource and schedule constraints, the decision was made to process with the 

CFD analysis for wing/body high-lift CRM configurations.

3.4 CFD Results for Wing/Body Configurations

Figure 13 illustrates the modifications to the conventional and simplified high-lift CRM configurations. The nacelles and pylons have been removed, and a transition element has been designed to fill in the gap between inboard and outboard leading edge devices that was created by 

removal of the pylon. This transition element introduces some waviness in the leading edge contour due to the differential positioning of the inboard and outboard slats and drooped leading 

edge settings for the conventional and simplified high-lift CRM configurations, respectively. 

Figure 13 also shows the difference in trailing edge flap settings for the conventional and simplified high-lift designs. For the conventional high-lift CRM configuration, the trailing edge flaps 

is deflected down and moved aft, creating a slot between main wing and flap. For the simplified 

high-lift CRM concept, the trailing edge flap is deflected down only without any slot-creating 

Fowler motion.

Figure 14 shows the variation of lift coefficient with angle of attack for the conventional 

CRM configuration with and without nacelle/pylon installations. Results for several slat settings 

are shown, all computed at wind-tunnel freestream conditions. Without the sub-optimal nacelle/wing installation, maximum lift increases by about five percent and the associated angle of 

attack by about a couple degrees. One could say that the trends were to be expected, but that the 

magnitude of the high-lift improvements are still falling short of expectations for a high-lift design intended to be representative of the high-lift performance of current fleet transonic transport 

aircraft. As illustrated by the surface flow visualization inserts, this shortfall is due to the premature stalling of the outboard wing sections. If anything, these non-representative wing stall characteristics appear to be exacerbated by the high-lift system when compared to the stall 

characteristics of the clean wing configuration (see Figure 10). This shortfall in maximum lift 

performance is expected to be cured by re-twisting and re-cambering the CRM wing which turns 

out to be a point design for transonic cruise only.

Figures 15 and 16 compare lift performance for the clean wing configuration with those for 

the conventional and simplified high-lift CRM configurations. Figure 15 shows these comparisons for the respective high-lift CRM concepts for a landing-approach rigging, and Figure 16

repeats these comparisons for takeoff settings. As before, lift increases for any give angle of attack when going from wind-tunnel to flight freestream conditions. As also is to be expected, the 

computed lift values for the simplified high-lift CRM configuration fall somewhere between 

those for the clean and the conventional high-lift CRM configurations.

While Figure 14 has already addressed the sensitivity of the high-lift performance of the conventional high-lift CRM concept to slat settings, the computed lift variations with angle of attack 

in Figure 17 show that the drooped leading edge setting for the simplified high-lift CRM concept 

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that was used for the comparisons in Figure 15 and 16 is close to one producing optimum highlift performance.

3.5 Performance Targets for AFC-Enhancements to Simplified High-Lift CRM Concepts

The purpose of the CFD-based high-lift assessments for the conventional and simplified 

high-lift CRM concepts is to quantify the decrements in high-lift performance for the simplified 

high-lift system that need to be recovered by the integration of AFC. Figure 18 compares the 

variation of incremental lift with angle of attack for the conventional and the simplified high-lift 

CRM wing/body concepts at wind-tunnel and flight freestream conditions. For any given angle 

of attack, the largest loss in lift is observed for the landing configurations at low Reynolds number (i.e., wind-tunnel) conditions. Note that these comparisons are for untrimmed lift conditions. 

Figures 19 and 20 help to assess the effect of trim on these lift decrements.

Figures 19 and 20 compare pitching moments for the clean wing configuration with those for 

the conventional and simplified high-lift CRM wing/body configurations. Figure 19 shows these 

comparisons for the respective high-lift CRM concepts for a landing-approach rigging, and Figure 20 repeats these comparisons for takeoff settings. In general, the Reynolds number effects 

manifest themselves by yielding higher pitching moment values for any given lift value when 

going from wind-tunnel to flight freestream conditions. The same observations hold true when 

comparing lift for any given pitching moment when going from wind-tunnel to flight freestream 

conditions. 

From Figures 19 and 20, it appears that the conventional high-lift CRM wing/body concept 

produces more nose-down pitching moment than the simplified CRM wing/body concept. This, 

in turn, requires more download from the horizontal tail for the conventional than for the simplified high-lift CRM concept. This tail download will reduce lift for the trimmed conventional 

high-lift CRM concept. In other words, the lift decrements shown in Figure 18 that need to be 

recovered through AFC deployment are conservative. They should be expected to be reduced 

for trimmed full CRM concepts (i.e., wing/body/nacelle/tail configurations).

4.0 AFC SIZING

Several AFC concepts have been explored over the years. Prevalent among them are pneumatic and synthetic AFC actuators. To take advantage of the data obtained in a NASA/Boeing 

wind-tunnel test of an AFC-enhanced full-scale Boeing-757 vertical tail at the National Full 

Scale Aerodynamic Complex (NFAC), the decision was made to consider pneumatic AFC actuators for recovering the high-lift performance of the simplified high-lift CRM wing/body concept 

to match that of its conventional high-lift counterpart. It should be noted here that public-domain 

documentation of this NFAC AFC wind-tunnel test is forthcoming but not in hand yet.

The AFC augmentation for the simplified high-lift CRM concept needs to be sized to match 

the performance of the conventional high-lift CRM at both landing-approach and takeoff conditions. There is some debate as to the ability of CFD to reliably predict the stall characteristics of 

high-lift transport configurations. To avoid having the present AFC sizing exercise contaminated by potentially contentious Navier-Stokes solutions at stall, so-called nominal landingapproach and takeoff conditions have been selected for estimating AFC requirements for a simplified high-lift CRM concept.

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Figure 21 compares computed surface pressures for the conventional and the simplified highlift CRM configurations along with offbody iso-surfaces that indicate the enveloped of separated 

flow. The iso-surfaces represent the location of off-surface flow with a zero axial velocity component. It can be seen that the both trailing-edge flaps of the simplified high-lift CRM are fully 

separated. This is the prime contributor to the loss of high-lift performance of the simplified 

high-lift CRM compared to its CRM counterpart with the conventional high-lift system. This 

comparison is illustrated in Figure 21 for a so-called nominal landing condition where the approach speed is large than 123% of the stall speed. The comparisons of lift variation with angle 

of attack in Figure 21 show that the so-called nominal approach condition occurs well within the 

linear range of the lift curves with a safe margin against maximum-lift or stall conditions.

The lift curves in Figure 21 are plotted for wind-tunnel and flight freestream conditions. To 

quantify the lift that needs to be recovered with AFC, first locate the intercept between the lines

indicating nominal lift coefficient with the lift curve for the conventional high-lift CRM concept. 

The difference in lift between the simplified and conventional high-lift CRM concept at the angle-of-attack associated with this intercept gives the lift that needs to be recovered by AFC. Applying this process at flight freestream conditions shows that the AFC performance needs to 

recover about 18% of the lift decrement to produce matching lift for both simplified and conventional CRM high-lift systems. Repeating this process for the lift traces computed at wind-tunnel 

freestream conditions indicates that lift decrement to be 30% going from the conventional to the 

simplified CRM high-lift system.

For a nominal takeoff condition (i.e., take off speed is greater than 113% of the stall speed), 

Figure 22 shows the variation of lift with angle of attack and the variation of lift with lift-to-drag 

ratio. The latter is important as the lift-to-drag ratio at takeoff is an important driver in the mission performance of commercial transport aircraft, as will be shown later in the net present value 

(NPV) analysis section. As before in the comparisons in Figure 21, choosing a nominal takeoff 

condition ensures that the AFC estimates are computed for the linear segments in the lift curves 

separated from stall conditions by a robust margin.

Applying the same analysis process as used with Figure 21, it turns out that AFC needs to recover about 15% in lift decrement at both wind-tunnel and flight freestream conditions to produce matching lift for both simplified and conventional high-lift systems. In addition, for 

takeoff, this needs to happen while maintaining the lift-to-drag ratio to avoid adverse impact of 

an AFC integration on the overall mission performance of an AFC-enhanced simplified high-lift 

CRM concept.

In surveying the lift decrements at landing-approach and takeoff conditions due to the mechanical simplification of the CRM high-lift system, it appears that an AFC sizing at wind-tunnel 

freestream conditions overall is more challenging than at flight freestream conditions. To produce conservative AFC estimates, the decision was made to size AFC at wind-tunnel freestream 

conditions.

4.1 Inviscid Performance Limit Concept for AFC Sizing

Previous studies have shown that optimum AFC sizing can be characterized as recovering in 

the limit inviscid flow which by definition is free of flow separation. Sizing AFC in excess of 

this so-called inviscid limit, flow control turns into circulation control which is still an effective 

while no longer the most efficient means of flow control. This is illustrated Figure 23. This fig9
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ure shows the variation of side force with momentum of the AFC discharge for the aforementioned Boeing-757 vertical tail. Without AFC, the flow over the deflected rudder is separated. 

With an AFC discharge of a specific momentum of about 0.11, the flow over the deflected rudder remains attached and produces a sideforce equal to that computed from an Euler solution for 

inviscid flow over the same Boeing-757 vertical tail configuration. If the specific momentum of 

the AFC discharge is further increased beyond 0.11, the slope of the side force trace drops significantly, becoming typical of side force sensitivities found in circulation control setups.

Applying these lessons learned from the Boeing-757 AFC sizing, first inviscid limit solutions 

were computed from Euler solutions for flow over the simplified high-lift CM configuration with 

dropped leading edges and trailing edge flaps at landing approach settings. The next step involves Navier-Stokes solutions over the same simplified high-loft CRM configuration where the 

AFC discharge was simulated by applying constant blowing boundary conditions. These AFC 

mass flow boundary conditions prescribe total pressure and direction of the velocity vector along 

strips just upstream of the trailing edge flaps.

4.2 CFD-Based AFC Sizing

Figure 24 summarizes the results from these initial AFC sizing estimates. This composite 

figure shows variation of lift with AFC mass flow in the right upper quadrant. This plot also 

gives the lift level of the conventional high-lift CRM configuration. This is the target lift level 

that needs to be recovered by AFC deployment. In addition the inviscid lift limit is indicated as 

well. Any AFC solution where the achieve lift exceeds this inviscid limit indicates that the dialed-in AFC mass flow coefficient is too high producing a blend of AFC and circulation control.

The plot in the lower right quadrant in Figure 24 shows spanload distributions. To maintain 

the lift-to-drag ratio of the conventional high-lift CRM configuration, AFC needs to be sized to 

resemble the spanload distribution of that conventional high-lift CRM configuration. Also in 

Figure 24 are three plots that give the computed surface pressures and iso-surfaces that indicate 

separated flow (see previous discussion of similar plot layouts for Figure 21). These flow visualizations correlate with the data points in the Cl-vs-Cq plot in the right upper quadrant. AFC applied to both inboard and outboard trailing flaps is twice as effective as AFC application to the 

inboard trailing-edge flap only. Subsequently the AFC mass flow over the outboard trailingedge flap was reduced to more closely match the spanload distribution (i.e., the L/D ratio) of the 

conventional high-lift CRM configuration.

This initial step in the current CFD-based AFC sizing process was followed up with a more 

detailed AFC model. As shown in Figure 25, backward-facing steps were carved into the upper 

trailing-edge flaps surfaces. The vertical extend of these steps is scaled from the AFC integration validated in the aforementioned NFAC Boeing-757 AFC test. Along the vertical faces of 

the backward facing steps, AFC is either modeled as blowing through a continuous strip or as 

discharge through discrete ports.

The discrete AFC port computer model with 19 ports along the inboard and 18 ports along 

the outboard flap is illustrated in Figure 26. Each port consists of a set of grids, meaning that the 

internal port flow is modeled. Trial and error is employed to guide the orientation of the discrete 

ducts to achieve the desired Coanda effect for energizing the boundary-layer flow over the flaps. 

An initial nozzle orientation at a steeper angle to the flap surface proved ineffective in re10
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attaching the flow over the trailing edge flap. A later iteration with the nozzle cut into the flap 

surface at a shallower angle proved effective.

Figure 27 adopts a layout similar to that of Figure 24. Whereas the results in Figure 24 were 

for continuous blowing across slots over the full spanwise extent of both inboard and outboard

trailing edge flaps, Figure 27 shows results for a discrete AFC model with constant blowing 

through a bank of 37 ports. The AFC mass flow rate was dialed in to roughly match the 

spanload distribution of the conventional high-lift CRM concept. This setup succeeds in simultaneously approximating the takeoff lift-to-drag characteristics of the conventional high-lift 

CRM concept while recovering its lift. At an aircraft systems level, this AFC setup translates 

into an air mass flow rate of 72 lbm/sec at an operating pressure of 80 psia for the full wing (i.e., 

left and right wing halves).

5.0 TRADE STUDY ON AFC-ENABLED SIMPLIFIED HIGH-LIFT SYSTEM

This section discusses results from a system-level trade study on the mechanically simplified 

high-lift system to accommodate a pneumatic AFC system defined and sized in the preceding 

section. The purpose of this trade study is to identify any potential weight and fuel savings from 

the mechanically simplified high-lift system. This requires a baseline transonic transport concept 

configuration to increment any such potential savings from. In this study, the NASA ERA-0003

concept was chosen to be that baseline concept. The ERA-0003 concept represents a twin-aisle 

275-passenger transonic transport aircraft with 1990-era technology projections [7]. Its high lift 

system was comprised of Krueger flaps and of single-element slotted trailing edge flaps. These 

features make the ERA-003 concept comparable to one that the CRM OML was defined for. 

Key dimensions for the ERA-0003 concept are provided along with a 3-view and a cabin layout 

in Figure 28.

Four different options have been considered for providing the pressurized air mass flow rate 

of 72 lbm/sec at an operating pressure of 80 psia that the CFD-based AFC estimation calls for:

Option 1: use existing tail-mounted APU

Option 2: install a dedicated APU system

Option 3: siphon off turbine bleed

Option 4: siphon off core bleed

Option 1 is derived from the AFC installation on the recent Boeing-757 NFAC full-scale experiment, where the AFC mass flow rate and operating pressure are consistent with the capability 

of a Boeing-757 APU. While the CFD-based estimates of the power requirements of the current 

AFC high-lift application were found to scale well with data from that recent Boeing-757 vertical 

tail AFC test at NFAC, this also means that the power requirements exceed the capabilities of 

APU systems that would be typical of commercial transonic transports. This situation gets exacerbated by the fact that one would have to provide for redundancy in case of an APU failure and 

by the long pressure lines from a tail-mounted APU to the wing-mounted AFC units. These considerations eliminated Option 1 from further consideration.

For option 2, it was assumed that engine core bleed would provide an air mass flow rate of 8 

pps at 80 psia. For the two-engine ERA-0003 concept, this means that two dedicated APUs 

would be required to provide the remaining 56 pps of air mass flow at 80 psia. There are APUs 

available that would satisfy this requirement. For instance, the Honeywell 331-500 engine seems 

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to hold the potential for producing 28 pps of air mass flow at the 80 psia after some modifications, such as the removal of its power shaft capability while adding a second stage to the air 

compressor. Two of these engines would satisfy the AFC high-lift power requirements.

The Honeywell 331-500 engine measures 80” in length, 35” in height, and 48” in width. To 

minimize pressure line losses, two of these APUs would be installed in the cargo hold close to 

the wing. Each engine weighs in at about 650 lbs. Their operation would require provisions for 

inlet (350 sq.in. / unit) and exhaust (200 sq.in / unit) areas. Installation of these dedicated APUs 

would add further weight due to requirements for engine mounts, firewalls, fire extinguishing 

gear, controllers, valves, insulation, pre-cooler, and inlet and exhaust apertures with doors. One 

can easily see that this AFC high-lift design solution combines increases in weight and complexity with likely adverse impact on the configuration aerodynamics due to the dedicated APU inlet 

and exhaust areas. Hence, option 2 was abandoned as being unlikely to lead to any of the desired 

aircraft OEW reductions or fuel savings.

The fan bleed design option (i.e., Option 3) is estimated to provide an air mass flow rate of 

100 pps, albeit at only 22 psia operating pressure. This alone disqualifies Option 3 from further 

consideration, apart from the fact that the duct diameter from the engine to the wing would exceed the pylon width and require unconventional propulsion/airframe integration solutions.

The core bleed option (i.e., Option 4) was considered for the system performance study in the 

remainder of this section. This option is considered to much more readily integrate with the 

overall ERA-0003 configuration than any of the other alternatives. It also provides the AFC operating pressure that the AFC power estimates call for. The operating pressure is considered 

more critical than the air mass flow rate for the AFC implementation. Thus, pursuing option 4 

implies that further AFC high-lift trade studies can address the mass flow rate requirements (e.g., 

fewer AFC units or efficiency gains by going to unsteady/intermittent operations rather than continuous blowing).

Figure 29 summarizes the mission performance for the ERA-0003 baseline configuration, 

and for an un-scaled (center column in Figure 29) and for a re-sized AFC-variant of the ERA0003 (far right column in Figure 29). The un-scaled AFC variant looks photographically like the 

baseline ERA-0003, accounts for a lower weight (750 lbs OEW reduction due to AFC) and a reduction in overall drag by 3.3 counts due to elimination of high-lift system fairings. As the core 

bleed required to feed the AFC system reduces maximum available engine thrust, TOFL increases by about 1,100 ft. The un-scaled AFC variant of the ERA-0003 concept shows a 1.7% gain in 

range compared to the baseline ERA-0003 design. This is partially due to the increase in fuel 

which was increased to have both baseline and un-scaled AFC variant of the ERA-0003 take off 

at the same TOGW. Re-running the performance analysis for same range for both baseline and 

un-scaled AFC variant of the ERA-0003 indicates fuel savings of 400 gals of fuel for the unscaled AFC variant.

The re-sized AFC variant of the ERA-0003 increases engine size to make up for the core 

bleed. The bigger engine and a slightly larger wing recover takeoff performance to match that of 

the baseline ERA-0003 concept while that overall still yielding mission fuel savings. These results are illustrated in the ‘thumbprints’ in Figures 30 and 31. These plots show lines of constant 

TOGW, block fuel, rejection field length, optimum altitude, and distance to climb, all plotted 

against engine thrust (vertical axis) and wing area (horizontal axis). For the baseline ERA-0003

concept in Figure 30, the mission calls for a climb in 200 nm to an initial cruise altitude (ICA) of 

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35,000 ft. The point where the lowest possible constant fuel burn trace touches the climb-to-ICA 

boundary sizes the ERA-0003 concept. For the re-sized AFC variant of the ERA-0003 concept

in Figure 31, engine thrust and wing area are sized by takeoff field length instead of rate of climb 

(i.e., thrust and wing area are identified by the point where the lowest possible constant fuel burn 

trace touches the TOD or takeoff distance constraint).

5.1 Net Present Value (NPV) Assessment

The NPV assessment is carried out for the un-sized AFC variant of the ERA-0003 concept. 

Flying the same mission as the baseline ERA-0003 concept, its AFC variant is estimated to save 

2,595 lbs or 399.2 gals of jet fuel per flight. Assuming an average flight duration of 16 hours per 

trip and major maintenance intervals of 10,000 flight hours, there are 625 flights between major 

maintenance stand downs. With 10 major maintenance cycle during an assumed 20-year operational life, the aircraft will complete about 6,000 flights in its lifetime or 300 flights for every 

year of operation. This translates to lifetime fuel savings of close to 2.4 million gallons of fuel.

The amount of cost savings over the lifetime of an AFC-enabled ERA-000-type aircraft depends on assumptions of fuel price and of a discount rate. The discount rate is a function of the 

weighted average cost of capital (WACC) of the operator of such an aircraft. The latter directly 

depends on the capital structure of the operator and its credit rating, and indirectly on macroeconomic factors such as market interest and inflation rates. For the sake of producing an NPV dollar estimate, let the fuel price be assumed at $2/gal on average for the 20-year operational flight, 

and a discount rate of 10%. This would translate into an NPV of $2.25M per AFC-variant of an 

ERA-0003-like concept over its 20-year operational life. While this dollar amount might be subject to debate due to the assumption about future fuel prices and operator WACC, the key takeaway here is that the NPV is positive, indicating that there might be value in considering AFCenabled mechanically simpler high-lift system for an ERA-0003-class aircraft.

6.0 ANALYSIS, CONCLUSIONS, AND RECOMMENDATIONS

The key findings of the current trade study on AFC-enhanced mechanically simplified highlift systems for commercial transports indicate platform performance improvements due to OEW 

savings (primarily due to the elimination of the mechanical systems facilitating the Fowler motion of the trailing-edge slotted flaps) and improved aerodynamics (i.e., elimination of high-lift 

systems fairings). These platform performance improvements translate into a positive NPV, indicating value to an operator to acquire such AFC-enhanced transport aircraft, although the actual NPV amount might be subject to debate.

These key findings hinge on the assumption that further AFC integration studies will identify 

options for significant reductions in the current AFC power requirements. First pass CFD-based 

AFC power estimates call for an air mass flow rate of 72 pps at an operating pressure of 80 psia, 

whereas core bleed from Boeing-777-type 1990’s-era turbofan engines provides an air mass flow 

rate of about 8 pps per engine at an operating pressure of 80 psia. This suggests an AFC performance improvement by a factor of four. In reality, the improvement targets need to be much 

more stringent, as the indicated core bleed of 8 pps mass flow at 80 psia requires running the engines at full throttle which is impractical during landing approach.

Several potential options have been identified for reducing the AFC air mass flow rate requirements. First off, the high-lift device setting schedules have thus far been adopted from con13
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ventional high-lift systems. They might have to be adjusted for AFC-enabled mechanically simplified high lift systems. Secondly, sensitivities of the AFC unit placements such as spacing of 

AFC units and the alignment of AFC jets with wing surface flow need to be better understood. 

Thus far only pneumatic AFC units with constant blowing capability have been considered. 

There might be gains in efficiency by going to sweeping jet-type AFC units or adopting different 

AFC technologies such as piezoelectric instead of pneumatic AFC units. Finally, the Reynolds 

number effects need to be better understood as it appears that they are immaterial for takeoff 

conditions yet relevant for landing-approach operations.

The results of the current AFC high-lift trade study are also somewhat clouded by CRM 

high-lift characteristics that are non-representative of current high-lift design practices. It is recommended to reintegrate the nacelle/pylon assembly with the CRM wing/body geometry to eliminate any premature flow separation in the mid-span section with the attendant loss in maximum 

lit capability. In addition, the CRM out wing section ought to be re-cambered and re-twisted to 

move the onset of wing stall toward the inboard section.

Once the high-lift baseline CRM configuration has been revised, and a suitable AFC integration for high-lift applications has been identified, future studies on AFC-enhanced simplified 

high-lift systems should shift their focus on validating this technology in large-scale wind-tunnel 

tests, followed by flight tests to make this technology ready for integration with future transporttype aircraft.

7.0 REFERENCES

[1] McLean, J.D., Crouch, J.D., Stoner, R.C., Sakarai, S., Seidel, G.E., Feifel, W.M., and Rush, 

H.M., “Study of the Application of Separation Control by Unsteady Excitation to Civil 

Transport Aircraft, NASA Contractor Report, NASA-CR-1999-209338, June 1999.

[2] DeSalve, M., Whalen, E., and Glezer, A.,”High-Lift Enhancement Using Active Flow Control,” AIAA Paper 2012-3245, June 2012.

[3] Kiedaisch, J., Demanett, B., and Nagib, H., “Active Flow Control Applied to High-Lift Airfoils Utilizing Simple Flaps,” AIAA Paper 2006-2856, June 2006.

[4] Harrison, N.A., Vassberg, J.C., DeHaan, M.A., and Gea, L.-M., “The Design and Test of a 

Swept Wing Upper Surface Blowing (USB) Concept, AIAA Paper 2013-1102, January 2013.

[5] Rumsey, C.L., and Slotnick, J.P., “Overview and Summary of the Second AIAA High Lift 

Prediction Workshop, AIAA Paper 2014-0747, January 2014.

[6] Vassberg, J.C., “Introduction: Drag Prediction Workshop,” Journal of Aircraft, Vol. 45, No. 

3, May – June 2008, p. 737.

[7] Bonet, J. T., Schellenger, H. G., Rawdon, B. K., Elmer, K. R., Wakayama, S. R., Brown, D., 

and Guo, Y. P., “Environmentally Responsible Aviation (ERA) Project – N+2 Advanced 

Vehicle Concepts Study and Conceptual Design of Subscale Test Vehicle (STV),” NASA

Contract Report 2013-216519, 2013

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Figure 1 NASA/Boeing Common Research Model (CRM) configuration was chosen 

over DRL-F11 concept for developing a public-domain high-lift research 

concept.

Figure 2 NASA/Boeing Common Research Model (CRM) configuration cruise wing 

loft was re-sheared for high lift system development.

8.0 FIGURES

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Figure 3 Conventional high lift CRM planform layout is representative of current fleet 

commercial transport aircraft.

CSLAT = 0.06 CPLAN

CSLAT = 0.25 CPLAN

Initial slat: trailing edge is a 

straight line in the FRP thru the 

endpoints at K = 0.10 and K = 1.0

CFLAP = 0.26 CPlan

Flap midspan definition K = 0.37

CFLAP = 0.26 CPLAN

CPLAN = 74.3032 inches

Flap inboard definition K = 0.15

CFLAP = 74.3032 inches

K = 0.72

K = 0.37

K = 0.15

K = 0.10

K = 1.0

26.1354

(Splr TE X/C=0.86)

40.0094

(Splr TE X/C=0.86)

40.0094

Flap LE

Spoiler / fixed wing TE

CSLAT = 0.17 CPLAN

K = 0.2980

K = 0.3116

K = 0.3453

K = 0.9878

RevA slat: outboard slat is a 

constant chord defined at the 

wing tip as CSLAT = 0.25 CPLAN

Figure 4 High-lift systems on current fleet aircraft trend toward combining slats with single-element slotted flaps.

 Distill high lift planform from a survey of 

current generation aircraft

 Choose representative device chord ratios and 

span extents

A300-600 A320-200 A330-200 A340-500 A380-800 ERJ-170 A350-900

B737-700 B747-8 B757-200 B767-200 B777-200 B787-8 B747-400

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Figure 5 CRM slat chord ratio is in Family with slat sizing on current fleet commercial 

transport aircraft.

Figure 6 CRM trailing-edge flap chord ratio is in family with sizing of similar trailing-edge devices on current fleet commercial transports.

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Figure 7 CRM spoiler-to-wing trailing edge chord ratio is in Family with control surface 

layouts on current fleet commercial transport aircraft.

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Figure 8 QFD assessment of simplified high-lift systems is illustrated by a house-of-quality.

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Figure 9 Conventional and simplified high lift system planform definitions are 

matched to promote meaning high-lift aerodynamics comparisons.

21.3598

K = 0.72

K = 0.10

K = 1.0

K = 0.3189

K = 0.3544

K = 0.9912

K = 0.1220

K = 0.37

36.7740

61.1823

61.3844

19.5989

Projection of drooped leading 

edge hinge line onto the wing 

upper surface

Projection of plain flap hinge line 

onto the wing upper surface

Figure 10 Variation of lift with angle of attack for a clean (i.e., all high-lift devices 

retracted) high-lift CRM wing/body/nacelle/pylon configuration..

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Figure 11 Lift-vs-pitching moment trace indicates progression in massive flow separation with increasing angle of attack.

Figure 12 Surface flow visualizations aid in the explanation of stall characteristics of 

conventional high-lift CRM.

Nacelle/pylon/wing interaction 

leads to premature wing stall 

across mid-wing region for the 

landing configuration

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Figure 13 Wing/body definitions of conventional and simplified high-lift CRM variants. 

Conventional High Lift System

Simplified High Lift System

Config GSLAT GFLAP

Takeoff 18° / 25° 15°

Landing 18° / 25° 30°

Config GLE GFLAP

Takeoff 20° / 25° 15°

Landing 20° / 25° 30°

Figure 14 Wing/body high-lift CRM configuration eliminates massive nacelle-induced 

midspan flow separation and improves stall characteristics.

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Figure 16 Lift curve comparisons – clean vs. takeoff high-lift CRM configurations 

(conventional & simplified).

conventional

simplified

clean wing

airfoil cut at Y = 650” (mid-span of OB flap)

• The 15o flap deflection of the 

takeoff configuration tends to 

close the gap between the 

conventional and simplified.

• Similar trends due to 

Reynolds number compared 

to the landing configuration.

Figure 15 Lift curve comparisons – clean vs. landing-approach high-lift CRM configurations (conventional & simplified).

conventional

simplified

clean wing

airfoil cut at Y = 650” (mid-span of OB flap)

• Simplified landing 

configuration lift curves seem 

to family reasonably well with 

clean wing and conventional 

landing results

• Stronger Reynolds number 

effect for the simplified 

configuration through linear 

range of lift curve

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Figure 17 Lift curve comparisons for simplified high-lift CRM wing/body configuration 

indicate 25-degree leading-edge droop to be near optimum setting.

Curve fit LED-00 data 

to estimate CLmax

Cross plot CLmax as a 

function of LED

2nd-order curve fit 

suggests optimum is 

near the 25 deg droop

Figure 18 Incremental-lift comparisons between conventional and simplified high-lift 

CRM configurations.

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Figure 19 Pitching-moment comparisons – clean vs. landing-approach high-lift CRM 

configurations (conventional & simplified).

Figure 20 Pitching-moment comparisons – clean vs. takeoff high-lift CRM configurations (conventional & simplified).

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Figure 22 Nominal lift condition for takeoff configuration is selected as reference condition for AFC integration trade studies

D (deg)

CL

0 2 4 6 8 10 12 14 16 18 20 22 24 0.6

0.8

1.0

1.2

1.4

1.6

1.8

2.0

2.2

2.4

2.6

conventional: slat 18o/25o, flap 15o, RN = 24.6 mil

conventional: slat 18o

/25o, flap 15o, RN = 3.3 mil

simplified: LED 20o

/25o, flap 15o, RN = 24.6 mil

simplified: LED 20o

/25o, flap 15o, RN = 3.3 mil

Nacelle/Pylon Off

Mach 0.2

Takeoff

CL

L/D

0.4 0.8 1.2 1.6 2 2.4 6

8

10

12

14

16

18

conventional: slat 18o/25o, flap 15o, RN = 24.6 mil

conventional: slat 18o

/25o, flap 15o, RN = 3.3 mil

simplified: LED 20o

/25o, flap 15o, RN = 24.6 mil

simplified: LED 20o

/25o, flap 15o, RN = 3.3 mil

Nacelle/Pylon Off

Mach 0.2 Takeoff

Nominal condition

Takeoff: V2 > 1.13 x VStall

Figure 21 Nominal lift condition for landing-approach configuration is selected as reference condition for AFC integration trade studies.

D (deg)

CL

0 2 4 6 8 10 12 14 16 18 20 22 24 0.6

0.8

1.0

1.2

1.4

1.6

1.8

2.0

2.2

2.4

2.6

conventional: slat 18o/25o, flap 30o, RN = 24.6 mil

conventional: slat 18o

/25o, flap 30o, RN = 3.3 mil

simplified: LED 20o

/25o, TED 30o, RN = 24.6 mil

simplified: LED 20o

/25o, TED 30o, RN = 3.3 mil

Landing

Nacelle/Pylon Off

Mach 0.2

Nominal condition

Landing: VApp > 1.23 x VStall

Į = 8°

Į = 8°

Conventional high-lift 

CRM at nominal 

landing condition

Simplified high-lift 

CRM at nominal 

landing condition

Nominal CL

26
AFC-Enabled High-Lift System Study

SMAAART Contract # NNL10AA05B

Figure 23 AFC is sized with inviscid limit solution 

Less efficient

actuation

757 vertical tail

70 ports

More efficient

actuation

Inviscid

Momentum

Tail side force

0 0.01 0.02 0.03 0.04 0.05

0.92

0.96

1.00

1.04

1.08

1.12

•Inviscid level is being used to gauge efficiency of AFC

• Desirable to have CL target < CL inviscid

– Guides the geometrical layout Inviscid

Baseline

Figure 24 CFD solutions with constant-blowing boundary conditions identify approximate extent and mass flow rate for AFC high-lift application

Inboard

Inboard+Outboard

Inboard+Outboard 

(reduced)

Cq = 0.001 | 32 pps

27
AFC-Enabled High-Lift System Study

SMAAART Contract # NNL10AA05B

Figure 25 CFD study of AFC integration with NASA B-757 vertical tail guides present 

AFC application study.

 AFC is placed on a backward facing step carved into the flap

 Vertical cut through the surface at jet location

 Step height is scaled from AFC integration validated in recent NFAC windtunnel test of NASA B-757 vertical tail with AFC

 AFC is applied at sections of the step

 Optional spanwise continuous jet or discretized jets

Figure 26 Convergent-divergent ducts are used to model fluidic oscillators for discrete-blowing AFC simulation.

 37 ports in total

 19 ports along inboard flap

 18 ports along outboard flap

 Each port consists of a set of grids

 Trial-&-error used to guide orientation 

of discrete ducts to achieve desired 

Coanda effect for energizing the 

boundary-layer flow 

Jet plume depicted as an iso-surface of M=0.7

Note alignment of jet plumes with flap curvature

g the

28
AFC-Enabled High-Lift System Study

SMAAART Contract # NNL10AA05B

Figure 27 Discrete-port AFC simulation yields most efficient AFC high-lift system.

37 port system

Cq = 0.001 | 32 pps

Figure 28 Trade studies are based on ERA-0003 concept since its mass properties 

are representative of CRM design.

29
AFC-Enabled High-Lift System Study

SMAAART Contract # NNL10AA05B

Figure 29 Both sized and un-scaled AFC variants of ERA-0003 concept indicate potential for overall performance improvements.

Baseline AFC Enabled AFC Enabled

ERA-003-PAX  ̈)1WR OE  ̈)1WR OE

Reference Required 1998 T&W Unscaled Sized FN & S

Perf. Mcr=0.85 Mcr=0.85 Mcr=0.85

RR TF RR TF RR TF

MTOGW 581820 581820 587066

OEW 268686 268084 275240

Payload 50,000 50000 50000 50000

Fuel at Max PL 263134 263736 261826

Range (nm) 8000 8000 8143 8000

Block Fuel Burned 238569 239260 237149

Ton-nm/lbf 0.84 0.85 0.84

Wing Area, aero reference (sqft) 4960 4960 5350

Aspect Ratio, aero reference 8.7222 8.7222 8.7222

Span, aero reference (ft) 208.0 208.0 216.0

TO Ref FN @ SL, 0.0, +27F (lbf/eng) 88000 88000 90950

Number of Engines 2 2 2

T/W 0.30 0.30 0.31

W/S 117 117 110

Cruise type Step-Cruise Step-Cruise Step-Cruise

Initial cruise altitude (ft) 35000 35000 35000 35000

Top of climb thrust (all engines, tau=1) 35122 35122 36299

Initial cruise Mach 0.85 0.85 0.85 0.85

Initial cruise L/D 19.372 19.650 19.66

Initial cruise SFC 0.5921 0.5921 0.5922

Initial cruise tau 0.833 0.821 0.802

Initial cruise CL 0.4536 0.4538 0.425

Distance to climb (nm) 200 199 193 179

Takeoff field length @ SL, 86F (ft) 10500 8657 9759 8656

2nd Segment Gradient (%) 2.4 2.527 2.444 2.452

Landing field length (ft) 5200 4168 4164 4049

Approach speed (ktas) 150 108 108 105

1.7% gain 

in range

0.6% lower 

fuel burn

1,110 ft TO 

field length 

penalty

Figure 30 Climb-to-cruise requirement drives performance of baseline ERA-0003.

TOGW BlkFuel RJFL optimumHp distToClimb

Sw ing

4500 4650 4800 4950 5100 5250 5400 5550 5700 5850 6000

Fn 100000

99000

98000

97000

96000

95000

94000

93000

92000

91000

90000

89000

88000

87000

86000

85000

577221

579638

582055

584472

586889

589306

591723

594140

596557

598974

601391

603808

606225

608642

238039

238873

239706

240539

241372

242206

243039

243872

244705

245539

246372

246372

247205 247205 248038

248038

248872

248872

35000

200

Wing Area (sqft)

Thrust per Engine (lbf)

Baseline

SW = 4960 sqft

FN = 88,000 lb

Range = 8000 nm

ICA=35,000 ft TO Dist = 8657 ft

Climb to ICA in 200 nm

30
AFC-Enabled High-Lift System Study

SMAAART Contract # NNL10AA05B

Figure 31 Takeoff field length requirement drives performance of AFC variant of ERA0003.

TOGW BlkFuel RJFL optimumHp distToClimb

Sw ing

4500 4650 4800 4950 5100 5250 5400 5550 5700 5850 6000

Fn 100000

99000

98000

97000

96000

95000

94000

93000

92000

91000

90000

89000

88000

87000

86000

85000

571657

574416

577176

579935

582695

585454

588214

590973

593733

596493

599252

602012 604771

607531

233644

234501

235357

236214

237071

237927

238784

239641

240497

241354

242210

243067

243067

243924

243924

244780

244780

8657

35000

200

Wing Area (sqft)

Thrust per Engine (lbf)

AFC-Enabled

SW = 5350 sqft

FN = 90,950 lb

Range = 8000 nm

TO Dist = 8657 ft

Baseline

SW = 4960 sqft

FN = 88,000 lb

ICA=35,000 ft

Climb to ICA in 200 nm

TOD=8657 ft

31
REPORT DOCUMENTATION PAGE Form Approved

OMB No. 0704-0188

2. REPORT TYPE 

Contractor Report

 4. TITLE AND SUBTITLE

AFC-Enabled Simplified High-Lift System Integration Study 

5a. CONTRACT NUMBER

NNL10AA05B

 6. AUTHOR(S)

Hartwich, Peter M.; Dickey, Eric D.; Sclafani, Anthony J.; Camacho, Peter; 

Gonzales, Antonio B.; Lawson, Edward L.; Mairs, Ron Y.; Shmilovich, 

Arvin

 7. PERFORMING ORGANIZATION NAME(S) AND ADDRESS(ES)

NASA Langley Research Center 

Hampton, Virginia 23681

 9. SPONSORING/MONITORING AGENCY NAME(S) AND ADDRESS(ES)

National Aeronautics and Space Administration

Washington, DC 20546-0001

 8. PERFORMING ORGANIZATION

 REPORT NUMBER

10. SPONSOR/MONITOR'S ACRONYM(S)

NASA

13. SUPPLEMENTARY NOTES

Langley Technical Monitor: John C. Lin

12. DISTRIBUTION/AVAILABILITY STATEMENT

Unclassified - Unlimited

Subject Category 34

Availability: NASA CASI (443) 757-5802

19a. NAME OF RESPONSIBLE PERSON

STI Help Desk (email: help@sti.nasa.gov)

14. ABSTRACT

The primary objective of this trade study report is to explore the potential of using Active Flow Control (AFC) for achieving lighter and mechanically simpler 

high-lift systems for transonic commercial transport aircraft. This assessment was conducted in four steps. First, based on the Common Research Model 

(CRM) outer mold line (OML) definition, two high-lift concepts were developed. One concept, representative of current production-type commercial transonic 

trans-ports, features leading edge slats and slotted trailing edge flaps with Fowler motion. The other CRM-based design relies on drooped leading edges and 

simply hinged trailing edge flaps for high-lift generation. The relative high-lift performance of these two high-lift CRM variants is established using 

Computational Fluid Dynamics (CFD) solutions to the Reynolds-Averaged Navier-Stokes (RANS) equations for steady flow. These CFD assessments identify 

the high-lift performance that needs to be recovered through AFC to have the CRM variant with the lighter and mechanically simpler high-lift system match 

the performance of the conventional high-lift system. Conceptual design integration studies for the AFC-enhanced high-lift systems were conducted with a 

NASA Environmentally Responsible Aircraft (ERA) reference configuration, the so-called ERA-0003 concept. These design trades identify AFC performance 

targets that need to be met to produce economically feasible ERA-0003-like concepts with lighter and mechanically simpler high-lift designs that match the 

performance of conventional high-lift systems. 

15. SUBJECT TERMS

Active flow control; Commercial transport; Common research model; High lift system; High lift integration; System 

integration

18. NUMBER

 OF 

 PAGES

39

19b. TELEPHONE NUMBER (Include area code)

(443) 757-5802

a. REPORT

U

c. THIS PAGE

U

b. ABSTRACT

U

17. LIMITATION OF 

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Prescribed by ANSI Std. Z39.18

Standard Form 298 (Rev. 8-98)

3. DATES COVERED (From - To)

5b. GRANT NUMBER

5c. PROGRAM ELEMENT NUMBER

5d. PROJECT NUMBER

5e. TASK NUMBER

NNL13AB65T

5f. WORK UNIT NUMBER

473452.02.07.03.02.01

11. SPONSOR/MONITOR'S REPORT

 NUMBER(S)

NASA/CR-2014-218521

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1. REPORT DATE (DD-MM-YYYY)

01- 09 - 2014